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Table 8.4-9: Reflood Test Matrix ............................................................................. 8.4-48

Table 8.5-1 Methodology Treatment of Important PIRT Phenomena .................... 8.5-27

Table 8.5-2 Summary of Evaluated Uncertainties of Important PIRT Parameters . 8.5-30

Table 8.5-3 Packing Factors and Sources ............................................................ 8.5-51

Table 8.5-4 Biases Used in Assessments ............................................................. 8.5-54

Table 8.5-5 Film Boiling Multiplier ......................................................................... 8.5-58

Table 8.5-6 Dispersed Flow Film Boiling Multiplier ................................................ 8.5-58

Table 8.6-1: Test Ranges for Film Boiling Heat Transfer Test Comparison ............ 8.6-18

Table 9-1 Large Break LOCA Nodalization .............................................................. 9-26

Table 9-2 Confidence Level and Sample Size ......................................................... 9-55

Table 9-3 Minimum Sample Size for Tolerance Regions Constructed from Sample .................................................................................................. 9-56

Table 9-4 NPP Parameters for Consideration in the Performance of a RLBLOCA Analysis ............................................................................... 9-58

Table 9-5 Relationship of PIRT to Operational Parameters ..................................... 9-59 Table A-1 Nodalization Numbering for Westinghouse PWR Plants ......................... A-10 Table A-2 Nodalization Numbering Differences for CE PWR Plants ........................ A-11 Table A-3 Required Sample Size ........................................................................... A-122

Table A-4 Film Boiling Multiplier ............................................................................. A-127

Table A-5 Dispersed Flow Film Boiling Multiplier ................................................... A-127

Table A-6 Uncertainty Parameters and PDFs ........................................................ A-129 Table A-7 Model Parameter Uncertainty Ranges ................................................... A-130

Table A-8 Phenomenological Model Parameters ..................................................... A-131 Table A-9 Example Statistical Evaluation Data ........................................................ A-134 Table A-10 Parameter Input for Offsite Power Determination ................................ A-138 Table A-11: Analysis Codes ..................................................................................... A-140 Table A-12: Automation Computer Codes ................................................................ A-140 Table A-13: Plant Parameter Uncertainty Variables Required ................................. A-143 Table A-14 Power History Data Files ..................................................................... A-148 Table A-15 Power History Data Values Read from File .......................................... A-148 Table A-16 Power History Calculated Values ......................................................... A-149 Table A-17: Power History Calculated Values .......................................................... A-151

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Table A-18: Power History Calculated Values .......................................................... A-154 Table A-19 S RELAP5 Rod Heat Structure ............................................................ A-158

Table A-20 Random Number Sequence ................................................................ A-162

Table A-21 Radial Power Peaking Factors ............................................................. A-177

Table A-22 Number of Outer Row Assemblies ....................................................... A-179

Table A-23 Minor Edit Requests ............................................................................ A-221

Table A-24 Single Case, Single Rod COPRE Calculation Input ............................. A-223

Table A-25 Single Case, Single Rod COPRE Calculation Output .......................... A-223

Table A-26 Single Case, Single Rod COPERNIC Calculation Input ...................... A-223

Table A-27 Single Case, Single Rod COPERNIC Calculation Output .................... A-224

Table A-28 Single Case Steady-State Calculation Input ........................................ A-224

Table A-29 Single Case Steady-State Calculation Output ..................................... A-224

Table A-30 Single Case Transient Calculation Input .............................................. A-225

Table A-31 Single Case Transient Calculation Output ........................................... A-225

Table A-32 Key Operational Parameter Plots ........................................................ A-227

Table A-33 Scatter Plot Parameters ....................................................................... A-228

Table A-34 Maximum Clad Surface Temperature Control Variables ...................... A-229

Table A-35 Limiting Case Plots .............................................................................. A-230

Table B-1 Identification of Heat Transfer Parameters during a Limiting LBLOCA Simulation ............................................................................................. B-13

Table B-2 Simulation and Application Space for CHF during Blowdown .................. B-14

Table B-3 Simulation and Application Space for Film Boiling Heat Transfer Including Thermal Radiation ................................................................. B-15

Table B-4 Simulation and Application Space for Transition Boiling Heat Transfer ... B-16

Table B-5 Simulation and Application Space for Nucleate Boiling Heat Transfer (late reflood) .......................................................................................... B-16

Table B-6 Summary of Full Range of Applicability ................................................... B-17

Table B-7 3-Loop Westinghouse Plant Operating Range Supported by the RLBLOCA Analysis ............................................................................... B-21

Table B-8 3-Loop Westinghouse Statistical Distribution Used for Process Parameters ........................................................................................... B-24

Table B-9 3-Loop Westinghouse Summary of Major Parameters for the MRMC Case ..................................................................................................... B-25

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Table B-10 3-Loop Westinghouse Compliance with 10 CFR 50.46 ......................... B-26

Table B-11 3-Loop Westinghouse Calculated Event Times for the MRMC Case ..... B-27

Table B-12 Westinghouse 3-Loop Heat Transfer Parameters for Limiting MRMC Case ..................................................................................................... B-28

Table B-13 Westinghouse 3-Loop Fuel Rod Rupture Ranges of Parameters for all [ ] cases ................................................................................ B-29

Table B-14 4-Loop Westinghouse Plant Operating Range Supported by the LOCA Analysis ...................................................................................... B-53

Table B-15 4-Loop Westinghouse Statistical Distribution Used for Process Parameters ........................................................................................... B-56

Table B-16 4-Loop Westinghouse Summary of Plant Major Parameters for for the MRMC Case ................................................................................... B-57

Table B-17 4-Loop Westinghouse Compliance with 10 CFR 50.46 ......................... B-58

Table B-18 4-Loop Westinghouse Calculated Event Times for the MRMC Case ..... B-59

Table B-19 Westinghouse 4-Loop Heat Transfer Parameters for the MRMC Case ..................................................................................................... B-60

Table B-20 Westinghouse 4-Loop Fuel Rod Rupture Ranges of Parameters for all [ ] Cases ................................................................................ B-61

Table B-21 CE 2x4 Plant Operating Range Supported by the LOCA Analysis ........ B-85

Table B-22 CE 2x4 Statistical Distribution Used for Process Parameters ................ B-88

Table B-23 CE 2x4 Summary of Major Parameters for the MRMC Case ................. B-89

Table B-24 CE 2x4 COPERNIC Compliance with 10 CFR 50.46 ............................. B-90

Table B-25 CE 2x4 Calculated Event Times for the MRMC Case ............................ B-91

Table B-26 CE 2x4 Heat Transfer Parameters for the MRMC Case ........................ B-92

Table B-27 CE 2x4 Fuel Rod Rupture Ranges of Parameters for all [ ] cases .................................................................................................... B-93

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Figure 8.5-3 Decay Heat Ratios, Finite Operation Over Infinite Operation U235, All Isotopes (0 to 10 seconds) ............................................................ 8.5-20

Figure 8.5-4 Decay Heat Ratios, Finite Operation Over Infinite Operation U235, All Isotopes (0 to 600 seconds) .......................................................... 8.5-21

Figure 8.5-5 Rupture Temperature Data and Dunn’s M5 Rupture Temperature Correlation from BAW-10227 (Ref. 3: p. K-35) .................................. 8.5-45

Figure 8.5-6 M5 Slow Ramp Correlations with Supporting Rupture Strain Data ... 8.5-47

Figure 8.5-7 M5 Fast Ramp Correlations with Support Rupture Strain Data1 ......... 8.5-47

Figure 8.5-8 Packing Factor Data and Fits ............................................................. 8.5-49

Figure 8.5-9 COPERNIC2 Cumulative Centerline Fuel Temperature Error Distribution ......................................................................................... 8.5-59

Figure 8.5-10 Temperature Distribution in the Vessel Wall – S-RELAP5 versus Exact Solution .................................................................................... 8.5-60

Figure 8.6-1 Data Based Nusselt Number versus Reynolds Number for FLECHT-SEASET Steam Cooling Tests Compared with Dittus-Boelter Correlation ........................................................................................... 8.6-9

Figure 9-1 Uncertainty Analysis Case Description ................................................... 9-27

Figure 9-2 Sample Loop Nodalization for NPP ......................................................... 9-28

Figure 9-3 Sample Steam Generator Secondary Nodalization for NPP ................... 9-29

Figure 9-4 Double-Ended Guillotine and Split Break Nodalization ........................... 9-30

Figure 9-5 Sample Reactor Vessel Nodalization for NPP ........................................ 9-31

Figure 9-6 Westinghouse/AREVA 3- and 4-Loop and CE 2x4 Plant Vessel Downcomer Configurations ................................................................... 9-32

Figure 9-7 NPP Core Nodalization ........................................................................... 9-33

Figure 9-8 Sample NPP Upper Plenum Nodalization - Axial Plane .......................... 9-34

Figure 9-9 Sample NPP Upper Plenum Nodalization - Cross-Sectional Plane ........ 9-35

Figure 9-10 Reactor Coolant Pump Showing Impeller Spill Height ............................ 9-36 Figure A-1 Uncertainty Analysis Case Description ................................................. A-232

Figure A-2 Loop Nodalization Example .................................................................. A-233

Figure A-3 Loop 1 Secondary Side Nodalization .................................................... A-234

Figure A-4 Reactor Vessel Nodalization Example (Downflow Baffle Case) ........... A-235

Figure A-5 Westinghouse 3- and 4-Loop and CE 2x4 Loop Plant Vessel Downcomer Configuration .................................................................. A-236

Figure A-6 Core Nodalization Example – Axial Plane (example) ........................... A-237

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Figure A-7 Core Nodalization – Cross-Sectional Plane .......................................... A-238

Figure A-8 Spacer and Node Locations Example for 23 Volume Core (example) .. A-239

Figure A-9 Upper Plenum Nodalization – Axial Plane (for Plants with Mixer Vanes/Standpipes) ............................................................................. A-240

Figure A-10 Upper Plenum Nodalization – Cross-Sectional Plane (for Plants with Mixer Vanes/Standpipes) .................................................................... A-241

Figure A-11 Upper Plenum Nodalization – Axial Plane (for Plants with UHI Columns) ............................................................................................ A-242

Figure A-12 Upper Plenum Nodalization – Axial Plane (for Plants without Mixer Vanes/Standpipes) ............................................................................. A-243

Figure A-13 Detailed Emergency Core Cooling System Nodalization Example ....... A-244

Figure A-14 Double-Ended Guillotine Break Nodalization ........................................ A-245

Figure A-15 Double-Ended Split Break Nodalization ................................................ A-246

Figure B-1 3-Loop Westinghouse Scatter Plot of Operational Parameters for all [ ] Cases ..................................................................................... B-30

Figure B-2 3-Loop Westinghouse PCT versus PCT Time Scatter Plot from the Cases within the 95/95 Range .............................................................. B-32

Figure B-3 3-Loop Westinghouse PCT versus Break Size Scatter Plot from the Cases within the 95/95 Range .............................................................. B-33

Figure B-4 3-Loop Westinghouse Maximum Local Oxidation versus PCT Scatter Plot from the Cases within the 95/95 Range ........................................ B-34

Figure B-5 3-Loop Westinghouse Total Core-Wide Oxidation versus PCT Scatter Plot from the Cases within the 95/95 Range ........................... B-35

Figure B-6 3-Loop Westinghouse Peak Cladding Temperature (Independent of Elevation) for the MRMC Case ............................................................. B-36

Figure B-7 3-Loop Westinghouse Break Flow for the MRMC Case ......................... B-37

Figure B-8 3-Loop Westinghouse Core Inlet Mass Flux for the MRMC Case .......... B-38

Figure B-9 3-Loop Westinghouse Core Outlet Mass Flux for the MRMC Case ........ B-39

Figure B-10 3-Loop Westinghouse Void Fraction at RCS Pumps for the MRMC Case ..................................................................................................... B-40

Figure B-11 3-Loop Westinghouse ECCS Flows (Includes Accumulator, Charging, SI and RHR) for the MRMC Case ........................................ B-41

Figure B-12 3-Loop Westinghouse Upper Plenum Pressure for the MRMC Case ... B-42

Figure B-13 3-Loop Westinghouse Collapsed Liquid Level in the Downcomer for the MRMC Case ................................................................................... B-43

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Figure B-14 3-Loop Westinghouse Collapsed Liquid Level in the Lower Plenum for the MRMC Case .............................................................................. B-44

Figure B-15 3-Loop Westinghouse Collapsed Liquid Level in the Core for the MRMC Case ......................................................................................... B-45

Figure B-16 3-Loop Westinghouse Containment and Loop Pressures for the MRMC Case ......................................................................................... B-46

Figure B-17 3-Loop Westinghouse Pressure Difference between Upper Plenum and Downcomer .................................................................................... B-47

Figure B-18 3-Loop Westinghouse Validation of BOCR Time using MPR CCFL Correlation ............................................................................................ B-48

Figure B-19 4-Loop Westinghouse Scatter Plot of Operational Parameters for all [ ] cases ...................................................................................... B-62

Figure B-20 4-Loop Westinghouse PCT versus PCT Time Scatter Plot from the Cases within the 95/95 Range .............................................................. B-64

Figure B-21 4-Loop Westinghouse PCT versus Break Size Scatter Plot from the Cases within the 95/95 Range ........................................................ B-65

Figure B-22 4-Loop Westinghouse Maximum Oxidation versus PCT Scatter Plot from the Cases within the 95/95 Range ................................................ B-66

Figure B-23 4-Loop Westinghouse Total Core-Wide Oxidation versus PCT Scatter Plot from the Cases within the 95/95 Range ............................ B-67

Figure B-24 4-Loop Westinghouse Peak Cladding Temperature (Independent of Elevation) for the MRMC Case ............................................................. B-68

Figure B-25 4-Loop Westinghouse Break Flow for the MRMC Case ....................... B-69

Figure B-26 4-Loop Westinghouse Core Inlet Mass Flux for the MRMC Case ........ B-70

Figure B-27 4-Loop Westinghouse Core Outlet Mass Flux for the MRMC Case ...... B-71

Figure B-28 4-Loop Westinghouse Void Fraction at RCS Pumps for the MRMC Case ..................................................................................................... B-72

Figure B-29 4-Loop Westinghouse ECCS Flows (Includes Accumulator, Charging, SI and RHR) for the MRMC Case ........................................ B-73

Figure B-30 4-Loop Westinghouse Upper Plenum Pressure for the MRMC Case ... B-74

Figure B-31 4-Loop Westinghouse Collapsed Liquid Level in the Downcomer for the MRMC Case ................................................................................... B-75

Figure B-32 4-Loop Westinghouse Collapsed Liquid Level in the Lower Plenum for the MRMC Case .............................................................................. B-76

Figure B-33 4-Loop Westinghouse Collapsed Liquid Level in the Core for the MRMC Case ......................................................................................... B-77

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Figure B-34 4-Loop Westinghouse Containment and Loop Pressures for the MRMC Case ......................................................................................... B-78

Figure B-35 4-Loop Westinghouse Pressure Difference between Upper Plenum and Downcomer .................................................................................... B-79

Figure B-36 4-Loop Westinghouse Validation of BOCR Time using MPR CCFL Correlation ............................................................................................ B-80

Figure B−37 CE 2x4 Scatter Plot of Operational Parameters for all [ ] cases .................................................................................................... B-94

Figure B−38 CE 2x4 PCT versus PCT Time Scatter Plot from the Cases within the 95/95 Range ................................................................................... B-96

Figure B−39 CE 2x4 PCT versus Break Size Scatter Plot from the Cases within the 95/95 Range ................................................................................... B-97

Figure B−40 CE 2x4 Maximum Local Oxidation versus PCT Scatter Plot from the Cases within the 95/95 Range ........................................................ B-98

Figure B−41 CE 2x4 Total Oxidation versus PCT Scatter Plot from the Cases within the 95/95 Range ......................................................................... B-99

Figure B−42 CE 2x4 Peak Cladding Temperature (Independent of Elevation) for the MRMC Case (COPERNIC) ........................................................... B-100

Figure B−43 CE 2x4 Break Flow for the MRMC Case ............................................ B-101

Figure B−44 CE 2x4 Core Inlet Mass Flux for the MRMC Case ............................. B-102

Figure B−45 CE 2x4 Core Outlet Mass Flux for the MRMC Case .......................... B-103

Figure B−46 CE 2x4 Void Fraction at RCS Pumps for the MRMC Case ................ B-104

Figure B−47 CE 2x4 ECCS Flows (Includes SIT, Charging, SI and RHR) for the MRMC Case ....................................................................................... B-105

Figure B−48 CE 2x4 Upper Plenum Pressure for the MRMC Case ...................... B-106

Figure B−49 CE 2x4 Collapsed Liquid Level in the Downcomer for the MRMC Case ................................................................................................... B-107

Figure B−50 CE 2x4 Collapsed Liquid Level in the Lower Plenum for the MRMC Case ................................................................................................... B-108

Figure B−51 CE 2x4 Collapsed Liquid Level in the Core for the MRMC Case ....... B-109

Figure B−52 CE 2x4 Containment and Loop Pressures for the MRMC Case ........ B-110

Figure B−53 CE 2x4 Pressure Difference between Upper Plenum and Downcomer for the MRMC Case ........................................................ B-111

Figure B−54 CE 2x4 Validation of BOCR Time using MPR CCFL Correlation ....... B-112

Figure D−1: Time Step Sensitivity of Westinghouse 3-Loop Analysis ........................ D-3

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Figure D−2: Variability of Westinghouse 3-Loop Analysis .......................................... D-4

Figure D−3: Time Step Sensitivity of Westinghouse 4-Loop Analysis ........................ D-5

Figure D−4: Variability of Westinghouse 4-Loop Analysis .......................................... D-6

Figure D−5: Time Step Sensitivity of CE Analysis ...................................................... D-7

Figure D−6: Variability of CE Analysis ........................................................................ D-8

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ABSTRACT

AREVA NP had developed and licensed a large-break loss-of-coolant accident (LOCA) methodology for Westinghouse 3- and 4-loop designs and Combustion Engineering 2x4 designs (April 2003). The licensed methodology uses a non-parametric statistical sampling approach for the uncertainty treatment following the Wilks’ method. In this method, uncertainty contributors are ranged individually to determine the expected peak cladding temperature (PCT), Maximum Local Oxidation (MLO) and total Whole Core Oxidation (WCO) response.

In this report, an alternative uncertainty methodology is developed resolving the multi-variant evaluation by implementing Tukey’s Tolerance Region approach. This methodology replaces the Revision 0 Wilks’ technique. In this technique, the results are ranked for PCT, MLO, and WCO and reported based on margin to the specific criteria. The new technique resolves the NRC staff’s previous concerns with AREVA’s statistical approach. The method is called the AREVA PWR Realistic LBLOCA (RLBLOCA) Revision 3.

This report documents a road map to the methodology, patterned after the Code Scaling, Applicability, and Uncertainty (CSAU) methodology. The thermal-hydraulic computational tool used in AREVA’s RLBLOCA Revision 3 is S-RELAP5. The models and correlations used in the code are also documented here. The sections describing the models and correlations are taken from the approved topical report with minimal revisions. Application to a Westinghouse and AREVA 3-loop and 4-loop plant designs along with a Combustion Engineering 2x4 plant design are given as a demonstration of RLBLOCA Revision 3, following the technical bases of the methodology changes.

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Section 4.0 Scenario Identification

This section describes the first two steps of the EMDAP, namely Element 1, Step 1

(E1.S1) Specify Analysis Purpose, Transient Class, and Power Plant Class and E1.S2

Specify Figures of Merit. The corresponding steps from the CSAU methodology are

Step 1 Scenario Specification (Reference 2-10, Section 2.1) and Step 2 Nuclear Power

Plant Selection (Reference 2-10, Section 2.2).

Section 5.0 Evaluation Model Requirements

This section presents the Evaluation Model requirements and summarizes them in the

Phenomena Identification and Ranking Table (PIRT), which provides the basis for

determining code applicability (does the code properly model the important

phenomena); establishing the assessment matrix (identifying test data that contain the

appropriate phenomena during each accident phase), and identifying phenomenological

parameters to be ranged and quantified for evaluating uncertainties. Section 5.0

corresponds to Element 1, Steps 3 and 4 (E1.S3 and E1.S4) of EMDAP and CSAU

Step 3, and it incorporates Section 2.3 of Reference 2-10.

Section 6.0 Assessment Data Base Description

This section provides the objectives for the assessment base, identifies the existing

data being used for EM development, presents an assessment matrix which lists the

test facilities, the actual tests analyzed from each test facility, and the associated

phenomena being examined. This section covers applicable steps from Element 2 of

EMDAP and corresponds to CSAU Step 7 (Section 3.1 of Reference 2-10).

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Section 7.0 Evaluation Model Computer Codes Description

This section describes in detail the S-RELAP5 code, including the COPERNIC and

ICECON modules used in the evaluation model. It provides detailed descriptions of the

models and correlations implemented in S-RELAP5 and relevant to the RLBLOCA EM.

This section addresses Steps 10 through 13 of the Element 3 and Steps 13, 15, 16, and

17 of Element 4 of EMDAP. It also covers Steps 4, 5, and 6 of the CSAU methodology.

The content provided in this section originates mainly from the relevant segments of

EMF-2100 S-RELAP5 Models and Correlations Code Manual (Reference 2-6), as well

as Sections 3.4, 3.5, 3.6 of Reference 2-10, and the relevant portions from the ICECON

theory and user's manual (References 2-7 and 2-8). New content has been added to

describe the revised model for Fuel Swelling, Rupture and Relocation (FSRR).

Section 8.0 Assessment Results

This section presents the results of the assessment process intended to demonstrate

the code capabilities for simulation of important phenomena primarily associated with

large-scale PWR systems LBLOCA. The material previously presented in Reference 2-

10, Sections 3.3 and 3.4 has been greatly expanded by including all the relevant

material from EMF-2102 S-RELAP5 Code Verification and Validation (Reference 2-9).

Section 8.0 addresses the remaining steps from Element 4 of EMDAP as well as Steps

9 and 10 of the CSAU.

Section 9.0 Evaluation Model Implementation

This section describes the implementation of the Evaluation Model, including the

nodalization definition and the sensitivity and uncertainty analysis, covering Step 9 and

11 through 14 of the CSAU. The relevant sections from Supplement 1 of Revision 2

(Reference 2-11) have also been incorporated into this section.

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1. Wong-Hochreiter Correlation – Revision 3

The Forslund-Rohsenow (F-R) correlation is no longer used in the core heat transfer

that determines the fuel cladding temperature. The F-R correlation is still used for

the passive metal heat structure heat transfer. For the dispersed flow film boiling

regime in the core, Wong-Hochreiter with enhancements replaces the use of

Sleicher-Rouse. This alteration was adopted as a model improvement. The

modified approach to flow film boiling in the core is presented in Section 7.6.7.2 and

is assessed in Sections 8.2.1, 8.2.4, 8.4.1 and 8.4.4. A temperature correction was

added to the Wong-Hochreiter heat transfer correlation and is also presented in

Section 7.6.7.2, while its impact is qualitatively assessed in Section 8.1.5.

2. Rod-to-Rod Radiation – Revisions 2 and 3

A rod-to-rod radiation model has been incorporated into the methodology and the

reflood heat transfer benchmarking has been redone. This upgrade was

incorporated to more accurately assess reflood heat transfer by recognizing the

individual components of the process. The alteration is presented in Section 7.6.8.2

and assessed in Sections 8.2.5, 8.5.2.4 and 8.6.2.1. The model was subsequently

revised in Revision 3 to implement separate radiation enclosures for each burned

rod, rather than one enclosure for all burned rods. The model change is presented in

Appendix A and its impact qualitatively assessed in Section 8.1.5.

3. Cold Leg Condensation Model – Transition Package and Revision 3

A cold leg condensation model, specific to both the accumulator and the pumped

injection period of the accident, has been incorporated. The revised model predicts

more accurately the cold leg condensation during the pumped ECC injection phase

resulting in near saturated fluid conditions at the downcomer entrance, thus

conservatively increasing the potential for downcomer boiling. This alteration is

presented in Sections 8.2.9, 8.2.10, and 8.5.1.14.

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4. Statistical Evaluation – Revision 2

The statistical evaluation has been upgraded, with the application of the Tukey

methodology, to provide a multi-variant evaluation. This alteration is presented in

Section 9.4.

5. COPERNIC Fuel Performance Code – Revision 2

This change has been applied in response to NRC concerns over thermal

conductivity degradation. The COPERNIC fuel performance code has replaced

RODEX3A as the source of fuel initial conditions. COPERNIC is NRC approved for

application to M5 cladding.

6. Second Cycle Fuel – Transition Program, Supplement 1

The methodology has been upgraded such that a direct calculation of second cycle

fuel performance is accomplished. This expands the range of evaluations and

ensures that fuel experiencing its second burn will be evaluated and, if limiting,

recognized as limiting. This alteration is presented in Appendix A.

7. Break Modeling – Transition Program

The break modeling was altered from EMF-2103, Revision 0 to concur with the

approach outlined in Regulatory Guide 1.157. The split versus double-ended break

type is no longer related to break area. This alteration is presented in Section

8.5.2.6.

8. Interfacial Drag Package – Revision 2

The interfacial drag package has been modified with improved logic for transition

between flow regimes to cover a wider range of experimental data. This serves to

update the state-of–the-art of S-RELAP5. The details of this alteration are

presented in Section 7.5.2.

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9. Reported Local Cladding Oxidation – Supplement 1

The accounting of the operational (pre-transient) and interior oxidation of the

cladding for compliance with the maximum local oxidation criteria of 10 CFR 50.46 is

presented in Section 7.9.3.5 and Appendix A, sub-section A.2.3.10.

10. Decay Heat Simulation – Supplement 1

The decay heat calculation, which in EMF-2103, Revision 0 had been sampled

according to the standard deviation presented in the 1979 ANS standard, has been

replaced by a fixed, non-sampled, application of the 1979 standard which bounds

possible decay heat values for Uranium Oxide fuel. The change, previously

presented in Supplement to Revision 2, provides assurance that the transient power

of the fuel rod is not undervalued. The fixed, non-sampled application of the 1979

standard bounds the best estimate method. The details of this change are

presented in Section 8.5.1.17.

11. Fuel Swelling Rupture and Relocation (FSRR) Modeling – Revision 3

A model for FSRR based on a statistical approach for geometry and the evaluation

of cooling for a fuel rod isolated from other ruptures has been added. This model

improves the evaluation of fuel rod rupture during LOCA through a mechanistic

approach and it includes a sub-channel cooling model. The details of this change

are presented in Section 7.9.3.3.

12. Clarification of Single Failure – Supplement 1

The documentation of the treatment of single failure within the evaluation model is

upgraded to clarify the approach. The revised documentation is provided in

Appendix A, sub-section A.2.4.1.1.

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13. Correction to the Steam Absorptivity – Revision 3

A correction to the steam absorptivity was made. In computing the vapor absorption

coefficient, the pressure is conservatively truncated at 150 psi. This alteration is

presented in Section 7.6.8.1 and its impact qualitatively assessed in Section 8.1.5.

14. Core Nodalization – Revision 3

The core nodalization has been slightly changed to align the node boundaries with

the bottom of the grid spacers, rather than the grid centerline. The change in the

core nodalization effectively changes the hydrodynamic volume boundaries such

that they are aligned with the bottom of the grid spacers, in support of the

implementation of the grid droplet shattering model (item #15 below). This alteration

is presented in Section 9.0 and Appendix A and its impact qualitatively assessed in

Section 8.1.5.

15. Grid Spacer Droplet Breakup Heat Transfer Enhancement – Revision 3

A model to increase the heat transfer downstream of a grid spacer due to droplet

breakup was added. The implementation of a model to increase the heat transfer

downstream of a grid spacer is expected to have an impact during reflood above the

mid-plane of the core. This alteration is presented in Section 7.5.4.10.1 and its

impact verified in Sections 8.2.3 and 8.4.1.

16. Interphase Heat Transfer – Revision 3

The interphase heat transfer for mist flow was modified to raise steam and cladding

temperatures and to obtain better agreement with test data from separate effects

reflood tests. The details of this change are presented in Section 7.5.4.

17. Steam generator Tube Inlet Interfacial Drag – Revision 3

An error correction to the level tracking model required modification of the steam

generator tube inlet drag. The model change is presented in Appendix A and its

impact is qualitatively assessed in Section 8.1.5.

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acceptable provided their technical basis is demonstrated with appropriate

data and analysis."

The post-CHF heat transfer model includes provisions for thermal radiation between

structures (rod-to-rod). This adds to the current model, which already includes thermal

radiation from structures to the fluid (rod-to-droplets and rod-to-steam). The rod-to-rod

radiation model is only applied to the hot rod because its power level is elevated

compared to its surroundings. Applying rod-to-rod radiation exclusively to the hot rod

logically leads to the development of separate heat transfer uncertainties for the hot rod

and the rest of the core.

The core wide heat transfer uncertainty was developed from code comparisons using

the FLECHT-SEASET reflood test data as discussed in Section 8.4.1. These

comparisons were used to derive the heat transfer multipliers that are applied to film

boiling (FILMBL) heat transfer and dispersed flow film boiling heat transfer (DFFBHTC). [

]

Assessment of this configuration is performed by using the same FLECHT-SEASET

reflood tests that were used to determine the heat transfer multipliers FILMBL and

DFFBHTC, as discussed in Section 8.4.1. Note that the post-CHF heat transfer

multipliers were re-evaluated for Revision 3, as presented in Section 8.4.1.4.

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3.1.3.4.2 Calculation of Post-Blowdown Thermal Hydraulics for Pressurized Water Reactors

The refill and reflood phases of the transient are calculated on a best-estimate basis,

taking into consideration the thermal and hydraulic characteristics of the core, the ECCS

performance, and important reactor systems. The distribution of water and steam in the

reactor vessel is calculated directly from the S-RELAP5 conservation equations, and

appropriate constitutive relations.

For the S-RELAP5/ICECON code interface, break flow junction variables (e.g.,

velocities, specific enthalpies, densities, void fractions) are transferred each time step

from S-RELAP5 to ICECON. These variables are then used in ICECON to generate a

new containment pressure which is transferred back to S-RELAP5 and used to alter the

pressure in the time-dependent volume or volumes which represent the containment in

the S-RELAP5 model. At each time step, S-RELAP5 performs the necessary data

transfers between the two codes and calls for execution of the external code. After

execution of the external code, control is returned to S-RELAP5, which continues

execution.

A series of sensitivity studies was performed using S-RELAP5 with the ICECON

interface to demonstrate the equivalence of the Tagami-Uchida and Uchida best

estimate condensation heat transfer formulations. These studies were performed using

best estimate S-RELAP5 and ICECON input models for a three-loop PWR with dry

containment. The simulated transient was a double-ended large break LOCA with the

break located in a reactor coolant pump discharge pipe. [

]

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3.1.4.2.2 Application of Methodology

The AREVA NP RLBLOCA methodology is a statistics-based methodology; therefore,

the application does not involve the evaluation of different deterministic calculations.

Instead, a minimum set of LOCA calculations, as detailed in Section 9.0, are performed

with the values of key parameters randomly varied over identified uncertainty ranges.

The methodology has the advantage of being able to treat a large number of

parameters by randomly varying each parameter in each single calculation. This

random selection process is repeated to define a large number of RLBLOCA

calculations, all of which are then run.

All criteria are shown to be met simultaneously with at least 95 percent probability and

95 percent confidence by comparing the peak cladding temperature, local oxidation and

core-wide oxidation values to their related criteria.

3.2 References 3-1 “Emergency Core Cooling Systems; Revisions to Acceptance Criteria,”

Federal Register, Vol. 53, No. 180, September 16, 1988, 10 CFR Part

50.

3-2 NUREG-1230, “Compendium of ECCS Research for Realistic LOCA

Analysis,” December 1988.

3-3 Regulatory Guide 1.157, “Best-Estimate Calculations of Emergency

Core Cooling System Performance,” U.S. NRC, May 1989.

3-4 NUREG/CR-5249, “Quantifying Reactor Safety Margins, Application of

Code Scaling, Applicability, and Uncertainty Evaluation Methodology to

a Large Break, Loss-of-Coolant Accident,” U.S. NRC, December 1989.

3-5 NUREG-0800, U.S. Nuclear Regulatory Commission Standard Review

Plan.

3-6 BAW-10231P-A, Revision 1, COPERNIC Fuel Rod Design Computer

Code, Framatome ANP, January 2004.

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Provided in Appendix B are sample problems for a Westinghouse 4-loop PWR design, a

Westinghouse 3-Loop PWR design, and a CE 2x4 PWR design. Table B-9, Table B-16,

and Table B-23 provide values for the most important NPP parameters. As illustrated, a

major difference in the important NPP parameters is the accumulator pressure for the

Westinghouse and AREVA designs, and the SITs in the CE designs. The impact of this

difference is shown in the sequence of events given in Table B-14, Table B-23, and

Table B-33, where the SIT flow initiation is delayed in the CE design until the pressure

in the cold legs drops below the SIT pressure. Taking into account this delay in the SIT

delivery, the sequence of events is similar for all three of the NPP types.

4.2 Figures of Merit

The figures of merit for the LOCA EM are derived from the first three acceptance criteria

of 10 CFR 50.46, as presented in Section 3.0. They are Peak Clad Temperature (PCT),

Maximum Local Oxidation (MLO) and Core-Wide Oxidation (CWO).

Complementary figures of merit are defined in the EM assessment process

(Section 8.0) for various benchmarks where the modeling constraints do not make it

possible to supply directly one of the principal figures of merit to match the physical

configuration or where the alternate figure of merit makes for an easier physical

interpretation. For instance, the ratio of the convective heat transfer coefficient to the

global heat transfer at the time of PCT is used as figure of merit in one of the FLECHT-

SEASET tests. If such alternate figures of merit are used, they are identified throughout

the report where appropriate.

4.3 References 4-1 NUREG-0800, “Standard Review Plan, Section 15.6.5 Loss-Of-Coolant

Accidents Resulting from Spectrum of Postulated Piping Breaks within

the Reactor Coolant Pressure Boundary,” Revision 3, U.S. NRC, March

2007.

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Table 7-1 Models Added to S-RELAP5 from COPERNIC

7.1.3 References 7-1. BAW-10231P-A, Revision 1, COPERNIC Fuel Rod Design Computer

Code, Framatome ANP, January 2004.

7-2. EMF-2102(P) Revision 1, S-RELAP5 Code Verification and Validation,

November 2010.

7-3. EMF-2100(P) Revision 16, S-RELAP5 Models and Correlations Code

Manual, December 2011.

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7.2 Provision of Complete Code Documentation

The documentation for the codes used in the development of this methodology is

provided in Reference 7-1 for COPERNIC, References 7-2, 7-3, 7-4, and 7-5 for the S-

RELAP5 code, and References 7-8 and 7-9 for the ICECON code. The documentation

describes the models and correlations used in the codes; defines the code inputs, and

provides a description of the code structure. These documents were verified against

the actual coding to ensure the documentation and coding are consistent (Section 8.0).

Revision 3 includes all the supporting documentation from these references that

pertains to the RLBLOCA EM in Section 7.0, including a comprehensive description of

the models and correlations.

The code validation for RLBLOCA from Reference 7-2, which compares the code

predictions to measured data in a number of SET and IET facilities, is included in

Section 8.0. All the benchmarks are identified in the assessment matrix (Section 6.0,

Table 6-1). In addition, AREVA NP has guidelines covering the development of

S-RELAP5 input for the NPP model and procedures for performing an actual analysis.

The input development and analysis guidelines are part of the EM implementation

description and are included in Section 9.0.

7.3 Determination of Code Applicability

The objective of the determination and code applicability step of CSAU is to

demonstrate that the selected codes are capable of modeling the chosen event for all

NPP types. This is accomplished by comparing the event and important phenomena

identified in the PIRT with the models and correlations documents for the selected

codes. Four attributes are needed to make this comparison:

• Field equations that provide code capability to address global processes.

• Closure (constitutive) equations that support the conservation equations by providing

code capability to model and scale specific phenomena or processes.

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• Code numerics that demonstrate code capability to perform calculations efficiently

and reliably.

• Structure and nodalization that address code capability to model the NPP geometry

and components, and to provide efficient and accurate NPP predictions.

These four attributes are discussed in the following sections.

7.3.1 Field Equations

The field equations (conservation of mass, momentum, and energy) must possess the

capability of simulating each of the distinct phases (blowdown, refill, and reflood) of a

LBLOCA. During the refill and reflood phases, countercurrent flow occurs at various

locations in the RCS, and subcooled liquid coexists with superheated steam in parts of

the reactor core. Therefore, for realistic analyses, the field equations must be non-

homogeneous (unequal velocity for each phase) and non-equilibrium (unequal

temperature for each phase). The presence of nitrogen in the accumulator requires an

additional field equation to model and track the movement of this noncondensable gas.

The required field equations are given in Table 7-2. The relationships to specific PIRT-

important phenomena along with references to specific models are provided in Table

7-3. As indicated in Table 7-2 and Table 7-3, the S-RELAP5 code has the required field

equations and models to address the important LBLOCA phenomena. A detailed

discussion of the fluid field equations and their numerical solutions is provided in

Section 7.4.

7.3.2 Closure Equations

Closure equations (constitutive models and correlations) are required to support the

basic field equations. The closure equations are essential for modeling the processes

and phenomena given in the PIRT (see Table 5-1). The S-RELAP5 constitutive models

and correlations are presented in Reference 7-3. A detailed discussion of the

constitutive models and correlations included in the EM is provided in Section 7.4. The

verification and validation of the code models and correlations are given in Reference 7-

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2 and reiterated in Section 8.0. These two components of the present document

together demonstrate that the S-RELAP5 code adequately simulates LBLOCA events

with a high level of confidence

The capability of the S-RELAP5 code closure equations to meet the requirements of the

PIRT (see Table 5-1) is summarized in Table 7-14. The closure equations address wall

friction, interphase friction, mass transfer (interphase heat transfer), wall-to-fluid heat

transfer, form-losses, and similar functions. The various models require flow regime

maps, boiling curves, state relationships, and fluid and material properties for

completeness. As indicated in Table 7-2, the S-RELAP5 code has the required closure

equations to address the important LBLOCA phenomena.

Table 7-2 Field Equations/Models in S-RELAP5

Scenario and PIRT Requirements

S-RELAP5 Model Existence Field Equations/Model

Non-equilibrium Two-phase Flow Yes Six equation unequal velocity, unequal

temperature Non-condensable Gas Flow Yes Gas mass balance in vapor flow field Multi-D Flow Capability Yes 2-D components available as required

Separation Due to Gravity Yes Gravity pressure differential in flow field equations

Interphase Exchange Terms Yes Mass and energy transfer between phases, vaporization, and condensation

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7.5.5 Wall Friction

The wall friction model computes the friction terms, αfρfFWFvf and αgρgFWGvg, in the

liquid and vapor momentum equations (see Equations (7.3) and (7.4)). In S-RELAP5,

the friction model is essentially the same as that of RELAP5/MOD2 (Reference 7-6

and 7-69), except that the RELAP5/MOD2 approximation to the Colebrook equation of

friction factor (Reference 7-92) is replaced by an explicit formula developed by Jain

(Reference 7-93).

The wall friction model consists of two main parts: (1) computing the overall two-phase

wall-friction pressure drop, and (2) apportioning the total wall friction into liquid and

vapor components. The two-phase friction multiplier approach, with the two-phase

multiplier calculated from the Heat Transfer and Fluid Flow Service (HTFS) modified

Baroczy correlation (Reference 7-94), is used to obtain the total wall-friction pressure

drop, which is independent of flow regimes. The phasic friction factor model, from

which Chisholm (Reference 7-95) developed a theoretical basis for the Lockhart-

Martinelli friction correlation (Reference 7-96), is used to develop the phasic partition

factors, which depend on flow regimes.

According to the two-phase friction multiplier approach, the overall wall-friction pressure

drop can be written in terms of the liquid-alone wall-friction pressure drop as

φ

∂ ∂= φ∂ ∂

2f

2 f

P P x x

(7.448)

or the vapor-alone wall-friction pressure drop

2g

2 g

P P x xφ

∂ ∂= φ∂ ∂

(7.449)

Here φ f and φ g are, respectively, the liquid-alone and vapor-alone friction multipliers.

The liquid- and vapor-alone friction pressure gradients are

2 22 2

ggg gfff f

f g

vP Pv , x 2D x 2D

′λ ρ′λ α∂ ρ ∂α= =∂ ∂

(7.450)

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Figure 7-10 Diagram of pre-CHF Heat Transfer Correlations Selection Logic

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Figure 7-11 Diagram of Critical Heat Flux Correlation Selection Logic

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7.6.5 Single-Phase Vapor

7.6.5.1 Single-Phase Vapor for Non-Core Heat Structures

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Figure 7-12 Diagram of Single Phase Vapor Heat Transfer Correlation Selection Logic

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Figure 7-13 Diagram of Transition Boiling Heat Transfer Calculation Logic

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Figure 7-14 Diagram of Film Boiling Heat Transfer Calculation Logic for Non-Core Heat Structures

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Thus, for the bundle model, the convective heat transfer coefficient, hcg, is given by

, ,nc ,wh ,db( , , , )cg cg lam cg cg cgh MAX h h h h Fψ ψ= (7.526)

where the laminar region is

h

g0.3333glamcg, D

k7.86Prh = (7.527)

and, as previously mentioned, the natural convection is given by Holman and is not

repeated here. From the Wong-Hochreiter correlation, the low Reynolds number region

that was fitted to steam cooling data is used:

g0.6774 0.3333cg,wh g g

hcf

kh 0.0797Re Pr T

D= (7.528)

The Dittus-Boelter correlation is modified to account for variable properties across the

boundary layer when large temperature differences exist:

g0.8 0.3333cg,db g cg

hf

kh 0.023Re Pr T

D= (7.529)

The form used is by Sleicher-Rouse and only considers the situation where the wall is

heating the steam. The temperature correction factor cfT has the same formulation as

given for the Sleicher-Rouse correlation and presented in Equation (7.508). Although

the turbulent two-phase enhancement by Drucker and Dhir was developed for wet and

dry walls, only the dry-wall form is used:

( ) 2g

3

gfg

2g

flimit

DgGr

ReGr3.25,1UMIN

−=

+=

(7.530)

Also, its influence is restricted by Ulimit to a maximum value of 5.0 and its affect is

diminished linearly as void fraction varies from 0.90 to 0.70, inclusive.

Finally, two optional grid spacer effects and a laminar flow heat transfer enhancement

are given by a three factor multiplier to the convective heat transfer. The first two

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factors, the optional grid spacer effects, are given by Yao, Hochreiter, and Leech

(Reference 7-185), while the third factor is Reynolds Number dependent and was

developed by Meholic et al. (Reference 7-186) using recent high void low flow data from

the Rod Bundle Heat Transfer (RBHT) facility at Pennsylvania State University. The

convective enhancement is given by:

1 2 3F FF F= (7.531)

where

x0.132 D

1F 1 5.55 e−

= + ε (7.532)

0.4x0.0342 2 D

2F 1 tan e−

= + β ϕ (7.533)

(7.534)

and where

ε is the blockage ratio of the spacer to flow channel when looking from upstream,

x is the axial distance from the downstream end of the spacer,

D is the hydraulic diameter of the flow channel;

is the fraction of area of the vane to the flow cross-section, viewing from

upstream,

ϕ is the angle of swirling vane with respect to axial flow direction.

The factors F1 and F2 are user input, while F3 is implemented into the S-RELAP5

coding and active for all calculations.

In the bundle model, the transition between dispersed flow film boiling and steady film

boiling is made by using a quadratic or cubic exponential decay to the Modified Bromley

<≤≤

<= −

g

g0.2788

g

g

3

Re54501.05450Re736Re11.008

736Re1.75F

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Figure 7-15 Diagram of Film Boiling Heat Transfer Calculation Logic for Active Core Heat Structures

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7-107. BAW-10231P-A, Revision 1, COPERNIC Fuel Rod Design Computer

Code, Framatome ANP, January 2004.

7-108. J. V. Cathcart, et al., Reaction Rate Studies, IV, Zirconium Metal-Water

Oxidation Kinetics, ORNL/NUREG-17, August 1977.

7-109. Fuel Material Manual Basic Properties of M5 Cladding Tube, FTG

Document Number 38-1287806-00, Internal Identification Number TFJC

DC 1241 E0.

7-110. D. L. Hagrman, G. A. Reymann, and R. E. Mason, MATPRO-11

(Revision 2) A Handbook of Material Properties for Use in the Analysis

of Light Water Reactor Fuel Rod Behavior, NUREG-CR-0497,

TREE-1280, Rev. 2, EG&G Idaho, Inc., Idaho Falls, ID, August 1981.

7-111. EMF-CC-39(P), Revision 3, ICECON: A Computer Program Used to

Calculate Containment Back Pressure for LOCA Analysis (Including Ice

Condenser Plants), Siemens Power Corporation, July 2004.

7-112. “Minimum Containment Pressure Model for PWR ECCS Performance

Evaluation,” U.S. Nuclear Regulatory Commission Branch Technical

Position, CSB 6-2, March 2007.

7-113. Pump Two-Phase Performance Program, EPRI NP-1556, Volumes 1-8,

September 1980.

7-114. The RELAP5 Development Team, RELAP5/MOD3 Code Manual

Volume 1: Code Structure, System Models, and Solution Methods,

NUREG/CR-5535, INEL-95/0174, June, 1995.

7-115. EG&G Idaho, Inc., RELAP4/MOD6: A Computer Program for Transient

Thermal-Hydraulic Analysis of Nuclear Reactors and Related Systems,

Users Manual, CDAP-TR-003, May 1978.

7-116. W. Kastner and G. J. Seeberger, “Pump Behavior and its Impact on a

Loss-of-Coolant Accident in a Pressurized Water Reactor,” Nuclear

Technology, Volume 60, February 1983, pp. 268-277.

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7.7 Hydrodynamic Process Models

Certain models in RELAP5 have been developed to simulate special hydrodynamic

processes. These models are presented in the following subsections.

7.7.1 Choked Flow

The choked-flow model is used to predict if the flow is choked at a break or nozzle and,

if it is, to establish the discharge boundary condition. Generally, the flow at the break or

nozzle is choked until the system pressure nears the containment pressure. In addition,

the choked-flow model can be used to predict existence of and calculate choked flow at

internal points in the system.

7.7.1.1 Choked Flow Theory

The RELAP5 choked flow model is based on the choking theory developed by Ransom

and Trapp (Reference 7-118). Choking is defined as the condition wherein the mass

flow rate becomes independent of the downstream conditions (that point at which

further reduction in the downstream pressure does not change the mass flow rate). The

fundamental reason that choking occurs is that acoustic signals can no longer

propagate upstream. This occurs when the fluid velocity equals or exceeds the

propagation velocity. The choked-flow model is based on a definition that is established

by a characteristic analysis using time-dependent differential equations.

Consider a system of n first-order, quasi-linear, partial differential equations of the form

A U Ut + B U U

x + C U = 0 (7.569)

The characteristic directions (or characteristic velocities) of the system are defined

(References 7-119 and 7-120) as the roots4, λi (i ≤ n), of the characteristic polynomial

( )A B 0λ − = (7.570)

4 The number n is the number of differential equations comprising the system defined by

Equation (7.569), and the number i designates any of the corresponding n roots.

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7.9.3.2 COPERNIC Gap Conductance Model

7.9.3.3 Clad Ballooning, Rupture and Area Adjustment Models

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Table 7-10 Packing Factors and Sources

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7.9.3.4 Material Thermal Properties

7.9.3.5 Zirconium/Steam Reaction Kinetics

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7.9.3.6.6 M5 Clad Radial Ballooning and Rupture Deformations

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In one-dimensional problems, boundary conditions are applied on the left and right

surfaces. In steady-state problems, a valid physical problem requires that A be nonzero

on at least one of the two boundary surfaces. If a transient or steady-state problem has

cylindrical or spherical geometry and a zero radius for the left surface (that is, a solid

cylinder or sphere), the left boundary condition is normally the symmetry condition,

0nT =

∂∂ . Under these conditions, if B is nonzero, the numerical technique forces the

symmetry boundary condition, even if it is not specified.

7.10.2 Mesh Point and Thermal Property Layout

Figure 7-43 illustrates the placement of mesh points at which temperatures are to be

calculated. The mesh point spacing for a rectangular problem is taken in the positive x

direction. For cylindrical and spherical problems, the mesh point spacing is in the

positive radial direction. Mesh points are placed on the external boundaries of the

problem, at the interfaces between different materials, and at desired intervals between

the interfaces, boundaries, or both.

Figure 7-43 Mesh Point Layout

Figure 7-44 represents three typical mesh points. The subscripts are space indexes

indicating the mesh point number; and l and r (if present) designate quantities to the left

and right, respectively, of the mesh point. The δs indicate mesh point spacings that are

not necessarily equal. Between mesh points, the thermal properties, k and ρ, and the

source term, S, are assumed spatially constant; but klm is not necessarily equal to krm

and similarly for ρ and S.

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8.0 ASSESSMENT RESULTS

8.1 Assessment Methodology Overview

8.1.1 Introduction

This section describes the verification and validation performed for the S-RELAP5 code.

The material included originates from the verification and validation report

(Reference 8.1-1), which is intended to be a generic compilation of assessments

supporting all S-RELAP5 based methodologies. The code assessments from

Reference 8.1-1 applicable to the RLBLOCA methodology are those discussed in

Section 6.0 and listed in the assessment matrix, Table 6-2. These assessments were

chosen to address the important PIRT phenomena identified in Table 5-1. The cross

correlation between assessments and PIRT phenomena is provided in Table 6-2.

Additionally, some assessments were chosen to address issues of code scalability;

these assessments, and the discussion with respect to scalability, are provided in

Section 8.6.

Appendix A discusses the appropriate nodalization to represent PWR system

components. The nodalization used in the assessments must be consistent with the

large-scale plant nodalization in the regions where the phenomena are being assessed

in order for the assessment results to apply to large scale PWRs. AREVA NP used the

plant nodalization and the RLBLOCA S-RELAP5 input guidelines as described in

Appendix A to the extent possible to derive assessment nodalizations, which are

consistent with the PWR application nodalization. Unique features of small-scale

facilities can require deviations from the guidelines. The assessment nodalizations are

generally consistent with the plant application, and where deviations were made, the

reasons for the deviations and the effects on results are discussed.

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The reported assessment results support the more detailed requirements of this

procedure because the AREVA NP Realistic LBLOCA methodology follows the Code

Scaling, Applicability, and Uncertainty (CSAU) methodology. The assessments were

chosen to address the important PIRT phenomena identified in Table 5-1. The cross

correlation between assessments and PIRT phenomena is provided in Table 6-2.

Consequently, the documented assessment results should provide quantitative

statements about the code’s capability to predict key parameters. NUREG 1737

Appendix C (Reference 8.1-4) suggests how this could be done by defining acceptance

criteria associated with levels of code-data agreement. In this document, five levels of

agreement are defined: (1) excellent agreement, (2) good agreement, (3) reasonable

agreement, (4) acceptable agreement, and (5) insufficient agreement. These criteria

move from the most desirable correlation of data to the code, to the unacceptable

prediction of the data by the code, and are defined as follows:

• Excellent Agreement: The calculation of major phenomena, with rare exception, will

lie within the scatter of the data including data uncertainty and any known data

biases. The trends in the major phenomena also will therefore be predicted within

the uncertainty of the data. For this criterion, the code-to-data comparison could be

represented with an uncertainty and perhaps a small bias.

• Good Agreement: While the correct trends are predicted, the calculation of major

phenomena will frequently lie outside the scatter of the data including data

uncertainty and any known data biases. However, the correct conclusions about

trends and phenomena would be reached if the code were used in similar

applications. For the code to demonstrate good agreement, the calculation of major

phenomena should, on average, not differ from the data by more than plus or minus

10%. For this criterion, the code-to-data comparison may be representable with a

bias and uncertainty. For application wherein the code is conservative for a specific

phenomenon, the code results could be used directly without determining a bias or

uncertainty for that phenomenon.

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In the rapid-cooling period, [

]

The FLECHT SEASET and FLECHT Skewed comparisons shown in this section

demonstrate the capacity of the S-RELAP5 code to predict LBLOCA reflood

phenomena. Data from these tests are combined with other reflood test data to

establish the bias and uncertainties to be used for RLBLOCA as given in Section 8.4.1.

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8.2.4.4 S-RELAP5 Model Description

The input model is based on the standard input model (discussed in Section 8.2.3.5 and

is shown in Figure 8.2-51,) [

] The initial conditions of pressure, inlet flow

rate, inlet flow temperature and power were taken from Reference 8.2-22, and are listed

in Table 8.2-10.

8.2.4.5 Discussion of Results

The cases were run to steady conditions, as shown in Figure 8.2-143. The resulting

temperature comparisons are shown in Figure 8.2-144. [

]

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For this analysis, [

] given in Section 8.4.1.3.

8.2.5.4 Discussion of Results

The amount of [

] Test 31504, documented in NUREG/CR-2256 (Reference 8.2-26). In that analysis, the

rod temperature distribution, steam temperature, component temperatures and liquid

droplet concentrations from measured data were used to compute the total effective

heat transfer from reverse conduction and the total effective radiation heat flux. From

that analysis, the researchers presented the ratio of convective heat transfer to total

heat transfer at 20 s intervals from 80 to 200 s.

The [

] Consequently, the ratio of convective heat transfer to total heat transfer at 100 s (time of

PCT) was determined to be the figure of merit for comparison. From preliminary

sensitivity studies, [

] The results from Test

31504 are shown in Figure 8.2-145. The data were estimated from Figure 6-12 in

Reference 8.2-26.

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Test data are available to establish a pressure boundary condition either at the lowest

part of the water separator, from sensor JEA05CP001, or at a site between the water

separator and the downcomer, from sensor JEC04CP21 (Reference [8.2-39]). [

]

Core Simulator Steam Injection

Component 660 serves as the source for steam injected into the core simulator. Test

data obtained from system pressure and temperature sensors show the steam in the

delivery lines to the core steam injection nozzles to be very nearly saturated, hence the

thermodynamic state of the steam in component 660 was specified to be saturated

steam with a specified temperature history.

Lower Plenum Drain Flow Boundary Condition

The lower plenum drain flow was activated only for simulation of Run 203 of Test 7, in

which drain flow from the bottom of the lower plenum was used to maintain the lower

plenum liquid level at an acceptable elevation. In the S-RELAP5 model, a time-

dependent junction was used to simulate the pumped drain flow. The flow rates

measured during the test was used to specify this boundary condition.

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Figure 8.2-280 through Figure 8.2-282 compare the S-RELAP5 predictions to the UPTF

experimental results for Test 10 Run 080. Figure 8.2-280 presents a plot of Kutateladze

parameters calculated from the S-RELAP5 results compared to the UPTF correlation.

The comparison shows that S-RELAP5 is correctly limiting liquid downflow as is shown

by the linear upper limit of . This is based [

] specified in the RLBLOCA guidelines (Appendix A). Figure

8.2-280 clearly shows that the S-RELAP5 calculation is conservative relative to the

UPTF correlation.

Figure 8.2-281 compares the S-RELAP5 upper plenum pressure calculation to the

measured UPTF upper plenum pressure. The two are shown to be in good agreement.

Figure 8.2-282 compares the S-RELAP5 calculated downflow to downflow calculated

using UPTF test data. The downflow in the UPTF test was derived from data sensors

for the test vessel level, the downcomer level, and the lower plenum drain rate. The

comparison shows that the overall downflow (countercurrent and cocurrent) calculated

by S-RELAP5 trends the UPTF data, but conservatively underpredicts the data

(adequate agreement). This prediction is based on a core to upper plenum junction

specification consistent with the UTP.

Figure 8.2-294 through Figure 8.2-296 compare the S-RELAP5 predictions to the UPTF

experimental results for Test 12 Run 014. Figure 8.2-294 presents a plot of Kutateladze

parameters calculated from the S-RELAP5 results compared to the UPTF correlation.

The comparison shows that S-RELAP5 is correctly limiting liquid downflow as is shown

by the linear upper limit of . The S-RELAP5 calculation is based [

] specified in the RLBLOCA guidelines

(Appendix A).

Figure 8.2-294 clearly shows that the S-RELAP5 calculation is conservative relative to

the UPTF correlation.

*gK

*gK

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8.2.9.6.2 Test Description (UPTF Tests 10B and 29B)

For Tests 10B and 29B, the UPTF system was configured to simulate the reflood phase

of a cold leg break PWR LBLOCA. For these tests, the lower plenum and lower

downcomer were filled with water to block steam flow directly from the core to the

downcomer and cold legs. A mixture of steam and water was injected into the core

simulator to simulate reflood steam generation and water entrainment. The injected

steam and entrained water then flowed to the hot legs via the upper core support plate

and upper plenum. From the hot legs, the steam/water mixture flowed into the steam

generator simulator inlet plenum and from there to the cyclone separators where water

was separated from the mixture. The separated water was stored and measured in

holding tanks, while the steam (and any unseparated water droplets) flowed onward

through the pump simulators, intact cold legs, upper downcomer and broken cold leg,

and flowed out the break into the containment simulator. Each test consisted of a

sequence of phases using different steam and water injection rates. Test 10B was a

300 second transient consisting of four different flow phases. The conditions for the

four phases of this test are given in Table 8.2-22 .

Test 29B Runs 211 and 212 were 900 second transients consisting of six different flow

phases. Each phase consisted of a period of constant steam and water flow rates,

followed by a period of no flow. The first two phases of Run 211 and last three phases

of Run 212 were flawed. Consequently, the S-RELAP5 predictions will be compared to

Run 212 data from Phases 1 and 2 (0 through 300 seconds), and Run 211 data from

Phases 3 through 6 (300 through 900 seconds). The test parameters for the six phases

in combined Run 212/211 are shown in Table 8.2-23: Test Phase Parameters for Test

29B.

The specific LBLOCA reflood phenomena addressed by UPTF Tests 10B and 29B

benchmarks are:

• Steam generator steam binding

• Upper plenum two-phase flow

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Test S2-AC1 differs from Test S2-SH1 in the accumulator core coolant (ACC) injection

rate and duration.

The S2-10 test is the SCTF-II forced-flood base case. In Test S2-10, ECC was injected

into the lower plenum only, with no hydraulic communication between the lower plenum

and the downcomer. The ECC injection rate was specified to match the core inlet flow

rate achieved in the gravity feed test.

Test S2-11 differs from S2-10 in that a high ACC flow rate was used.

• Effect of Radial Nodalization (Phase II). In this assessment phase, two tests were

chosen to study the effect of radial nodalization and radial power split on reflooding

behavior. These tests are S2-17 and S2-18. Test S2-17 represents a flat radial

core power distribution. Two radial nodalizations were used for Test S2-18: a

nominal nodalization, and [ ] See

Section 8.2.12.4.6 for additional details.

Table 8.2-23 shows the test conditions for each of the test configurations examined.

8.2.12.4 S-RELAP5 Model Description

Figure 8.2-335 is a schematic of the nodalization used for the SCTF-II facility. The

nodalization scheme and the model flags were selected to be consistent with the

S-RELAP5 RLBLOCA, Revision 3 methodology. The modeling of the downcomer,

lower plenum, core, upper plenum, and upper head is summarized here.

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• [

]

To study the radial power distribution effect, Test S2-18 was also simulated using [

]

Specific modeling changes implemented to the S-RELAP5 test models are discussed in

the following sections. The heat structure modeling scheme representing the different

bundle configurations is summarized in Table 8.2-34.

8.2.12.4.1 Test S2-11 Input Model

• The initial temperature distribution for the HR region is approximated by the initial

temperature data measurements, which are shown in Table 8.2-35. All other heat

structures are assigned an initial temperature of 411 K (average core wall

temperature, Reference 8.2-53, Table 8.2-35).

• The ECC injection is via the lower plenum. The injection flow rate and temperature

are given in Table 8.2-36: Injection started at 122.5 seconds and terminated at 155.0

seconds. It then was switched to low pressure coolant injection (LPCI) at 155.0

seconds, which lasted to the end of the transient.

• Initial water level in the lower plenum is 1.58 m. The lower plenum initial conditions

are shown in Table 8.2-37.

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Each heated rod consisted of a nichrome heating element, magnesium oxide (MgO)

and boron nitride (BN) insulators, and an Inconel-600 sheath (clad), as shown in Figure

8.2-436. The heated length, the outer diameter (O.D.) and the rod pitch were,

respectively, 3.66 m (144 inch), 10.7 mm (0.422 inch) and 14.3 mm (0.563 inch), which

are identical to the corresponding dimensions of actual PWR fuel rods. The heating

element was a helical coil with a varying pitch to generate a 17-step chopped cosine

axial power profile with a peaking factor of 1.40, as shown in Figure 8.2-437.

The unheated rods were either stainless steel pipes or solid bars of 13.8 mm (0.543

inch) O.D. All pipes were used for installation of instruments and all bars were used for

carrying the assembly loads. The heater rods and the unheated rods were held in radial

position by grid spacers that were located at six elevations along the axial length, as

shown in Figure 8.2-437. ECCS consisted of an ACC (Accumulator) and LPCI (Low

Pressure Coolant Injection). The injection points are at each cold leg and at the lower

plenum. The upper plenum and downcomer injection system were available for other

alternative ECCS tests.

The instrumentation was divided into two groups: the USNRC-supplied instruments, and

the JAERI-supplied instruments. The USNRC-supplied instruments were the advanced

instrumentation for the two-phase flow measurement. The JAERI-supplied instruments

measured the temperatures, absolute pressures, differential pressures, water levels,

and flow rates. Examples of instrumentation arrangements and notations are shown in

Figure 8.2-438 for temperature measurements in the core and lower plenum, and in

Figure 8.2-439 for differential pressure measurements in the vessel.

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The heater rod surface temperatures were measured at several elevations in various

bundles (see Figure 8.2-438). The maximum surface temperatures from all

measurements in the high-power core region are plotted in Figure 8.2-469, along with

the calculated peak temperatures as a function of elevation. The spread of measured

temperature rises in the range of 10 K to 70 K depending on the elevation and the

number of thermocouples. (Note that a single data point means that only one

measurement was taken at that elevation). The calculated points are seen to be

reasonably close to the data range. This shows that the code-calculated maximum

temperature behavior, including PCT, slightly underpredict the experiment.

8.2.13.5.6 Test Run 54 Discussion of Results

According to the PIRT (Table 5-1), the most important reflood phenomena are: core

post-CHF and reflood heat transfer, vapor generation/distribution and entrainment/de-

entrainment in the core, entrainment/de-entrainment in the upper plenum and in the hot

legs, steam binding in steam generator, pump p, hot wall effects in the downcomer

and the lower plenum, non-condensable gases in cold leg/accumulator, loop flow

oscillations, decay heat, and oxidation and gas conductance for fuel rods.

Except for the fuel rod oxidation and conductance and cold leg/accumulator non-

condensable gases, all the important reflood phenomena were observed in the CCTF

Test Run 54 and were calculated reasonably well by S-RELAP5.

As mentioned in previous sections, CCTF Test Run 54 power decay simulated ANS

Standard 1.0 plus actinide decay heat. This power decay curve was treated in

S-RELAP5 by tabulated input values (see Figure 8.2-474). The pump/loop p was

simulated in the test by an orifice plate and in S-RELAP5 by a form loss factor . Figure

8.2-452 shows that the pump/loop p is well calculated.

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With the chopped-cosine axial power profile, the peak temperature rise is expected to

occur in the neighborhood just above the mid-plane. Figure 8.2-492 and Figure 8.2-493

show the calculated rod surface temperatures at 1.83 m and 2.035 m elevations, and

the data with the highest temperature rise for the high-power core. The measured PCT

of 1132 K (1578 °F) was recorded by the thermocouple at 1.83 m elevation in bundle

number 30 (TE30Y37). The maximum surface temperature calculated at 1.83 m is

1106 K (1531 °F), while the calculated PCT is 1116 K (1548 °F) at 2.235 m elevation,

showing good agreement with the data. Thus, the calculated temperature rise is higher

above the mid-plane, and the PCT point is shifted to a higher elevation. The PCT time

is 154 seconds from the measurement and is at 235 seconds in the calculation. The

difference is attributable to the observed top-down cooling proceeding at a greater rate

than the calculation.

The heater rod surface temperatures were measured at several elevations in various

bundles (see Figure 8.2-438). The maximum surface temperatures from all

measurements in the high-power core region are plotted along with the calculated peak

temperatures as a function of elevation in Figure 8.2-499. The spread of measured

temperature rises is in the range of 10 K to 70 K, depending on the elevation and the

number of thermocouples. (Note that a single data point means that only one

measurement was taken at that elevation). The calculated points are seen to be at the

high end of the data range or above it. This shows that the code calculated maximum

temperature behavior, including PCT, is in good agreement with the experimental

behavior.

8.2.13.6.6 Test Run 62 Discussion of Results

According to the PIRT (Table 5-1), the most important reflood phenomena are core

reflood heat transfer, void generation/distribution and entrainment/de-entrainment in the

core, entrainment and de-entrainment in the upper plenum and in the hot legs, steam

binding in the steam generator, pump p, hot wall in the downcomer and lower plenum,

non-condensable gases in the cold leg/accumulator, loop flow oscillations, decay heat,

and oxidation and gas conductance for fuel rods.

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With the chopped-cosine axial power profile, the peak temperature rise is expected to

occur in the neighborhood just above the mid-plane. Figure 8.2-522 and Figure 8.2-523

show the calculated rod surface temperatures at 1.83 m and 2.035 m elevations, and

the data with the highest temperature rise for the high power core. The measured PCT

of 1143 K (1598 °F) was recorded by the thermocouple at the 1.83 m elevation. The

maximum surface temperature was calculated at 1.83 m is 1180 K (1664 °F). The

calculated PCT of 1238 K (1769 °F) occurs at the 2.645 m elevation. Thus, the

calculated temperature rise is higher above the mid-plane and the PCT point is shifted

to a higher elevation. The PCT time is 164 seconds from the measurement, and 385

seconds from the calculation. The difference is attributable to the elevation difference of

the PCT locations and over- and undercooling in the early reflood period.

The heater-rod surface temperatures were measured at several elevations in various

bundles (see Figure 8.2-438). The maximum surface temperatures from all

measurements in the high-power core region are plotted along with the calculated peak

temperatures, as a function of elevation in Figure 8.2-529. The calculated points are

seen to be at the high end of the data range or above it. This shows that the code

calculated maximum temperature behavior, including PCT, shows conservative

acceptable agreement with the experiment.

8.2.13.7.6 Test Run 67 Discussion of Results

The most important reflood phenomena, according to the PIRT (Table 5-1), are: core

reflood heat transfer, void generation/distribution and entrainment/ de entrainment in the

core, entrainment/de-entrainment in the upper plenum and in the hot legs, steam

binding in steam generator, pump p, hot wall in the downcomer and lower plenum,

noncondensable gases in the cold leg/accumulator, loop flow oscillations, decay heat,

and oxidation and gas conductance for fuel rods. Except for the fuel rod oxidation and

conductance and cold leg/accumulator noncondensable gases, all the important reflood

phenomena were observed in the CCTF Test Run 67, and generally were calculated

with reasonable or acceptable agreement by S-RELAP5.

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With the chopped-cosine axial power profile, the peak temperature rise is expected to

occur in the neighborhood just above the mid-plane. Figure 8.2-552 and Figure 8.2-553

show the calculated rod surface temperatures at 2.035 m and 1.83 m elevations, and

the data with the highest temperature rise for the high-power core. The measured PCT

of 1122 K (1560 °F) was recorded by the thermocouple at 1.83 m elevation. The

maximum surface temperature calculated at 1.83 m is 1081 K (1486 °F). The

calculated PCT of 1123 K (1561 °F) occurs at the 2.44 m elevation. Thus, the

calculated temperature rise is higher above the mid-plane and the PCT point is shifted

to a higher elevation than measured. The PCT time is 144 seconds from the

measurement and is 246 seconds from the calculation.

The heater rod surface temperatures were measured at several elevations in various

bundles (see Figure 8.2-438). The maximum surface temperatures from all

measurements in the high-power core region are plotted, along with the calculated peak

temperatures, as a function of elevation in Figure 8.2-559.

This comparison shows good agreement with the data for lower and mid-core elevations

and overprediction by the calculated values for higher coer elevation, consistent with the

behavior exhibited by S-RELAP5.

8.2.13.8.6 Test Run 68 Discussion of Results

The most important reflood phenomena according to the PIRT (Table 5-1) are core

reflood heat transfer, void generation/distribution and entrainment/de-entrainment in the

core, entrainment and de-entrainment in the upper plenum and in the hot legs, steam

binding in steam generator, pump p, hot wall in the downcomer and lower plenum,

noncondensable gases in the cold leg/accumulator, loop flow oscillations, decay heat,

and oxidation and gas conductance for fuel rods. Except for the fuel rod oxidation and

conductance and cold leg/accumulator noncondensable gases, all the important reflood

phenomena were observed in the CCTF Test Run 68, and the agreement between

S-RELAP5 calculated and measured data is reasonable to good.

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Figure 8.2-440 CCTF Facility Nodalization

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Figure 8.2-441 CCTF Vessel Nodalization

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Figure 8.2-442 CCTF Downcomer Azimuthal Nodalization

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Figure 8.2-443 CCTF Core Nodalization

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8.2.16.3 Facility Description

The Ferrell-McGee test loop is shown in Figure 8.2-614, and the adiabatic test section is

shown in Figure 8.2-615. Figure 8.2-615 shows the pressure tap locations, the location

of the abrupt area change, and lengths of the upper and lower pipes. The test was

executed by first establishing the desired flow rate, pre-heat the fluid to desired

subcooling, and apply power to the heated section (located upstream of the test section)

to establish the desired test conditions. The test loop was allowed to come up to test

conditions over a period of one hour before data was recorded.

8.2.16.4 S-RELAP5 Model Description

The Ferrell-McGee Test 2C-7 was simulated using S-RELAP5. The Revision 3

RLBLOCA input development guidelines (Section 9.0) were not followed since this

assessment validates S-RELAP5 two-phase frictional flow in general. The test

conditions were obtained from Reference 8.2-64 and are shown in Table 8.2-69. The

test section geometry modeled is shown in Table 8.2-70.

For the inlet conditions, a time dependent volume was set to a pressure of 122.163 psia

and quality of 0.0243 to approximate the station 1 pressure and void fraction from the

measurements (Reference 8.2-64), the flow area was set to the lower test section flow

area, and a time dependent junction was used to specify the inlet flow rates discussed

above. The test section was represented by the pipe component which was initialized

to saturated liquid at 118 psia. The outlet junction was set to the upper test section flow

area, as was the outlet volume. The outlet volume was initialized to saturated steam at

117.97 psia, to represent the test section outlet pressure.

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8.2.17.3 Facility Description

The Moby Dick test facility’s primary loop has five main components; a pump, a pre-

heater, a nitrogen injection system, a test section, and a condenser. A simplified

schematic of the facility as represented by the TRACE computer code is shown in

Figure 8.2-618. Flow is directed vertically upward in the test section. The outlet of the

vertical test section is located inside the condenser. Test flow conditions were obtained

by maintaining constant inlet conditions to the test section and lowering the downstream

pressure in the condenser to atmospheric pressure. Reference 8.2-66 describes the

Moby Dick facility and provides the test data for the experiments. Figure 8.2-619 is a

sketch of the actual Moby Dick facility, and Figure 8.2-620 depicts the vertical test tube,

both are extracted from Reference 8.2-66.

Nitrogen is injected into the pipe at a location 0.985 meters upstream of the expansion.

The gas is injected through four porous screens surrounding the flow pipe. Pressure

measurements are taken at the various positions labeled with P1-P22, on Figure B.2-2

of Reference 8.2-66. Water temperature was measured at the inlet to the test section,

and the temperature of the injected nitrogen was also measured. The accuracy of the

temperature measurement is indicated to be ±0.2 °C. The water mass flow rate was

also measured.

Table B.2.1 of Reference 8.2-66 provides the dimensions of the test section. These are

further shown here in Table 8.2-71.

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A steady-state initialization calculation was performed to reach the desired steady-state

conditions for initiating the LOCA calculation. Table 8.3-5 compares the calculated initial

conditions with the test conditions measured before break initiation. The calculated and

measured initial conditions agree within the measurement uncertainty. Table 8.3-5

shows that the desired steady-state conditions were achieved and that the calculation

reached the L2-5 test initial conditions.

Event setpoints and boundary conditions that have an impact after the start of the

transient portion of the Test L2-5 simulation are given in Table 8.3-6. [

]

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Initial and Boundary Conditions, Test LP-02-6

[

]

A steady-state initialization calculation was performed to reach the desired steady-state

conditions for initiating the LOCA calculation. Table 8.3-8 compares the calculated

initial conditions with the test conditions measured before break initiation. [

]

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The vessel in the MOD-3 system consists of the upper plenum with internals required to

represent guide and support tubes, upper head, 25-rod electrically heated core, and an

external single pipe downcomer. The active intact loop is scaled to represent three

loops of a PWR, and the active broken loop is scaled to represent a single loop of a

PWR. The intact loop contains a pump and the short Type I steam generator, and is

connected to the pressurizer. The broken loop contains the taller Type II steam

generator in addition to a pump and break simulators or rupture assemblies connected

to a blowdown suppression system. The blowdown suppression system simulates

containment pressure.

The 25-rod electrically heated core is characterized by fuel pin pitch (0.563 inch) and

outside diameter (0.42 inch) typical of a PWR. The heated length (12 ft) of the MOD-3

core is identical to the 4-loop PWR core.

Test S-07-1 was performed to establish the baseline performance of the MOD-3 system

during a blowdown with cold-leg ECC injections. It was conducted to obtain core heat

transfer and departure from nucleate boiling (DNB) characteristics of the heater rods.

The MOD-3 system was initialized in the experiment to a primary pressure of 15.9 MPa,

total-loop flow of 9.4 kg/s and cold-leg temperatures of 559 K for the intact loop and 557

K for the broken loop at a core power level of 2.027 MW nominal (see Table 8.3-16 and

Reference 8.3-21). The system was subjected to a double-ended cold-leg break

through a rupture assembly and two non-communicative nozzles (Reference 8.3-22).

8.3.2.8 S-RELAP5 Model Description for Semiscale MOD-3

The Semiscale MOD-1 model described here was used as the starting point for the

development of the MOD-3 input model. The features of the MOD-1 and MOD-3

facilities are compared in Table 8.3-17.

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8.3-5 EG&G Idaho, Inc., OECD/LOFT-T-3504, Quick-Look Report on OECD

LOFT Experiment LP-LB-1, February 1984.

8.3-6 Idaho National Engineering Laboratory, NUREG/CR-0247, TREE-

1208, LOFT System and Test Description (5.5 ft Nuclear Core 1

LOCEs), July 1978.

8.3-7 ASME, 74-WA/HT-53, Examination of LOFT Scaling, 1974.

8.3-8 EG&G Idaho, Inc., NUREG/CR-3214, Summary of Nuclear Regulatory

Commission’s LOFT Program Experiments, July 1983.

8.3-9 EG&G Idaho, Inc., NUREG/CR-3005, Summary of Nuclear Regulatory

Commission’s LOFT Program Research Findings, June 1983.

8.3-10 J. P. Adams et al., Influence of LOFT PWR Transient Simulations on

Thermal-Hydraulic Aspects of Commercial PWR Safety, Nuclear

Safety, 27-2:179–192, April 1986.

8.3-11 EG&G Idaho Inc., OECD/LOFT-T-3907, An Account of the OECD

LOFT Project, May 1990.

8.3-12 AlChE, Vol. 70 No. 138, A Correlation For The Minimum Film Boiling

Temperature.

8.3-13 Nuclear Regulatory Commission, NUREG/CR-6061, Determination of

the Bias in LOFT Fuel Peak Cladding Temperature Data from the

Blowdown Phase of Large-Break LOCA Experiments, May 1993.

8.3-14 BAW-10231P-A, Revision 1, COPERNIC Fuel Rod Design Computer

Code, Framatome ANP, January 2004.

8.3-15 AREVA NP, 2A4-COPERNIC, Fuel Rod Computer Code; User’s

Manual, November 2004.

8.3-16 Siemens Power Corporation, ECJ-83-87, Transmittal of Requested

Input Decks. August 1987, Memo to H. Chow on August 6th, 1987.

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Table 8.4-1: Summary of Evaluated Biases and Uncertainties of Important Code Related PIRT Parameters

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8.4.1 Film Boiling Heat Transfer

Film boiling heat transfer is significant to the outcome of the RLBLOCA methodology.

The assessment of the film boiling heat transfer and its uncertainty are based on the

separate effects experiment: Full-Length Emergency Cooling Heat Transfer, System

Effects, and Separate Effects Tests (FLECHT-SEASET) (Section 8.2.3). The FLECHT-

SEASET tests were chosen because they exclusively represent low pressure industry-

accepted and -evaluated experiments applicable to pressurized water reactor (PWR)

post-critical heat flux (CHF) heat transfer. Test results from the Oak Ridge National

Laboratory/Thermal Hydraulic Test Facility (ORNL/THTF) (Section 8.2.1) are used for

confirmation purposes.

In Reference 8.4-1, the complete process of the calculation of the film boiling heat

transfer coefficient bias, and uncertainty associated with the S-RELAP5 determinations

of the film boiling heat transfer coefficients, was presented and a condensed version is

presented later in this section.

The discussion presented in Section 8.4.1.2 shows the determination of the heat

transfer coefficient uncertainties and multipliers for FILMBL and DFFBHTC under low

pressure reflood conditions. The multiplier FILMBL is applied to the inverted annular

film boiling heat transfer regime, which is considered (in these discussions) as the low

void regime. The multiplier DFFBHTC is applied to the dispersed flow film boiling heat

transfer regime, which is considered as the high void regime. In the RLBLOCA plant

model, the values from the uncertainty determination [

] are applied to the hot assembly, high power region, average

power region, and the core periphery.

5 The FILMBL and DFFBHTC multipliers discussed above were developed using Rev.2 methodology. Modifications to the Rev. 2 methodology were added in the Rev. 3 methodology. The reevaluated multipliers are given in Appendix A.2.3.6.5.

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[

] The bundle is not effected by the rod-to-rod radiation

model (rod–to-rod is applied only to the hot rod which is a supplemental heat structure),

and its role is to give a best estimate of the coolant properties. [

]

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Table 8.4-5: Defining Distributions for DFFBHTC

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UPTF Tests 10, Run 081 (Test 10B), and 29, Runs 211 and 212 (Test 29B) were

analyzed to provide specific S-RELAP5 input modeling guidelines for the hot leg and

STGR inlet plenum regions to ensure that S-RELAP5 properly predicts the liquid

entrainment to steam generator (STGR) tube region, and to limit countercurrent flow at

the upper core tie plate (UTP) of a PWR during the LBLOCA reflood phase. These

tests were the separate effect tests specifically designed under the 2D/3D program to

investigate water mass distribution in the reactor vessel upper plenum, hot-legs, STGR

inlet plenums, and in the STGR tube region during the reflood phase of a LBLOCA

transient. UPTF Tests 10B and 29B, the S-RELAP5 input model and simulation results

were discussed in Section 8.2.9.6.

The simulation results of Tests 10B and 29B show that S-RELAP5, with the RLBLOCA

plant nodalization and the specific guidelines described below, conservatively over-

predicts the carryover of liquid to the steam generators.

7 [

]

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8.4.7 Tmin

The primary objective of this analysis is to quantify the bias (if present) and uncertainty

associated with the TMINK parameter of the S-RELAP5 code. The S-RELAP5 heat

transfer package does not have an explicit model for the minimum stable film boiling

temperature (Tmin) to determine the wall surface temperature at which transition from

film to nucleate boiling can begin. Instead, for the heat transfer regime to be considered

to be in transition boiling, the following criteria have to be met:

• The surface heat flux calculated using the transition boiling correlation must be

greater than that calculated for film boiling.

• The wall surface temperature must be less than the parameter TMINK.

• The void fraction must be less than 0.95 in the hydrodynamic volume.

In conclusion, the proposed value for TMINK was shown to contain a considerable

degree of conservatism because of its neglect of high pressure and use of stainless

steel data to represent zirconium alloys and zirconium oxide effects.

The overall approach used to determine these values is summarized as follows:

• Develop a distribution for the quench temperature using the low-pressure database.

• Assume that the minimum film boiling temperature is essentially equal to the quench

temperature.

• Develop a relationship between Tmin and the code parameter TMINK.

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As discussed in Reference 8.4-11, this low-pressure bounding distribution is clearly

conservative even when compared to the high-pressure data with the lowest values of

the quench temperature (ROSA/TPTF), and is highly conservative compared to the data

from the ORNL/THTF and the Westinghouse G1/G2 tests. Discussion in Reference

8.4-11 also concludes that a substantial degree of conservatism has been built into the

proposed distribution (relative to zirconium alloy including M5® clad rods) by basing it on

only low pressure data for stainless steel clad rods. These conclusions remain valid for

Revision 3.

A relationship between the effective value of Tmin and the specified value of the code

parameter TMINK was developed. These two not being equal was attributed to the

manner in which the boiling curve is evaluated for fine mesh nodes. A limiting normal

distribution was developed for the difference between Tmin and TMINK with a mean

value of 9 K and a standard deviation of 10 K.

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8.4.8 COPERNIC Uncertainties

8.4.8.1 Background

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8.4.9 Metal-Water Reaction

Two LOCA ECCS criteria limit the extent of local and core-wide reaction of the

zirconium-based cladding with water; therefore, metal-water reaction is an important

parameter to be calculated for LOCA analyses. For realistic calculation of PWR large

break LOCA, S-RELAP5 uses the metal-water reaction rate equation from Cathcart and

Pawel (Reference 8.4-13). Two constants in this rate equation are derived from

experimental data and have uncertainties as determined by the authors, the reaction

rate constant and a constant in the temperature dependent exponential term. These

uncertainties are shown in Table 8.4-1 and are applied in the realistic LOCA analyses.

The Cathcart-Pawel equation is applicable to all zirconium based cladding alloys

currently used, including M5® cladding.

8.4.10 Hot Wall (CHF Multiplier)

During the LOCA energy is added to the fluid due to sensible heat transfer from hot

metal components such as the rector vessel walls. To assure that the maximum energy

is transferred, the hot wall CHF value is sampled from a discrete, binary 50/50

distribution with one possible outcome the nominal calculated value and the other

possible outcome a value of 1000.0 times the normal calculated value, as shown in

Table 8.4-1. This assures nucleate boiling is calculated for these walls in 50 percent of

the cases studied.

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8.4-11 AREVA NP (Framatome ANP), EMF-2102(P) Revision 0. S-RELAP5:

Code Verification and Validation. August 2001.

8.4-12 BAW-10231P-A, Revision 1, COPERNIC Fuel Rod Design Computer

Code, Framatome ANP, January 2004.

8.4-13 Oak Ridge National Laboratory, ORNL/NUREG-17. Zirconium Metal-

Water Oxidation Kinetics: IV. Reaction Rate Studies. August 1977.

8.4-14 Electric Power Research Institute, EPRI NP-2013, NUREG/CR-2256,

WCAP-9891, PWR FLECHT SEASET Unblocked Bundle, Forced and

Gravity Reflood Task Data Evaluation and Analysis Report, February

1982.

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8.5.1.8 Pump Differential Pressure Loss

The pump differential pressure loss is addressed in the methodology strictly as a best-

estimate model. The S-RELAP5 code has the ability to input the pump-specific

homologous curves for the NPP being analyzed, and this option is used. The

homologous curves for the specific NPP pumps are obtained from the utility and, if plant

data are available, a pump coast down benchmark is performed to ensure the behavior

is consistent.

8.5.1.9 Noncondensible Transport

The treatment of noncondensibles in the S-RELAP5 code was demonstrated to be

conservative through the assessment of the ACHILLES ISP 25. The rod thermocouples

in the test all clearly showed a reduction in temperature following the introduction of

nitrogen into the system. The S-RELAP5 code conservatively underpredicted this

cooldown, as shown in Figure 8.2-606 through Figure 8.2-611. Figure 8.2-613 shows

the calculated increase in system pressure is lower than the data, which also potentially

reduces the core cooling because of the effect of system pressure on steam binding.

The impact of the nitrogen injection following the accumulator emptying of water will

therefore be conservatively predicted in the NPP analysis.

8.5.1.10 Downcomer Entrainment

The S-RELAP5 code prediction of the ECC bypass during the refill phase of a LOCA

was demonstrated to be conservative through the assessment of UPTF Tests 6 and 7

(Section 8.2.9.3). Additionally, a CCFL correlation developed by MPR Associates is

used in the sample plant cases given in Appendix B to demonstrate S-RELAP5

conservatively calculates the bottom of core recovery (or beginning of core reflood)

time. The MPR correlation is described in Section 8.6.2.2.7. Acceptable downcomer

entrainment during the reflood phase was demonstrated for the CCTF benchmarks

discussed in Section 8.5.1.12.

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Table 8.5-1 Methodology Treatment of Important PIRT Phenomena

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Table 8.5-2 Summary of Evaluated Uncertainties of Important PIRT Parameters

1 The FILMBL and DFFBHTC multipliers discussed above were developed using Rev.2 methodology. Modifications to the Rev. 2 methodology were added in the Rev. 3 methodology. The reevaluated multipliers are given in Appendix A.2.3.6.5.

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8.5.2 Important PIRT Phenomena Treated Statistically

A summary, giving the parameter bias and uncertainty, and how they are to be applied

in the methodology, is provided in this section. The determination of code or physical

phenomena uncertainties is presented in Section 8.4. Other parameters treated

statistically are discussed in detail, including background information, justification of the

statistical approach and explanation of the objective of the statistical treatment.

Table 8.5-2 presents a summary of the key parameters treated statistically in the

AREVA RLBLOCA methodology. The table lists the biases and provides a description

of the statistical treatment of uncertainty for each key parameter.

8.5.2.1 Stored Energy

Revision 3 of the RLBLOCA methodology incorporates COPERNIC2 (Reference 8.5-12)

as the fuel performance code, from which the initial fuel conditions and the fuel thermal

mechanical correlations are determined. This code is used for Uranium oxide fuel

pellets with and without Gadolinia.

The analysis of stored energy uncertainty was performed in Section 8.4.8 by assessing

COPERNIC2 predictions for centerline fuel temperature relative to data (see data

discussion in Reference 8.5-13). The assessment was established as a bias and an

uncertainty in the form of the difference of measured and predicted temperatures

ratioed to the predicted temperatures. For the development in Section 8.4.8 , the form

was:

( )Predicted Measured

Predicted

TT

T−

This gives an adjustment proportional to the magnitude of the predicted centerline fuel

temperature, and is easy to apply within a code structure. The (TPredicted-TMeasured)

means the negative of the adjustment is provided.

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COPERNIC2 is an NRC-approved current generation fuel performance code. The

assessment database used to develop the bias and uncertainty for the RLBLOCA

methodology was incorporated in the code approval. The approval resulted in the

assignment of a zero bias and, for deterministic evaluations, a 71 °C increase in the

centerline fuel temperature to achieve a 95/95 prediction. This adjustment is an

absolute and not dependent on the magnitude of the prediction. For RLBLOCA, it is

replaced with a proportional adjustment of the form (TPredicted-TMeasured)/TPredicted. [

]

In line with the realistic treatment of uncertainty, the adjustment is sampled separately

for each member analysis of the case set and is sampled as a positive and a negative

adjustment. Figure 8.5-9 gives the uncertainty used in the methodology as a

cumulative distribution in comparison to the actual cumulative distribution of the

benchmarked database. Within the range of negative adjustments to temperature, the

adjustment is somewhat less than the data would justify making the methodology

slightly conservative.

8.5.2.2 Oxidation

Energy released through the oxidation of cladding is calculated using the Cathcart-

Pawel correlation (Reference 8.5-3) for oxide layer growth:

2 35890·0.01126·

2R Teφδ −

=,

where R is the universal gas constant (1.987 cal/mole-K) and T is clad temperature.

This is given in Section 7.9.3.5 as:

180620.0000022522

Tr

et r

φ

φ

δδ

−Δ=

Δ

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8.5.2.4 Core Post-CHF Heat Transfer

The post-CHF heat transfer model now includes provisions for thermal radiation

between structures (rod-to-rod). This adds to the current model which already includes

thermal radiation from structures to the fluid (rod-to-droplets and rod-to-steam). The

rod-to-rod radiation model is only applied to the hot rod since its power level is elevated

compared to it surroundings. Applying rod-to-rod radiation exclusively to the hot rod

logically leads to the development of separate heat transfer uncertainties for the hot rod

and the rest of the core.

The core wide heat transfer uncertainty was developed from code comparisons using

the FLECHT-SEASET reflood test data as discussed in Section 8.4.1. These

comparisons were used to derive the heat transfer multipliers that are applied to film

boiling (FILMBL) heat transfer and dispersed flow film boiling heat transfer (DFFBHTC).

[

]

2 The FILMBL and DFFBHTC multipliers discussed above were developed using Rev.2 methodology. Modifications to the Rev. 2 methodology were added in the Rev. 3 methodology. The reevaluated multipliers are given in Appendix A.2.3.6.5.

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8.5.3 Application of Code Biases

This section summarizes the biases applied to the assessments presented in the

previous sections. The biases were developed from uncertainty analyses performed on

the SETs. In most instances, each bias developed has an uncertainty associated with

it, but the uncertainties were not included in the assessments.

The biases listed below were taken from Table 8.5-2:

3 [ ] 4 [

]

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Table 8.5-4 Biases Used in Assessments

5 [ ] 6 [ ] 7 [

] 8 [

]

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8.5-10 NEA/CSNI/R(2010)6, Benchmark Calculations on Halden IFA-650

LOCA Test Results, Organization for Economic Co-Operation and

Development (OECD) – Nuclear Energy Agency (NEA), November

2010.

8.5-11 B. C. Oberlander, M. Espeland, N. O. Solum, H. K. Jenssen, LOCA

IFA650-4: Fuel Relocation Study, LOCA Workshop/HPG Meeting,

Prague, September 2007.

8.5-12 BAW-10231P-A, Revision 1, COPERNIC Fuel Rod Design Computer

Code, Framatome ANP, January 2004.

8.5-13 EPRI-294-2. Mixing of ECC Water with Steam: 1/3 Scale Test and

Summary. June 1975.

8.5-14 NUREG-1230. Compendium of ECCS Research for Realistic LOCA

Analysis. December 1988.

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Table 8.5-5 Film Boiling Multiplier

Table 8.5-6 Dispersed Flow Film Boiling Multiplier

9 The FILMBL and DFFBHTC multipliers discussed above were developed using Rev.2 methodology. Modifications to the Rev. 2 methodology were added in the Rev. 3 methodology. The reevaluated multipliers are given in Appendix A.2.3.6.5.

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Code scaling evaluation will focus on those items identified by the sensitivity studies of

PIRT phenomena as having the greatest impact on LBLOCA. Table 6-1 shows the

results of sensitivity studies on the PIRT phenomena in a PWR LBLOCA. The models,

related to these phenomena, and the scalability of each of these models, are discussed

in the following paragraphs.

Items related to fuel rod performance are not affected by scaling, because the basis for

the fuel-stored energy and dynamic response are based on COPERNIC2

(Reference 8.6-2), which was benchmarked to data from actual fuel rods.

8.6.2.1 Post-CHF and Reflood Heat Transfer

When heat flux from the fuel rods and any other metal masses exceeds the CHF, the

heat transfer is calculated using correlations specific to the heat transfer regimes. The

single-phase vapor, transition boiling and film boiling regimes constitute the post-CHF

heat transfer regimes. For each of these regimes, the effects of radiation heat transfer

also are considered.

Single-phase vapor heat transfer is the maximum of the Wong-Hochreiter correlation for

forced flow regimes (turbulent and laminar) and the turbulent natural convection heat

transfer recommended by Holman, as described in Sections 7.7.5 and 7.7.8. The

Wong-Hochreiter correlation generally determines the heat transfer.

The natural convection heat transfer model is based on data from the flow between

vertical plates. If the boundary layer is small compared to the diameter of the rod, then

heat transfer through this layer would be very similar to that through the boundary layer

on a plate. With the Prandtl number near unity and the rod diameter large compared to

the boundary layer, the Holman formulation for natural convection heat transfer used in

S-RELAP5 applies as long as

0.2535·G

LrD −≥

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8.6.2.2.5 Pump Scaling

The S-RELAP5 code uses normalized single-phase homologous curves for a full-scale

reactor coolant pump (RCP) as code input. The use of full-scale data for the pump

makes code scaling moot. These homologous curves are set to applicable values by

entering plant-specific values for rated head, torque, moment of inertia, etc. The

coastdown of the pump is driven by the torque and moment of inertia of the rotating

mass. The torque includes the effects of friction and back EMF (pump torque), and of

the loop pressure losses (hydraulic torque). Although the two-phase degradation of

RCP performance is not considered a phenomenon of significance (Table 5-1), the

single-phase pump head and torque curves are adjusted for two-phase effects based on

the EPRI two-phase degradation data (Section 7.9.2.3). The pumps in the EPRI test

program are similar to PWR coolant pumps and the data represents a best estimate

approximation of both the single phase and two phase performance.

8.6.2.2.6 Cold Leg Condensation

As discussed in Section 8.4.2, several EPRI 1/3 scale tests, in combination with UPTF

Test 8 Phase A (Run 111), and Phase B (Run 112), and Test 25, were simulated using

S-RELAP5. The simulation results were used to develop the biases (multipliers) on the

liquid-side (CONMAS) and vapor-side (CONMSG) interphase heat transfer coefficients.

The tests selected generally cover both the accumulator and pumped injection period of

the LOCA transient. Additionally, further EPRI tests were simulated using S-RELAP5,

and the results are discussed in Section 8.2.10. The UPTF is close to a full-scale

facility and the EPRI test facility is a 1/3 scaled facility.

Correlations based on the Stanton numbers are used to calculate the interphase

condensation. These correlations are generally insensitive to geometry as

demonstrated by the EPRI and UPTF benchmark results. The interphase heat transfer

correlations used in S-RELAP5 are discussed in detail in Section 7.6.4.

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The details of this CCFL correlation are given in Reference 8.6-6. In this correlation, *,effgJ is the net steam flow rate available to entrain the ECC liquid to the break and its

value determines the potential for ECC bypass. If *,effgJ is zero or negative, the steam

flow is insufficient to entrain liquid, and bypass will not occur (complete end-of-bypass).

If *,effgJ is positive, then partial or full bypass occurs.

Since the correlation is normalized using the downcomer flow area and circumference, it

is directly applicable to calculate the ECC bypass in the plant during the refill phase.

This correlation has already been approved by the NRC to calculate the complete end-

of- bypass time ( *,effgJ < 0.0) as part of AREVA NP’s Appendix K-based Recirculating

Steam Generator LOCA Evaluation Model (Reference 8.6-7).

The correlation is used in the sample plant cases discussed in Appendix B to estimate

the beginning of core reflood time in order to demonstrate S-RELAP5 will appropriately

calculate the beginning of core reflood time. To estimate the beginning of core reflood

time, the correlation is used to calculate the complete end-of-bypass time. At this time,

the liquid volume in the lower head, lower plenum, and downcomer below the active

core region is determined. Knowing this time, the beginning of core refloood time can

be estimated by the ECC injection rates in the intact cold legs and the remaining fluid

volume below the active core region that need to be filled with water. The results for the

sample cases (Figure B-18, Figure B-36 and Figure B-54) demonstrate S-RELAP5

calculates the beginning of core reflood time appropriately.

The highly separated flow behavior observed in the full-scale UPTF tests (see Figure

4.1-3 in Reference 8.6-8) were not observed in scaled facilities like LOFT and

Semiscale. The tests conducted in these scaled facilities therefore cannot be used to

determine code scalability of ECC bypass, and the multi-dimensional flow phenomena

that will occur in the downcomer and lower plenum during the refill phase.

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In summary, from the simulation of the above tests, it can be concluded that S-RELAP5

will calculate the acceptable loop oscillations during a large break LOCA in a PWR.

Table 8.6-1: Test Ranges for Film Boiling Heat Transfer Test Comparison

Parameter Maximum Minimum

Tests LBLOCA Tests LBLOCA Pressure (MPa) 8.2 10.8 0.13 0.22 Mass Flux Vapor (kg/s-m2) 907 367 0 0 Mass Flux Liquid (kg/s-m2) 4254 945 0 0 Void Fraction 1 1 0.13 0.13 Saturation Temperature (K) 570 589 381 390 Vapor Temperature (K) 1294 1160 384 391 Wall Temperature (K) 1525 1400 390 396 Quality 1 1 -0.11 0

8.6.3 References 8.6-1 NUREG/CR-5249, Quantifying Reactor Safety Margins, Application of

Code Scaling, Applicability, and Uncertainty Evaluation Methodology to a

Large Break, Loss-of-Coolant Accident, December 1989.

8.6-2 BAW-10231P-A, Revision 1, COPERNIC Fuel Rod Design Computer

Code, Framatome ANP, January 2004.

8.6-3 NUREG/CR-0410, Comparisons of Thermal-Hydraulic Phenomena

During Isothermal Loss-Of-Coolant Experiments and Effect of Scale in

LOFT and SEMISCALE MOD-1, December 1978.

8.6-4 NUREG/CR-1533, Analysis of the FLECHT-SEASET Unblocked Bundle

Steam Cooling and Boil-off Tests, January 1981.

8.6-5 Dittus, F. W. and L. M. K. Boelter, Heat Transfer in Automobile Radiators

of the Tubular Type, Publications in Engineering, Volume 2, pp. 443-461.

University of California, Berkeley, 1930.

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8.6-6 MPR Report, Summary of Results from the UPTF Downcomer Separate

Effects Tests, Comparisons to Previous Scaled Tests, and Application to

U.S. Pressurized Water Reactors, MPR-1163, July 1190.

8.6-7 BAW-10168-A, Revision 3, RSG LOCA – BWNT Loss-of-Coolant

Accident Evaluation Model for Recirculating Steam Generator Plants,

Volume I – Large Break, December 1996.

8.6-8 Damerell, P. S., Simsons, J. W., Reactor Safety Issues Resolved by the

2D/3D Program, NUREG/IA-0127, July 1993.

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The plant base model is fixed for the evaluation and serves as the model to which each

of individual cases makes modification before case execution. The cases differ from the

base model by the values of the sampled parameters assigned for each case. These

parameters comprise the set of highly important parameters identified within the PIRT

and selected plant operational parameters, which may be identified as beneficial to

include. The PIRT identified parameters are listed in Table 5-1.

A representative sample of plant parameters is shown in Table B-9, Table B-16, and

Table B-23 for each of the sample problem calculations in Appendix B, respectively.

9.1 Base Model

The base model establishes the special discretization, noding, with which the

components of the reactor system are modeled, and associated modeling of flow paths

and heat structures, including the subparts of each of these, and initial conditions. After

some general guidance discussion, each of these is described, in turn, in the following

subsections.

9.1.1 Nodalization Methodology

Reference 9-1 ("Quantifying Reactor Safety Margins") makes the following statements

regarding nodalization:

The plant model must be nodalized finely enough to represent both the important phenomena and design characteristics of the NPP but coarsely enough to remain economical. Thus, the preferred path is to establish a standard NPP nodalization for the subsequent analysis. This minimizes or removes nodalization, and the freedom to manipulate noding, as a contributor to uncertainty. Therefore, a nodalization selection procedure defines the minimum noding needed to capture the important phenomena. This procedure starts with analyst experience in previous code assessment and application studies and any documented nodalization studies. Next, nodalization studies are performed during the simulation of separate- and integral-effects code data comparisons. Finally, an iterative process using the NPP model is employed to determine sufficiency of the NPP model nodalization.

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Determination of which process parameters to treat statistically begins with identifying

the relationship a particular parameter has to any PIRT phenomenon. Table 9-3 lists

process parameters determined to be important based on their potential influence to the

moderate-to-high ranked phenomena given in the PIRT, Table 5-1.

A refinement of the conclusions presented in Table 9-2, based on sensitivity studies, is

within the precepts of the methodology. Such studies can be used to adopt a bias over

an uncertainty distribution for process parameters or to assist in the quantification of an

uncertainty range or distribution.

Other process parameters are considered to be of lower importance, and are generally

treated on a nominal basis. As with any parameter, there is no prohibition to treating

these parameters on a statistical basis.

9.3.1.2 Quantifying Uncertainty for Process Parameters

To treat a parameter statistically, the parameter uncertainty must be quantified in terms

of biases and distributions. Quantifying this uncertainty with plant data is the best

approach. At most plants, histories of parameters values, such as RCS flow rate, core

inlet temperature, pressurizer condition, accumulator parameters, and containment

temperature are maintained and useable for quantifying RLBLOCA analysis

uncertainties. Operational uncertainty is defined as the true fluctuation of a parameter

during normal operation. Setting the uncertainty distribution for a process parameter

requires addressing the impact of measurement uncertainty for the parameter.

The choice of distribution may be influenced by how a utility manages a given process

parameter. For example, using a uniform distribution may properly reflect the control

provided for a parameter, if that control is random within a range. A uniform distribution

is also considered a conservative approach in that equal likelihood is given for values at

the limits of the distribution where the strongest influence is expected. However, if the

there is an expectation that the true distribution is substantially non-uniform, the actual

distribution can be used.

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• A minimum of [ ] fuel “rods” are used to model the active core: [

] The heat structure modeling conventions for fuel rods described in

Section A.1.3.6.1.4 are used for the fuel rods. Axially, each rod has [ ] heat

structures of equal length per hydraulic volume [

] This information must also be present in the COPERNIC model. [

] Additional guidelines for the fuel rod heat structures are as

follows:

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A.1.3.6.4 Control Variables and Trips

This section describes system monitoring control variables, steady-state initialization

aids, and reactor trips, which are needed or are useful in LOCA calculations. As a

methodology based on the principles of “Code Scaling, Uncertainty and Applicability”

(Reference A-1), an AREVA RLBLOCA analysis distinguishes uncertainty between

phenomenological influences (e.g., film boiling heat transfer and stored energy), and

plant operational influences (e.g., pressurizer level and accumulator/SIT pressure). In

these separate domains, the uncertainty will be treated differently in RLBLOCA

analyses. The emphasis of any CSAU methodology is focused on the assessment of

phenomenological uncertainty. In addition, a CSAU-based methodology recognizes a

hierarchical relationship among phenomena influencing the key figure-of-merit (i.e., the

peak clad temperature). During development of the AREVA RLBLOCA methodology,

this hierarchical relationship was first established by the development of a Phenomena

Identification and Ranking Table (PIRT) (Table 5-1), and then quantified by a series of

sensitivity studies performed in EMF-2103 Revision 0 (Reference A-4).

For those phenomena identified as important, code bias and uncertainty were evaluated

using experimental test data from a diverse set of experiments. Instrument uncertainty

was inherent in all tests was, but no effort was made to remove this uncertainty to

improve code-to-data agreement. Instead, this uncertainty represents a component of

the final uncertainty values determined for the statistically ranged parameters applied in

the methodology. A benefit of accepting this additional uncertainty is that the population

of plant states and operational ranges into the computer models does not need to

include measurement uncertainty. While it is acceptable to explicitly apply this

uncertainty in parameter ranges, to do so is double accounting for measurement

uncertainty and, hence, conservative.

A.1.3.6.4.1 System Monitoring Control Variables

The following control variables can be present for monitoring of various parameters, or

to provide derived values for trip functions (others are certainly permitted):

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Material property tables for the fuel, gap, and cladding must be present even through

the COPERNIC fuel model is to be incorporated into the calculation. The accuracy of

these tables is to be confirmed for a particular analysis. If additional materials are

required, new tables can be appended to the current list.

For the COPERNIC fuel model, the properties of fuel rod materials (UO2, M5®, and gap

fill gas) are evaluated from the COPERNIC calculation and then read by S-RELAP5. In

this case the material properties entered by users are used only for the unheated

portion of the fuel rods if they are modeled, or they may be needed to satisfy certain

input requirements.

A.1.3.6.6 Steady-State Initialization

Certain plant specific parameters must be approximated in the base input file prior to

steady-state initialization. These include bypass flow rates, upper head temperature,

and steam generator secondary steam heat transfer rate (via feedwater temperature,

pressure, liquid level, and recirculation ratio). In addition, best-estimate pressure drop

information may be used to refine form-loss values. This is often done to validate steam

generator, core, and vessel pressure drops.

Geometry and inherent limits of finite difference computer codes to model heat

exchanger heat transfer make steam generator initialization a particular challenge.

Analysts are expected to make an effort to best match plant data on the key steam

generator parameters: steam dome pressure, main steam flows, main feedwater flows,

and steam generator mass. [

]

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Also required for a transient is Word 7 on Card 30000001. This parameter is set to -1.0

to apply the neutron capture correction factor which is a time dependent ratio applied to

decay heat. Using a negative number prescribes S-RELAP5 to use the hardwired

correction factor table. Note that this parameter must not be set in a steady-state

calculation.

A.1.3.7 Instructions for COPERNIC Input

The static COPERNIC parameters (i.e., not time- or burnup-dependent) are based

primarily on AREVA mechanical specifications of the fuel. For this reason, best-

estimate values should be available for these parameters. Nonetheless, while PCTs

are strongly influenced by burnup, scoping studies have shown that within the normal

range of uncertainty, variation of most COPERNIC parameters have only a small effect

on PCT.

COPERNIC is the preferred fuel rod code to be used for all Zirconium clad fuel. This

guideline section highlights the unique code input requirements for performing an

RLBLOCA analysis using COPERNIC.

Application of the COPERNIC code is for the generation of best-estimate fuel rod

properties to be used in the S-RELAP5 RLBLOCA calculation. The COPERNIC Theory

and Users Manual (Reference A-8) should be consulted for a description of input. [ ] much of the input must be

generated dynamically. Instruction for creating this input is presented in the RLBLOCA

Analysis Guideline (Section A.2.5.16.1 below). Nonetheless, it is necessary to have a

base COPERNIC model to initialize the S-RELAP5 model. Automation tools are

provided to create this input.

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A.2.3.6.3 Gaussian (Normal) PDF

A Gaussian PDF is the natural limit to the convolution of many random events. Using

the floating point random number generator, this PDF is defined using the Box-Muller

transform (Reference A-16):

( ) ( )random2**2cos * 1randomln*2* π−σ+η

Where η is the mean and σ is the standard deviation.

A.2.3.6.4 Log-Normal PDF

The log-normal PDF provides a distribution for variables whose natural logarithm is

normally distributed. Using the floating point random number generator, the log-normal

PDF is defined as:

( ) ( )random2**2cos*)random1ln*2*exp( π−σ+η

Where η is the mean and σ is the standard deviation.

A.2.3.6.5 Film Boiling Multipliers

[

]

A.2.3.6.5.1 FILMBL and DFFBHTC

[

]

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[

]

A.2.3.6.5.2 PDF of FILMBL and DFFBHTC

[

]

[

]

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Table A-6 Uncertainty Parameters and PDFs

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A.2.3.7.1 Model Parameter Ranges

Table A-7 summarizes the model parameters that must be applied in every RLBLOCA

analysis.

Table A-7 Model Parameter Uncertainty Ranges

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[

]

[

]

A.2.3.11 Statistical Evaluation

As previously explained, the methodology has the advantage of being able to treat a

large number of parameters by randomly varying each parameter in each single

calculation. This random selection process is repeated to define a large number of

RLBLOCA calculations, all of which are then run.

The key results are read for each of the RLBLOCA calculations and the values for PCT,

Maximum Local Oxidation, and Core-Wide Oxidation are saved for each case. The

values for PCT, % Maximum Local Oxidation and % Core-Wide Oxidation are then set

to the PCTi, MLOi, and CWOi for each case respectively.

[

]

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A.2.5.15 Radial Power Factor Calculation

The radial power peaking factors are used to partition the power among the modeled

fuel rod regions defined in Table A-21.

Table A-21: Radial Power Peaking Factors

The number of fuel rods must be calculated for each region, so that the power in each

region is normalized to the number of rods in the core. Table A-21 identifies the

variable used for the number of rods in each region, for the equations used to calculate

the radial power fractions for each region in the sections that follow.

The total number of assemblies in the core (AssyCore) is read from the plant database

record with component core_design and parameter num_assy. The number of fuel rods

per assembly (RodsAssy) is read from the plant database record with component

assembly_* and parameter assy_num_rods_fuel.

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A.2.5.19 Containment Model Input

The S-RELAP5/ICECON containment input file is processed during (and only during)

the S-RELAP5 transient calculation. A distinct containment model input file is also

required for each transient calculation. [

] A.2.5.20 Calculations

As described in Section A.2.3.2, calculations for the preferred set of [ ] unique

cases will be performed for the uncertainty analysis. For each case, the fuel rod code is

executed first, followed by the S-RELAP5-SS calculation and concludes with the

S-RELAP5-TR calculation.

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A detailed assessment of the S-RELAP5 computer code was made through comparisons to

experimental data, as documented in Section 8.0. These assessments were used to develop

quantitative estimates of the ability of the code to predict key physical phenomena in a PWR

LBLOCA. The final step of the best-estimate methodology is to combine all the uncertainties

related to the code and plant parameters and estimate the PCT at 95 percent probability and 95

percent confidence. The steps taken to derive the PCT uncertainty estimate are summarized

below:

1. Base Plant Input File Development

First, base COPERNIC and S-RELAP5 input files for the plant (including the containment input

file) are developed. Code input development guidelines documented in Appendix A are applied

to ensure that the model nodalization is consistent with the model nodalization used in the code

validation.

2. Sampled Case Development

The statistical approach requires that many “sampled” cases be created and processed. For

every set of input created, each “key LOCA parameter” is randomly sampled over a range

established through code uncertainty assessment or expected operating limits (provided by

plant technical specifications or data). Those parameters considered "key LOCA parameters"

are listed in Table A-6. This list includes both parameters related to LOCA phenomena (based

on the PIRT provided in Section 5.0) and to plant operating parameters. The uncertainty ranges

associated with each of the model parameters are provided in Table A-7.

3. Determination of Adequacy of ECCS

The RLBLOCA methodology uses a non-parametric statistical approach to determine that the

first three criteria of 10 CFR 50.46 (PCT < 2200 °F, local oxidation < 17 %, and core-wide

oxidation < 1 %) are met with a probability higher than 95 percent with 95 percent confidence.

B.1.3 GDC-35 Limiting Condition Determination

GDC-35 states that the plant shall be able to mitigate design basis accidents with or without off

site power available. The methodology does this by determining the most severe condition

between these two configurations and then performing the RLBLOCA statistical analysis for the

plant with off site power availability set to the most severe condition.

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To determine the limiting assumption, a sensitivity study of two LBLOCA cases is performed

with and without offsite power available. The plant conditions incorporated in this study are set

to those expected to challenge the ECCS capability, such that the validity of the result is

established for conditions expected to be representative of those that will eventually determine

the LBLOCA results which will be compared to the 10 CFR 50.46 criteria. A detailed description

of the sensitivity study is provided in Appendix A of this report, sub-section A.2.4.2.

The study is performed with and without off site power available. As mentioned previously, the

conditions assumed will be based on those considered to be representative of the LBLOCA

results that be compared to the 10 CFR 50.46 criteria. The statistical case set, set of LBLOCA

sample events, is then run under the assumption that off-site power is always either available or

unavailable according to the study result.

B.1.4 Overall Statistical Compliance to Criteria

For the RLBLOCA analyses the determination of compliance to the criteria is treated as a

[ ] with all of the first three criteria of 10 CFR 50.46 using non-parametric

statistics. The approach is outlined in detail in Section 9.4 of this report. [

] Generally, the minimum margins for each

of the three parameters of interest will be established by different cases. For the sample

evaluations presented in this appendix, a case set size of [ ] was selected. At this size,

the 95/95 metric value is provided by the [ ] for the criterion of interest.

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B.2 Westinghouse 3-Loop PWR

B.2.1 Summary

The parameter specification for this analysis is provided in Table B-9. [

] The

analysis addresses typical operational ranges or technical specification limits (whichever is

applicable) with regard to pressurizer pressure and level; accumulator pressure, temperature

(containment temperature), and level; core inlet temperature; core flow; containment pressure

and temperature; and refueling water storage tank temperature. [

]

B.2.2 Plant Description and Summary of Analysis Parameters

The plant analysis presented in this section is a Westinghouse designed PWR, having three

loops, each with a hot leg, a U-tube steam generator, and a cold leg with a RCP. The RCS also

includes a pressurizer. The ECCS comprises three accumulators, one per loop, and one full

train of LHSI and HHSI injection (after applying the single failure assumption). The HHSI and

LHSI feed into common headers (cross connected) that are connected to the accumulator lines.

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The S-RELAP5 model explicitly describes the RCS, reactor vessel, pressurizer, and ECCS back

to the common LHSI header and accumulators. This model also describes the secondary-side

steam generator that is instantaneously isolated (closed MSIV and feedwater trip) at the time of

the break.

As described in Appendix A, many parameters associated with LBLOCA phenomenological

uncertainties and plant operation ranges are sampled. A summary of those parameters

sampled is given in Table A-6. The LBLOCA phenomenological uncertainties are provided in

Table A-7. Values for process or operational parameters, including ranges of sampled process

parameters, and fuel design parameters used in the analysis are given in Table B-7 Plant data

are analyzed to develop uncertainties for the process parameters sampled in the analysis. Table

B-8 presents a summary of the uncertainties used in the analysis. Two parameters (RWST

temperature and diesel start time) are set at conservative bounding values for all calculations.

Where applicable, the sampled parameter ranges are based on technical specification limits.

Plant data are used to define range boundaries for loop flow (high end) and containment

temperature (low end).

B.2.3 Realistic Large Break LOCA Results

[

]

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Table B-9 is a summary of the major input parameters for the MRMC case.The results for the

MRMC case are presented in Table B-10. The fraction of total hydrogen generated was not

directly calculated; however, it is conservatively bounded by the calculated total percent

oxidation, which is well below the 1-percent limit. The event times for the MRMC case can be

found in Table B-11 and the heat transfer parameter range for the limiting margin case is

provided in Table B-12. [

]

The analysis plots are shown in Figure B-1 through Figure B-18. Figure B-1 shows linear scatter

plots of the key parameters sampled for the cases that lie in the 95/95 range. Parameter labels

appear to the left of each individual plot. These figures illustrate the parameter ranges used in

the analysis.

Figure B-2 and Figure B-3 show PCT scatter plots versus the time of PCT and versus break

size from the set of cases (LOCA events) that lie within the 95/95 range. The scatter plots for

the maximum local oxidation and total core-wide oxidation for the set of cases that lie within the

95/95 range are shown in Figure B-4 and Figure B-5, respectively. Figure B-6 through Figure B-

17 show key parameters from the S-RELAP5 calculations for the MRMC case. Figure B-6 is the

plot of PCT, independent of elevation. Figure B-18 compares the beginning of core recovery

times for the set of cases that lie within the 95/95 range to the BOCR time predicted using the

MPR CCFL correlation. Note that Figure B-18 uses the total break area, while previous plots

used break area per side.

B.2.4 Conclusions

[

]

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Table B-7 3-Loop Westinghouse Plant Operating Range Supported by the RLBLOCA Analysis (continued)

Event Operating Range 3.0 Accident Boundary Conditions

a) Break location

b) Break type

c) Break size (each side, relative to cold leg pipe area)

d) Worst single-failure e) Offsite power

f) ECCS pumped injection temperature

g) HHSI pump delay

h) LHSI pump delay

i) Containment pressure

j) Containment temperature

k) Containment sprays delay

l) Containment spray water temperature

m) LHSI Flow

1 This is determined prior to the execution of the set of [ ] cases.

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Table B-9 3-Loop Westinghouse Summary of Major Parameters for the MRMC Case

Parameter Value Time in Cycle (hrs) Burnup (GWd/mtU) Core Power (MWt) Core Peaking (Fq)Radial Peak (F H)

Axial Offset Local Peaking (Fl)

Break Type Break Size (ft2/side)

Offsite Power Availability

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Table B-10 3-Loop Westinghouse Compliance with 10 CFR 50.46

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Table B-11 3-Loop Westinghouse Calculated Event Times for the MRMC Case

Event Time (sec)

Begin Analysis Break Opens RCP Trip SIAS Issued Start of Broken Loop Accumulator Injection Start of Intact Loop Accumulator Injection (Loop 2 and 3 respectively)

Start of HHSI Start of Charging Beginning of Core Recovery (Beginning of Reflood) LHSI Available PCT Occurred Broken Loop LHSI Delivery Began Intact Loops LHSI Delivery Began (Loop 2 and 3 respectively)

Broken Loop HHSI Delivery Began Intact Loops HHSI Delivery Began (Loop 2 and 3 respectively)

Broken Loop Accumulator Emptied Intact Loop Accumulator Emptied (Loop 2 and 3 respectively)

Transient Calculation Terminated

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Table B-12 Westinghouse 3-Loop Heat Transfer Parameters for Limiting MRMC Case

Time (s)

LOCA Phase Early Blowdown Blowdown 1 Refill Reflood Quench Long Term

Cooling

Heat Transfer Mode

Heat Transfer Correlations

Maximum LHGR (kW/ft)

Pressure (psia)

Core Inlet Mass Flux (lbm/s-ft22)

Vapor Reynolds Number3

Liquid Reynolds Number

Vapor Prandtl Number

Liquid Prandtl Number

Vapor Superheat4

(°F)

1 End of Blowdown considered as beginning of refill. 2 Conservatively biased parameter 3 Not important in pre-CHF heat transfer 4 Vapor superheat is meaningless during blowdown and system depressurization

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Table B-13 Westinghouse 3-Loop Fuel Rod Rupture Ranges of Parameters for all [ ] cases

Parameter Name Minimum Value Maximum Value

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Figure B-1 3-Loop Westinghouse Scatter Plot of Operational Parameters for all [ ] Cases

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Figure B-1 3-Loop Westinghouse Scatter Plot of Operational Parameters for all [ ] Cases (continued)

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Figure B-2 3-Loop Westinghouse PCT versus PCT Time Scatter Plot from the Cases within the 95/95 Range

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Figure B-3 3-Loop Westinghouse PCT versus Break Size Scatter Plot from the Cases within the 95/95 Range

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Figure B-4 3-Loop Westinghouse Maximum Local Oxidation versus PCT Scatter Plot from the Cases within the 95/95 Range

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Figure B-5 3-Loop Westinghouse Total Core-Wide Oxidation versus PCT Scatter Plot from the Cases within the 95/95 Range

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Figure B-6 3-Loop Westinghouse Peak Cladding Temperature (Independent of Elevation) for the MRMC Case

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Figure B-7 3-Loop Westinghouse Break Flow for the MRMC Case

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Figure B-8 3-Loop Westinghouse Core Inlet Mass Flux for the MRMC Case

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Figure B-9 3-Loop Westinghouse Core Outlet Mass Flux for the MRMC Case

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Figure B-10 3-Loop Westinghouse Void Fraction at RCS Pumps for the MRMC Case

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Figure B-11 3-Loop Westinghouse ECCS Flows (Includes Accumulator, Charging, SI and RHR) for the MRMC Case

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Figure B-12 3-Loop Westinghouse Upper Plenum Pressure for the MRMC Case

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Figure B-13 3-Loop Westinghouse Collapsed Liquid Level in the Downcomer for the MRMC Case

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Figure B-14 3-Loop Westinghouse Collapsed Liquid Level in the Lower Plenum for the MRMC Case

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Figure B-15 3-Loop Westinghouse Collapsed Liquid Level in the Core for the MRMC Case

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Figure B-16 3-Loop Westinghouse Containment and Loop Pressures for the MRMC Case

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Figure B-17 3-Loop Westinghouse Pressure Difference between Upper Plenum and Downcomer

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Figure B-18 3-Loop Westinghouse Validation of BOCR Time using MPR CCFL Correlation

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B.3 Westinghouse 4-Loop PWR

B.3.1 Summary

The parameter specification for this analysis is provided in Table B-14. [

] This analysis also

addresses typical operational ranges or technical specification limits (which ever is applicable)

with regard to pressurizer pressure and level; accumulator pressure, temperature (containment

temperature), and level; core inlet temperature; core flow; containment pressure and

temperature; and refueling water storage tank temperature. The analysis explicitly analyzes

fresh and once-burned fuel assemblies. [

]

B.3.2 Plant Description and Summary of Analysis Parameters

The plant analysis presented in this appendix is a Westinghouse designed pressurized water

reactor (PWR), which has four loops, each with a hot leg, a U-tube steam generator, and a cold

leg with a RCP. The RCS also includes one pressurizer. The ECCS includes one charging and

one accumulator/SI/RHR injection path per RCS loop (after applying the single failure

assumption). The SI and RHR feed into common headers which are connected to the

accumulator lines. The charging pumps are also cross-connected.

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The S-RELAP5 model explicitly describes the RCS, reactor vessel, pressurizer, and

accumulator lines. The charging injection flows are connected to the RCS, and the SI and RHR

injection flows are connected to the accumulator lines. This model also describes the

secondary-side steam generator that is instantaneously isolated (closed MSIV and feedwater

trip) at the time of the break.

As described in Appendix A, many parameters associated with LBLOCA phenomenological

uncertainties and plant operation ranges are sampled. A summary of those parameters

sampled is given in Table A-6. The LBLOCA phenomenological uncertainties are provided in

Table A-7. Values for process or operational parameters, including ranges of sampled process

parameters, and fuel design parameters used in the analysis are given in Table B-14. Plant

data is analyzed to develop uncertainties for the process parameters sampled in the analyses.

Table B-15 presents a summary of the uncertainties used in the analysis. Two parameters

(RWST temperature and diesel start time) are set at conservative bounding values for all

calculations.

Where applicable, the sampled parameter ranges are based on technical specification limits.

Plant data are used to define range boundaries for loop flow (high end) and containment

temperature (low end).

B.3.3 Realistic Large Break LOCA Results

[

]

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Table B-16 is a summary of the major parameters input parameters for the MRMC case. The

results for the MRMC case are presented in Table B-17. The fraction of total hydrogen

generated was not directly calculated; however, it is conservatively bounded by the calculated

total percent oxidation, which is well below the 1-percent limit. The event times for the MRMC

case can be found in Table B-18 and the heat transfer parameter range for is provided in

Table B-19. [

].

The analysis plots are shown in Figure B-19 through Figure B-35. Figure B-19 shows linear

scatter plots of the key parameters sampled for the 95/95 case set. Parameter labels appear to

the left of each individual plot. These figures illustrate the parameter ranges used in the

analysis.

Figure B-20 and Figure B-21 show PCT scatter plots versus the time of PCT and versus break

size from the set of cases (LOCA events) that lie within the 95/95 range. The scatter plots for

the maximum local oxidation and total core-wide oxidation for the set of cases that lie within the

95/95 range are shown in Figure B-22 and Figure B-23, respectively. Figure B-24 through

Figure B-35 show key parameters from the S-RELAP5 calculations for the MRMC case.

Figure B-24 is the plot of PCT, independent of elevation. [

] Note that Figure B-36 uses the total break area while previous

plots used break area per side.

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B.3.4 Conclusions

[

]

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Table B-14 4-Loop Westinghouse Plant Operating Range Supported by the LOCA Analysis (continued)

Event Operating Range

f) Accumulator pressure g) Accumulator liquid volume

h) Accumulator temperature

i) Accumulator fL/D j) Minimum ECCS boron 3.0 Accident Boundary Conditions

a) Break location b) Break type c) Break size (each side, relative to cold leg pipe area) d) Worst single-failure e) Offsite power

f) ECCS pumped injection temperature

g) Charging pump delay

h) SI pump delay

i) RHR pump delay

j) Containment pressure

k) Containment upper compartment temperature

l) Containment lower compartment temperature

m) Containment sprays delay

1 This is determined prior to the execution of the set of [ ] cases.

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Table B-16 4-Loop Westinghouse Summary of Plant Major Parameters for for the MRMC Case

Parameter Value

Time in Cycle (hrs) Burnup (GWd/mtU) Core Power (MWt) Core Peaking (Fq) Radial Peak (F h)

Axial Offset Local Peaking (Fl)

Break Type Break Size (ft2 / side)

Offsite Power Availability

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Table B-17 4-Loop Westinghouse Compliance with 10 CFR 50.46

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Table B-18 4-Loop Westinghouse Calculated Event Times for the MRMC Case

Event Time (sec)

Begin Analysis Break Opens RCP Trip SIAS Issued Start of Broken Loop Accumulator Injection Start of Intact Loop Accumulator Injection (Loop 2, 3, and 4 respectively)

Start of SI Start of CC Beginning of Core Recovery (Beginning of Reflood) RHR Available PCT Occurred (1921oF) Broken Loop RHR Delivery Began Intact Loops RHR Delivery Began (Loop 2, 3, and 4 respectively)

Broken Loop SI Delivery Began Intact Loops SI Delivery Began (Loop 2, 3, and 4 respectively) Broken Loop Accumulator Emptied Intact Loop Accumulator Emptied (Loop 2, 3, and 4 respectively)

Transient Calculation Terminated

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Table B-19 Westinghouse 4-Loop Heat Transfer Parameters for the MRMC Case

Time (s)

LOCA Phase Early Blowdown Blowdown 1 Refill Reflood Quench Long Term

Cooling

Heat Transfer Mode

Heat Transfer Correlations

Maximum LHGR (kW/ft)

Pressure (psia)

Core Inlet Mass Flux (lbm/s-ft2)

Vapor3 Reynolds Number

Liquid Reynolds Number

Vapor Prandtl Number

Liquid Prandtl Number

Vapor Superheat

(°F)

1 End of Blowdown considered as beginning of refill. 2 Conservatively biased parameter, as per the methodology 3 Not important in pre-CHF heat transfer.

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Table B-20 Westinghouse 4-Loop Fuel Rod Rupture Ranges of Parameters for all [ ] Cases

Parameter Name Minimum Value Maximum Value

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Figure B-19 4-Loop Westinghouse Scatter Plot of Operational Parameters for all [ ] cases

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Figure B-19 4-Loop Westinghouse Scatter Plot of Operational Parameters for all [ ] Cases(continued)

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Figure B-20 4-Loop Westinghouse PCT versus PCT Time Scatter Plot from the Cases within the 95/95 Range

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Figure B-21 4-Loop Westinghouse PCT versus Break Size Scatter Plot from the Cases within the 95/95 Range

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Figure B-22 4-Loop Westinghouse Maximum Oxidation versus PCT Scatter Plot from the Cases within the 95/95 Range

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Figure B-23 4-Loop Westinghouse Total Core-Wide Oxidation versus PCT Scatter Plot from the Cases within the 95/95 Range

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Figure B-24 4-Loop Westinghouse Peak Cladding Temperature (Independent of Elevation) for the MRMC Case

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Figure B-25 4-Loop Westinghouse Break Flow for the MRMC Case

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Figure B-26 4-Loop Westinghouse Core Inlet Mass Flux for the MRMC Case

ID:32978 16Sep2014 06:38:12 R5DMX

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Figure B-27 4-Loop Westinghouse Core Outlet Mass Flux for the MRMC Case

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Figure B-28 4-Loop Westinghouse Void Fraction at RCS Pumps for the MRMC Case

ID:32978 16Sep2014 06:38:12 R5DMX

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Figure B-29 4-Loop Westinghouse ECCS Flows (Includes Accumulator, Charging, SI and RHR) for the MRMC Case

ID:32978 16Sep2014 06:38:12 R5DMX

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Figure B-30 4-Loop Westinghouse Upper Plenum Pressure for the MRMC Case

ID:32978 16Sep2014 06:38:12 R5DMX

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Figure B-31 4-Loop Westinghouse Collapsed Liquid Level in the Downcomer for the MRMC Case

ID:32978 16Sep2014 06:38:12 R5DMX

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Figure B-32 4-Loop Westinghouse Collapsed Liquid Level in the Lower Plenum for the MRMC Case

ID:32978 16Sep2014 06:38:12 R5DMX

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Figure B-33 4-Loop Westinghouse Collapsed Liquid Level in the Core for the MRMC Case

ID:32978 16Sep2014 06:38:12 R5DMX

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Figure B-34 4-Loop Westinghouse Containment and Loop Pressures for the MRMC Case

ID:32978 16Sep2014 06:38:12 R5DMX

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Figure B-35 4-Loop Westinghouse Pressure Difference between Upper Plenum and Downcomer

ID:32978 16Sep2014 06:38:12 R5DMX

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Figure B-36 4-Loop Westinghouse Validation of BOCR Time using MPR CCFL Correlation

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B.4 CE 2x4 PWR

B.4.1 Summary

The parameter specification for this analysis is provided in Table B-21. [

] This

analysis also addresses typical operational ranges or technical specification limits (whichever is

applicable) with regard to pressurizer pressure and level; SIT pressure, temperature

(containment temperature), and level; core inlet temperature; core flow; containment pressure

and temperature; and refueling water storage tank temperature. [

]

For the sample analysis [

]

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B.4.2 Plant Description and Summary of Analysis Parameters

The plant analysis presented in this report is for a CE-designed PWR, which has 2X4-loop

arrangement. There are two hot legs each with a U-tube steam generator and four cold legs

each with a RCP. The RCS includes one Pressurizer connected to a hot leg. The core contains

217 thermal-hydraulic compatible AREVA HTP 14X14 fuel assemblies with 2, 4, 6 and 8 weight

percent gadolinia pins. The ECCS includes one high pressure safety injection (HPSI), one LPSI

and one SIT injection path per RCS loop. The break is modeled in the same loop as the

pressurizer, as directed by the RLBLOCA methodology. The RLBLOCA transients are of

sufficiently short duration that the switchover to sump cooling water (i.e., RAS) for ECCS

pumped injection need not be considered.

The S-RELAP5 model explicitly describes the RCS, reactor vessel, Pressurizer, and ECCS. The

ECCS includes a SIT path and a LPSI/HPSI path per RCS loop. The HPSI and LPSI feed into a

common header that connects to each cold leg pipe downstream of the RCP discharge. The

ECCS pumped injection is modeled as a table of flow versus backpressure. This model also

describes the secondary-side steam generator that is instantaneously isolated (closed MSIV

and feedwater trip) at the time of the break.

As described in Appendix A, many parameters associated with LBLOCA phenomenological

uncertainties and plant operation ranges are sampled. A summary of those parameters sampled

is given in Table A-6. The LBLOCA phenomenological uncertainties are provided in Table A-7.

Values for process or operational parameters, including ranges of sampled process parameters,

and fuel design parameters used in the analysis are given in Table B-21. Plant data are

analyzed to develop uncertainties for the process parameters sampled in the analyses.

Table B-22 presents a summary of the uncertainties used in the analysis. Two parameters

(RWST temperature and diesel start time) are set at conservative bounding values for all

calculations.

Where applicable, the sampled parameter ranges are based on technical specification limits.

Plant data are used to define range boundaries for loop flow (high end) and containment

temperature (low end).

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B.4.3 Realistic Large Break LOCA Results

[

]

Table B-23 is a summary of the major parameters for the MRMC case. The results for the

MRMC case are presented in Table B-24. The fraction of total hydrogen generated was not

directly calculated; however, it is conservatively bounded by the calculated total percent

oxidation, which is well below the 1-percent limit. The event times for the MRMC case can be

found in Table B-25. [

]

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The heat transfer parameter range for the limiting margin case is provided in Table B-26.

Table B-27 [

]

The analysis plots for the minimum margin case are shown in Figure B−37 through Figure B−54.

Figure B−37 shows linear scatter plots of the key parameters sampled for the cases that lie in

the 95/95 range. Parameter labels appear to the left of each individual plot. These figures

illustrate the parameter ranges used in the analysis.

Figure B−38 and Figure B−39 show PCT scatter plots versus the time of PCT and versus break

size from the set of cases (LOCA events) that lie within the 95/95 range. The scatter plots for

the maximum local oxidation and total core-wide oxidation are shown in Figure B−40 and

Figure B−41, respectively. Figure B−42 through Figure B−53 show key parameters from the S-

RELAP5 calculations for the MRMC case. Figure B−42 is the plots of PCT, independent of

elevation. Figure B−54 compares the beginning of core recovery times for the set of cases that

lie within the 95/95 range to the BOCR time predicted using the MPR CCFL correlation. Note

that Figure B−54 uses the total break area, while previous plots used break area per side.

B.4.4 Conclusions

[

]

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Table B-21 CE 2x4 Plant Operating Range Supported by the LOCA Analysis (continued)

Event Operating Range 3.0 Accident Boundary Conditions a) Break location b) Break type c) Break size (each side, relative to cold leg

pipe area) d) Worst single-failure e) Offsite power f) ECCS pumped injection temperature g) HPSI pump delay

h) LPSI pump delay

i) Containment pressure j) Containment temperature k) Containment sprays delay l) Containment spray water temperature m) LPSI Flow

1 Determined prior to the execution of the set of [ ] cases. 2 Nominal containment pressure range is -0.7 to 0.5 psig. For RLBOCA, a reasonable value between this range is acceptable.

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Table B-23 CE 2x4 Summary of Major Parameters for the MRMC Case

Parameter Value Time in Cycle (hrs) Burnup (GWd/mtU) Core Power (MWt)

LHGR (kW/ft) Core Peaking (Equivalent Fq)

Radial Peak (FΔH) Axial Shape Index Local Peaking (Fl)

Break Type Break Size (ft2 / side)

Offsite Power Availability

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Table B-24 CE 2x4 COPERNIC Compliance with 10 CFR 50.46

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Table B-25 CE 2x4 Calculated Event Times for the MRMC Case

Event Time (sec)

Begin Analysis

Break Opens

RCP Trip

SIAS Issued

Start of Broken Loop SIT Injection

Start of Intact Loop SIT Injection (Loop 2, 3 and 4 respectively)

PCT Occurred

Start of HPSI

Start of Charging

Beginning of Core Recovery (Beginning of Reflood)

LPSI Available

Broken Loop LPSI Delivery Began

Intact Loops LPSI Delivery Began (Loop 2, 3, and 4 respectively)

Broken Loop HPSI Delivery Began

Intact Loops HPSI Delivery Began (Loop 2, 3, and 4 respectively)

Broken Loop SIT Emptied

Intact Loop SIT Emptied (Loop 2, 3, and 4 respectively)

Transient Calculation Terminated

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Table B-26 CE 2x4 Heat Transfer Parameters for the MRMC Case

Time (s)

LOCA Phase Early Blowdown Blowdown 1 Refill Reflood Quench Long Term

Cooling Heat

Transfer Mode

Heat Transfer

Correlations

Maximum LHGR (kW/ft)

Pressure (psia)

Core Inlet Mass Flux (lbm/s-ft2)

Vapor Reynolds Number3

Liquid Reynolds Number

Vapor Prandtl Number

Liquid Prandtl Number Vapor

Superheat4 (°F)

1 End of Blowdown considered as beginning of refill. 2 Conservatively biased per the methodology 3 Not important in pre-CHF heat transfer. 4 Vapor superheat is meaningless during blowdown and system depressurization.

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Table B-27 CE 2x4 Fuel Rod Rupture Ranges of Parameters for all [ ] cases

Parameter Name Minimum Value Maximum Value

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Figure B−37 CE 2x4 Scatter Plot of Operational Parameters for all [ ] cases

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Figure B−37 CE 2x4 Scatter Plot of Operational Parameters for all [ ] cases (continued)

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Figure B−38 CE 2x4 PCT versus PCT Time Scatter Plot from the Cases within the 95/95 Range

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Figure B−39 CE 2x4 PCT versus Break Size Scatter Plot from the Cases within the 95/95 Range

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Figure B−40 CE 2x4 Maximum Local Oxidation versus PCT Scatter Plot from the Cases within the 95/95 Range

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Figure B−41 CE 2x4 Total Oxidation versus PCT Scatter Plot from the Cases within the 95/95 Range

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Figure B−42 CE 2x4 Peak Cladding Temperature (Independent of Elevation) for the MRMC Case (COPERNIC)

ID:27305 3Oct2014 03:00:42 R5DMX

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Figure B−43 CE 2x4 Break Flow for the MRMC Case

ID:27305 3Oct2014 03:00:42 R5DMX

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Figure B−44 CE 2x4 Core Inlet Mass Flux for the MRMC Case

ID:27305 3Oct2014 03:00:42 R5DMX

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Figure B−45 CE 2x4 Core Outlet Mass Flux for the MRMC Case

ID:27305 3Oct2014 03:00:42 R5DMX

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Figure B−46 CE 2x4 Void Fraction at RCS Pumps for the MRMC Case

ID:27305 3Oct2014 03:00:42 R5DMX

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Figure B−47 CE 2x4 ECCS Flows (Includes SIT, Charging, SI and RHR) for the MRMC Case

ID:27305 3Oct2014 03:00:42 R5DMX

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Figure B−48 CE 2x4 Upper Plenum Pressure for the MRMC Case

ID:27305 3Oct2014 03:00:42 R5DMX

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Figure B−49 CE 2x4 Collapsed Liquid Level in the Downcomer for the MRMC Case

ID:27305 3Oct2014 03:00:42 R5DMX

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Figure B−50 CE 2x4 Collapsed Liquid Level in the Lower Plenum for the MRMC Case

ID:27305 3Oct2014 03:00:42 R5DMX

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Figure B−51 CE 2x4 Collapsed Liquid Level in the Core for the MRMC Case

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Figure B−52 CE 2x4 Containment and Loop Pressures for the MRMC Case

ID:27305 3Oct2014 03:00:42 R5DMX

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Figure B−53 CE 2x4 Pressure Difference between Upper Plenum and Downcomer for the MRMC Case

ID:27305 3Oct2014 03:00:42 R5DMX

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Figure B−54 CE 2x4 Validation of BOCR Time using MPR CCFL Correlation

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APPENDIX D TIME STEP SENSITIVITY

For the AREVA RLBLOCA methodology, solution convergence is demonstrated by

performing sensitivity studies in which the calculation time step is varied for three

appropriate plant designs. This approach demonstrates solution convergence while

recognizing that a certain degree of variability is to be expected. This sensitivity study

was performed in an earlier revision of EMF-2103 (Reference D-1), but the results and

conclusions are equally applicable to EMF-2103, Revision 3.

This sensitivity study was performed by first regenerating steady-state plant analysis

decks for three types of plants appropriate for this methodology, i.e., 3- and 4-loop

Westinghouse designs, and a CE design. These decks were then brought to typical

steady-state conditions, and a transient initiated with a DEG break with nominal

parameters, other than decay heat. Each transient used 120 percent of nominal decay

heat to drive the temperatures sufficiently high that code models would be challenged.

The recommended time step selection strategy is to set a single maximum time step

during the portions of the transient of most significance to safety, that is, the blowdown,

refill, and early reflood phases. The requested time step should then be increased

during late reflood when the flooding phenomena are reasonably stable. This approach

was found to provide a reasonable compromise between optimal numerical stability and

run time. It should be noted that the time step requested by the user is actually the

maximum time step allowed by the code for that time period, and that in fact the code

will reduce the requested time step should instability be detected. The nominal or base

case used a requested time step of 0.002 seconds from 0 to 400 seconds, and then

0.004 seconds from 400 to 600 seconds, 0.008 seconds from 600 to 800 seconds and

0.010 seconds beyond 800 seconds. Code convergence and stability at the nominal

time step of 0.002 seconds were demonstrated by incrementally varying the time step

from 0 to 400 seconds over a range from the nominal time step to an order of

magnitude smaller.

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[

] The nominal case for each of the designs

noted in the time step sensitivity study was repeated with this new time step and it was

determined that the code continued to proceed through the analysis with the requested

time steps, indicating code stability, with a minor deviation at the time of quench at the

core hot spot.

Figure D-1, Figure D-3, and Figure D-5 show the calculated PCTs from the 3-loop, 4-

loop, and CE studies, respectively. S-RELAP5 shows stability and convergence for all

design types during the blowdown period. During refill and early reflood, there is some

noticeable divergence in the results; however this has little impact on the PCT.

Figure D-2, Figure D-4, and Figure D-6 show the variability about the mean PCT from

the 3-loop, 4-loop, and CE studies, respectively. The data for these figures were

generated by averaging the calculated PCTs for each design, and then calculating the

maximum deviation, whether it is above or below the mean. As shown in these figures,

the nominal variability for the 3-loop design is approximately 15 K (27 °F), the 4-loop

design is approximately 12 K (21 °F), and the CE design is approximately 15 K (27 °F).

[

]

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Figure D-1: Time Step Sensitivity of Westinghouse 3-Loop Analysis

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Figure D-2: Variability of Westinghouse 3-Loop Analysis

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Figure D-3: Time Step Sensitivity of Westinghouse 4-Loop Analysis

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Figure D-4: Variability of Westinghouse 4-Loop Analysis

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Figure D-5: Time Step Sensitivity of CE Analysis

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Figure D-6: Variability of CE Analysis

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