Text

LLC, Submittal of Topical Report Loss-of-Coolant Accident Evaluation Model, TR-0516-49422, Revision 3
	ML23008A002
Person / Time
Site:	99902078, 05200050
Issue date:	01/08/2023
From:	Shaver M; NuScale
To:	; Office of Nuclear Reactor Regulation, Document Control Desk
Shared Package
ML23008A001	List: ML23008A002;
References
	LO-133399 TR-0516-49422-NP, Rev 3
	Download: ML23008A002 (1)
	v • d • e

LO-133399 January 8, 2023 Docket No.52-050 U.S. Nuclear Regulatory Commission ATTN: Document Control Desk One White Flint North 11555 Rockville Pike Rockville, MD 20852-2738

SUBJECT:

NuScale Power, LLC Submittal of Topical Report Loss-of-Coolant Accident Evaluation Model, TR-0516-49422, Revision 3

REFERENCES:

1. Letter from NuScale to NRC, NuScale Power, LLC Submittal of the Approved Version of NuScale Topical Report, Loss-of-Coolant Accident Evaluation Model, TR-0516-49422, Revision 2, dated July 7, 2020 (ML20189A644)

NuScale Power, LLC (NuScale) hereby submits Revision 3 of the Loss-of-Coolant Accident Evaluation Model, TR-0516-49422. The purpose of this submittal is to request that the NRC review and approve the changes from the previously approved version in Reference 1.

NuScale respectfully requests that the acceptance review be completed in 60 days from the date of transmittal. contains the proprietary version of the report entitled Loss-of-Coolant Accident Evaluation Model, TR-0516-49422, Revision 3. NuScale requests that the proprietary version be withheld from public disclosure in accordance with the requirements of 10 CFR § 2.390. The enclosed affidavit (Enclosure 3) supports this request. Enclosure 1 has also been determined to contain Export Controlled Information. This information must be protected from disclosure per the requirements of 10 CFR § 810. Enclosure 2 contains the nonproprietary version of the report.

This letter makes no regulatory commitments and no revisions to any existing regulatory commitments.

If you have any questions, please contact Thomas Griffith at 541-452-7813 or tgriffith@nuscalepower.com.

Sincerely, Mark W. Shaver Manager, Licensing NuScale Power, LLC NuScale Power, LLC 1100 NE Circle Blvd., Suite 200 Corvallis, Oregon 97330 Office 541.360.0500 Fax 541.207.3928 www.nuscalepower.com

LO-133399 Page 2 of 2 01/08/23 Distribution: Michael Dudek, NRC Getachew Tesfaye, NRC Bruce Bavol, NRC Enclosure 1: Loss-of-Coolant Accident Evaluation Model, TR-0516-49422, Revision 3, Proprietary Version Enclosure 2: Loss-of-Coolant Accident Evaluation Model, TR-0516-49422, Revision 3, Nonproprietary Version Enclosure 3: Affidavit of Mark W. Shaver, AF-133400 NuScale Power, LLC 1100 NE Circle Blvd., Suite 200 Corvallis, Oregon 97330 Office 541.360.0500 Fax 541.207.3928 www.nuscalepower.com

LO-133399 :

Loss-of-Coolant Accident Evaluation Model, TR-0516-49422, Revision 3, Proprietary Version NuScale Power, LLC 1100 NE Circle Blvd., Suite 200 Corvallis, Oregon 97330 Office 541.360.0500 Fax 541.207.3928 www.nuscalepower.com

LO-133399 :

Loss-of-Coolant Accident Evaluation Model, TR-0516-49422, Revision 3, Nonproprietary Version NuScale Power, LLC 1100 NE Circle Blvd., Suite 200 Corvallis, Oregon 97330 Office 541.360.0500 Fax 541.207.3928 www.nuscalepower.com

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Licensing Topical Report Loss-of-Coolant Accident Evaluation Model December 2022 Revision 3 Docket: 52-050 NuScale Power, LLC 1100 NE Circle Blvd., Suite 200 Corvallis, Oregon 97330 www.nuscalepower.com

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Licensing Topical Report COPYRIGHT NOTICE This report has been prepared by NuScale Power, LLC and bears a NuScale Power, LLC, copyright notice. No right to disclose, use, or copy any of the information in this report, other than by the U.S. Nuclear Regulatory Commission (NRC), is authorized without the express, written permission of NuScale Power, LLC.

The NRC is permitted to make the number of copies of the information contained in this report that is necessary for its internal use in connection with generic and plant-specific reviews and approvals, as well as the issuance, denial, amendment, transfer, renewal, modification, suspension, revocation, or violation of a license, permit, order, or regulation subject to the requirements of 10 CFR 2.390 regarding restrictions on public disclosure to the extent such information has been identified as proprietary by NuScale Power, LLC, copyright protection notwithstanding. Regarding nonproprietary versions of these reports, the NRC is permitted to make the number of copies necessary for public viewing in appropriate docket files in public document rooms in Washington, DC, and elsewhere as may be required by NRC regulations.

Copies made by the NRC must include this copyright notice and contain the proprietary marking if the original was identified as proprietary.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Licensing Topical Report Department of Energy Acknowledgement and Disclaimer This material is based upon work supported by the Department of Energy under Award Number DE-NE0008928.

This report was prepared as an account of work sponsored by an agency of the United States Government. Neither the United States Government nor any agency thereof, nor any of their employees, makes any warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, completeness, or usefulness of any information, apparatus, product, or process disclosed, or represents that its use would not infringe privately owned rights.

Reference herein to any specific commercial product, process, or service by trade name, trademark, manufacturer, or otherwise does not necessarily constitute or imply its endorsement, recommendation, or favoring by the United States Government or any agency thereof. The views and opinions of authors expressed herein do not necessarily state or reflect those of the United States Government or any agency thereof.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table of Contents Abstract . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1 Executive Summary. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3 1.0 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6 1.1 Purpose . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6 1.2 Scope . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6 1.3 Abbreviations and Definitions. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8 2.0 Background . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11 2.1 Loss-of-Coolant Accident Evaluation Model Roadmap . . . . . . . . . . . . . . . . . . . . . . . . . 11 2.2 Regulatory Requirements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 2.2.1 10 CFR 50.46 Loss-of-Coolant Accident Acceptance Criteria . . . . . . . . . . . . . . 16 2.2.2 NuScale Loss-of-Coolant Accident Evaluation Model Acceptance Criteria . . . . 16 2.2.3 10 CFR 50 Appendix K. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 17 2.2.4 Other Requirements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 34 2.2.5 Compliance with Containment Response Analysis Related Regulatory Requirements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 34 3.0 NuScale Power Module Description and Operations . . . . . . . . . . . . . . . . . . . . . . . . 48 3.1 General Plant Design . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 48 3.2 Plant Operation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 51 3.3 Safety-Related System Operation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 52 3.3.1 Emergency Core Cooling System . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53 3.3.2 Decay Heat Removal System . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 54 4.0 Phenomena Identification and Ranking. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55 4.1 Phenomena Identification and Ranking Process . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55 4.2 Loss-of-Coolant Accident Scenarios . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 56 4.3 Figures of Merit . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 58 4.4 Definitions of Importance and Knowledge Level Rankings . . . . . . . . . . . . . . . . . . . . . . 58 4.5 Systems, Structures, and Components . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 59 4.6 High-Ranked Phenomena . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 59 4.6.1 Discussion of LOCA Phenomena Ranked High Importance . . . . . . . . . . . . . . . 63 4.6.2 Discussion of IORV Phenomena Ranked High Importance . . . . . . . . . . . . . . . . 67 4.6.3 Treatment of Boron Transport Phenomena Originating in Phase 1a, 1b . . . . . . 69 4.7 Phenomena Identification and Ranking Table Summary . . . . . . . . . . . . . . . . . . . . . . . . 69

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table of Contents 4.8 LOCA Containment Response PIRT Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 70 4.9 Non-Loss-of-Coolant Accident Event Phenomena Identification and Ranking Table Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 70 5.0 Evaluation Model Description. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 71 5.1 NRELAP5 Loss-of-Coolant Accident Model for the NuScale Power Module . . . . . . . . . 71 5.1.1 General Model Nodalization . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 71 5.1.2 Reactor Coolant System . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 75 5.1.3 Helical Coil Steam Generators . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 80 5.1.4 Containment Vessel and Reactor Pool . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 81 5.1.5 Chemical and Volume Control System. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 82 5.1.6 Secondary System . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 82 5.1.7 Decay Heat Removal System . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 83 5.1.8 NRELAP5 Modeling Options . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 83 5.1.9 Time Step Size Control . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 86 5.2 Analysis Setpoints and Trips . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 87 5.3 Initial Plant Conditions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 88 5.4 Loss-of-Coolant Accident Break Spectrum . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 89 5.4.1 Break Location . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 89 5.4.2 Break Configuration and Size. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 89 5.4.3 Single Failures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 91 5.4.4 Loss of Power. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 91 5.4.5 Decay Heat Removal System Availability . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 92 5.5 Containment Response Analysis Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 93 5.6 Containment Response Analysis M&E Release Model . . . . . . . . . . . . . . . . . . . . . . . . . 93 5.6.1 NRELAP5 Primary Release Event Analysis Model . . . . . . . . . . . . . . . . . . . . . . 93 5.6.2 NRELAP5 Secondary System Mass and Energy Release Analysis Model . . . 101 5.7 Sensitivity Studies . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 108 6.0 NRELAP5 Code Description . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 109 6.1 Quality Assurance Requirements. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 110 6.2 NRELAP5 Hydrodynamic Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 111 6.2.1 Field Equations. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 111 6.2.2 State Relations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 114

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table of Contents 6.2.3 Flow Regime Maps. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 115 6.2.4 Momentum Closure Relations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 118 6.2.5 Heat Transfer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 124 6.3 Heat Structure Models . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 126 6.4 Point Reactor Kinetics Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 128 6.5 Trips and Control System Models . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 130 6.6 Special Solution Techniques . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 130 6.6.1 Choked Flow . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 130 6.6.2 Abrupt Area Change. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 134 6.6.3 Counter Current Flow Limitation. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 135 6.7 Helical Coil Steam Generator Component . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 137 6.7.1 Helical Coil Tube Friction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 137 6.7.2 Helical Coil Tube Heat Transfer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 139 6.8 Wall Heat Transfer and Condensation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 140 6.8.1 NRELAP5 Solution Approach for Wall Condensation Heat Transfer . . . . . . . . 141 6.8.2 Wall Condensation Correlation. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 145 6.9 Interfacial Drag in Large Diameter Pipes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 147 6.10 Fission Decay Heat and Actinide Models. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 148 6.11 Critical Heat Flux Models . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 151 6.11.1 ((2(a),(c) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 152 6.11.2 Implementation of Critical Heat Flux correlations. . . . . . . . . . . . . . . . . . . . . . . 154 6.11.3 (( }}2(a),(c) . . . . . . . . . . . . . . . . . . . . . . . . . 154 6.11.4 (( }}2(a),(c) . . . . . . . . . . . . . . . . . . 158 6.11.5 NSPN-1 Critical Heat Flux Correlation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 159 7.0 NRELAP5 Assessments . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 177 7.1 Assessment Methodology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 177 7.2 Legacy Test Data . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 179 7.2.1 Ferrell-McGee . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 179 7.2.2 GE Level Swell (1 ft). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 183 7.2.3 GE Level Swell (4 ft). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 191 7.2.4 KAIST . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 197 © Copyright 2022 by NuScale Power, LLC vi

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table of Contents 7.2.5 FRIGG . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 205 7.2.6 FLECHT-SEASET . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 212 7.2.7 SemiScale (S-NC-02 and S-NC-10). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 218 7.2.8 Wilson Bubble Rise . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 223 7.2.9 Marviken Jet Impingement Test (JIT) 11 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 227 7.2.10 Bankoff Perforated Plate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 231 7.2.11 Marviken Critical Flow Test 22 and 24 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 233 7.3 NuScale Stern Critical Heat Flux Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 241 7.3.1 Facility Description . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 242 7.3.2 Experimental Procedure. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 244 7.3.3 Phenomenon Addressed . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 244 7.3.4 Parameter Ranges Assessed. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 244 7.3.5 Special Analysis Techniques . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 245 7.3.6 Assessment Results. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 245 7.4 NuScale SIET Steam Generator Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 246 7.4.1 SIET Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 246 7.4.2 SIET Fluid-Heated Test . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 256 7.5 NuScale NIST Test Assessment Cases. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 266 7.5.1 Test Facility Description . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 267 7.5.2 Facility NRELAP5 Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 274 7.5.3 Facility Test Matrix . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 275 7.5.4 Separate Effect High Pressure Condensation Tests (NIST-1) . . . . . . . . . . . . . 277 7.5.5 Natural Circulation Test at Power. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 291 7.5.6 RCS Discharge Line Loss-of-Coolant Accident Integral Effects Tests (NIST-1) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 296 7.5.7 NIST-1 Pressurizer Spray Supply Line Loss-of-Coolant Accident Integral Effects Test . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 316 7.5.8 Spurious Reactor Vent Valve Opening Tests. . . . . . . . . . . . . . . . . . . . . . . . . . 324 7.5.9 Spurious Reactor Recirculation Valve Opening Integral Effects Test . . . . . . . 340 7.5.10 NIST-2 LOCA Integral Effects Test Series . . . . . . . . . . . . . . . . . . . . . . . . . . . . 347 7.5.11 NIST-2 IORV Integral Effects Test Series . . . . . . . . . . . . . . . . . . . . . . . . . . . . 379 7.6 Containment Response Methodology Assessment . . . . . . . . . . . . . . . . . . . . . . . . . . . 391 © Copyright 2022 by NuScale Power, LLC vii

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table of Contents 8.0 Assessment of Evaluation Model Adequacy. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 392 8.1 Adequacy Demonstration Overview. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 392 8.2 Evaluation of Models and Correlations (Bottom-Up Assessment) . . . . . . . . . . . . . . . . 392 8.2.1 Important Models and Correlations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 394 8.2.2 ((

                               }}2(a),(c) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 403 8.2.3     ((
                               }}2(a),(c) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 408 8.2.4     ((                                                                                              }}2(a),(c). . . . . 410 8.2.5     ((                                                                                       }}2(a),(c) . . . . . . . . 414 8.2.6     ((                                                                                }}2(a),(c). . . . . . . . . . . . 415 8.2.7     ((                                            }}2(a),(c) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 420 8.2.8     ((                                                       }}2(a),(c) . . . . . . . . . . . . . . . . . . . . . . . . 421 8.2.9     ((                                                              }}2(a),(c). . . . . . . . . . . . . . . . . . . . . 427 8.2.10 Flashing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 428 8.2.11 ((                                        }}2(a),(c) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 429 8.2.12 ((                                                                                   }}2(a),(c) . . . . . . . . . . . . 434 8.2.13 ((                                                                                         }}2(a),(c). . . . . . . . . 435 8.2.14 ((                                                                         }}2(a),(c) . . . . . . . . . . . . . . . . . 436 8.2.15 ((
                           }}2(a),(c) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 439 8.2.16 ((                                                                               }}2(a),(c) . . . . . . . . . . . . . . 440 8.2.17 ((                                               }}2(a),(c) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 440 8.2.18 ((                    }}2(a),(c) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 443 8.2.19 ((                                }}2(a),(c) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 445 8.2.20 ((                                                                  }}2(a),(c) . . . . . . . . . . . . . . . . . . . . 448 8.2.21 ((                                                                                   }}2(a),(c) . . . . . . . . . . . . 450 8.2.22 ((                                                                            }}2(a),(c) . . . . . . . . . . . . . . . 451 8.3      Evaluation of Integral Performance (Top-Down Assessment) . . . . . . . . . . . . . . . . . . . 452 8.3.1     Review of Code Governing Equations and Numerics . . . . . . . . . . . . . . . . . . . 453 8.3.2     NuScale Facility Scaling. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 455 8.3.3     Assessment of NuScale Facility Integral Effects Test Data . . . . . . . . . . . . . . . 477

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table of Contents 8.3.4 Evaluation of NuScale Integral Effects Tests Distortions and NRELAP5 Scalability . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 479 8.3.5 Calculation of Peak CNV pressure. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 488 8.4 Summary of Adequacy Findings . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 489 8.4.1 Findings from Bottom-Up Evaluation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 489 8.4.2 Findings from Top-Down Evaluation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 500 8.4.3 Summary of Biases and Uncertainties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 506 9.0 Loss-of-Coolant Accident Calculations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 507 9.1 Loss-of-Coolant Accident Progression in the NuScale Power Module . . . . . . . . . . . . 507 9.1.1 Liquid Space Break Phase 0 Analysis (NPM-20) . . . . . . . . . . . . . . . . . . . . . . . 508 9.1.2 Steam Space Break Phase 0 Analysis. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 523 9.1.3 Phase 1 LOCA Results Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 530 9.2 Break Size . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 549 9.3 Decay Heat Removal System . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 558 9.4 Power Availability . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 558 9.5 Single Failure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 561 9.6 Sensitivity Studies . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 568 9.6.1 Model Nodalization. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 568 9.6.2 Time-Step Size Selection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 572 9.6.3 Counter Current Flow Limitation Behavior on Pressurizer Baffle Plate . . . . . . 575 9.6.4 Emergency Core Cooling System Valve Parameters (NPM-160 only) . . . . . . 575 9.6.5 Initial Reactor Pool Temperature . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 578 9.6.6 Core Power Distribution . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 580 9.6.7 DHRS Capacity . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 584 9.7 Primary and Secondary System Release Scenario Containment Response Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 589 9.7.1 Analysis Approach . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 589 9.7.2 Base Case Analysis and Sensitivity Results . . . . . . . . . . . . . . . . . . . . . . . . . . 590 9.7.3 Primary Release Scenario Pressure and Temperature Results. . . . . . . . . . . . 591 9.7.4 Main Steamline Break Containment Pressure and Temperature Results . . . . 629 9.7.5 Feedwater Line Break Containment Pressure and Temperature Results . . . . 636 © Copyright 2022 by NuScale Power, LLC ix

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table of Contents 9.8 Inadvertent Opening of Reactor Pressure Vessel Valves (IORV) . . . . . . . . . . . . . . . . 643 9.8.1 Phase 0 IORV Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 644 9.8.2 Phase 1 IORV Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 656 9.9 Loss-of-Coolant Accident and Inadvertent Opening of Reactor Pressure Vessel Valve Calculation Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 682 10.0 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 684 11.0 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 688 © Copyright 2022 by NuScale Power, LLC x

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Tables Table 1-1 Abbreviations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8 Table 1-2 Definitions. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10 Table 2-1 Evaluation Model Development and Assessment Process Steps and the Associated Sections in this Document . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19 Table 2-3 Compliance with Design Specific Review Standard Section 6.2.1. . . . . . . . . . . 36 Table 2-4 Compliance with Design Specific Review Standard Section 6.2.1.1.A . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 38 Table 2-5 Compliance with Design Specific Review Standard Section 6.2.1.3 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 40 Table 2-6 Compliance with Design Specific Review Standard Section 6.2.1.4 . . . . . . . . . 45 Table 4-1 Importance Rankings . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 59 Table 4-2 Knowledge Levels . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 59 Table 4-3 Systems, Structures, and Components . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 59 Table 4-4 LOCA High-Ranked Phenomena . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 60 Table 4-5 Inadvertent Opening of a Reactor Valve Newly High-Ranked Phenomena Relative to LOCA PIRT . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 62 Table 5-1 Default Junction Options for the NRELAP5 Loss-of-Coolant Accident Model . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 84 Table 5-2 Default Volume Options for the NRELAP5 Loss-of-Coolant Accident Model. . . 85 Table 5-3 Typical NuScale Power Module Safety-Related System Measurement Parameters. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 87 Table 5-4 Plant Initial Conditions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 88 Table 5-5 Summary of Analyzed Break Sizes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 90 Table 5-6 NuScale Power Module Valve Fail-Safe Positions with Loss of DC Power . . . . 92 Table 5-7 LOCA Versus Containment Response Models . . . . . . . . . . . . . . . . . . . . . . . . . 93 Table 5-8 Containment Vessel and Reactor Pool Heat Transfer Modeling . . . . . . . . . . . . 96 Table 5-9 Primary System Initial Conditions. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 97 Table 5-10 Containment Vessel and Reactor Pool Initial Conditions . . . . . . . . . . . . . . . . . . 98 Table 5-11 Containment Vessel Peak Pressure and Temperature Calculation Assumptions. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 99 Table 5-12 Secondary System Initial Conditions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 105 Table 5-13 Boundary Conditions for the Main Steam Line Break Containment Response Analysis Methodology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 106 Table 6-1 Extended Shah Dimensionless Vapor Velocity Transition Criteria. . . . . . . . . . 147 © Copyright 2022 by NuScale Power, LLC xi

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Tables Table 6-2 Extended Shah Condensation Heat Transfer Coefficients Dependent on Regime . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 147 Table 6-3 ANS 1973 11-Group Fission Decay Constants . . . . . . . . . . . . . . . . . . . . . . . . 148 Table 6-4 ANS-79 Actinide Model Constants . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 150 Table 6-5 Coefficient of Revised Pressure Correction Term in Equation 6-108. . . . . . . . 156 Table 6-6 (( }}2(a),(c) Critical Heat Flux Correlation Application Range. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 158 Table 6-7 NSPN-1 CHF Correlation Coefficients . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 163 Table 6-8 NSPN-1 Range of Conditions. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 166 Table 6-9 Validation Set Statistics for the 3-Fold Cross-Validation . . . . . . . . . . . . . . . . . 169 Table 6-10 Subsets of NuFuel-HTP2' Data for NSPN-1 . . . . . . . . . . . . . . . . . . . . . . . . . 170 Table 6-11 Correlation Limit by Sub-Region for NSPN-1 CHF Correlation . . . . . . . . . . . . 170 Table 7-1 NRELAP5 Loss-of-Coolant Accident Assessment Matrix . . . . . . . . . . . . . . . . 177 Table 7-2 Summary of Ferrell-McGee Experimental Test Data Range . . . . . . . . . . . . . . 181 Table 7-3 Summary of GE 1 ft. Vessel Level Swell Experiments. . . . . . . . . . . . . . . . . . . 186 Table 7-4 Range of KAIST Test Data. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 201 Table 7-5 Range of Stern Steady State Critical Heat Flux Data. . . . . . . . . . . . . . . . . . . . 245 Table 7-6 Facility High Priority Tests for NRELAP5 Code Validation. . . . . . . . . . . . . . . . 275 Table 7-7 LOCA Characteristic State-Points Considered for Runs 1 and 3 . . . . . . . . . . . 350 Table 7-8 LOCA ECCS Actuation Sequences . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 350 Table 7-9 LOCA Characteristic State-Points Considered for Runs 4, 6, and 7 . . . . . . . . 351 Table 7-10 LOCA ECCS Actuation Sequences . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 351 Table 7-11 LOCA Run 1 Sequence of Events . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 352 Table 7-12 LOCA Run 2 Sequence of Events . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 363 Table 7-13 LOCA Run 5 Sequence of Events. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 363 Table 7-14 LOCA Run 3 Sequence of Events . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 368 Table 7-15 LOCA Run 4 Sequence of Events . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 371 Table 7-16 LOCA Run 6 Sequence of Events . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 372 Table 7-17 LOCA Run 7 Sequence of Events . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 372 Table 8-1 Dominant NRELAP5 Models and Correlations. . . . . . . . . . . . . . . . . . . . . . . . . 394 Table 8-2 Process Parameter Ranges for EM Applicability Evaluation . . . . . . . . . . . . . . 397 Table 8-3 Geometric Parameters for NPM-160 and NPM-20. . . . . . . . . . . . . . . . . . . . . . 400 Table 8-4 Marviken Range of Parameters Compared to the NuScale Power Module . . . 405 © Copyright 2022 by NuScale Power, LLC xii

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Tables Table 8-5 Ferrell-McGee Range of Parameters Compared to the NuScale Power Module . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 409 Table 8-6 Dimensions of NuScale Power Module, NIST-1 and Bankoff Pressurizer Plate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 412 Table 8-7 Range of Riser Interphase Friction - Separate Effects Tests and NuScale Power Module . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 418 Table 8-8 Ranges of Key Parameters for Core Interphase Friction - Separate Effects Tests and Plant . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 442 Table 8-9 Range of Key Parameters for Core Flow - Separate Effects Tests and Plant . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 444 Table 8-10 Range of Key Parameters for Core Boiling - Separate effects tests and plant . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 447 Table 8-11 Range of Key Parameters for Subcooling Boiling and Separate Effects Tests and Plant . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 451 Table 8-12 Scaling Factors for NIST Facility . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 462 Table 8-13 Mass Flow Paths for NPM and NIST (RCS and CNV) . . . . . . . . . . . . . . . . . . . 471 Table 8-14 Heat Flow Paths for RCS in NPM and NIST . . . . . . . . . . . . . . . . . . . . . . . . . . 471 Table 8-15 Heat Flow Paths for Containment in NPM and NIST . . . . . . . . . . . . . . . . . . . . 472 Table 8-16 Description of Groups for the RCS Mass/Energy Balance. . . . . . . . . . . . . . . . 473 Table 8-17 Description of Groups for the Containment Mass/Energy Balance . . . . . . . . . 474 Table 8-18 Summary of Bottom-Up Evaluation of NRELAP5 Models and Correlations. . . 491 Table 8-19 Applicability Summary for High-Ranked Phenomena . . . . . . . . . . . . . . . . . . . 500 Table 9-1 Discharge Line Break MCHFR Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 510 Table 9-2 Injection Line Break MCHFR Results. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 510 Table 9-3 High Point Vent and Pressurizer Spray Line Break MCHFR Results. . . . . . . . 523 Table 9-4 Summary of Phase 1 Limiting LOCA Parameters by Break Type . . . . . . . . . . 531 Table 9-5 RCS Injection Line Sequence of Events (Limiting Collapsed Level Case). . . . 534 Table 9-6 RCS High Point Vent Line Sequence of Events - Limiting Collapsed Level Case . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 544 Table 9-7 Number of Volumes in Reactor Pressure Vessel and Containment Vessel Nodalization . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 569 Table 9-8 Primary System Mass and Energy Release Scenarios . . . . . . . . . . . . . . . . . . 589 Table 9-9 Initial Conditions for Mass and Energy Release Event Analyses. . . . . . . . . . . 591 Table 9-10 Case 1 Sequence of Events - Reactor Coolant System Discharge Line Break Loss-of-Coolant Accident. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 593 © Copyright 2022 by NuScale Power, LLC xiii

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Tables Table 9-11 Spectrum of Discharge Line Breaks Evaluated for Peak CNV Temperature and Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 594 Table 9-12 Spectrum of Injection Line Breaks Evaluated for Peak CNV Temperature and Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 595 Table 9-13 Spectrum of High Point Vent Line Breaks Evaluated for Peak CNV Temperature and Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 596 Table 9-14 Spectrum of Reactor Vent Valve Opening Events Evaluated for Peak CNV Temperature and Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 597 Table 9-15 Spectrum of Reactor Recirculation Valve Opening Events Evaluated for Peak CNV Temperature and Pressure. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 598 Table 9-16 Spectrum of Inadvertent ECCS Actuation Events Evaluated for Peak CNV Temperature and Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 599 Table 9-17 Spectrum of Main Steam Line Break Events Evaluated for Peak CNV Temperature and Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 631 Table 9-18 Limiting Secondary Break Containment Peak Pressure and Temperature (Case SLB-3) - Sequence of Events . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 631 Table 9-19 Spectrum of Feedwater Line Break Events Evaluated for Peak CNV Temperature and Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 638 Table 9-20 Representative Results of IORV Event Analyses. . . . . . . . . . . . . . . . . . . . . . . 645 Table 9-21 Sequence of Events for Inadvertent RVV Opening for NPM-20. . . . . . . . . . . . 658 Table 9-22 Sequence of Events for Inadvertent RRV Opening for NPM-20 . . . . . . . . . . . 659 Table 9-23 Summary of Phase 1 Limiting IORV Parameters . . . . . . . . . . . . . . . . . . . . . . . 659 © Copyright 2022 by NuScale Power, LLC xiv

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 2-1 Evaluation Model Development and Assessment Process . . . . . . . . . . . . . . . . 13 Figure 3-1 A Single NuScale Power Module during Normal Operation . . . . . . . . . . . . . . . . 49 Figure 3-2 Schematic of NuScale Power Module Decay Heat Removal System and Emergency Core Cooling System during Operation . . . . . . . . . . . . . . . . . . . . . 51 Figure 4-1 Timing of LOCA Phases. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 58 Figure 5-1 Noding Diagram of NRELAP5 Loss-of-Coolant Accident Input Model for the NPM-160 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 73 Figure 5-2 Noding Diagram of NRELAP5 Loss-of-Coolant Input Model for the NPM-20 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 74 Figure 5-3 Typical Non-LOCA Nodalization Diagram . . . . . . . . . . . . . . . . . . . . . . . . . . . . 102 Figure 6-1 Schematic of Vertical Flow-Regime Map Indicating Transitions . . . . . . . . . . . 117 Figure 6-2 NRELAP5 Boiling and Condensing Curves . . . . . . . . . . . . . . . . . . . . . . . . . . . 125 Figure 6-3 (( }}2(a),(c) . . . . . . . 142 Figure 6-4 NRELAP5 ANS 1973 Implemented Fission Decay Heat Curve . . . . . . . . . . . . 149 Figure 6-5 NRELAP5 ANS-79 Implemented Actinide Heat Curve. . . . . . . . . . . . . . . . . . . 151 Figure 6-6 (( }}2(a),(c) . . . . . . . . . . . . . . . . . . . . . . . . . 153 Figure 6-7 ((

                                              }}2(a),(c) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 157 Figure 6-8         Comparison of VIPRE-01 and NRELAP5 Quality and Mass Flux at the Measured CHF Location . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 165 Figure 6-9         Pressure, Mass Flux, and Quality Domain with Local CHF Contours . . . . . . . 166 Figure 6-10        Design Limit Statistical Methods Flow Chart . . . . . . . . . . . . . . . . . . . . . . . . . . 168 Figure 6-11        NSPN-1 Predicted vs. Measured CHF Binned by Test ID . . . . . . . . . . . . . . . . 171 Figure 6-12        Pressure Bias Plot for NSPN-1 CHF Correlation . . . . . . . . . . . . . . . . . . . . . . . 172 Figure 6-13        Mass Flux Bias Plot for NSPN-1 CHF Correlation . . . . . . . . . . . . . . . . . . . . . . 172 Figure 6-14        Quality Bias Plot for NSPN-1 CHF Correlation. . . . . . . . . . . . . . . . . . . . . . . . . 173 Figure 6-15        Cold Wall Factor Bias Plot for NSPN-1 CHF Correlation . . . . . . . . . . . . . . . . . 173 Figure 6-16        Boiling Length Bias Plot for NSPN-1 CHF Correlation . . . . . . . . . . . . . . . . . . . 174 Figure 6-17        Inlet Enthalpy Bias Plot for NSPN-1 CHF Correlation . . . . . . . . . . . . . . . . . . . 174 Figure 6-18        Pressure, Mass Flux, and Quality Domain Sorted by Correlation Performance. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 175 Figure 7-1         Schematic of the Ferrell-McGee Test Section . . . . . . . . . . . . . . . . . . . . . . . . . 180 Figure 7-2         Predicted Versus Measured Pressure Drop for Selected Contraction Tests. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 183 Figure 7-3         Schematic of the GE 1 ft. Blowdown Vessel . . . . . . . . . . . . . . . . . . . . . . . . . . 184

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 7-4 GE Level Swell 1 ft. Vessel Pressure Versus Time . . . . . . . . . . . . . . . . . . . . . 187 Figure 7-5 GE Level Swell 1 ft. Vessel Void Fraction Versus Elevation at 10 Seconds . . 188 Figure 7-6 GE Level Swell 1 ft. Vessel Void Fraction Versus Elevation at 40 Seconds . . 189 Figure 7-7 GE Level Swell 1 ft. Vessel Void Fraction Versus Elevation at 100 Seconds . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 190 Figure 7-8 GE Level Swell 1 ft. Vessel Void Fraction Versus Elevation at 160 Seconds . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 191 Figure 7-9 Schematic of the GE 4 ft. Blowdown Vessel . . . . . . . . . . . . . . . . . . . . . . . . . . 192 Figure 7-10 GE Level Swell 4-ft Vessel Pressure Versus Time . . . . . . . . . . . . . . . . . . . . . 194 Figure 7-11 GE Level Swell 4-ft Vessel Void Fraction Versus Elevation at 5 Seconds. . . . 195 Figure 7-12 GE Level Swell 4-ft Vessel Void Fraction Versus Elevation at 10 Seconds. . . 196 Figure 7-13 GE Level Swell 4-ft Vessel Void Fraction Versus Elevation at 20 Seconds. . . 197 Figure 7-14 Schematic of KAIST Test Facility . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 199 Figure 7-15 Schematic of the KAIST Test Section . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 200 Figure 7-16 Measured versus predicted heat transfer coefficient . . . . . . . . . . . . . . . . . . . . 202 Figure 7-17 KAIST and NRELAP5 Axial Heat Transfer Coefficient. . . . . . . . . . . . . . . . . . . 203 Figure 7-18 KAIST and NRELAP5 Axial Inner Wall Temperature . . . . . . . . . . . . . . . . . . . . 204 Figure 7-19 KAIST and NRELAP5 Axial Liquid Mass Flow Rate . . . . . . . . . . . . . . . . . . . . 205 Figure 7-20 FRIGG-4 Experimental Loop . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 207 Figure 7-21 FRIGG-4 36 Rod Test Section . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 208 Figure 7-22 FRIGG-4 Zones for Evaluation of Radial Void Distribution . . . . . . . . . . . . . . . 209 Figure 7-23 FRIGG Mean Void Data of NRELAP5 Versus Test 613123 Data . . . . . . . . . . 210 Figure 7-24 FRIGG Mean Void Data of NRELAP5 Versus Test 613130 Data . . . . . . . . . . 211 Figure 7-25 FRIGG Mean Void Data of NRELAP5 Versus Test 613010 data. . . . . . . . . . . 211 Figure 7-26 FRIGG Mean Void Data of NRELAP5 Versus Test 613118 Data . . . . . . . . . . 212 Figure 7-27 FLECHT-SEASET Experimental Facility . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 213 Figure 7-28 FLECHT-SEASET Level 1 Void Fraction Versus Time - Test 35557. . . . . . . . 214 Figure 7-29 FLECHT-SEASET Level 2 Void Fraction Versus Time - Test 35557. . . . . . . . 215 Figure 7-30 FLECHT-SEASET Level 3 Void Fraction Versus Time - Test 35557. . . . . . . . 215 Figure 7-31 FLECHT-SEASET Level 4 Void Fraction Versus Time - Test 35557. . . . . . . . 216 Figure 7-32 FLECHT-SEASET Level 1 Collapsed Water Level Versus Time - Test 35557 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 216 Figure 7-33 FLECHT-SEASET Level 2 Collapsed Water Level Versus Time - Test 35557 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 217 © Copyright 2022 by NuScale Power, LLC xvi

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 7-34 FLECHT-SEASET Level 3 Collapsed Water Level Versus Time - Test 35557 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 217 Figure 7-35 FLECHT-SEASET Level 4 Collapsed Water Level Versus Time - Test 35557 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 218 Figure 7-36 Semiscale Mod-2A Single (Intact) Loop Test Facility Configuration . . . . . . . . 219 Figure 7-37 S-NC-2 30 kW Average Mass Flow Rate Versus Percent Inventory . . . . . . . . 221 Figure 7-38 S-NC-2 60 kW Average Mass Flow Rate Versus Percent Inventory . . . . . . . . 222 Figure 7-39 S-NC-10 100 kW Average Mass Flow Rate Versus Percent Inventory . . . . . . 222 Figure 7-40 Schematic of Wilson Bubble Rise Test Facility . . . . . . . . . . . . . . . . . . . . . . . . 223 Figure 7-41 NRELAP5 and Wilson Void Fraction Versus Superficial Velocity at 600 psig (4.14 MPa) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 225 Figure 7-42 NRELAP5 and Wilson Void Fraction Versus Superficial Velocity 1,000 psig (6.89 MPa) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 225 Figure 7-43 NRELAP5 and Wilson Void Fraction Versus Superficial Velocity 2,000 psig (13.8 MPa) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 226 Figure 7-44 Predicted Versus Measured Area Averaged Void Fraction (all cases) . . . . . . 226 Figure 7-45 Marviken Jet Impingement Test Facility . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 228 Figure 7-46 Marviken Jet Impingement Test 11 Flowrate . . . . . . . . . . . . . . . . . . . . . . . . . . 230 Figure 7-47 Marviken Jet Impingement Test 11 Density . . . . . . . . . . . . . . . . . . . . . . . . . . . 231 Figure 7-48 Schematic of Bankoff Counter Current Flow Apparatus (from Reference 68). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 232 Figure 7-49 Superficial Vapor Velocity Versus Superficial Liquid Velocity . . . . . . . . . . . . . 233 Figure 7-50 Schematic of the Marviken Pressure Vessel . . . . . . . . . . . . . . . . . . . . . . . . . . 234 Figure 7-51 Discharge Pipe Dimensions and Instrument Locations . . . . . . . . . . . . . . . . . . 235 Figure 7-52 Measured Versus Calculated Mass Flow Rate for Marviken Critical Flow Test 22 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 238 Figure 7-53 Marviken Critical Flow Test 22 Comparison to Calculated Mixture Density . . . 239 Figure 7-54 Measured Versus Calculated Mass Flow Rate for Marviken Critical Flow Test 24 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 240 Figure 7-55 Marviken Critical Flow Test 24 Mixture Density and Calculated Mixture Density . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 241 Figure 7-56 U1 & C1 (left) versus U2 (right) radial layout . . . . . . . . . . . . . . . . . . . . . . . . . . 242 Figure 7-57 Stern Test Section Axial Layout . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 243 Figure 7-58 Predicted Versus Measured Stern Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . 246 Figure 7-59 SIET Electrically-Heated Test Instrumentation Diagram . . . . . . . . . . . . . . . . . 248 Figure 7-60 Time Averaged Wall Temperature Profile for Coil 2 Test TD0015 . . . . . . . . . . 250 © Copyright 2022 by NuScale Power, LLC xvii

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 7-61 Time averaged wall temperature profile for coil 2 test TD0003 . . . . . . . . . . . . 251 Figure 7-62 SIET Electrically-Heated Test Differential Pressure for all Coil 1 Diabatic Tests. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 252 Figure 7-63 SIET Electrically-Heated Test Differential Pressure for all Coil 2 Diabatic Tests. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 253 Figure 7-64 SIET Electrically-Heated Test Differential Pressure for all Coil 3 Diabatic Tests. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 254 Figure 7-65 SIET Electrically-Heated Test Fluid Temperatures for all Coil 1 Diabatic Tests. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 255 Figure 7-66 SIET Electrically-Heated Test Wall Temperature for all Coil 1 Diabatic Tests. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 256 Figure 7-67 SIET Fluid-Heated Test Adiabatic Primary Differential Pressure . . . . . . . . . . . 259 Figure 7-68 SIET Fluid-Heated Test Diabatic Test Primary Differential Pressure. . . . . . . . 260 Figure 7-69 SIET Fluid-Heated Test Diabatic Test Primary Temperature. . . . . . . . . . . . . . 261 Figure 7-70 Comparison of Wall Temperatures in TD0001 (Case 1A) . . . . . . . . . . . . . . . . 262 Figure 7-71 Comparison of Wall Temperatures in TD0005 (Case 1A) . . . . . . . . . . . . . . . . 263 Figure 7-72 Comparison of Wall Temperatures in TD0015 (Case 1A) . . . . . . . . . . . . . . . . 264 Figure 7-73 Comparison of Primary and Secondary Side Fluid Temperatures in TD0001 (Case 1A) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 265 Figure 7-74 Comparison of Primary and Secondary Side Fluid Temperatures in TD0005 (Case 1A) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 266 Figure 7-75 Schematic of NuScale Integral Test Facility and NRELAP5 Nodalization . . . . 269 Figure 7-76 HP-02 Run 1 Containment Vessel Pressure Response. . . . . . . . . . . . . . . . . . 280 Figure 7-77 HP-02 Run 1 Containment Vessel Collapsed Level Response . . . . . . . . . . . . 281 Figure 7-78 HP-02 Run 1 Upper Containment Vessel Fluid Temperature Response (in vapor space). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 282 Figure 7-79 HP-02 Run 1 Upper Cooling Pool Vessel Temperature Response . . . . . . . . . 283 Figure 7-80 HP-02 Run 2 Containment Vessel Pressure Response. . . . . . . . . . . . . . . . . . 284 Figure 7-81 HP-02 Run 2 Containment Vessel Collapsed Level Response . . . . . . . . . . . . 285 Figure 7-82 HP-02 Run 2 Upper Containment Vessel Fluid Temperature Response (in vapor space). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 286 Figure 7-83 HP-02 Run 2 Upper Cooling Pool Temperature Response . . . . . . . . . . . . . . . 287 Figure 7-84 HP-02 Run 3 Containment Vessel Pressure Response. . . . . . . . . . . . . . . . . . 288 Figure 7-85 HP-02 Run 3 Containment Vessel Collapsed Level Response . . . . . . . . . . . . 289 Figure 7-86 HP-02 Run 3 Upper Containment Vessel Fluid Temperature Response (in vapor space). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 290 © Copyright 2022 by NuScale Power, LLC xviii

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 7-87 HP-02 Run 3 Upper Cooling Pool Temperature Response . . . . . . . . . . . . . . . 291 Figure 7-88 HP-05 NIST-1 Averaged Mass Flowrate and NRELAP5 Results. . . . . . . . . . . 294 Figure 7-89 HP-05 NIST-1 Averaged Core Inlet Temperature and NRELAP5 Results. . . . 295 Figure 7-90 HP-05 NIST-1 Averaged Core Outlet Temperature and NRELAP5 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 296 Figure 7-91 NIST-1 HP-06 NRELAP5 Chemical and Volume Control System Discharge Line Break Mass Flow Rate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 299 Figure 7-92 NIST-1 HP-06 Break Orifice Differential Pressure . . . . . . . . . . . . . . . . . . . . . . 300 Figure 7-93 NIST-1 HP-06 Primary Mass Flow Rate. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 301 Figure 7-94 NIST-1 HP-06 Pressurizer Level Comparison . . . . . . . . . . . . . . . . . . . . . . . . . 302 Figure 7-95 NIST-1 HP-06 Reactor Pressure Vessel Level Comparison . . . . . . . . . . . . . . 303 Figure 7-96 NIST-1 HP-06 Containment Vessel Level Comparison . . . . . . . . . . . . . . . . . . 304 Figure 7-97 HP-06 Containment Vessel Pressure Comparison . . . . . . . . . . . . . . . . . . . . . 305 Figure 7-98 NIST-1 HP-06 Containment Vessel Pressure Comparison . . . . . . . . . . . . . . . 306 Figure 7-99 NIST-1 HP-06 Primary Pressure Comparison . . . . . . . . . . . . . . . . . . . . . . . . . 307 Figure 7-100 Comparison of Core Power in HP-06 and HP-06b Tests with the NuScale Power Module Decay Power after Reactor Trip (scaled) . . . . . . . . . . . . . . . . . 308 Figure 7-101 NIST-1 HP-06b Primary Pressure Comparison . . . . . . . . . . . . . . . . . . . . . . . . 309 Figure 7-102 NIST-1 HP-06b Containment Vessel Pressure Comparison . . . . . . . . . . . . . . 310 Figure 7-103 NIST-1 HP-06b Reactor Pressure Vessel Level Comparison . . . . . . . . . . . . . 311 Figure 7-104 NIST-1 HP-06b Containment Vessel Level Comparison . . . . . . . . . . . . . . . . . 312 Figure 7-105 Comparison of NIST-1 HP-06 and HP-06b Reactor Pressure Vessel Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 313 Figure 7-106 Comparison of NIST-1 HP-06 and HP-06b Containment Vessel Pressure . . . 314 Figure 7-107 Comparison of NIST-1 HP-06 and HP-06b Reactor Pressure Vessel Level. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 315 Figure 7-108 Comparison of NIST-1 HP-06 and HP-06b Containment Vessel Level . . . . . . 316 Figure 7-109 Comparison of Core Power in HP-07 with the NuScale Power Module Power (fission and decay) after Reactor Trip (scaled) . . . . . . . . . . . . . . . . . . . 318 Figure 7-110 NIST-1 HP-07 Pressurizer Spray Supply Line Break Discharge Mass Flow Rate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 319 Figure 7-111 NIST-1 HP-07 Primary Mass Flow Rate. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 320 Figure 7-112 NIST-1 HP-07 Reactor Pressure Vessel Level Response Comparison with Data . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 321 Figure 7-113 NIST-1 HP-07 Containment Vessel Level Response. . . . . . . . . . . . . . . . . . . . 322 © Copyright 2022 by NuScale Power, LLC xix

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 7-114 NIST-1 HP-07 Containment Vessel Pressure Comparison . . . . . . . . . . . . . . . 323 Figure 7-115 NIST-1 HP-07 Primary Pressure Comparison . . . . . . . . . . . . . . . . . . . . . . . . . 324 Figure 7-116 Comparison of HP-09 Core Power with the Scaled NuScale Power Module Fission and Decay Power. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 326 Figure 7-117 NIST-1 HP-09 Valve Mass Flow Rate . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 327 Figure 7-118 NIST-1 HP-09 Pressurizer Pressure Comparison . . . . . . . . . . . . . . . . . . . . . . 328 Figure 7-119 NIST-1 HP-09 Pressurizer Pressure Comparison - 500 Seconds . . . . . . . . . . 329 Figure 7-120 NIST-1 HP-09 Containment Vessel Pressure Comparison . . . . . . . . . . . . . . . 330 Figure 7-121 NIST-1 HP-09 Containment Vessel Pressure Comparison - 500 Seconds . . . 331 Figure 7-122 NIST-1 HP-09 Pressurizer Level Comparison . . . . . . . . . . . . . . . . . . . . . . . . . 332 Figure 7-123 NIST-1 HP-09 Reactor Pressure Vessel Level Comparison . . . . . . . . . . . . . . 333 Figure 7-124 NIST-1 HP-09 Reactor Pressure Vessel Level Comparison - 500 Seconds . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 334 Figure 7-125 HP-43 Transient Short-Term Pressurizer Pressure Comparison . . . . . . . . . . . 335 Figure 7-126 HP-43 Transient Short-Term Pressurizer Level Code-to-Data Comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 336 Figure 7-127 HP-43 Transient Short-Term RPV Level Code-to-Data Comparison . . . . . . . . 337 Figure 7-128 HP-43 Transient Short-Term CNV Pressure Code-to-Data Comparison . . . . . 338 Figure 7-129 HP-43 Transient Short-Term Spurious RVV Orifice Mass Flow Rate Code-to-Data Comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 339 Figure 7-130 HP-43 Transient Short-Term CNV Level Code-to-Data Comparison. . . . . . . . 340 Figure 7-131 NIST-1 HP-49 Spurious Orifice Differential Pressure. . . . . . . . . . . . . . . . . . . . 343 Figure 7-132 NIST-1 HP-49 Primary Mass Flow Rate. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 344 Figure 7-133 NIST-1 HP-49 Pressurizer Level Comparison . . . . . . . . . . . . . . . . . . . . . . . . . 344 Figure 7-134 NIST-1 HP-49 Reactor Pressure Vessel Level Comparison . . . . . . . . . . . . . . 345 Figure 7-135 NIST-1 HP-49 Containment Vessel Level Comparison . . . . . . . . . . . . . . . . . . 345 Figure 7-136 NIST-1 HP-49 Containment Vessel Peak Pressure Comparison. . . . . . . . . . . 346 Figure 7-137 NIST-1 HP-49 Containment Vessel Pressure Comparison . . . . . . . . . . . . . . . 346 Figure 7-138 NIST-1 HP-49 Primary Pressure Comparison . . . . . . . . . . . . . . . . . . . . . . . . . 347 Figure 7-139 Run 1 Predicted RPV/CNV Pressure Comparison with Data. . . . . . . . . . . . . . 353 Figure 7-140 Run 1 Predicted RPV Pressure Comparison with Data for Full HTP Surface Area . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 354 Figure 7-141 Run 1 Predicted CNV Pressure Comparison with Data for Full HTP Surface Area . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 354 Figure 7-142 Run 1 Predicted RPV Level Comparison with Data . . . . . . . . . . . . . . . . . . . . . 355 © Copyright 2022 by NuScale Power, LLC xx

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 7-143 Run 1 Predicted CNV Level Comparison with Data . . . . . . . . . . . . . . . . . . . . . 356 Figure 7-144 Run 1 Predicted RPV Level Comparison with Data for Full HTP Surface Area . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 356 Figure 7-145 Run 1 Predicted CNV Level Comparison with Data for Full HTP Surface Area . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 357 Figure 7-146 Run 1 Predicted RPV Shell Energy Comparison . . . . . . . . . . . . . . . . . . . . . . . 358 Figure 7-147 Run 1 Predicted CNV Shell Energy Comparison . . . . . . . . . . . . . . . . . . . . . . . 358 Figure 7-148 Run 1 Predicted HTP Energy Transfer Comparison . . . . . . . . . . . . . . . . . . . . 359 Figure 7-149 Run 1 Predicted HTP Energy Transfer Comparison - Full HTP Surface Area . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 359 Figure 7-150 Run 1 Level 1 CPV Fluid Temperature Comparison . . . . . . . . . . . . . . . . . . . . 360 Figure 7-151 Run 1 Level 5 CPV Fluid Temperature Comparison . . . . . . . . . . . . . . . . . . . . 361 Figure 7-152 Run 1 Level 6 CPV Fluid Temperature Comparison . . . . . . . . . . . . . . . . . . . . 361 Figure 7-153 Run 1 CPV Fluid Level Comparison. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 362 Figure 7-154 Run 2 RPV-CNV Pressure Comparison. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 364 Figure 7-155 Run 2 RPV Level Comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 365 Figure 7-156 Run 2 CNV Level Comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 365 Figure 7-157 Run 5 RPV-CNV Pressure Comparison. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 366 Figure 7-158 Run 5 RPV Level Comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 367 Figure 7-159 Run 5 CNV Level Comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 367 Figure 7-160 Run 3 RPV-CNV pressure comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 369 Figure 7-161 Run 3 RPV level comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 370 Figure 7-162 Run 3 CNV Level Comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 371 Figure 7-163 Run 4 RPV-CNV pressure comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 373 Figure 7-164 Run 6 RPV-CNV Pressure Comparison. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 374 Figure 7-165 Run 7 RPV-CNV Pressure Comparison. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 374 Figure 7-166 Run 7 RPV-CNV Pressure Comparison. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 375 Figure 7-167 Run 4 CNV Level Comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 376 Figure 7-168 Run 6 RPV Level Comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 376 Figure 7-169 Run 6 CNV Level Comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 377 Figure 7-170 Run 7 RPV Level Comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 377 Figure 7-171 Run 7 CNV Level Comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 378 Figure 7-172 IORV Run 3 Pressurizer Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 383 Figure 7-173 IORV Run 6 Pressurizer Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 384 © Copyright 2022 by NuScale Power, LLC xxi

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 7-174 IORV Run 3 Core Differential Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 385 Figure 7-175 IORV Run 6 Core Differential Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 386 Figure 7-176 IORV Run 6 Baffle Plate Total Mass Flow Rate Comparison . . . . . . . . . . . . . 387 Figure 7-177 IORV Run 6 Baffle Plate Total Energy Flow Rate Comparison . . . . . . . . . . . . 388 Figure 7-178 IORV Run 6 Break Baffle Plate Differential Pressure. . . . . . . . . . . . . . . . . . . . 389 Figure 7-179 IORV Run 6 Break Total Mass Flow Rate Comparison . . . . . . . . . . . . . . . . . . 390 Figure 7-180 IORV Run 6 Break Total Energy Flow Rate Comparison. . . . . . . . . . . . . . . . . 391 Figure 8-1 CNV Wall Heat Transfer Modes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 423 Figure 8-2 Thermal Resistance Network Between CNV and UHS . . . . . . . . . . . . . . . . . . 423 Figure 8-3 Transient Void Fraction in Node 5 for the GE 4-ft Level Swell Test . . . . . . . . . 431 Figure 8-4 Transient Void Fraction in Node 4 for the GE 4-ft Level Swell Test . . . . . . . . . 432 Figure 8-5 Transient Void Fraction in Node 6 for the GE 1-ft Level Swell Test . . . . . . . . . 432 Figure 8-6 Transient Void Fraction in Node 5 for the GE 1-ft Level Swell Test . . . . . . . . . 433 Figure 8-7 General Design Framework for the NuScale Integral System Test Facility . . . 457 Figure 8-8 Flow Diagram for the Hierarchical, Two-Tiered Scaling Analysis (NUREG/CR-5809). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 459 Figure 8-9 NuScale System Breakdown into Hierarchical Levels and Primary Operational Modes to be Scaled . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 460 Figure 8-10 Scaling Analysis Flow Diagram for Single-Phase Primary Loop Natural Circulation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 463 Figure 8-11 Scaling Analysis Flow Diagram for Reactor Coolant System Depressurization . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 465 Figure 8-12 Scaling Analysis Flow Diagram for Containment Pressurization . . . . . . . . . . . 466 Figure 8-13 Scaling Analysis Flow Diagram for Long-Term Recirculation Cooling Mode . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 467 Figure 8-14 Scaling Analysis Flow Diagram for Reactor Building Pool Heat-Up. . . . . . . . . 469 Figure 8-15 Comparison of HP-05 Feedwater Flow to Test Data . . . . . . . . . . . . . . . . . . . . 481 Figure 8-16 Comparison of HP-05 Reactor Pressure Vessel Flow to Test Data. . . . . . . . . 482 Figure 8-17 Comparison of HP-05 Upper Riser Inlet Temperature to Test Data. . . . . . . . . 483 Figure 8-18 Comparison of HP-05 Core Inlet Temperature to Test Data . . . . . . . . . . . . . . 484 Figure 9-1 Reactor Power for Discharge Line Break Spectrum - AC and EDAS Power Available. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 511 Figure 9-2 Pressurizer Pressure for Discharge Line Break Spectrum - AC and EDAS Power Available . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 511 © Copyright 2022 by NuScale Power, LLC xxii

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 9-3 RCS Flow for Discharge Line Break Spectrum - AC and EDAS Power Available. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 512 Figure 9-4 CHFR for Discharge Line Break Spectrum - AC and EDAS Power Available. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 512 Figure 9-5 Reactor Power for Injection Line Break Spectrum - AC and EDAS Power Available. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 513 Figure 9-6 Pressurizer Pressure for Injection Line Break Spectrum - AC and EDAS Power Available . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 513 Figure 9-7 RCS Flow for Injection Line Break Spectrum - AC and EDAS Power Available. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 514 Figure 9-8 CHFR for Injection Line Break Spectrum - AC and EDAS Power Available. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 514 Figure 9-9 Reactor Power for Discharge Line Break Spectrum - Loss of AC Power. . . . . 515 Figure 9-10 Pressurizer Pressure for Discharge Line Break Spectrum - Loss of AC Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 515 Figure 9-11 RCS Flow for Discharge Line Break Spectrum - Loss of AC Power . . . . . . . . 516 Figure 9-12 CHFR for Discharge Line Break Spectrum - Loss of AC Power . . . . . . . . . . . 516 Figure 9-13 Reactor Power for Injection Line Break Spectrum - Loss of AC Power . . . . . . 517 Figure 9-14 Pressurizer Pressure for Injection Line Break Spectrum - Loss of AC Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 517 Figure 9-15 RCS Flow for Injection Line Break Spectrum - Loss of AC Power . . . . . . . . . . 518 Figure 9-16 CHFR for Injection Line Break Spectrum - Loss of AC Power . . . . . . . . . . . . . 518 Figure 9-17 Reactor Power for Discharge Line Break Spectrum - Loss of EDAS Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 519 Figure 9-18 Pressurizer Pressure for Discharge Line Break Spectrum - Loss of EDAS Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 519 Figure 9-19 RCS Flow for Discharge Line Break Spectrum - Loss of EDAS Power . . . . . . 520 Figure 9-20 CHFR for Discharge Line Break Spectrum - Loss of EDAS Power . . . . . . . . . 520 Figure 9-21 Reactor Power for Injection Line Break Spectrum - Loss of EDAS Power. . . . 521 Figure 9-22 Pressurizer Pressure for Injection Line Break Spectrum - Loss of EDAS Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 521 Figure 9-23 RCS Flow for Injection Line Break Spectrum - Loss of EDAS Power . . . . . . . 522 Figure 9-24 CHFR for Injection Line Break Spectrum - Loss of EDAS Power . . . . . . . . . . 522 Figure 9-25 Reactor Power High Point Vent and Pressurizer Spray Spectrum - AC and EDAS Power Available. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 524 Figure 9-26 Pressurizer Pressure High Point Vent and Pressurizer Spray Spectrum - AC and EDAS Power Available . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 525 © Copyright 2022 by NuScale Power, LLC xxiii

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 9-27 RCS Flow High Point Vent and Pressurizer Spray Spectrum - AC and EDAS Power Available. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 525 Figure 9-28 CHFR High Point Vent and Pressurizer Spray Spectrum - AC and EDAS Power Available . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 526 Figure 9-29 Reactor Power High Point Vent and Pressurizer Spray Spectrum - Loss of AC Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 526 Figure 9-30 Pressurizer Pressure High Point Vent and Pressurizer Spray Spectrum - Loss of AC Power. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 527 Figure 9-31 RCS Flow High Point Vent and Pressurizer Spray Spectrum - Loss of AC Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 527 Figure 9-32 CHFR High Point Vent and Pressurizer Spray Spectrum - Loss of AC Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 528 Figure 9-33 Reactor Power High Point Vent and Pressurizer Spray Spectrum - Loss of EDAS Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 528 Figure 9-34 Pressurizer Pressure High Point Vent and Pressurizer Spray Spectrum - Loss of EDAS Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 529 Figure 9-35 RCS Flow High Point Vent and Pressurizer Spray Spectrum - Loss of EDAS Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 529 Figure 9-36 CHFR High Point Vent and Pressurizer Spray Spectrum - Loss of EDAS Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 530 Figure 9-37 Break Flow and Net ECCS Valve Flow for 100 percent RCS Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 535 Figure 9-38 RPV and CNV Pressure for 100 percent RCS Injection Line Break. . . . . . . . . 535 Figure 9-39 Collapsed Liquid Levels for 100 percent RCS Injection Line Break . . . . . . . . . 536 Figure 9-40 Recirculation Flow Rate and Core Flow Distribution for 100 percent RCS Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 536 Figure 9-41 Core Flow Transient for 100 percent RCS Injection Line Break. . . . . . . . . . . . 537 Figure 9-42 Components of Reactor Power for 100 percent RCS Injection Line Break . . . 537 Figure 9-43 Post-Scram MCHFR during 100 percent RCS Injection Line Break . . . . . . . . 538 Figure 9-44 Peak Clad and Fuel Centerline Temperature during 100 percent RCS Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 538 Figure 9-45 Comparison of RPV and CNV Pressure Response 100 percent RCS Injection and Discharge Line Breaks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 539 Figure 9-46 Comparison of Collapsed Riser Level 100 percent RCS Injection and Discharge Line Breaks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 539 Figure 9-47 Comparison of Break Flows between 100 percent RCS Injection and Discharge Line Breaks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 540 © Copyright 2022 by NuScale Power, LLC xxiv

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 9-48 NPM-160 RPV and CNV Pressure for Base Case RCS Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 540 Figure 9-49 NPM-20 RPV and CNV Pressure for Base Case RCS Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 541 Figure 9-50 NPM-160 Collapsed Liquid Levels for Base Case RCS Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 541 Figure 9-51 NPM-20 Collapsed Liquid Levels for Base Case RCS Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 542 Figure 9-52 NPM-20 RPV and CNV Pressure Comparing Injection Line Break to High Point Vent Line Break. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 544 Figure 9-53 Break and ECCS Flow for 100 percent High Point Vent Line Break . . . . . . . . 545 Figure 9-54 Collapsed Liquid Levels for 100 percent High Point Vent Line Break . . . . . . . 545 Figure 9-55 Post-Scram MCHFR for 100 percent High Point Vent Line Break . . . . . . . . . . 546 Figure 9-56 NPM-160 RPV and CNV Pressure Base Case High Point Vent Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 547 Figure 9-57 NPM-20 RPV and CNV Pressure Base Case High Point Vent Line Break . . . 547 Figure 9-58 NPM-160 Collapsed Liquid Levels Base Case High Point Vent Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 548 Figure 9-59 NPM-20 Collapsed Liquid Levels Base Case High Point Vent Line Break . . . 548 Figure 9-60 RPV Pressure Transient for RCS Injection Line Breaks . . . . . . . . . . . . . . . . . 551 Figure 9-61 CNV Pressure Transient for RCS Injection Line Breaks . . . . . . . . . . . . . . . . . 551 Figure 9-62 Collapse Liquid Level above TAF for RCS Injection Line Breaks . . . . . . . . . . 552 Figure 9-63 RPV Pressure Transient for RCS Discharge Line Breaks . . . . . . . . . . . . . . . . 552 Figure 9-64 CNV Pressure Transient for RCS Discharge Line Breaks . . . . . . . . . . . . . . . . 553 Figure 9-65 Collapse Liquid Level above TAF for RCS Discharge Line Breaks . . . . . . . . . 553 Figure 9-66 RPV Pressure Transient for High Point Vent Line Breaks . . . . . . . . . . . . . . . . 554 Figure 9-67 CNV Pressure Transient for High Point Vent Line Breaks . . . . . . . . . . . . . . . . 554 Figure 9-68 Collapse Liquid Level above TAF for High Point Vent Line Breaks . . . . . . . . . 555 Figure 9-69 Post-Scram MCHFR for RCS Injection Line Breaks . . . . . . . . . . . . . . . . . . . . 555 Figure 9-70 Post-Scram MCHFR for High Point Vent Line Breaks . . . . . . . . . . . . . . . . . . . 556 Figure 9-71 Pressure Differential across the RVV at Actuation by Break Size and Location . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 556 Figure 9-72 ECCS Valve Actuation Time by Break Size and Location . . . . . . . . . . . . . . . . 557 Figure 9-73 CNV Peak Pressure by Break Size and Location . . . . . . . . . . . . . . . . . . . . . . 557 Figure 9-74 Minimum Collapsed Liquid Level by Break Size and Location. . . . . . . . . . . . . 558 © Copyright 2022 by NuScale Power, LLC xxv

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 9-75 Effect of Power Availability on Peak CNV Pressure for RCS Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 559 Figure 9-76 Effect of Power Availability on Peak CNV Pressure for High Point Vent Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 560 Figure 9-77 Effect of Power Availability on Collapsed Liquid Level for RCS Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 560 Figure 9-78 Effect of Power Availability on Collapsed Liquid Level for High Point Vent Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 561 Figure 9-79 Effect of Single Failures on Peak CNV Pressure for RCS Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 562 Figure 9-80 Effect of Single Failures on Peak CNV Pressure for High Point Vent Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 563 Figure 9-81 Effect of Single Failures on Collapsed Liquid Level for RCS Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 563 Figure 9-82 Effect of Single Failures on Collapsed Liquid Level for High Point Vent Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 564 Figure 9-83 Effect of Single Failures on Collapsed Liquid Level for RCS Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 564 Figure 9-84 Effect of Single Failures on RRV Flow for RCS Injection Line Break. . . . . . . . 565 Figure 9-85 Effect of Single Failures on Collapsed Liquid Level for RCS Discharge Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 565 Figure 9-86 Effect of Single Failures on RRV Flow for RCS Discharge Line Break . . . . . . 566 Figure 9-87 Effect of Single Failures on Collapsed Liquid Level for High Point Vent Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 566 Figure 9-88 Effect of Single Failures on RRV Flow for High Point Vent Line Break . . . . . . 567 Figure 9-89 Peak CNV Pressure with DHRS Operation for Break Locations with Different Power Availability and Single Failures. . . . . . . . . . . . . . . . . . . . . . . . 567 Figure 9-90 Collapsed Liquid Level with DHRS Operation for Break Locations with Different Power Availability and Single Failures. . . . . . . . . . . . . . . . . . . . . . . . 568 Figure 9-91 Reactor Pressure Vessel and Containment Vessel Pressure (Left) and Collapsed Level Above Top of Active Fuel (Right) for 100 Percent Reactor Coolant System Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 570 Figure 9-92 Reactor Pressure Vessel and Containment Vessel Pressure (Left) and Collapsed Level above Top of Active Fuel (Right) for 10 Percent Reactor Coolant System Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 570 Figure 9-93 Reactor Pressure Vessel and Containment Vessel Pressure (Left) and Collapsed Level above Top of Active Fuel (Right) for 100 Percent High Point Vent Line Break. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 571 © Copyright 2022 by NuScale Power, LLC xxvi

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 9-94 Hot Channel Core Flow (Left) and Core Critical Heat Flux Ratio (right) During 100 Percent Reactor Coolant System Injection Line Break . . . . . . . . . 572 Figure 9-95 Time-Step Size Sensitivity on Reactor and Containment Vessel Pressures and Reactor Pressure Vessel Collapsed Liquid Level for 100 Percent Reactor Coolant System Injection Line Break.. . . . . . . . . . . . . . . . . . . . . . . . . 573 Figure 9-96 Time-Step Size Sensitivity on Hot Assembly Flow and Minimum Critical Heat Flux Ratio for 100 Percent Reactor Coolant System Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 573 Figure 9-97 Time-Step Size Sensitivity on Reactor and Containment Vessel Pressures and Reactor Pressure Vessel Collapsed Liquid Level for 100 Percent High Point Vent Line Break. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 574 Figure 9-98 Time-Step Size Sensitivity on Hot Assembly Flow and Minimum Critical Heat Flux Ratio for 100 Percent High Point Vent Break. . . . . . . . . . . . . . . . . . 574 Figure 9-99 Effect of Counter Current Flow Limitation Line Slope on Levels for 100 Percent High Point Vent Line Break. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 575 Figure 9-100 Effect of Inadvertent Actuation Block Release Pressure on Peak Containment Vessel Pressure and Minimum Collapsed Liquid Level above Top of Active Fuel as a Function of Break Size for Reactor Coolant System Injection Line Break. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 577 Figure 9-101 Effect of Reactor Recirculation Valve Size on Peak Containment Vessel Pressure and Minimum Collapsed Liquid Level for Reactor Coolant System Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 577 Figure 9-102 Effect of Reactor Vent Valve Size on Peak Containment Vessel Pressure and Minimum Collapsed Liquid Level for Reactor Coolant System Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 578 Figure 9-103 Effect of Initial Reactor Pool Temperature on Peak Containment Vessel Pressure and Minimum Collapsed Liquid Level above Top of Active Fuel for Reactor Coolant System Injection Line Break. . . . . . . . . . . . . . . . . . . . . . . 579 Figure 9-104 Containment Vessel to Pool Energy Transfer at Different Initial Pool Temperatures for 100 Percent (Left) and 10 Percent (Right) Reactor Coolant System Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 580 Figure 9-105 Generic Normalized Axial Power Shapes . . . . . . . . . . . . . . . . . . . . . . . . . . . . 581 Figure 9-106 Effect of Axial Power Shape on Reactor Pressure Vessel and Containment Pressures and Collapsed Liquid Level above Top of Active Fuel for Reactor Coolant System Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . 582 Figure 9-107 Effect of Axial Power Shape on Peak Containment Vessel Pressure and Minimum Collapsed Liquid Level above Top of Active Fuel for Reactor Coolant System Injection Line Break . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 583 Figure 9-108 Effect of Axial Power Shape on Hot Assembly Flow and Minimum Critical Heat Flux Ratio during Reactor Coolant System Injection Line Break . . . . . . . 583 © Copyright 2022 by NuScale Power, LLC xxvii

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 9-109 RPV Pressure for Injection Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) . . . . . . . . . . . . . . . . . 585 Figure 9-110 CNV Pressure for Injection Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) . . . . . . . . . . . . . . . . . 585 Figure 9-111 Collapsed Liquid Level for Injection Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) . . . . . . . . . 586 Figure 9-112 CNV Pressure for Injection Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) . . . . . . . . . . . . . . . . . 586 Figure 9-113 RPV Pressure for High Point Vent Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) . . . . . . . . . 587 Figure 9-114 CNV Pressure for High Point Vent Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) . . . . . . . . . 587 Figure 9-115 Collapsed Liquid Level for High Point Vent Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) . . . 588 Figure 9-116 CNV Pressure for High Point Vent Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) . . . . . . . . . 588 Figure 9-117 Discharge Line Break - RPV Pressure Response . . . . . . . . . . . . . . . . . . . . . . 600 Figure 9-118 Discharge Line Break - CNV Pressure Response . . . . . . . . . . . . . . . . . . . . . . 601 Figure 9-119 Discharge Line Break - RPV Level above Top of Active Fuel . . . . . . . . . . . . . 602 Figure 9-120 Discharge Line Break - CNV Level. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 603 Figure 9-121 Discharge Line Break - CNV Wall Temperature Response . . . . . . . . . . . . . . . 604 Figure 9-122 Injection Line Break - RPV Pressure Response. . . . . . . . . . . . . . . . . . . . . . . . 605 Figure 9-123 Injection Line Break - CNV Pressure Response . . . . . . . . . . . . . . . . . . . . . . . 606 Figure 9-124 Injection Line Break - RPV Level above Top of Active Fuel. . . . . . . . . . . . . . . 607 Figure 9-125 Injection Line Break - CNV Level . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 608 Figure 9-126 Injection Line Break - CNV Wall Temperature Response . . . . . . . . . . . . . . . . 609 Figure 9-127 High Point Vent Line Break - RPV Pressure Response . . . . . . . . . . . . . . . . . . 610 Figure 9-128 High Point Vent Line Break - CNV Pressure Response. . . . . . . . . . . . . . . . . . 611 Figure 9-129 High Point Vent Line Break - RPV Level above Top of Active Fuel . . . . . . . . . 612 Figure 9-130 High Point Vent Line Break - CNV Level . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 613 Figure 9-131 High Point Vent Line Break - CNV Wall Temperature Response. . . . . . . . . . . 614 Figure 9-132 Inadvertent RVV Opening - RPV Pressure Response . . . . . . . . . . . . . . . . . . . 615 Figure 9-133 Inadvertent RVV Opening - CNV Pressure Response . . . . . . . . . . . . . . . . . . . 616 Figure 9-134 Inadvertent RVV Opening - RPV Level above Top of Active Fuel . . . . . . . . . . 617 Figure 9-135 Inadvertent RVV Opening - CNV Level . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 618 © Copyright 2022 by NuScale Power, LLC xxviii

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 9-136 Inadvertent RVV Opening - CNV Wall Temperature Response. . . . . . . . . . . . 619 Figure 9-137 Inadvertent RRV Opening - RPV Pressure Response . . . . . . . . . . . . . . . . . . . 620 Figure 9-138 Inadvertent RRV Opening - CNV Pressure Response. . . . . . . . . . . . . . . . . . . 621 Figure 9-139 Inadvertent RRV Opening - RPV Level above Top of Active Fuel . . . . . . . . . . 622 Figure 9-140 Inadvertent RRV Opening - CNV Level . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 623 Figure 9-141 Inadvertent RRV Opening - CNV Wall Temperature Response. . . . . . . . . . . . 624 Figure 9-142 Inadvertent ECCS Actuation Signal - RPV Pressure Response . . . . . . . . . . . 625 Figure 9-143 Inadvertent ECCS Actuation Signal - CNV Pressure Response . . . . . . . . . . . 626 Figure 9-144 Inadvertent ECCS Actuation Signal - RPV Level above Top of Active Fuel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 627 Figure 9-145 Inadvertent ECCS Actuation Signal - CNV Level . . . . . . . . . . . . . . . . . . . . . . . 628 Figure 9-146 Inadvertent ECCS Actuation Signal - CNV Wall Temperature Response . . . . 629 Figure 9-147 Main Steam Line Break - RPV Pressure Response. . . . . . . . . . . . . . . . . . . . . 632 Figure 9-148 Main Steam Line Break - CNV Pressure Response. . . . . . . . . . . . . . . . . . . . . 633 Figure 9-149 Main Steam Line Break - RPV Level above Top of Fuel . . . . . . . . . . . . . . . . . 634 Figure 9-150 Main Steam Line Break - CNV Level . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 635 Figure 9-151 Main Steam Line Break - CNV Wall Temperature . . . . . . . . . . . . . . . . . . . . . . 636 Figure 9-152 Main Feedwater Line Break - RPV Pressure Response . . . . . . . . . . . . . . . . . 639 Figure 9-153 Main Feedwater Line Break - CNV Pressure Response . . . . . . . . . . . . . . . . . 640 Figure 9-154 Main Feedwater Line Break - RPV Level above Top of Fuel . . . . . . . . . . . . . . 641 Figure 9-155 Main Feedwater Line Break - CNV Level . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 642 Figure 9-156 Main Feedwater Line Break - CNV Wall Temperature . . . . . . . . . . . . . . . . . . . 643 Figure 9-157 Reactor Power for IORV Event Spectrum - Electric Power Available . . . . . . . 645 Figure 9-158 Pressurizer Pressure for IORV Event Spectrum - Electric Power Available. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 646 Figure 9-159 Reactor Coolant System Flow for IORV Event Spectrum - Electric Power Available. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 647 Figure 9-160 Critical Heat Flux Ratio for IORV Event Spectrum - Electric Power Available. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 648 Figure 9-161 Reactor Power for IORV Event Spectrum - Loss of AC Power . . . . . . . . . . . . 649 Figure 9-162 Pressurizer Pressure for IORV Event Spectrum - Loss of AC Power . . . . . . . 650 Figure 9-163 Reactor Coolant System Flow for IORV Event Spectrum - Loss of AC Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 651 Figure 9-164 Critical Heat Flux Ratio for IORV Event Spectrum - Loss of AC Power . . . . . . 652 © Copyright 2022 by NuScale Power, LLC xxix

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 List of Figures Figure 9-165 Reactor Power for IORV Event Spectrum - Loss of EDAS Power . . . . . . . . . . 653 Figure 9-166 Pressurizer Pressure for IORV Event Spectrum - Loss of EDAS Power . . . . . 654 Figure 9-167 Reactor Coolant System Flow for IORV Event Spectrum - Loss of EDAS Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 655 Figure 9-168 Critical Heat Flux Ratio for IORV Event Spectrum - Loss of EDAS Power . . . 656 Figure 9-169 NPM-20 RPV/CNV Pressure Response for Single ECCS Valve Opening. . . . 660 Figure 9-170 NPM-20 Collapsed Liquid Level above TAF for Single ECCS Valve Opening . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 661 Figure 9-171 NPM-20 Pressurizer Level Response for Single ECCS Valve Opening . . . . . 662 Figure 9-172 NPM-20 Hot Channel Mass Flux for Single ECCS Valve Opening . . . . . . . . . 663 Figure 9-173 NPM-20 Phase 1 MCHFR for Single ECCS Valve Opening . . . . . . . . . . . . . . 664 Figure 9-174 NPM-20 Phase 1 Reactor Vessel Pressure for Inadvertent ECCS Signal (Two RVV Opening) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 665 Figure 9-175 NPM-20 Phase 1 Containment Pressure for Inadvertent ECCS Signal (Two RVV Opening) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 666 Figure 9-176 NPM-20 Phase 1 Collapse Liquid Level over TAF for Inadvertent ECCS Signal (Two RVV Opening) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 667 Figure 9-177 NPM-20 Phase 1 MCHFR for Inadvertent ECCS Signal (Two RVV Opening). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 668 Figure 9-178 NPM-160 System Pressure - Inadvertent RVV Opening . . . . . . . . . . . . . . . . . 669 Figure 9-179 NPM-20 System Pressure - Inadvertent RVV Opening . . . . . . . . . . . . . . . . . . 670 Figure 9-180 NPM-160 System Level - Inadvertent RVV Opening . . . . . . . . . . . . . . . . . . . . 671 Figure 9-181 NPM-20 System Level - Inadvertent RVV Opening . . . . . . . . . . . . . . . . . . . . . 672 Figure 9-182 NPM-160 System Pressure - Inadvertent RRV Opening . . . . . . . . . . . . . . . . . 673 Figure 9-183 NPM-20 System Pressure - Inadvertent RRV Opening . . . . . . . . . . . . . . . . . . 674 Figure 9-184 NPM-160 System Level - Inadvertent RRV Opening . . . . . . . . . . . . . . . . . . . . 675 Figure 9-185 NPM-20 System Level - Inadvertent RRV Opening . . . . . . . . . . . . . . . . . . . . . 676 Figure 9-186 Reactor Pressure Vessel and Containment Pressure for Inadvertent RVV Opening - Loss of Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 677 Figure 9-187 Riser Collapsed Level for Inadvertent RVV Opening - Loss of Power . . . . . . . 678 Figure 9-188 Core MCHFR for Inadvertent RVV Opening - Loss of Power . . . . . . . . . . . . . 679 Figure 9-189 Reactor Pressure Vessel and Containment Pressure for Inadvertent RRV Opening - Loss of Power . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 680 Figure 9-190 Riser Collapsed Level for Inadvertent RRV Opening - Loss of Power . . . . . . . 681 Figure 9-191 Core MCHFR for Inadvertent RRV Opening - Loss of Power . . . . . . . . . . . . . 682 © Copyright 2022 by NuScale Power, LLC xxx

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Abstract NuScale Power, LLC (NuScale) has developed a small modular reactor (SMR) that supports operation of multiple NuScale Power Modules (NPM) at a specific site. An NPM is an advanced, light-water, integrated pressurized water reactor (PWR) using natural circulation for primary coolant flow. Each NPM has an independent nuclear steam supply system (NSSS), a standard steam power conversion system, and a compact steel containment vessel (CNV). In an NPM design, all primary components are integral to the reactor pressure vessel (RPV), which eliminates most of the reactor piping found on conventional PWRs, thereby reducing the possibility of a pipe rupture that would result in a loss-of-coolant accident (LOCA). NuScale is requesting Nuclear Regulatory Commission (NRC) review and approval to use the LOCA evaluation model (EM) described in this report for analyses of design-basis LOCA events, inadvertent reactor valve opening events, and CNV peak pressure and temperature analyses resulting from LOCA and non-LOCA events. The NuScale LOCA EM is developed using the evaluation model development and assessment process (EMDAP) of "Transient and Accident Analysis Methods," Regulatory Guide (RG) 1.203 (Reference 1), and it adheres to the applicable requirements of "ECCS Evaluation Models," 10 CFR 50 Appendix K (Reference 2), and "Acceptance Criteria for Emergency Core Cooling Systems for Light-Water Nuclear Power Reactors," 10 CFR 50.46 (Reference 3). This topical report is not intended to provide final design values or results; rather, example values for the various evaluations are provided for illustrative purposes in order to aid the reader's understanding of the context of the application of the NuScale LOCA EM. The LOCA EM uses the proprietary NRELAP5 systems analysis computer code as the computational engine, derived from the Idaho National Laboratory (INL) RELAP5-3D© computer code. The models and correlations used by NRELAP5 were reviewed and, where appropriate, modified for use within the NuScale LOCA EM. Validation and verification of the LOCA EM and NRELAP5 code has been performed in accordance with the EMDAP. A phenomena identification and ranking table (PIRT), which identifies the important phenomena and processes occurring in an NPM during a LOCA, was developed by gathering and ranking expert evaluations of phenomena that could occur in an NPM during a LOCA. Over 20 phenomena were identified as important to capture in the NuScale LOCA EM. Extensive NRELAP5 code validation was performed to ensure that the LOCA EM is applicable for important phenomena and processes over the range encountered in an NPM LOCA. The validation suite includes many legacy separate effects tests (SETs) and integral effects tests (IETs), as well as many SETs and IETs developed and run specifically for the NPM application. The EMDAP requires an applicability demonstration of the NRELAP5 code and tests. A unique aspect of the demonstration provided for the NPM is the comparison of NRELAP5 simulations of LOCA to NuScale Integral System Test Facility (NIST) test data and NRELAP5 simulation of the same LOCA in an NPM. The reasonable-to-excellent agreement obtained by these comparisons establishes the applicability of NRELAP5 to accurately predict LOCA phenomena at both the NIST and NPM scales. © Copyright 2022 by NuScale Power, LLC 1

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 This topical report provides an example application of the LOCA EM in order to aid the reader's understanding of the context of the application of the NuScale LOCA EM. Calculations are presented for break spectra that cover a range of break locations, break sizes, single failures, equipment unavailability, and initial and boundary conditions. Example calculations are also provided for inadvertent opening of reactor valve (IORV) events and peak CNV pressure and temperature. The methodology in this report to supports analyses for Non-LOCA events, and extended passive cooling evaluation. This report presents the containment response analysis methodology. The NRELAP5 simulation model used for the containment response analysis methodology is similar to the NRELAP5 simulation models used for LOCA, reactor valve opening events and non-LOCA methodologies, which are presented in this report. The PIRT developed for the LOCA and non-LOCA methodologies are applicable to the containment response analysis methodology. The qualification of the LOCA and non-LOCA methodologies, in particular the comparisons to separate effects tests and integral effects tests, applicable to the containment response analysis methodology are presented in this report. © Copyright 2022 by NuScale Power, LLC 2

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Executive Summary NuScale Power, LLC (NuScale) has developed a small modular reactor that supports operation multiple NuScale Power Modules (NPMs) at each site. An NPM is an advanced, light-water, integral pressurized water reactor (PWR) that uses a high-pressure containment vessel (CNV) partially immersed in a reactor pool coupled with simple, redundant, passive safety-related systems. The design ensures safe plant shutdown and cooldown in the event of a loss-of-coolant accident (LOCA). Each NPM has an independent nuclear steam supply system (NSSS) that includes a nuclear core, helical-coil steam generator (SG), integral pressurizer, and a compact, high-pressure steel CNV that contains the NSSS. The secondary system includes a traditional steam-power conversion system including a steam turbine generator, condenser, and feedwater system. The integral PWR design eliminates most of the reactor piping found on conventional PWRs, thereby reducing the possibility of a pipe rupture that would result in a LOCA. Piping in an NPM containment that potentially can break is limited to the reactor coolant system (RCS) injection line, RCS discharge line, pressurizer spray supply line, and pressurizer high point vent line. The RCS injection line is supplied by the chemical and volume control system (CVCS) and the discharge line returns to the CVCS. The NPM is designed to reduce the consequences of design basis LOCAs by using redundant, simplified, passive safety-related systems that eliminate the need for emergency core cooling system (ECCS) pumps, accumulators, and water storage tanks found on conventional PWRs. During operation, flow through the reactor is driven by natural circulation resulting from the thermal driving head produced by the temperature difference between the core and the heat sink afforded by the SG. Natural circulation flow increases reliability by eliminating primary coolant pumps that can fail or lock up. This topical report revision presents the updated NuScale evaluation model (EM). This topical report revision also integrates the methodology to evaluate containment pressure and temperature and other events involving the opening of reactor valves. This LOCA EM is developed following the guidelines in the evaluation model development and assessment process (EMDAP) of "Transient and Accident Analysis Methods," Regulatory Guide (RG) 1.203 (Reference 1), and adheres to the applicable requirements of "ECCS Evaluation Models," 10 CFR 50 Appendix K (Reference 2) and "Acceptance Criteria for Emergency Core Cooling Systems for Light-Water Nuclear Power Reactors," 10 CFR 50.46 (Reference 3). Multiple layers of conservatism are incorporated in the LOCA EM to ensure that a conservative analysis result is obtained. These conservatisms stem from application of the modeling requirements of 10 CFR 50 Appendix K and through a series of conservative modeling features. The LOCA EM uses the proprietary NRELAP5 systems analysis computer code as the computational engine, derived from the Idaho National Laboratory (INL) RELAP5-3D© computer code. RELAP5-3D© was procured and as part of the procurement process commercial grade dedication was performed by NuScale to establish the baseline NRELAP5 code for development. Subsequently, features were added and changes made to NRELAP5 to address the unique aspects of the NPM design and licensing methodology. NRELAP5 includes all of the necessary models for characterization of the NPM hydrodynamics, heat transfer between structures and fluids, modeling of fuel, reactor kinetics models, and control systems. The models and correlations used by NRELAP5 have been reviewed and, where appropriate, modified for use within the LOCA methodology. Code changes for the NuScale application include new helical coil SG heat transfer and pressure drop models, core critical heat flux (CHF) models, and interfacial © Copyright 2022 by NuScale Power, LLC 3

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 drag models for large-diameter pipes. The fuel CHF models were selected based on full-scale fuel bundle performance tests. Validation and verification of the EM and NRELAP5 code were conducted in accordance with the EMDAP process. A phenomena identification and ranking table (PIRT), which identifies the important phenomena and processes occurring in an NPM during a LOCA event, was developed by gathering and ranking expert evaluations of phenomena that could occur in an NPM during a LOCA. Phenomena and process ranking was performed in relation to specified figures of merit (FOMs) as described by RG 1.203. The PIRT also established a knowledge ranking for each phenomenon identified. Using these FOMs, over twenty phenomena were identified as important to capture in the NuScale LOCA EM. Extensive NRELAP5 code validation was performed to ensure that the LOCA EM is applicable for all important phenomena and processes over the range encountered in the NPM LOCA. The validation suite includes many legacy separate effects tests (SETs) and integral effects tests (IETs), as well as many SETs and IETs developed and run specifically for the NPM application. The SETs run for the NPM application were performed at the Societ Informazioni Esperienze Termoidrauliche (SIET) facility on a model helical coil SG, and at the Stern facility to obtain CHF data on a full-scale rod bundle test section. The IETs were performed at the Oregon State University NuScale Integral System Test (NIST) facility, a scaled representation of the complete NPM primary and secondary systems, as well as the reactor pool. The EMDAP requires an applicability demonstration of the NRELAP5 code and tests. A unique aspect of the demonstration provided for the NPM is the comparison of NRELAP5 simulations of LOCA events to NIST test data and NRELAP5 simulation of the same LOCA event in an NPM. In the comparisons, the NPM results are scaled down to the NIST size using the scaling ratios used to design the NIST facility. The reasonable-to-excellent agreement obtained by these comparisons establishes the applicability of NRELAP5 to accurately predict LOCA phenomena at both the NIST and NPM scales. This topical report provides example applications of the LOCA EM in order to aid the reader's understanding of the context of the application of the NuScale LOCA EM. These calculations are presented for break spectra that cover a range of break locations, break sizes, single failures, equipment unavailability and initial and boundary conditions. Nodalization and time-step sensitivity required by 10 CFR 50 Appendix K are also performed. The LOCA analyses demonstrate that the NPM retains sufficient water inventory in the primary system such that the core does not uncover, the fuel does not experience a CHF condition and the containment design pressure is not challenged. Peak cladding temperature (PCT) is shown to occur at the beginning of the LOCA event and cladding temperature decreases as the transient evolves. Because no fuel heat-up occurs for any design-basis LOCA, the following regulatory acceptance criteria from 10 CFR 50.46 are met:

1. Peak cladding temperature remains below 2,200 degrees Fahrenheit (1,204 degrees Celsius).

2. Maximum fuel oxidation is less than 0.17 times total cladding thickness before oxidation.

3. Maximum hydrogen generation is less than 0.01 times that generated if all cladding were to react.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3

4. Coolable geometry is retained.

NuScale requests Nuclear Regulatory Commission (NRC) review and approval to use the LOCA EM described in this report for analyses of design basis LOCA events, containment pressure and temperature analyses and reactor valve opening events for an NPM. The NuScale LOCA EM includes the following components: LOCA PIRT NRELAP5 code with NuScale-specific modifications assessment of the NRELAP5 code against experimental data demonstration of the applicability of the NRELAP5 code to LOCA analysis input model of the NPM This LOCA EM uses a conservative bounding approach to analyzing LOCA transients that follows the guidance provided in RG 1.203 and satisfies the applicable requirements of 10 CFR 50 Appendix K. Results show that its application to the NPM demonstrates acceptable performance based upon the acceptance criteria of 10 CFR 50.46. The methodology in this report is also used to support other analyses including:

1. events as described in Topical Report TR-0516-49416-P, "Non-Loss of Coolant Accident Methodology,"

2. containment peak pressure analysis for the NPM-160 approved design as described in Technical Report TR-0516-49084-P, Containment Response Analysis Methodology,

3. long term cooling as described in the NPM-160 approved design Technical Report, TR-0919-51299-P, "Long-Term Cooling Methodology," and

4. extended passive cooling as described in Topical Report, Extended Passive Cooling and Reactivity Control Methodology, TR-124587-P.

The NRELAP5 simulation model used for the containment response analysis methodology is similar to the NRELAP5 simulation models used for the LOCA, valve opening event and non-LOCA methodologies. The PIRT developed for the LOCA and non-LOCA methodologies are applicable to the containment response analysis methodology. The qualification of the LOCA and non-LOCA methodologies, in particular the comparisons to separate effects tests and integral effects tests, applicable to the containment response analysis methodology are presented in this report and in the non-LOCA topical report TR-0516-49416-P. © Copyright 2022 by NuScale Power, LLC 5

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 1.0 Introduction 1.1 Purpose The purpose of this report is to present the NuScale evaluation model (EM) used to evaluate emergency core cooling system (ECCS) performance in a NuScale Power Module (NPM) for design-basis loss-of-coolant accidents (LOCAs) and other reactor valve opening events. The LOCA EM follows the guidance provided in "Transient and Accident Analysis Methods," Regulatory Guide (RG) 1.203 (Reference 1) and satisfies the applicable requirements of "ECCS Evaluation Models," 10 CFR 50 Appendix K (Reference 2). NuScale requests U.S. Nuclear Regulatory Commission (NRC) approval to use the EM described in this report for analyses of design-basis LOCA events in an NPM. This report also presents the NuScale Power, LLC, methodology used to analyze the mass and energy (M&E) release into the containment vessel (CNV) for the spectrum of design-basis transients and accidents and the resulting pressure and temperature response of the CNV. NPM limiting peak pressure and temperature results determined using this method are presented in applications referencing this report. 1.2 Scope This report summarizes the following: NPM design and operation NuScale LOCA phenomena identification and ranking table (PIRT) NRELAP5 input model for the NPM NRELAP5 code features and modifications assessment of NRELAP5 against separate effects tests (SETs) and integral effects tests (IETs) applicability evaluation to determine the adequacy of NRELAP5 for NPM LOCA analyses and reactor valve opening events NRELAP5 model application to CNV peak pressure and temperature response This report also provides LOCA analyses at several locations and over a spectrum of break sizes to demonstrate the application of the EM to the NPM design. Additionally, the results of sensitivity calculations performed in accordance with the applicable requirements of 10 CFR 50 Appendix K are summarized. The scope of the NuScale LOCA EM is as follows: The EM is applicable to a nuclear power plant that follows the general description of the NuScale Power Plant design in Section 3.0 and that falls within the range of conditions in Table 8-2 evaluated for EM adequacy. Applicability of the EM is based on the NuScale LOCA PIRT, which identifies and ranks those phenomena the EM must be qualified to model during a LOCA in an NPM. © Copyright 2022 by NuScale Power, LLC 6

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The EM does not have restrictions concerning operating setpoints or loss of offsite-power conditions as long as the phenomena that occur during the progression of a LOCA have been identified by the PIRT process. This topical report is not intended to provide final design values or results; rather, example values for the various evaluations are provided for illustrative purposes in order to aid the reader's understanding of the context of the application of the NuScale LOCA EM. The acronym NPM refers to the generic NuScale Power Module design. Individual examples may be for specific design configurations. When referring to a specific design, a numerical designator is added. For example, the NPM-160 refers to the 160 MWt design and the NPM-20 refers to the module capable of housing a core up to 250 MWt rated thermal power. The acronym NIST refers generically to the NuScale Integral System Test facility. This facility has been modified to test specific NPM configurations. The NIST-1 configuration was utilized from 2015 to 2018 and the NIST-2 configuration was used from 2019 to 2022. The EM is qualified for thermal-hydraulic conditions that span normal operating conditions down to atmospheric pressure. Initially, the containment is at low absolute pressure conditions (subatmospheric). During a LOCA, the containment response depends primarily on the mass and energy release and, secondarily, on heat transfer processes on and within the containment shell. The mass and energy release does not depend on downstream (containment) conditions until the containment pressure is above atmospheric pressure. Hence, the lower limit for models and correlations used in the LOCA analysis is atmospheric pressure. The EM requires that certain checks be made and conservative assumptions be taken when building the model. This includes the generation and application of a bounding power shape and the selection of a set of thermal-mechanical properties that bounds all times in cycle. Application of the EM demonstrates that fuel does not experience CHF conditions, collapsed water level remains above the top of the active fuel, and containment remains intact and pressure and temperature remain below design limits. This assures that no fuel failure occurs and the acceptance criteria of 10 CFR 50.46 (Reference 3), excluding long-term cooling, are satisfied. The EM described in this document addresses ECCS performance in an NPM up to the time when a recirculation flow is established. Recirculation flow is considered established when pressure and level in containment and the RPV approach a stable equilibrium condition (i.e., flow is recirculating through the reactor recirculation valves (RRVs)), core heat is removed by boiling in the core, and steam exits through the reactor vent valves (RVVs). This EM does not assess radiological impacts, boron precipitation, or boron dilution. These aspects are assessed by separate methodologies. Long term cooling is addressed in the NuScale technical Report, "Long Term Cooling," TR-0916-51299 (Reference 11) (NPM-160) and topical report "Extended Passive Cooling and Reactivity Control Methodology," TR-124587 (Reference 13) for applications referencing that report. © Copyright 2022 by NuScale Power, LLC 7

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Pipe breaks inside containment are considered to be LOCA. Pipe breaks outside containment and failures in reactor pressure vessel (RPV) appurtenances, e.g., control rod drive mechanism housings and RPV nozzles and flanges, are not evaluated as part of the LOCA definition. Inadvertent opening of valves on the RPV leading to a decrease in reactor coolant system (RCS) inventory are not included in the LOCA definition per 10 CFR 50.46. However, the LOCA EM has been extended to model inadvertent RPV valve opening transients. The LOCA EM methodology is used to evaluate the mass and energy (M&E) release from the spectrum of primary system and secondary system design basis transients and accidents and the resulting CNV pressure and temperature response. The duration of the analyses is sufficient to establish the CNV peak pressure and peak temperature for all events, and to demonstrate the decrease in pressure to one-half of the peak value within 24 hours. 1.3 Abbreviations and Definitions Table 1-1 Abbreviations Term Definition ABWR advanced boiling water reactor AC alternating current AOO anticipated operational occurrence BOL beginning-of-life BWR boiling water reactor CCFL counter current flow limitation CFT critical flow test CHF critical heat flux CHFR critical heat flux ratio CPV cooling pool vessel CVCS chemical and volume control system CNV containment vessel DACS data acquisition and control system DC direct current DHRS decay heat removal system DSM direct substitution method DSRS design-specific review standard ECCS emergency core cooling system EM evaluation model EMDAP evaluation model development and assessment process EOL end-of-life FLECHT full length emergency cooling heat transfer FOM figure of merit FWIV feedwater isolation valve GDF general design framework HBM heat balance method HPCF high pressure core flooder system (for BWRs and ABWRs) HPSI high pressure safety injection (for conventional PWRs) HTFS heat transfer and fluid flow service © Copyright 2022 by NuScale Power, LLC 8

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 1-1 Abbreviations (Continued) Term Definition HTP heat transfer plate H2TS hierarchical two-tiered scaling IAB inadvertent actuation block ID inner diameter IET integral effects test INL Idaho National Laboratory IORV inadvertent opening of a reactor pressure vessel valve JIT jet impingement test KATHY Karlstein Thermal-Hydraulic test facility L/D length-to-diameter LLS linear least squares LOCA loss-of-coolant accident LP lower plenum LPFL low pressure core flooder system (for ABWRs) MASLWR Multi-Application Small Light Water Reactor MCHFR minimum critical heat flux ratio MSIV main steam isolation valve MPS module protection system NIST NuScale Integral System Test NRC U.S. Nuclear Regulatory Commission NSSS nuclear steam supply system NPM NuScale Power Module PCT peak cladding temperature PIRT phenomena identification and ranking table PWR pressurized water reactor QAPD Quality Assurance Program Description RCIC reactor core isolation cooling system (for BWRs and ABWRs) RCS reactor coolant system RG Regulatory Guide RHR residual heat removal system (conventional plants) RPV reactor pressure vessel RRV reactor recirculation valve RSV reactor safety valve RVV reactor vent valve SET separate effects test SG steam generator SIET Societ Informazioni Esperienze Termoidrauliche SRV safety relief valve SSC Structures, Systems and Components TAF top of active fuel UCP upper core plate © Copyright 2022 by NuScale Power, LLC 9

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 1-2 Definitions Term Definition Figure of merit A parameter selected to characterize the plant accident response. "Excellent" agreement One of the acceptance criteria defined in RG 1.203. "Excellent" agreement applies when the code exhibits no deficiencies in modeling a given behavior. Major and minor phenomena and trends are correctly predicted. The calculated results are judged to agree closely with the data. The calculation, with few exceptions, lies within the specified or inferred uncertainty bands of the data. The code may be used with confidence in similar applications. "Reasonable" agreement One of the acceptance criteria defined in RG 1.203. "Reasonable" agreement applies when the code exhibits minor deficiencies. Overall, the code provides an acceptable prediction. All major trends and phenomena are correctly predicted. Differences between calculation and data are greater than deemed necessary for excellent agreement. The calculation frequently lies outside but near the specified or inferred uncertainty bands of the data. However, the correct conclusions about trends and phenomena would be reached if the code was used in similar applications. "Minimal" agreement One of the acceptance criteria defined in RG 1.203. "Minimal" agreement applies when the code exhibits significant deficiencies. Overall, the code provides a prediction that is only conditionally acceptable. Some major trends or phenomena are not predicted correctly and some calculated values lie considerably outside the specified or inferred uncertainty bands of the data. Incorrect conclusions about trends and phenomena may be reached if the code were to be used in similar applications and an appropriate warning needs to be issued to users. Selected code models and facility model noding need to be reviewed, modified, and assessed before the code can be used with confidence in similar applications. © Copyright 2022 by NuScale Power, LLC 10

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 2.0 Background This topical report provides a description of the LOCA EM, developed following the guidelines in the EMDAP of RG 1.203. Six basic principles are identified in RG 1.203 as important in the process of developing and assessing an EM. Four of the principles (corresponding to the 20 steps identified in the EMDAP process) are addressed in this report. They include: determining the requirements for the EM. developing an assessment base consistent with the determined requirements. developing the EM. assessing the adequacy of the EM. The remaining principles related to establishing an appropriate quality assurance program and providing comprehensive, accurate, up-to-date documentation are addressed outside this report as part of "NuScale Topical Report: Quality Assurance Program Description," MN-122626 Rev. 0 (Reference 4). The NuScale LOCA EM specifically addresses the application of the EM to the NPM and how the EM meets the applicable requirements of 10 CFR 50 Appendix K. This report also demonstrates how the NuScale LOCA EM can be applied to evaluate ECCS performance to meet 10 CFR 50.46 acceptance criteria. The containment response analysis methodology and CNV peak pressure and temperature results are compared to applicable regulatory guidance, including the Design Specific Review Standard for NuScale Small Modular Reactor (SMR) Design, Section 6.2.1 (Reference 110). A spectrum of M&E release events is analyzed that bounds all of the LOCAs and valve-opening transients in the primary system and all secondary-system pipe-break accidents. This EM uses the NRELAP5 code that was developed from the Idaho National Laboratory (INL) RELAP5-3D© computer code. This report discusses the code modifications and modeling requirements needed to address the unique features and phenomena of an NPM design, as well as those required to comply with the applicable requirements of 10 CFR 50 Appendix K. The EM presented in this report is consistent with the applicable TMI Action Items (Reference 5) as described in the Design-Specific Review Standard for NuScale, Section 15.6.5 (Reference 6). 2.1 Loss-of-Coolant Accident Evaluation Model Roadmap Figure 2-1 shows various elements of the EMDAP as defined in RG 1.203 and provides a roadmap that relates the sections of this report to the elements and steps of the EMDAP. The EMDAP establishes the adequacy of a methodology for evaluating complex events © Copyright 2022 by NuScale Power, LLC 11

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 that are postulated to occur in nuclear power plant systems. The EMDAP described here has been developed for analyzing postulated LOCAs in an NPM. © Copyright 2022 by NuScale Power, LLC 12

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 2-1 Evaluation Model Development and Assessment Process Element 1 Establish Requirements for Evaluation Model Capability

Element 2 Element 3 Develop Assessment Base Develop Evaluation Model

Element 4 Assess Evaluation Model Adequacy ClosureRelations(Bottom-up) IntegratedEM(Top-down)

No Yes

Adequacy Decision

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-1 Evaluation Model Development and Assessment Process Steps and the Associated Sections in this Document EMDAP Description EM Section Step Element 1, Establish Requirements for Evaluation Model Capability The purpose of the LOCA EM is described in Section 1.1. Section 2.0 briefly describes the background of the process followed to develop Specify analysis the LOCA EM and the principal software used. purpose, transient Section 3.0 provides an overview of the NPM design and operation. 1 class, and power plant This includes the safety-related systems, the system logic, and class. operational phases that could occur in the NPM. The regulatory requirements with which the EM is designed to comply are described in Section 2.2. Specify figures of merit Section 4.3 discusses the FOMs which are used for the development 2 (FOMs). of the NuScale LOCA PIRT. Identify systems, components, phases, Systems, components, phases, and processes are identified as a 3 geometries, fields, and part of the NuScale LOCA PIRT discussed in Section 4.0. processes that should be modeled. Identify and rank Section 4.0 summarizes the PIRT that has been established for this 4 phenomena and EM. processes. Element 2, Develop Assessment Base Specify objectives for Section 7.0 describes objectives of the benchmarks selected for the 5 assessment base. assessment of NRELAP5 against SETs and IETs. A scaling analysis has been performed for the NPM based on the Perform scaling NuScale Integral System Test (NIST) facility. The results of the 6 analysis and identify scaling analysis are discussed in Section 8.3.2 to address the EM similarity criteria. applicability to the NPM LOCA analysis. Identify existing data Section 7.2 through Section 7.5 provide the results of the NRELAP5 and/or perform IETs validation against the SETs and IETs. In Section 8.0 these results 7 and SETs to complete are evaluated relative to NRELAP5 modeling of the high-ranked database. phenomena identified in the NuScale LOCA PIRT. Evaluate effects of IET The SET scale-up capability is evaluated in Section 8.2. NIST IET 8 distortions and SET distortions are evaluated in Section 8.3. These results justify the scale-up capability. applicability of the EM to NPM LOCA analysis. Determine experimental Section 7.0 covers experimental uncertainties for NRELAP5 9 uncertainties. assessments against the SETs and IETs. Element 3, Develop Evaluation Model The NRELAP5 development plan includes programming standards Establish EM 10 and procedures, quality assurance procedures, and configuration development plan. control, which are summarized in Section 6.1. The final structure of the LOCA EM is described in Section 5.0. The 11 Establish EM structure. NRELAP5 code description and new model features are discussed in Section 6.0. © Copyright 2022 by NuScale Power, LLC 14

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-1 Evaluation Model Development and Assessment Process Steps and the Associated Sections in this Document (Continued) EMDAP Description EM Section Step A full description of the closure models and the associated equations used in the LOCA EM is provided in the NRELAP5 theory and users Develop or incorporate 12 manuals. Section 6.2 provides a summary of NRELAP5 models and closure models. correlations. The applicability evaluation in Section 8.0 also provides further discussion of the NRELAP5 code models and correlations. Element 4, Assess Evaluation Model Adequacy Closure Relations (Bottom-up) Determine model Bottom-up assessments presented in Section 8.2 include discussion pedigree and of pedigree and applicability of dominant NRELAP5 models and 13 applicability to simulate correlations that are essential to simulate high-ranked PIRT physical processes. phenomena. Section 7.2 through Section 7.5 summarize the results of comparison Prepare input and of NRELAP5 against the selected SETs and IETs, including perform calculations to 14 evaluation of code fidelity and accuracy. These results are assess model fidelity considered in Section 8.2 to address the applicability of the EM to and accuracy. NPM LOCA analysis. Section 8.2 includes discussion of scalability of dominant NRELAP5 Assess scalability of 15 models and correlations that are essential to simulate high-ranked models. PIRT phenomena. Element 4, Assess Evaluation Model Adequacy Integrated EM (Top-down) Determine capability of field equations and NRELAP5 field equations and the numeric solution scheme are 16 numeric solutions to discussed in Section 6.2 and evaluated for their applicability to NPM represent processes LOCA in Section 8.0. and phenomena. Determine applicability The applicability of the EM to simulate the NPM system and 17 of EM to simulate components is demonstrated by assessment of NRELAP5 against system components. NuScale design-specific SETs and IETs in Section 8.3.1. Prepare input and perform calculations to Section 7.0 summarizes the results of the assessment of NRELAP5 18 assess system against NIST IET data. These results are considered in Section 8.3 interactions and global to address the applicability of the EM to NPM LOCA analysis. capability. Assess scalability of Section 8.3 provides an evaluation of scaling distortions between the 19 integrated calculations NIST IET data and the NPM design. The scalability of EM to and data for distortions. represent NPM LOCA phenomena and processes is presented. Determine EM biases This step is not required per RG 1.203 for safety analyses that 20 and uncertainties. implement 10 CFR 50 Appendix K. 2.2 Regulatory Requirements This section discusses the regulatory acceptance criteria for ECCS performance and the manner in which they are satisfied by application of the NuScale LOCA EM. © Copyright 2022 by NuScale Power, LLC 15

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 2.2.1 10 CFR 50.46 Loss-of-Coolant Accident Acceptance Criteria 10 CFR 50.46 requires that light water nuclear reactors fueled with uranium oxide pellets within cylindrical zircaloy cladding be provided with an ECCS that is designed in such a way that their calculated core cooling performance after a postulated LOCA conforms to certain criteria specified in 10 CFR 50.46(b). The five acceptance criteria are the following:

1. The calculated maximum fuel element cladding temperature shall not exceed 2200 degrees Fahrenheit (1,204 degrees Celsius).

2. The calculated total oxidation of the cladding shall nowhere exceed 0.17 times the total cladding thickness before oxidation.

3. The calculated total amount of hydrogen generated from the chemical reaction of the cladding with water or steam shall not exceed 0.01 times the hypothetical amount that would be generated if all the metal in the cladding cylinders surrounding the fuel, excluding the cladding surrounding the plenum volume, were to react.

4. Calculated changes in core geometry shall be such that the core remains amenable to cooling.

5. After any calculated successful initial operation of the ECCS, the calculated core temperature shall be maintained at an acceptably low value and decay heat shall be removed for the extended period of time required by the long-lived radioactivity remaining in the core.

The LOCA EM addresses the first four criteria as described in Section 2.2.2. The EM described in this document addresses ECCS performance in the NPM up to the time when a recirculation flow is established, pressures and levels in containment and the RPV approach a stable equilibrium condition (i.e., flow is recirculating in through the RRVs), core heat is removed by boiling in the core, and steam exits through the RVVs. 2.2.2 NuScale Loss-of-Coolant Accident Evaluation Model Acceptance Criteria An NPM is designed so that there is no core uncovery or heatup for a design-basis LOCA. As a result, peak cladding temperature (PCT) is well within the LOCA EM acceptance criterion of 2,200 degrees Fahrenheit (1,204 degrees Celsius). The parameters of interest are the collapsed liquid water level above the top of active fuel (TAF) and minimum critical heat flux ratio (MCHFR). Maintaining primary inventory and ensuring the core does not go into post-critical heat flux (CHF) heat transfer ensures that the 10 CFR 50.46(b) limitations for PCT, oxidation, and hydrogen production are protected. There is no oxidation of the cladding as a result of a LOCA. There is no hydrogen generated from the chemical reaction of the cladding with water or steam because fuel temperatures are not high enough to initiate this chemical reaction. There are no changes in core geometry resulting from a LOCA that would prevent the core from being amenable to cooling. Therefore, the first four acceptance criteria are met when © Copyright 2022 by NuScale Power, LLC 16

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 the collapsed liquid level is above the top of the active fuel and MCHFR is greater than the analysis limit for the entire time period covered by this EM (Section 7.3.6). The fifth criterion is also met during the shorter period this EM addresses. The longer-term evaluation for the fifth criterion is addressed by other NuScale methodologies (Reference 11 and Reference 13). In summary, the NuScale LOCA EM acceptance criteria are:

1. Collapsed liquid level (Section 5.1.2.6) remains above the top of the active fuel, and

2. MCHFR is greater than analysis limits (Section 6.11.5 and Section 7.3.6) 2.2.3 10 CFR 50 Appendix K The ECCS performance is calculated in conformance with the required and acceptable features of ECCS EMs specified in 10 CFR 50 Appendix K, and is calculated for a number of cases to provide assurance that the most severe postulated LOCAs are identified. 10 CFR 50.46 provides two options for an acceptable LOCA EM. Paragraph 50.46(a)(1)(i) allows for a best-estimate approach to be followed and Paragraph 50.46.(a)(ii) allows for the conservative deterministic approach detailed in 10 CFR 50 Appendix K. In view of the large safety margins in the NPM, the deterministic bounding approach in Paragraph 50.46(a)(1)(ii) is used by NuScale.

An NPM is designed to reduce the consequences of design-basis LOCAs compared to existing light water reactors for which 10 CFR 50 Appendix K was developed. Consequently, many of the phenomena that are the subject of 10 CFR 50 Appendix K requirements are not encountered in design-basis NPM LOCAs in the NPM. That is, certain phenomena have been designed out of an NPM and, therefore, a number of requirements are satisfied by design rather than by analysis. Examples of phenomena and processes that can occur during a typical pressurized water reactor (PWR) LOCA that do not occur during an NPM LOCA include: loop seal clearing pump coastdown two-phase pump performance entry of significant amounts of non-condensable gases into the system core uncovery core refilling core reflooding cladding swelling and rupture metal-water reaction post-CHF heat transfer © Copyright 2022 by NuScale Power, LLC 17

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 cladding rewet ECCS bypass Hence, only a subset of the phenomena that are addressed in 10 CFR 50 Appendix K is encountered in the design-basis NPM LOCAs and thus relevant to the LOCA EM. Table 2-2 lists each required and acceptable feature of the EM specified in 10 CFR 50 Appendix K and describes the manner in which the LOCA EM addresses each feature. The LOCA EM includes model features required by Appendix K that are relevant to the NPM. Features required by 10 CFR 50 Appendix K that are not relevant to the LOCA EM are identified in Table 2-2 as either "satisfied by design" or "excluded from model." A feature "satisfied by design" means that a 10 CFR 50 Appendix K required feature is expressly or impliedly conditional on the presence of process or phenomena in the design or analysis. Because such process or phenomena does not exist for an NPM, the required feature is not applicable and not included in the LOCA EM. For example, there are no reactor coolant pumps in an NPM. Therefore the phenomena that are the subject of 10 CFR 50 Appendix K Requirement I.C.6 "Pump Modeling" are not encountered because of the design of the NPM, and thus the required model features are "satisfied by design." A feature "excluded" from the EM means that 10 CFR 50 Appendix K directly requires the feature, without condition on the presence of a process or phenomena, but that the feature is not relevant to the LOCA EM. Table 2-2 technically justifies the exclusion of such feature from the model. However, an applicant or licensee referencing this report will be required to address regulatory compliance with 10 CFR 50.46 and 10 CFR 50 Appendix K (e.g., by seeking an exemption from that required feature). Similarly, an "acceptable alternative" model feature is technically justified by Table 2-2, but does not strictly meet the 10 CFR 50 Appendix K required feature, and thus an applicant or licensee referencing this report will be required to address regulatory compliance. Historically, RELAP5 has been applied to evaluate post-CHF fuel conditions for events in LWRs. While these features have been retained in NRELAP5, the application of the LOCA EM to predict fuel temperature response is limited to pre-CHF heat transfer regimes. In the LOCA EM, applicable closure models or correlations required by 10 CFR 50 Appendix K are used. The LOCA EM also uses appropriate closure models or correlations in addition to those required in 10 CFR 50 Appendix K. All closure models and correlations are verified and validated for use within their range of applicability. © Copyright 2022 by NuScale Power, LLC 18

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance 10 CFR 50 Appendix K Required and Acceptable NuScale LOCA EM Feature I.A Sources of heat during the LOCA: For the heat sources listed in paragraphs I.A.1 to 4 of this appendix, it must be assumed that the reactor has been operating continuously at a power level at The initial power level is set at 102 percent of rated least 1.02 times the licensed power level (to allow for power. The maximum radial peaking factor is used in instrumentation error) with the maximum peaking the hot assembly to bound all possible power factor allowed by the technical specifications. An peaking. (Section 9.6.6). Sensitivity calculations are assumed power level lower than the level specified performed with different axial power shapes that in this paragraph (but not less than the licensed bound maximum axial power peaking. power level) may be used provided the proposed alternative value has been demonstrated to account Further discussion on core power distribution is for uncertainties due to power level instrumentation provided in Section 5.1.2.2.3. error. A range of power distribution shapes and peaking factors representing power distributions that Therefore, the required features of I.A are included may occur over the core lifetime must be studied. in the LOCA EM. The selected combination of power distribution shape and peaking factor should be the one that results in the most severe calculated consequences for the spectrum of postulated breaks and single failures that are analyzed. I.A.1 The Initial Stored Energy in the Fuel: Based on the burn-up dependent fuel performance The steady-state temperature distribution and stored analysis, it is determined that choosing end-of-life energy in the fuel before the hypothetical accident (EOL) fuel thermal conductivity and beginning-of-life shall be calculated for the burn-up that yields the (BOL) volumetric heat capacity and fuel-cladding highest calculated cladding temperature (or, gap conductance maximizes the initial stored energy optionally, the highest calculated stored energy.) To in the fuel. An additional 15 percent bias is applied to accomplish this, the thermal conductivity of the UO2 both volumetric heat capacity and thermal shall be evaluated as a function of burn-up and conductivity to maximize the initial stored energy. temperature, taking into consideration differences in initial density, and the thermal conductance of the Further discussion on selection of fuel rod gap between the UO2 and the cladding shall be mechanical property input is provided in evaluated as a function of the burn-up, taking into Section 5.1.2.2.4. consideration fuel densification and expansion, the composition and pressure of the gases within the Therefore, the required features of I.A.1 are included fuel rod, the initial cold gap dimension with its in the LOCA EM. tolerances, and cladding creep. © Copyright 2022 by NuScale Power, LLC 19

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance (Continued) 10 CFR 50 Appendix K Required and Acceptable NuScale LOCA EM Feature A point kinetics model is used to calculate fission I.A.2 Fission Heat: power. Credit is taken for reactor trip. A conservative control rod insertion curve is used along with a Fission heat shall be calculated using reactivity and minimum rod worth and conservative delay in reactor kinetics. Shutdown reactivities resulting from initiation of rod insertion. The most reactive control temperatures and voids shall be given their minimum rod is assumed to be stuck out of the core. Doppler plausible values, including allowance for and moderator density coefficients are calculated uncertainties, for the range of power distribution conservatively (Section 5.1.2.2.5). shapes and peaking factors indicated to be studied above. Rod trip and insertion may be assumed if Therefore, the required features of I.A.2 are included they are calculated to occur. in the LOCA EM. The 1979 ANS actinide decay heat standard is I.A.3 Decay of Actinides: applied which includes the decay of neptunium and plutonium (Section 5.1.2.2.5). The heat from the radioactive decay of actinides, including neptunium and plutonium generated during The actinide decay heat assumes infinite operating operation, as well as isotopes of uranium, shall be time to maximize actinide concentration. This calculated in accordance with fuel cycle calculations assumption results in the highest calculated fuel and known radioactive properties. The actinide temperature during the LOCA. decay heat chosen shall be that appropriate for the time in the fuel cycle that yields the highest Therefore, the required features of I.A.3 are included calculated fuel temperature during the LOCA. in the LOCA EM. I.A.4 Fission Product Decay: The heat generation rates from radioactive decay of fission products shall be assumed to be equal to 1.2 times the values for infinite operating time in the ANS Standard (Proposed American Nuclear Society The 1973 ANS decay heat standard (Reference 44) Standards--Decay Energy Release Rates Following is used with a 20 percent uncertainty added to the Shutdown of Uranium-Fueled Thermal Reactors. base value. A bounding form of the 1973 ANS Approved by Subcommittee ANS-5, ANS Standards standard in NRELAP5 meets the intent of the 10 Committee, October 1971). This standard has been CFR 50 Appendix K requirement (Section 5.1.2.2.5). approved for incorporation by reference by the Director of the Federal Register. A copy of the Therefore, the LOCA EM includes an acceptable standard is available for inspection at the NRC alternative to the requirement of I.A.4. Library, 11545 Rockville Pike, Rockville, Maryland 20852-2738. The fraction of the locally generated gamma energy that is deposited in the fuel (including the cladding) may be different from 1.0; the value used shall be justified by a suitable calculation. © Copyright 2022 by NuScale Power, LLC 20

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance (Continued) 10 CFR 50 Appendix K Required and Acceptable NuScale LOCA EM Feature I.A.5 Metal-Water Reaction Rate: The rate of energy release, hydrogen generation, and cladding oxidation from the metal-water reaction shall be calculated using the Baker-Just equation (Baker, L., Just, L.C., Studies of Metal Water Calculated cladding temperatures for design basis Reactions at High Temperatures, III. Experimental LOCAs are well below the level where cladding and Theoretical Studies of the Zirconium-Water oxidation occurs on a time scale of a LOCA event for Reaction, ANL-6548, page 7, May 1962). This the NPM (results of LOCA break spectrum publication has been approved for incorporation by calculations in Section 9.0). Therefore, this reference by the Director of the Federal Register. A requirement is not relevant to an NPM, which copy of the publication is available for inspection at precludes fuel temperature reaching CHF and any the NRC Library, 11545 Rockville Pike, Two White significant fuel cladding heatup. For the NuScale Flint North, Rockville, Maryland 20852-2738. The LOCA EM, core coverage and an MCHFR greater reaction shall be assumed not to be steam limited. than the analysis limit precludes the occurrence of For rods whose cladding is calculated to rupture cladding oxidation. during the LOCA, the inside of the cladding shall be assumed to react after the rupture. The calculation Therefore, the required features of I.A.5 are of the reaction rate on the inside of the cladding shall excluded from the LOCA EM. also follow the Baker-Just equation, starting at the time when the cladding is calculated to rupture, and extending around the cladding inner circumference and axially no less than 1.5 inches each way from the location of the rupture, with the reaction assumed not to be steam limited. The NRELAP5 plant model explicitly represents all major reactor internal heat structures. Heat structures are also included for the primary and I.A.6 Reactor Internals Heat Transfer: secondary system pressure boundary materials. Section 5.1.2 contains details of the internal heat Heat transfer from piping, vessel walls, and non-fuel structures represented in the NuScale LOCA EM. internal hardware shall be taken into account. Therefore, the required features of I.A.6 are included in the LOCA EM. Heat transfer through the steam generator (SG) I.A.7 Pressurized Water Reactor tubes is included in the EM. The model is validated Primary-to-Secondary Heat Transfer: using experimental data from Societ Italiana Esperienze Termoidrauliche (SIET) tests (see Heat transferred between primary and secondary Section 7.4) and NIST tests (see Section 7.5). systems through heat exchangers (steam generators) shall be taken into account. (Not Therefore, the required features of I.A.7 are included applicable to boiling water reactors (BWRs).) in the LOCA EM. © Copyright 2022 by NuScale Power, LLC 21

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance (Continued) 10 CFR 50 Appendix K Required and Acceptable NuScale LOCA EM Feature I.B Swelling and Rupture of the Cladding and Fuel Rod Thermal Parameters: Each evaluation model shall include a provision for predicting cladding swelling and rupture from consideration of the axial temperature distribution of Calculated cladding temperatures for design basis the cladding and from the difference in pressure LOCAs in the NPM are well below the threshold for between the inside and outside of the cladding, both cladding swelling and rupture (see the results of as functions of time. To be acceptable the swelling LOCA break spectrum calculations in Section 9.0). and rupture calculations shall be based on Peak cladding temperatures in the NPM occur at applicable data in such a way that the degree of steady state normal operation. Because swelling swelling and incidence of rupture are not and rupture do not occur during normal operation, underestimated. The degree of swelling and rupture they do not occur in a NPM LOCA event. Therefore, shall be taken into account in calculations of gap this requirement is not relevant for the LOCA EM as conductance, cladding oxidation and embrittlement, core coverage precludes the occurrence of cladding and hydrogen generation. swelling and rupture. The calculations of fuel and cladding temperatures Therefore, the required features of I.B are excluded as a function of time shall use values for gap from the LOCA EM. conductance and other thermal parameters as functions of temperature and other applicable time-dependent variables. The gap conductance shall be varied in accordance with changes in gap dimensions and any other applicable variables. I.C Blowdown Phenomena A complete spectrum of break sizes and locations is analyzed up to the largest penetrations in the RPV I.C.1.a Break Characteristics and Flow: including the double-ended guillotine break where appropriate. The size of the pipes precludes the In analyses of hypothetical LOCAs, a spectrum of impact of longitudinal split breaks in the NPM possible pipe breaks shall be considered. This design. Therefore, the requirement for analyzing the spectrum shall include instantaneous double-ended effect of longitudinal split break is not relevant to the breaks ranging in cross-sectional area up to and LOCA EM. including that of the largest pipe in the primary coolant system. The analysis shall also include the Further discussion of break spectrum analysis is effects of longitudinal splits in the largest pipes, with provided in Section 5.4. The break spectrum the split area equal to the cross-sectional area of the calculation results are available in Section 9.0. pipe. Therefore, the required features of I.C.1.a are included or satisfied by design in the LOCA EM. © Copyright 2022 by NuScale Power, LLC 22

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance (Continued) 10 CFR 50 Appendix K Required and Acceptable NuScale LOCA EM Feature The required Moody critical flow is used when the I.C.1.b Discharge Model: break flow is calculated to be two-phase flow and the (( }}2(a),(c) model is used to calculate For all times after the discharging fluid has been single-phase choked flow (( calculated to be two-phase in composition, the discharge rate shall be calculated by use of the }}2(a),(c) For the NPM, single phase flow Moody model (F.J. Moody, Maximum Flow Rate of a through the break may recur after the transition to Single Component, Two-Phase Mixture, Journal of two-phase flow. The (( Heat Transfer, Trans American Society of Mechanical Engineers, 87, No. 1, February, 1965). }}2(a),(c) The calculation shall be conducted with at least model is conservative for single-phase break flow. three values of a discharge coefficient applied to the See Section 6.6.1 for details. postulated break area, these values spanning the range from 0.6 to 1.0. If the results indicate that the The range of postulated break sizes in the break maximum cladding temperature for the hypothetical analysis covers the 10 CFR 50 Appendix K required accident is to be found at an even lower value of the range of discharge coefficient, as discussed in discharge coefficient, the range of discharge Section 5.4. coefficients shall be extended until the maximum cladding temperatures calculated by this variation Therefore, the required features of I.C.1.b, including has been achieved. an acceptable alternative feature, are included in the LOCA EM. © Copyright 2022 by NuScale Power, LLC 23

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance (Continued) 10 CFR 50 Appendix K Required and Acceptable NuScale LOCA EM Feature I.C.1.c End of Blowdown. (Applies Only to Pressurized Water Reactors): For postulated cold leg breaks, all emergency cooling water injected into the inlet lines or the reactor vessel during the bypass period shall in the calculations be subtracted from the reactor vessel calculated inventory. This may be executed in the calculation during the bypass period, or as an alternative the amount of emergency core cooling water calculated to be injected during the bypass For an NPM, there are no cold legs and hence no period may be subtracted later in the calculation cold leg breaks. All of the coolant that exits the break from the water remaining in the inlet lines, remains in the containment and is available to return downcomer, and reactor vessel lower plenum after when the RRVs are opened. Emergency core the bypass period. This bypassing shall end in the cooling system bypass cannot occur in the NPM, so calculation at a time designated as the end of this requirement is not relevant to the LOCA EM. bypass, after which the expulsion or entrainment mechanisms responsible for the bypassing are Therefore, the required features of I.C.1.c are calculated not to be effective. The end-of-bypass satisfied by design. definition used in the calculation shall be justified by a suitable combination of analysis and experimental data. Acceptable methods for defining end of bypass include, but are not limited to, the following: (1) Prediction of the blowdown calculation of downward flow in the downcomer for the remainder of the blowdown period; (2) Prediction of a threshold for droplet entrainment in the upward velocity, using local fluid conditions and a conservative critical Weber number. Noding sensitivity studies have been conducted to I.C.1.d Noding Near the Break and the ECCS demonstrate that the calculated conditions in the Injection Points: vicinity of the break locations, RVVs, and RRVs are reliable. The noding in the vicinity of and including the broken or split sections of pipe and the points of ECCS The results of the noding sensitivity studies are injection shall be chosen to permit a reliable analysis discussed in Section 9.6.1. of the thermodynamic history in these regions during blowdown. Therefore, the required features of I.C.1.d are included in the LOCA EM. © Copyright 2022 by NuScale Power, LLC 24

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance (Continued) 10 CFR 50 Appendix K Required and Acceptable NuScale LOCA EM Feature I.C.2 Frictional Pressure Drops: The frictional losses in pipes and other components including the reactor core shall be calculated using models that include realistic variation of friction factor with Reynolds number, and realistic two-phase friction multipliers that have been adequately verified by comparison with experimental Friction losses in pipes and components are data, or models that prove at least equally calculated using Reynolds number-dependent conservative with respect to maximum cladding friction factors. The NRELAP5 wall friction model is temperature calculated during the hypothetical based on a two-phase multiplier approach accident. The modified Baroczy correlation (Section 6.2.4). The models used in NRELAP5 have (Baroczy, C. J., A Systematic Correlation for been validated for the range of conditions Two-Phase Pressure Drop, Chem. Enging. Prog. encountered in design-basis LOCAs as shown by Symp. Series, No. 64, Vol. 62, 1965) or a assessment against SETs and IETs in Section 7.0. combination of the Thom correlation (Thom, J.R.S., Prediction of Pressure Drop During Forced Therefore, the required features of I.C.2 are included Circulation Boiling of Water, Int. J. of Heat & Mass in the LOCA EM. Transfer, 7, 709-724, 1964) for pressures equal to or greater than 250 psia and the Martinelli-Nelson correlation (Martinelli, R. C. Nelson, D.B., Prediction of Pressure Drop During Forced Circulation Boiling of Water, Transactions of ASME, 695-702, 1948) for pressures lower than 250 psia is acceptable as a basis for calculating realistic two-phase friction multipliers. I.C.3 Momentum Equation: All of the momentum equation effects required by The following effects shall be taken into account in Section I.C.3 are included in NRELAP5 the conservation of momentum equation: (1) (Section 6.2.1 and Section 6.2.4). Benchmarks for temporal change of momentum, (2) momentum the NIST facility and other assessments show that convection, (3) area change momentum flux, (4) simulations made by NRELAP5 are acceptable, momentum change due to compressibility, (5) based on reasonable-to- excellent agreement with pressure loss resulting from wall friction, (6) experimental data (Section 7.0). pressure loss resulting from area change, and (7) gravitational acceleration. Any omission of one or Therefore, the required features of I.C.3 are included more of these terms under stated circumstances in the LOCA EM. shall be justified by comparative analyses or by experimental data. © Copyright 2022 by NuScale Power, LLC 25

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance (Continued) 10 CFR 50 Appendix K Required and Acceptable NuScale LOCA EM Feature The initial blowdown is evaluated using the (( }}2(a),(c) discussed in Section 6.11. The remainder of the event uses two CHF correlations to monitor for CHF occurrence, I.C.4.a (Critical Heat Flux): (( Correlations developed from appropriate steady-state and transient-state experimental data }}2(a),(c) are acceptable for use in predicting the CHF during Section 6.11 contains a description of the LOCA transients. The computer programs in which correlations. Section 7.3 describes the assessment these correlations are used shall contain suitable against the NuScale CHF data that bounds the checks to ensure that the physical parameters are range of LOCA parameters. The LOCA EM checks within the range of parameters specified for use of to ensure that the physical parameters are within the the correlations by their respective authors. range of parameters specified for use of the correlations. Therefore, the required features of I.C.4.a are included in the LOCA EM. I.C.4.b (Critical Heat Flux): I.C.4.b identifies acceptable, but not required, EM Steady-state CHF correlations acceptable for use in features. The LOCA EM includes an acceptable LOCA transients include, but are not limited to, the steady-state CHF correlation as addressed by following: [six acceptable CHF correlations are I.C.4.a. identified in 10 CFR 50 Appendix K, I.C.4.b]. I.C.4.c (Critical Heat Flux): Correlations of appropriate transient CHF data may be accepted for use in LOCA transient analyses if comparisons between the data and the correlations are provided to demonstrate that the correlations I.C.4.c identifies acceptable, but not required, EM predict values of CHF which allow for uncertainty in features. The LOCA EM does not use a transient the experimental data throughout the range of CHF correlation. I.C.4.a. parameters for which the correlations are to be used. Where appropriate, the comparisons shall use statistical uncertainty analysis of the data to demonstrate the conservatism of the transient correlation. I.C.4.d (Critical Heat Flux): I.C.4.d identifies acceptable, but not required, EM Transient CHF correlations acceptable for use in features. The LOCA EM does not use a transient LOCA transients include, but are not limited to, the CHF correlation. I.C.4.a. following: (GE transient CHF correlation is listed in 10 CFR 50 Appendix K, I.C.4.d.) © Copyright 2022 by NuScale Power, LLC 26

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance (Continued) 10 CFR 50 Appendix K Required and Acceptable NuScale LOCA EM Feature I.C.4.e (Critical Heat Flux): After CHF is first predicted at an axial fuel rod CHF does not occur in an NPM for LOCAs. Heat location during blowdown, the calculation shall not transfer beyond CHF is not a phenomenon use nucleate boiling heat transfer correlations at that encountered during a design-basis LOCA. location subsequently during the blowdown even if The LOCA methodology does not calculate heat the calculated local fluid and surface conditions transfer beyond CHF in the core. would apparently justify the reestablishment of nucleate boiling. Heat transfer assumptions Therefore, this requirement is satisfied by a design characteristic of return to nucleate boiling (rewetting) that has a margin to CHF for LOCA events. shall be permitted when justified by the calculated local fluid and surface conditions during the reflood portion of a LOCA. I.C.5.a (Post-CHF Heat Transfer Correlations): Correlations of heat transfer from the fuel cladding to the surrounding fluid in the post-CHF regimes of CHF does not occur in an NPM for LOCAs. Heat transition and film boiling shall be compared to transfer beyond CHF is not a phenomenon applicable steady-state and transient-state data encountered during a design-basis LOCA. using statistical correlation and uncertainty analyses. Therefore, this requirement is not relevant to the Such comparison shall demonstrate that the LOCA EM. correlations predict values of heat transfer co-efficient equal to or less than the mean value of Therefore, the required features of I.C.5.a are the applicable experimental heat transfer data excluded from the LOCA EM. throughout the range of parameters for which the correlations are to be used. The comparisons shall quantify the relation of the correlations to the statistical uncertainty of the applicable data. © Copyright 2022 by NuScale Power, LLC 27

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance (Continued) 10 CFR 50 Appendix K Required and Acceptable NuScale LOCA EM Feature I.C.5.b (Post-CHF Heat Transfer Correlations): The Groeneveld flow film boiling correlation (equation 5.7 of D.C. Groeneveld, An Investigation of Heat Transfer in the Liquid Deficient Regime, AECL-3281, revised December 1969) and the Westinghouse correlation of steady-state transition boiling (Proprietary Redirect/Rebuttal Testimony of Westinghouse Electric Corporation, USNRC Docket RM-50-1, page 25-1, October 26, 1972) are acceptable for use in the post-CHF boiling regimes. In addition, the transition boiling correlation of McDonough, Milich, and King (J.B. McDonough, W. Milich, E.C. King, An Experimental Study of Partial Film Boiling Region with Water at Elevated Pressures in a Round Vertical Tube, Chemical Engineering Progress Symposium Series, Vol. 57, I.C.5.b identifies acceptable, but not required, EM No. 32, pages 197-208, (1961) is suitable for use features. The LOCA methodology does not calculate between nucleate and film boiling. Use of all these heat transfer beyond CHF in the core. Therefore, correlations is restricted as follows: these acceptable correlations are not relevant to the EM. I.C.5.a. (1) The Groeneveld correlation shall not be used in the region near its low-pressure singularity, (2) The first term (nucleate) of the Westinghouse correlation and the entire McDonough, Milich, and King correlation shall not be used during the blowdown after the temperature difference between the cladding and the saturated fluid first exceeds

300F, (3) Transition boiling heat transfer shall not be reapplied for the remainder of the LOCA blowdown, even if the cladding superheat returns below 300F, except for the reflood portion of the LOCA when justified by the calculated local fluid and surface conditions.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance (Continued) 10 CFR 50 Appendix K Required and Acceptable NuScale LOCA EM Feature I.C.5.c (Post-CHF Heat Transfer Correlations): Evaluation models approved after October 17, 1988, which make use of the Dougall-Rohsenow flow film boiling correlation (R.S. Dougall and W.M. Rohsenow, Film Boiling on the Inside of Vertical Tubes with Upward Flow of Fluid at Low Qualities, MIT Report Number 9079 26, Cambridge, Massachusetts, September 1963) may not use this correlation under conditions where nonconservative predictions of heat transfer result. Evaluation models that make use of the Dougall-Rohsenow correlation I.C.5.c identifies acceptable, but not required, EM and were approved prior to October 17, 1988, features. The LOCA methodology does not calculate continue to be acceptable until a change is made to, heat transfer beyond CHF in the core. Therefore, or an error is corrected in, the evaluation model that these acceptable correlations are not relevant to the results in a significant reduction in the overall EM. I.C.5.a. conservatism in the evaluation model. At that time continued use of the Dougall-Rohsenow correlation under conditions where nonconservative predictions of heat transfer result will no longer be acceptable. For this purpose, a significant reduction in the overall conservatism in the evaluation model would be a reduction in the calculated peak fuel cladding temperature of at least 50F from that which would have been calculated on October 17, 1988, due either to individual changes or error corrections or the net effect of an accumulation of changes or error corrections. I.C.6 Pump Modeling: The characteristics of rotating primary system pumps (axial flow, turbine, or centrifugal) shall be derived from a dynamic model that includes momentum transfer between the fluid and the There are no primary system coolant pumps, so the rotating member, with variable pump speed as a requirements related to pump models are not function of time. The pump model resistance used relevant to the LOCA EM, as shown in Section 3.0. for analysis should be justified. The pump model for the two-phase region shall be verified by applicable Therefore, the required features of I.C.6 are satisfied two-phase pump performance data. For BWRs after by design. saturation is calculated at the pump suction, the pump head may be assumed to vary linearly with quality, going to zero for one percent quality at the pump suction, so long as the analysis shows that core flow stops before the quality at pump suction reaches one percent. © Copyright 2022 by NuScale Power, LLC 29

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance (Continued) 10 CFR 50 Appendix K Required and Acceptable NuScale LOCA EM Feature The core is represented by three channels: hot channel represents hot assembly, average channel represents rest of the core assemblies, and total core bypass. Crossflow is enabled between the core I.C.7.a Core Flow Distribution During Blowdown. regions. Section 6.11.5 describes modeling used to (Applies only to pressurized water reactors): result in an appropriately conservative flow distribution during the initial blowdown. The flow rate through the hot region of the core during blowdown shall be calculated as a function of Cladding swelling or rupture does not occur because time. For the purpose of these calculations the hot the fuel does not encounter a CHF event and region chosen shall not be greater than the size of because the core remains covered throughout the one fuel assembly. Calculations of average flow and LOCA event. Therefore, cross flows are not flow in the hot region shall take into account cross impacted by geometrical changes in the fuel during flow between regions and any flow blockage the transient. calculated to occur during blowdown as a result of cladding swelling or rupture. The calculated flow Due to the mild nature of natural circulation flow shall be smoothed to eliminate any calculated rapid during blowdown, rapid oscillations during the LOCA oscillations (period less than 0.1 seconds). transient with a period less than 0.1 second do not occur. Therefore, smoothing is not necessary. Therefore, the required features of I.C.7.a are included in the LOCA EM. The intention of the I.C.7.b requirement is to ensure that LOCA EMs that assessed the hot channel separately would use the correct thermal-hydraulic I.C.7.b Core Flow Distribution During Blowdown. boundary conditions. (Applies only to pressurized water reactors)]: For the LOCA EM, the active core is represented by A method shall be specified for determining the (( enthalpy to be used as input data to the hot channel }}2(a),(c) The hot heatup analysis from quantities calculated in the channel is not analyzed in a separate code, but is blowdown analysis, consistent with the flow included in the NPM model. (Section 5.1.2.2 for core distribution calculations. nodalization.) Therefore, the required features of I.C.7.b are included in the LOCA EM. I.D Post-Blowdown Phenomena; Heat Removal by the ECCS Safety-related system single failures considered for I.D.1 Single Failure Criterion: break spectrum calculations are discussed in Section 5.4.3. An evaluation of ECCS failure modes An analysis of possible failure modes of ECCS has been performed. Sensitivity studies were equipment and of their effects on ECCS conducted to determine the limiting single failure for performance must be made. In carrying out the each type of LOCA. The results of break spectrum accident evaluation the combination of ECCS calculations are discussed in Section 9.0. subsystems assumed to be operative shall be those available after the most damaging single failure of Therefore, the required features of I.D.1 are included ECCS equipment has taken place. in the LOCA EM. © Copyright 2022 by NuScale Power, LLC 30

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance (Continued) 10 CFR 50 Appendix K Required and Acceptable NuScale LOCA EM Feature The NPM containment design is intended to equilibrate RCS and containment vessel (CNV) I.D.2 Containment Pressure: pressure when ECCS has been actuated. Condensed effluent is then returned to the RCS in The containment pressure used for evaluating natural circulation flow. Although there are no active cooling effectiveness during reflood and spray pressure-reducing systems, the CNV is immersed in cooling shall not exceed a pressure calculated the reactor pool, resulting in significant condensation conservatively for this purpose. The calculation shall and cooling of effluent prior to returning to the RPV. include the effects of operation of all installed pressure-reducing systems and processes. Therefore, the required features of I.D.2 are satisfied by design. I.D.3 Calculation of Reflood Rate for Pressurized Water Reactors: The refilling of the reactor vessel and the time and rate of reflooding of the core shall be calculated by an acceptable model that takes into consideration the thermal and hydraulic characteristics of the core and of the reactor system. The primary system coolant pumps shall be assumed to have locked impellers if this assumption leads to the maximum calculated cladding temperature; otherwise the pump rotor shall be assumed to be running free. The Refilling or reflooding is not required for the NuScale ratio of the total fluid flow at the core exit plane to the design as in a conventional PWR, because there is total liquid flow at the core inlet plane (carryover no core uncovery (results of LOCA break spectrum fraction) shall be used to determine the core exit flow calculations in Section 9.0). This requirement is not and shall be determined in accordance with relevant to the NuScale LOCA EM. applicable experimental data (for example, PWR FLECHT (Full Length Emergency Cooling Heat There are no primary system coolant pumps, so the Transfer) Final Report, Westinghouse Report requirements related to pump models are satisfied WCAP-7665, April 1971; PWR Full Length by an NPM. Also, there are no accumulators, so Emergency Cooling Heat Transfer (FLECHT) Group requirements related to accumulator discharge are I Test Report, Westinghouse Report WCAP-7435, satisfied by being designed out of the NPM. January 1970; PWR FLECHT (Full Length Emergency Cooling Heat Transfer) Group II Test Report, Westinghouse Report WCAP-7544, September 1970; PWR FLECHT Final Report Supplement, Westinghouse Report WCAP-7931, October 1972). The effects on reflooding rate of the compressed gas in the accumulator which is discharged following accumulator water discharge shall also be taken into account. © Copyright 2022 by NuScale Power, LLC 31

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance (Continued) 10 CFR 50 Appendix K Required and Acceptable NuScale LOCA EM Feature Refilling or reflooding is not required for the NuScale design as in a conventional PWR, because there is I.D.4 Steam Interaction with Emergency Core no core uncovery (results of LOCA break spectrum Cooling Water in Pressurized Water Reactors: calculations in Section 9.0). Traditional concerns regarding steam interaction with injected ECCS The thermal-hydraulic interaction between steam water are not a factor in the NuScale design, and all emergency core cooling water shall be taken although the phenomenon of non-equilibrium into account in calculating the core reflooding rate. conditions existing between steam and subcooled During refill and reflood, the calculated steam flow in liquid does occur. For the NuScale design, such unbroken reactor coolant pipes shall be taken to be interactions could occur in either the CNV or in the zero during the time that accumulators are downcomer when subcooled containment liquid discharging water into those pipes unless enters from the RRVs. While I.D.4 is not relevant to experimental evidence is available regarding the the LOCA EM, the intent of this requirement is realistic thermal-hydraulic interaction between the addressed by the capability of NRELAP5 to model steam and the liquid. In this case, the experimental thermal non-equilibrium states and by the NPM data may be used to support an alternate design which minimizes these phenomena. assumption. Therefore, the required features of I.D.4 are satisfied by design. © Copyright 2022 by NuScale Power, LLC 32

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-2 10 CFR 50 Appendix K Required and Acceptable Features Compliance (Continued) 10 CFR 50 Appendix K Required and Acceptable NuScale LOCA EM Feature I.D.5.a (Refill and Reflood Heat Transfer for Pressurized Water Reactors): For reflood rates of one inch per second or higher, reflood heat transfer coefficients shall be based on applicable experimental data for unblocked cores including FLECHT results (PWR FLECHT (Full Length Emergency Cooling Heat Transfer) Final Report, Westinghouse Report WCAP-7665, April 1971). The use of a correlation derived from FLECHT data shall be demonstrated to be conservative for the transient to which it is applied; Refilling or reflooding is not required for the NuScale presently available FLECHT heat transfer design as in a conventional PWR, because there is correlations (PWR Full Length Emergency Cooling no core uncovery (results of LOCA break spectrum Heat Transfer (FLECHT) Group I Test Report, calculations in Section 9.0). This requirement is not Westinghouse Report WCAP-7544, September relevant to the LOCA EM. 1970; PWR FLECHT Final Report Supplement, Westinghouse Report WCAP-7931, October 1972) Therefore, the required features of I.D.5.a are are not acceptable. Westinghouse Report satisfied by design. WCAP-7665 has been approved for incorporation by reference by the Director of the Federal Register. A copy of this report is available for inspection at the NRC Library, 11545 Rockville Pike, Rockville, Maryland 20852-2738. New correlations or modifications to the FLECHT heat transfer correlations are acceptable only after they are demonstrated to be conservative, by comparison with FLECHT data, for a range of parameters consistent with the transient to which they are applied. I.D.5.b (Refill and Reflood Heat Transfer for Pressurized Water Reactors): Refilling or reflooding is not required for the NuScale design as in a conventional PWR, because there is During refill and during reflood when reflood rates no core uncovery (results of LOCA break spectrum are less than one inch per second, heat transfer calculations in Section 9.0). This requirement is not calculations shall be based on the assumption that relevant to the LOCA EM. cooling is only by steam, and shall take into account any flow blockage calculated to occur as a result of Therefore, the required features of D.5.b are cladding swelling or rupture as such blockage might satisfied by design. affect both local steam flow and heat transfer. I.D.6.Convective Heat Transfer Coefficients for The NuScale plant is not a BWR and does not have Boiling Water Reactor Fuel Rods Under Spray core spray cooling. Therefore, this requirement is not Cooling. applicable to the LOCA EM. The NuScale plant is not a BWR and does not have I.D.7 The Boiling Water Reactor Channel Box Under channel boxes. Therefore, this requirement is not Spray Cooling. applicable to the LOCA EM. © Copyright 2022 by NuScale Power, LLC 33

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 2.2.4 Other Requirements Per the Design-Specific Review Standard for NuScale SMR Design, Section 4.4 (Reference 7), the thermal-hydraulic design should account for the effects of crud in the CHF calculations in the core or in the pressure drop throughout the RCS. NuScale requires that the fuel supplied for the NPM be supported with a qualified and approved product that supports this regulatory requirement. It is, however, acknowledged that crud deposition is driven by factors beyond fuel design, such as operating conditions and RCS chemistry. In order to evaluate the impact of crud on the LOCA FOMs, NuScale has evaluated the effect of the changes in thermal properties of the maximum credible crud thickness on fuel centerline and cladding temperatures during a LOCA. This evaluation determines that although the initial stored energy increases as a result of crud, there is no significant impact on the LOCA response. 2.2.4.1 Containment Response Analysis Regulatory Requirements The Nuclear Regulatory Commission (NRC) regulations and regulatory guidance applicable to the containment response analysis methodology are described in this section. The elements of the containment response analysis methodology that address each of these regulations and requirements are discussed. 2.2.5 Compliance with Containment Response Analysis Related Regulatory Requirements 2.2.5.1 10 CFR 50 Appendix A - General Design Criteria for Nuclear Power Plants The General Design Criteria (GDC) for Nuclear Power Plants, Appendix A to 10 CFR 50 (Reference 109), include the NRC regulations applicable to the containment response methodology. Compliance with GDC 16 and 50 and PDC 38 is as follows: General Design Criterion 16 - The analyses performed per the containment response analysis methodology are used to establish the limiting CNV pressure and temperature conditions resulting from the spectrum of design-basis primary system and secondary system M&E releases resulting from pipe breaks and valve actuations. The CNV is designed to ensure that the design pressure and temperature limits are not exceeded as demonstrated by the analysis results. Principal Design Criterion 38 - The analyses performed per the containment response analysis methodology establish the performance of NPM containment heat removal and demonstrate that the containment peak pressure and temperature are rapidly reduced. The methodology addresses LOCAs, valve-opening events and secondary pipe breaks. Following containment isolation and opening of the ECCS valves, the containment heat removal function is passive and does not require electric power. The requirement to rapidly reduce the containment pressure and temperature is © Copyright 2022 by NuScale Power, LLC 34

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 demonstrated by the peak pressure decreasing to less than 50 percent of the peak value consistent with Design Specific Review Standard (DSRS) Section 6.2.1.1.A (Reference 111). Potential single failures have been considered in the methodology, and the results of the analyses show that the safety functions can be performed including the limiting single failure. General Design Criterion 50 - The analyses performed per the containment response analysis methodology demonstrate that sufficient margin to the CNV design pressure and temperature is maintained. The methodology explicitly models all energy sources including energy in the steam generators (SGs). However, the energy from the post-LOCA oxidation of the cladding that is typical of light water reactors is not applicable to the NuScale design and is not included. Calculated cladding temperatures for design basis LOCAs are below the level where cladding oxidation occurs on a time scale of a LOCA event for the NPM. Therefore, this requirement is satisfied by the design that precludes fuel temperature reaching critical heat flux and any significant fuel cladding heatup. For the NPM loss-of-coolant accident evaluation model core coverage and a minimum critical heat flux ratio are significantly greater than the safety limit, which precludes the occurrence of cladding oxidation. The NRELAP5 code and model have been assessed by comparing predictions to experimental data to demonstrate the capability to reliably simulate the scenarios of interest. Conservative values for initial conditions and boundary conditions ensure an overall conservative analysis result. 2.2.5.2 Regulatory Guide 1.203 Regulatory Guide 1.203, Transient and Accident Analysis Methods (Reference 2), describes a process that the NRC staff considers acceptable for industry use to develop and assess evaluation models used to analyze transient and accident behavior that is within the design basis of a nuclear power plant. An evaluation model is the calculational framework for evaluating the behavior of the reactor system during a postulated transient or design basis accident. The containment response analysis methodology is an extension of the NuScale LOCA, valve opening event and non-LOCA methodologies developed following the guidance of Regulatory Guide 1.203. This report identifies and justifies the differences in the containment response methodology inputs when compared to LOCA and reactor valve opening events. 2.2.5.3 Design Specific Review Standard for NuScale Small Modular Reactor Design The NRC has issued "Design-Specific Review Standard for NuScale SMR Design" to guide the NRC staff review of the NuScale FSAR. This document replaces NUREG-0800, "Standard Review Plan." The NRC staff has specified the DSRS as an acceptable method for evaluating whether an application complies with NRC regulations for NuScale SMR applications, provided that the application does not deviate significantly from the design and siting assumptions made by the © Copyright 2022 by NuScale Power, LLC 35

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 NRC staff while preparing the DSRS. The DSRS is used by NuScale as a guide to ensure that the containment response analysis methodology addresses all of the elements that NRC has included. Sections 2.2.5.3.1 through 2.2.5.3.4 describe how the containment response analysis methodology is consistent with the applicable DSRS guidelines, justify differences, or indicate non-applicability. 2.2.5.3.1 Design Specific Review Standard 6.2.1 Containment Functional Design The DSRS Section 6.2.1, "Containment Functional Design" (Reference 110), includes a high-level summary of an acceptable approach and content for a containment response analysis methodology, and references the lower-tier subsections with additional detail about the approach and contents. The comparison of the containment response analysis methodology to applicable content in DSRS Section 6.2.1 is provided in Table 2-3: Table 2-3 Compliance with Design Specific Review Standard Section 6.2.1 DSRS Section 6.2.1, p. 1 Containment Response Analysis Methodology The containment structure must be capable of The containment response analysis methodology withstanding, without loss of function, the pressure addresses LOCAs resulting from postulated limiting and temperature conditions resulting from postulated breaks, valve-opening events, main steam line break loss- of-coolant (LOCA), steam line, or feedwater (MSLB) accidents, and feedwater line break (FWLB) line break accidents. accidents. A conservative approach to modeling the full spectrum of break and valve sizes and locations is included. The limiting results are less than the CNV design pressure and temperature. The containment design basis includes the effects of The containment response analysis methodology stored energy in the reactor coolant system, decay includes all primary system and secondary energy energy, and energy from other sources such as the sources that contribute to the M&E release. The secondary system, and metal-water reactions energy from the post-LOCA oxidation of the cladding including the recombination of hydrogen and that is typical of light water reactors is not applicable oxygen. to the NuScale design and is not included as discussed by Section 2.2.5.1. The subsequent thermodynamic effects in the The containment response analysis methodology containment resulting from the release of the coolant uses the NRELAP5 system thermal-hydraulic mass and energy are determined from a solution of analysis code. NRELAP5 solves the time-the incremental space and time-dependent energy, dependent conservation equations for mass, mass, and momentum conservation equations. momentum, and energy. DSRS Section 6.2.1, p. 2 Containment Response Analysis Methodology GDC 50, among other things, requires that The containment response analysis methodology consideration be given to the potential models engineered safety features including NPM consequences of degraded engineered safety containment heat removal and the ECCS with features, such as the containment heat removal conservative assumptions. Postulated single failures system and the ECCS, the limitations in defining are considered. Initial and boundary conditions are accident phenomena, and the conservatism of selected to maximize containment pressure and calculation models and input parameters in temperature response. Margin is maintained assessing containment design margins. between the analysis results and the CNV design pressure and temperature limits. © Copyright 2022 by NuScale Power, LLC 36

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-3 Compliance with Design Specific Review Standard Section 6.2.1 (Continued) The regulation in 10 CFR 50 Appendix K.I.A The containment response analysis methodology provides the sources of energy that are required and includes all of the sources of energy required in acceptable to be included in determining the mass Appendix K.I.A with the following exceptions to Items and energy release from loss-of-coolant accidents 4 and 5: and secondary systems pipe ruptures.

4) Fission Product Decay: The American Nuclear Society (ANS)-5.1- 1979 decay heat standard with a two- sigma uncertainty is used rather than 120 percent of the 1971 American Nuclear Society (ANS) standard.

Consistent with DSRS 6.2.1.3, Section II, Acceptance Criterion 1.C.v, the ANS- 5.1-1979 standard is equal to the decay heat model given in Standard Review Plan (SRP) Section 9.2.5.

5) Metal-Water Reaction: The energy from the post-LOCA oxidation of the cladding that is typical of light water reactors is not applicable to the NuScale design and is not included.

DSRS Section 6.2.1, p. 4 Containment Response Analysis Methodology The temperature and pressure profiles provided in Methodology for simulation of the M&E release and the applicant's technical submittal for the spectrum CNV response that is used for establishing the of LOCA and main steam line break accidents are equipment qualification pressure and temperature acceptable for use in equipment qualification (i.e., envelopes, and to demonstrate the long-term cooling there is reasonable assurance that the actual capabilities of the NPM, are outside of the scope of temperatures and pressures for the postulated this report. accidents will not those profiles anywhere within the specified environmental zones, except the break zone). 2.2.5.3.2 Design Specific Review Standard 6.2.1.1.A Containment The DSRS Section 6.2.1.1.A, "Containment" (Reference 111), includes content related to containment design, including some elements that are associated with the capability to withstand M&E releases. The comparison of the containment response analysis methodology to applicable content in DSRS Section 6.2.1.1.A is provided in Table 2-4: © Copyright 2022 by NuScale Power, LLC 37

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-4 Compliance with Design Specific Review Standard Section 6.2.1.1.A DSRS Section 6.2.1.1.A, p. 1 Containment Response Analysis Methodology The temperature and pressure conditions in the The containment response analysis methodology containment due to a spectrum (including break size includes the spectrum of primary release events and location) of postulated loss-of-coolant accidents resulting from postulated limiting breaks (LOCAs) (LOCAs) (i.e., reactor coolant system pipe breaks) and valve openings, MSLB accidents, and FWLB and secondary system steam and feedwater line accidents. The limiting results are less than the CNV breaks design pressure and temperature. The effectiveness of static (passive) and active heat The containment response analysis methodology removal mechanisms. includes conservative modeling of passive heat removal systems (there are no active heat removal systems in the NuScale design). Specifically, conservatisms are employed in conservative assumed initial and boundary conditions, including the reactor pool to ensure a bounding peak CNV peak pressure and temperature following events involving release of mass and energy into the CNV. The performance of these systems is shown to be effective in limiting the CNV pressure and temperature response to within acceptable design limits. DSRS Section 6.2.1.1.A, p. 4 Containment Response Analysis Methodology To satisfy the requirements of GDC 16 and 50 For applications referencing the containment regarding sufficient design margin, for plants in the response analysis methodology the limiting event design stage (i.e., at the construction permit (CP) or scenarios must be less than the CNV design design certification (DC) stage) of review, the pressure and temperature. Sample results for the containment design pressure should provide at least NPM-20 are provided in Section 9.7. a 10% margin above the accepted peak calculated containment pressure following a LOCA, or a steam or feedwater line break. Design margins of less than 10% may be sufficient, provided appropriate justification is provided. For plants at the operating license (OL) or COL stage of review, the peak calculated containment pressure following a LOCA, or a steam or feedwater line break, should be less than the containment design pressure. To satisfy the requirements of GDC 38 to rapidly The containment response analysis methodology is reduce the containment pressure, the containment applicable to the initial CNV response and pressure should be reduced to less than 50% of the demonstrates that the peak pressure and peak calculated pressure for the design basis LOCA temperature are within the CNV design limits. The within 24 hours after the postulated accident. If methodology also demonstrates that the CNV analysis shows that the calculated containment pressure decreases to less than 50 percent of the pressure may not be reduced to 50% of the peak peak pressure within 24 hours to satisfy the calculated pressure within 24 hours, the organization requirements of Principal Design Criterion 38 for responsible for DSRS Section 15.0.3 should be rapid reduction of containment pressure. notified. © Copyright 2022 by NuScale Power, LLC 38

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-4 Compliance with Design Specific Review Standard Section 6.2.1.1.A (Continued) DSRS Section 6.2.1.1.A, p. 5 Containment Response Analysis Methodology To satisfy the requirements of GDC 38 and 50 with The containment response analysis methodology respect to the containment heat removal capability models engineered safety features involving the and design margin, the LOCA analysis should be containment heat removal function, DHRS and the based on the assumption of loss of offsite power and ECCS. Conservative assumptions regarding safety the most severe single failure in the emergency feature performance, in conjunction with power system (e.g., a diesel generator failure), the conservative initial and boundary conditions, ensure containment heat removal systems (e.g., a fan, that the CNV peak pressure and temperature pump, or valve failure), or the core cooling systems analysis results following a primary system release (e.g., a pump or valve failure). The selection made are bounding. A limiting single failure is considered. should result in the highest calculated containment Sensitivity cases considering the availability of pressure. power are performed to ensure that assumptions associated with availability of these systems ensure

4. To satisfy the requirements of GDC 38 and 50 limiting peak pressure and temperature results.

with respect to the containment heat removal There are no emergency diesel generators capability and design margin, the containment associated with the NPM design. Margin is response analysis for postulated secondary system maintained between the analysis results and the pipe ruptures should be based on the most severe CNV design pressure and temperature limits for the single failure of the secondary system isolation limiting cases. provisions (e.g., main steam isolation valve failure or feedwater line isolation valve failure). The analysis The containment response analysis methodology should also be based on a spectrum of pipe break models engineered safety features including NPM sizes and reactor power levels. The accident containment heat removal, DHRS and the ECCS conditions selected should result in the highest with conservative assumptions that maximize calculated containment pressure or temperature containment pressure and temperature following a depending on the purpose of the analysis. secondary system pipe rupture. For postulated Acceptable methods for the calculation of the secondary system pipe ruptures, a limiting single containment environmental response to main steam failure is considered, including main steam isolation line break accidents are found in NUREG-0588, valve or feedwater isolation valve (FWIV) failure. For "Interim Staff Position on Environmental the NuScale design, full power and the maximum Qualification of Safety- Related Electrical break size at each break location are the limiting Equipment." conditions. Initial and boundary conditions are selected to maximize containment pressure and temperature response. Margin is maintained between the analysis results and the CNV design pressure and temperature limits. The longer-term response for equipment qualification is not in the scope of this report. 2.2.5.3.3 Design Specific Review Standard 6.2.1.3 Mass and Energy Release Analysis for Postulated Loss-of-Coolant Accidents The DSRS Section 6.2.1.3, "Mass and Energy Release Analysis for Postulated Loss-of-Coolant Accidents (LOCAs)" (Reference 111), includes the details of an acceptable approach and content for an M&E methodology for LOCAs. As noted, a comparison of NPM design reveals that some of the DSRS content is based on pressurized water reactor (PWR) large-break LOCA phenomena that are not applicable to the NuScale design. The © Copyright 2022 by NuScale Power, LLC 39

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 comparison of the M&E methodology to applicable content in DSRS Section 6.2.1.3 is provided in Table 2-5: Table 2-5 Compliance with Design Specific Review Standard Section 6.2.1.3 DSRS Section 6.2.1.3, p. 3 Containment Response Analysis Methodology A. Sources of Energy. The containment response analysis methodology The sources of stored and generated energy that includes reactor power; decay heat; stored energy in should be considered in analyses of LOCAs include: the core; stored energy in the reactor coolant system reactor power; decay heat; stored energy in the (RCS) metal, including the reactor vessel and core; stored energy in the reactor coolant system reactor vessel internals; and stored energy in the (RCS) metal, including the reactor vessel and secondary system, including the SG tubing and reactor vessel internals; metal-water reaction secondary water. Metal-water reaction energy is not energy; and stored energy in the secondary system, included in the containment response analysis including the steam generator tubing and secondary methodology. water. The containment response analysis methodology Calculations of the energy available for release from models available energy sources in accordance with the above sources should be done in general the requirements of 10 CFR Part 50, Appendix K, accordance with the requirements of paragraph I.A. paragraph I.A, with the exception of 1) metal-water in Appendix K to 10 CFR Part 50, "Sources of Heat reaction energy is not included, and 2) the during the LOCA." However, additional conservatism ANS-5.1-1979 decay heat standard with a should be included to maximize the energy release two-sigma uncertainty is used rather than a factor of to the containment during the blowdown and 1.2 with the 1971 ANS standard. Consistent with subsequent phases of a LOCA. An example of this DSRS 6.2.1.3, Section II, Acceptance Criterion would be accomplished by maximizing the sensible 1.C.v, the ANS-5.1-1979 standard is equal to the heat stored in the RCS and steam generator metal decay heat model given in SRP Section 9.2.5. and increasing the RCS and steam generator secondary mass to account for uncertainties and The containment response analysis methodology thermal expansion. model of initial stored energy in the fuel is consistent with Paragraph I.A.1 of Appendix K to The requirements of paragraph I.B in Appendix K to 10 CFR Part 50. Fuel rods are initialized at the 10 CFR Part 50, "Swelling and Rupture of the maximum initial stored energy condition as Cladding and Fuel Rod Thermal Parameters," determined by the fuel performance analysis. The concerning the prediction of fuel clad swelling and fuel heat capacity values are conservatively rupture should not be considered. This will maximize increased to 115 percent of their nominal values to the energy available for release from the core. maximize fuel stored energy. The fuel thermal conductivity values are conservatively decreased to 85 percent of their nominal values to maximize fuel stored energy. The containment response analysis methodology includes conservative elements that maximize the energy release including sensible heat stored in primary and secondary metal structures, and increasing the RCS mass to account for uncertainties and thermal expansion. The secondary mass is not a significant contributor and a nominal value is used. © Copyright 2022 by NuScale Power, LLC 40

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-5 Compliance with Design Specific Review Standard Section 6.2.1.3 (Continued) A (Continued) The containment response analysis methodology does not consider the fuel cladding swelling and rupture prediction requirements of paragraph I.B in Appendix K to 10 CFR Part 50. Calculated cladding temperatures for design basis LOCAs are below the threshold for cladding swelling and rupture. DSRS Section 6.2.1.3, p. 4 Containment Response Analysis Methodology B. Break Size and Location The containment response analysis methodology i The staff's review of the applicant's choice of break includes consideration of a spectrum of break types. locations and types is discussed in SRP Break locations are chosen such that M&E releases Section 3.6.2. to containment are maximized. ii Of several breaks postulated, the break selected (( as the reference case should yield the highest mass and energy release rates, consistent with the criteria for establishing the break location and area. iiiContainment design basis calculations should be performed for a spectrum of possible pipe break sizes and locations to assure that the worst case has been identified.

                                                                       }}2(a),(c)

C. Calculations The containment response analysis methodology In general, calculations of the mass and energy focuses on determining the maximum post-accident release rates for a LOCA should be performed in a containment pressure and temperature. The manner that conservatively establishes the methodology employs conservative elements to containment internal design pressure (i.e., ensure an overall conservative result. maximizes the post-accident containment pressure response). The criteria given below for each phase of the accident indicate the conservatism that should exist.

i. Containment Analysis The M&E release determined by the containment The analytical approach used to compute the mass response analysis methodology is based on the and energy release profile will be accepted if both NRELAP5 computer code, and the modeling the computer program and volume noding of the approach is similar to the NuScale LOCA evaluation reactor, piping and containment systems are similar model, presented in this report, that complies with to those of an approved ECCS analysis. The the applicable portions of 10 CFR 50 Appendix K.

computer programs that are currently acceptable Specific changes to the LOCA evaluation model include CRAFT-2, and RELAP5, when a flow required to convert it to a conservative methodology multiplier of 1.0 is used with the applicable choked to model primary system mass release events are flow correlation. An alternate approach, which is also described in Chapter 3. The Moody critical flow acceptable, is to assume a constant blowdown model with a discharge coefficient of 1.0 is used for profile using the initial conditions with an acceptable saturated two-phase critical flow. choked flow correlation. © Copyright 2022 by NuScale Power, LLC 41

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-5 Compliance with Design Specific Review Standard Section 6.2.1.3 (Continued) ii. Initial Blowdown Phase Containment Design Basis The containment response analysis methodology The initial mass of water in the reactor coolant assumes an initial power level of 1.02 times the system should be based on the RCS volume rated power level. Initial RCS volume and mass are calculated for the temperature and pressure consistent with that power level. The initial RCS conditions assuming that the reactor has been volume conservatively includes an allowance for operating continuously at a power level at least 1.02 RCS thermal expansion. times the licensed power level (to allow for instrumentation error). An assumed power level The containment response analysis methodology lower than the level specified (but not less than the uses the conservative Moody critical flow model for licensed power level) may be used provided the two-phase saturated fluid conditions consistent with proposed alternative value has been demonstrated Appendix K. For subcooled fluid conditions the to account for uncertainties due to power level (( instrumentation error. }}2(a),(c) Sections 8.2.2 and 8.2.3 demonstrate the adequacy of the LOCA evaluation model Mass release rates should be calculated using a two-phase and single-phase choked and un-choked model that has been demonstrated to be flow models for predictions of M&E release based on conservative by comparison to experimental data. assessments of comparisons of NRELAP5 mass flow predictions to experimental data. Calculations of heat transfer from surfaces exposed to the primary coolant should be based on nucleate The containment response analysis methodology boiling heat transfer. For surfaces exposed to steam, uses the heat transfer correlation package in the heat transfer calculations should be based on forced NRELAP5 computer code. The LOCA evaluation convection. model report demonstrates these correlations are applicable to the NPM design. The local fluid Calculations of heat transfer from the secondary conditions and the local heat structure surface coolant to the steam generator tubes should be temperatures determine the heat transfer mode. based on natural convection heat transfer for tube Nucleate boiling and forced convection are included surfaces immersed in water and condensing heat in the code and are selected if the local conditions transfer for the tube surfaces exposed to steam. are appropriate. Calculations of heat transfer to the containment wall The containment response analysis methodology from released reactor steam should be such that the uses the heat transfer correlation package in the heat removal from containment is conservatively NRELAP5 computer code. The LOCA evaluation underestimated so that the containment pressure is model report demonstrates these correlations are maximized. In regions where steam jetting occurs, applicable to the NPM design. The local fluid heat transfer correlations that are based on jetting of conditions and the local heat structure surface coolant (e.g. based on forced convection) may be temperatures determine the heat transfer mode. used as appropriate. Correlations should be Forced convection, natural convection, appropriately conservative in regions away from condensation, conduction, and nucleate boiling are jetting phenomena (e.g. based on natural included in the code and are selected if the local convection, as appropriate). All heat transfer conditions are appropriate. correlations used should be justified. Initial and boundary conditions are selected to Calculations of heat transferred from condensed maximize containment pressure and temperature reactor water in the containment sump into the response. Steam jetting effects are not modeled. containment wall and from the reactor vessel wall into the pooled sump water should be based on appropriate heat transfer regimes for the conditions present in containment. Heat transfer through the containment vessel wall into the Reactor Building pool should be demonstrated to conservatively underestimate heat transfer to the pool © Copyright 2022 by NuScale Power, LLC 42

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-5 Compliance with Design Specific Review Standard Section 6.2.1.3 (Continued) DSRS Section 6.2.1.3, p. 5 Containment Response Analysis Methodology iii. Postblowdown Recirculation Phase (Cold Leg The containment response analysis methodology RRV Penetration Breaks Only) After initial blowdown uses the NRELAP5 code that has been determined through a failed RRV, which includes the period from to be capable of modeling all of the phases of the the accident initiation (when the reactor is in a primary system release events for the NPM design. steady-state full power operation condition) to the NRELAP5 predicts the evolution of the primary time that the RCS equalizes to the containment system release event scenario, which includes the pressure, the water remaining in the reactor vessel time of pressure equalization and the time at which should be assumed to be saturated. Justification flow of condensed water through the RRVs into the should be provided for the duration of the reactor vessel occurs. The containment response recirculation period, which is the time from the end of analysis methodology models applicable the blowdown to the time when flow from the phenomena that contribute to maximizing the M&E condensed water in the containment vessel sump release and the resulting containment pressure and comes back through the RRVs into the reactor temperature. vessel. The "refill rate" is only applicable to large PWRs. As Calculations of the refill rate should be based on the discussed by the LOCA evaluation model report, the ECCS operating condition following the blowdown NPM design precludes core uncovery. phase, where energy is released to the RCS primary system by the RCS metal, core decay heat, and the The containment response analysis methodology steam generators. The calculated ECCS conditions models applicable phenomena that contribute to should conservatively maximize the containment maximizing the M&E release and the resulting pressure. containment pressure and temperature. Calculations of liquid entrainment, (i.e., the carryout The concept of carryout rate fraction that is rate fraction), which is the mass ratio of liquid exiting applicable to large PWRs is not applicable to the the core to the liquid entering the core, should be NuScale design. As discussed by the LOCA based on the NuScale full length emergency cooling evaluation model report, the NPM design precludes heat transfer experiments or conservatively core uncovery, so there is no reflooding phase. scaled-up test results from subscale test. The containment response analysis methodology The assumption of steam quenching should be models applicable phenomena that contribute to justified by comparison with applicable experimental maximizing the M&E release and the resulting data. Liquid entrainment calculations should containment pressure and temperature. consider the effect on the carryout rate fraction of the increased core inlet water temperature caused by The concept of steam quenching (that occurs from steam quenching assumed to occur from mixing with mixing with ECCS water) that is applicable to large the ECCS water. PWRs is not applicable to the NuScale design because ECCS water is not injected into the core. Steam leaving the steam generators should be assumed to be superheated to the temperature of The containment response analysis methodology the secondary coolant. models applicable phenomena that contribute to maximizing the M&E release and the resulting containment pressure and temperature. © Copyright 2022 by NuScale Power, LLC 43

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-5 Compliance with Design Specific Review Standard Section 6.2.1.3 (Continued) iii (Continued) The superheating effect described is a pressurized water reactor LOCA phenomenon that has minimal applicability to the NuScale design. For the NPM design, flow of primary steam over the SG tubes results in heat transfer based on the NRELAP5 heat transfer correlation package. This allows for superheating of the steam as determined by the local conditions. DSRS Section 6.2.1.3, p. 6 Containment Response Analysis Methodology iv. Post-Recirculation Phase The stored energy is distributed as predicted by the All remaining stored energy in the primary and NRELAP5 modeling of heat transfer to and from the secondary systems should be removed during the primary and secondary systems. The duration of the post-recirculation phase. analysis is consistent with the LOCA evaluation Steam quenching on the containment vessel walls, model and the applicable figures-of-merit. The due to pressure equalization between the reactor containment response analysis methodology vessel and the containment vessel, should be considers steam condensation on the CNV walls. justified by comparison with applicable experimental The NRELAP5 code and model have been justified data. by comparison to applicable experimental data. The results of post-recirculation analytical models should be compared to applicable experimental data. 2.2.5.3.4 Design Specific Review Standard 6.2.1.4 Mass and Energy Release Analysis for Postulated Secondary Pipe Ruptures The DSRS Section 6.2.1.4, "Mass and Energy Release Analysis for Postulated Secondary Pipe Ruptures" (Reference 112), includes the details of an acceptable approach and content for a M&E methodology for MSLBs and FWLBs. The comparison of the M&E methodology to applicable content in DSRS Section 6.2.1.4 is provided in Table 2-6: © Copyright 2022 by NuScale Power, LLC 44

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-6 Compliance with Design Specific Review Standard Section 6.2.1.4 DSRS Section 6.2.1.4, p. 4 Containment Response Analysis Methodology

1. Sources of Energy. The containment response analysis methodology The sources of energy that should be considered in includes all of the sources of energy stored in the the analyses of steam and feedwater line break fluid and structures that contribute to the secondary accidents include the stored energy in the affected line break scenarios. This includes energy stored in helical coil SG's metal, including the vessel tubing, fluid contained in piping systems connected to the feedwater line, and steam line; stored energy in the break flowpath into the CNV.

water contained within the affected helical coil SG; stored energy in the feedwater transferred to the The containment response analysis methodology affected helical coil SG before closure of the considers a spectrum of pipe break sizes and isolation valves in the feedwater line; stored energy various plant conditions. However, the limiting initial in the steam from the unaffected helical coil SG conditions are at 102 percent rated power as the before the closure of the isolation valves in the effect of SG liquid mass inventory and feedwater helical coil SG crossover lines; and energy flows is greatest at full power. (( transferred from the primary coolant to the water in the affected helical coil SG during blowdown to include energy transferred to the draining DHRS heat exchanger water. The steam line break accident should be analyzed for a spectrum of pipe break sizes and various plant conditions from hot standby to 102 percent of full power. The applicant need only analyze the 102- }}2(a),(c) percent power condition if it can demonstrate that the feedwater flows and fluid inventory are greatest at full power.

2. Mass and Energy Release Rate The containment response analysis methodology In general, calculations of the mass and energy maximizes the CNV peak pressure and temperature.

release rates during a steam or feedwater line break The Moody critical flow model with a discharge accident should be performed in a conservative coefficient of 1.0 is used for saturated two-phase manner from a containment response standpoint fluid conditions. For subcooled and superheated (i.e., the postaccident containment pressure and fluid conditions the (( temperature are maximized). The following criteria }}2(a),(c) A discharge coefficient of 1.0 indicate the degree of conservatism that is desired: is used. A. Mass release rates should be calculated using the Moody model (Reference 6) for saturated conditions or a model that is demonstrated to be equally conservative. © Copyright 2022 by NuScale Power, LLC 45

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-6 Compliance with Design Specific Review Standard Section 6.2.1.4 (Continued) B. Calculations of heat transfer to the water in the The containment response analysis methodology affected helical coil SG should be based on nucleate uses the heat transfer correlation package in the boiling heat transfer. NRELAP5 computer code. The non-LOCA evaluation model report demonstrates these correlations are applicable to the NPM design. The local fluid conditions and the local heat structure surface temperatures determine the heat transfer mode. Nucleate boiling heat transfer is included in the code and is selected if the local conditions are appropriate. For the helical coil SG, other heat transfer modes exist as the coolant enters as subcooled liquid and exits as superheated steam. Initial and boundary conditions are selected to maximize containment pressure and temperature response. C. Calculations of mass release should consider the The containment response analysis methodology water in the affected helical coil SG and feedwater includes the water inventory stored in piping systems line, feedwater transferred to the affected helical coil connected to the break flowpath into the CNV. The SG before the closure of the isolation valves in the closure of isolation valves, with consideration of a feedwater lines and upon flooding with the DHRS single failure, determines which sources of water heat exchanger inventory in the affected loop, and contribute to the M&E release to ensure limiting CNV steam in the helical coil SG. peak pressure and temperature results. D. If liquid entrainment is assumed in the steam line The containment response analysis methodology breaks, experimental data should support the uses the two-phase flow and heat transfer models in predictions of the liquid entrainment model. A the NRELAP5 code. The depressurization of the SG spectrum of steam line breaks should be analyzed, secondary causes flashing in addition to the beginning with the double-ended break (DEB) and increase in primary-to-secondary heat transfer. The decreasing in area until no entrainment is calculated initial liquid inventory in the SG secondary boils and to occur. This will allow selection of the maximum flash, and additional inventory results from continued release case. feedwater flow and from liquid in connecting pipes. The net effect may include some liquid entrainment If no liquid entrainment is assumed, a spectrum of in the break flow that is time dependent. An the steam line breaks should be analyzed beginning interfacial drag multiplier is available as a junction with the DEB and decreasing in area until it has component option in NRELAP5 to minimize liquid been demonstrated that the maximum release rate entrainment. has been considered. © Copyright 2022 by NuScale Power, LLC 46

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 2-6 Compliance with Design Specific Review Standard Section 6.2.1.4 (Continued) E. Feedwater flow to the affected helical coil SG The containment response analysis methodology should be calculated considering the diversion of includes the water inventory stored in piping systems flow from the other helical coil SG between the two connected to the break flowpath into the CNV. The feedwater pipes to the common header with inlets to increase in feedwater flow due to the the helical coil SG on opposite sides of the reactor depressurization of the helical coil SG is considered. vessel, feedwater flashing, and increased feedwater The closure of isolation valves with consideration of pump flow caused by the reduction in helical coil SG a single failure determines which sources of water pressure. An acceptable method for computing contribute to the M&E release. The net feedwater feedwater flow is to assume all feedwater travels to addition is calculated using conservative modeling the helical coil SG at the pump run-out rate before assumptions. isolation. After isolation, the unisolated feedwater mass should be added to the available inventory in the helical coil SG. DSRS Section 6.2.1.4, p. 5 The containment response analysis methodology considers single failures that affect the isolation of iii. Single-Failure Analyses the main steam lines and feedwater lines. Steam and feedwater line break analyses should Non-safety valves are credited for isolation as a assume a single active failure in the steam or backup. feedwater line isolation provisions to maximize the containment peak pressure and temperature. For the assumed failure of a safety-related steam or feedwater line isolation valve, operation of nonsafety-related equipment may be relied upon as a backup to the safety-related equipment. © Copyright 2022 by NuScale Power, LLC 47

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 3.0 NuScale Power Module Description and Operations 3.1 General Plant Design The NuScale Power Plant consists of one or more NPMs, each of which is a small, passive PWR. The NPM consists of the nuclear steam supply system (NSSS), which includes the nuclear core, the helical coil SGs and the pressurizer, within a single pressure vessel and the compact steel CNV that houses the NSSS. Unique features of the NuScale plant design include the following: reduced core size natural circulation reactor coolant flow (i.e., no reactor coolant pumps) integrated SG and a pressurizer inside the RPV. As a result, there is no piping connecting the SG or pressurizer with the reactor simplified passive safety-related systems that do not rely on ECCS pumps, accumulators, and water storage tanks (e.g., core makeup tank, in-containment refueling water storage tank) high-pressure steel containment containment partially immersed in a water-filled pool providing an effective passive heat sink for emergency cooling The NPM is designed to operate efficiently at full-power conditions using natural circulation as the means of providing core coolant flow, eliminating the need for reactor coolant pumps. As shown in Figure 3-1, the reactor core is located inside a shroud connected to the hot leg riser. The reactor core heats reactor coolant, decreasing its density, causing the coolant to flow upward through the riser. When the heated reactor coolant exits the riser, it passes across the tubes of the helical coil SG, which acts as a heat sink. As the reactor coolant passes over the SG tubes, it cools, increases in density, and naturally circulates down the downcomer to the reactor core where the cycle begins again. The NPMs are partially immersed in a reactor pool and protected by passive safety-related systems. Each NPM has a dedicated ECCS, chemical and volume control system (CVCS), and decay heat removal system (DHRS). NuScale has achieved a substantial improvement in safety over existing plants through simplicity of design, reliance on passive safety-related systems, and small fuel inventory. The definition of a LOCA in 10 CFR 50.46(c)(1) addresses the geometry of a typical PWR, in which reactor coolant piping connects the RPV to primary system components external to the RPV. In the NuScale Power Plant design, all primary components are integral to the RPV, eliminating external coolant loops and pressurizer piping, which significantly reduces the number of possible LOCA scenarios. © Copyright 2022 by NuScale Power, LLC 48

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 3-1 A Single NuScale Power Module during Normal Operation The potential break sizes included in the LOCA EM include the possible spectrum of breaks that can result in a break flow that exceeds the capability of the CVCS. The NPM piping break locations are few (when compared to conventional PWR designs), and consist of the RCS injection and discharge lines, pressurizer spray supply line, and pressurizer high point vent line. These connections can be grouped into penetrations that are high on the RPV (pressurizer steam space) and low on the RPV (penetrate into an area which is normally in a liquid condition). All of the penetrations in the NPM design are at an elevation above the top of the core. The NPM was designed with the intent of reducing the impact of a LOCA event. All LOCAs result in the actuation of both the ECCS and the DHRS. As shown in Figure 3-2, the ECCS consists of independent RVVs and independent RRVs. The ECCS is initiated by opening the RVVs located at the top of the RPV and the RRVs located in the downcomer region (above the core elevation). Opening the valves allows the RPV and the CNV pressure to equalize and creates a natural circulation path to remove decay heat from the core. Water that is vaporized in the core leaves as steam through the RVVs, is condensed and collected in the CNV, and is then returned to the downcomer region inside the RPV through the RRVs by natural circulation. © Copyright 2022 by NuScale Power, LLC 49

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The CNV is sized such that the displacement of liquid from the RPV into containment results in the liquid level being above the RRVs (which are located above the core) establishing a natural circulation loop. By the time the natural circulation pattern forms, the outside of the RPV cools enough that boiling on the outside of the RPV is relatively limited and the liquid level in the containment has minimum swelling. The natural circulation loop removes decay heat from the core and RPV, and deposits it in containment. Heat deposited in containment is transferred by conduction and convection to the water in the reactor pool. Following actuation of the ECCS, heat removal through the CNV rapidly reduces reactor and containment pressures and temperatures, and maintains them at acceptably low levels for extended periods of time. Because the CNV is evacuated to a low absolute pressure during normal operation (i.e., vacuum), only a small amount of non-condensable gas is present inside the CNV at the beginning of the event. The DHRS provides additional capacity to remove decay heat during the initial blowdown period of a LOCA. The DHRS provides secondary-side reactor cooling when normal feedwater is not available. The system, as shown in Figure 3-1, is a closed-loop, two-phase natural circulation cooling system. Two trains of decay heat removal equipment are provided, one attached to each SG loop. Each train is independently capable of removing 100 percent of the decay heat load and can cool the reactor primary-side inventory. Each train has a passive condenser submerged in the reactor pool. The condensers are maintained with sufficient water inventory for stable operation. Analyses using the EM described in this report show that LOCAs do not challenge the safety of an NPM (results of representative LOCA break spectrum calculations are in Section 9.0). © Copyright 2022 by NuScale Power, LLC 50

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 3-2 Schematic of NuScale Power Module Decay Heat Removal System and Emergency Core Cooling System during Operation 3.2 Plant Operation This LOCA EM initiates the analyses of an NPM with 102 percent full-rated power operation (as required by 10 CFR 50 Appendix K). This assumption represents the uncertainty in the initial power. Pressurizer heaters and a spray system are used to maintain nominal operating pressure similar to conventional PWRs. The reactor coolant is driven by natural circulation. At nominal full-power conditions, the flow rate is dependent on the fluid density differences through the loop, the losses incurred along the loop, and the elevation difference between the core and the SG. During nominal full-power conditions, the control rods are retracted up to or above their insertion limits. Borated water is used as the primary coolant and the CVCS regulates the boron concentration to maintain criticality. The CVCS provides reactor inventory make-up through the RCS injection line in the riser and inventory let-down through a separate RCS discharge line in the downcomer region. © Copyright 2022 by NuScale Power, LLC 51

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The secondary side is operated such that the SGs remove the heat generated by the reactor core. The DHRS heat exchangers are isolated from the steam line and do not remove heat during normal operation. The containment is evacuated during normal operation to provide an insulated barrier between the reactor and containment; no physical RPV insulation is present inside containment. 3.3 Safety-Related System Operation The NuScale Module Protection System (MPS) is composed primarily of the reactor trip system and the engineered safety features actuation system. The MPS protection functions are limited to automated safety responses to off-normal conditions. The MPS functional response to an initiating event is a reactor trip; isolation (as necessary) of main feedwater, main steam, CVCS, and containment; followed by an integrated safety actuation of one or more of the passive safety-related systems (DHRS and ECCS). Containment isolation is achieved by closing of the following containment isolation valves: RCS isolation valves

             -    RCS injection line
             -    RCS discharge line
             -    pressurizer spray supply line
             -    pressurizer high point vent line, reactor component cooling water system isolation valves main steam system isolation valves feedwater system isolation valves containment flood and drain system isolation valves containment evacuation system isolation valves Dual safety-related isolation valves are installed on piping for the CVCS, containment evacuation system, containment flood and drain system, and reactor component cooling water system. There is one safety-related containment isolation valve in the main steam and feedwater piping penetrating containment with a redundant nonsafety-related isolation valve for each safety-related valve.

The reactor trip system consists of four independent separation groups with independent measurement channels to monitor plant parameters that can generate a reactor trip. Each measurement channel trips when the parameter exceeds a predetermined setpoint. The engineered safety features actuation system also consists of four independent separation groups with independent measurement channels that monitor plant parameters that activate the operation of the engineered safety features. © Copyright 2022 by NuScale Power, LLC 52

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ECCS is actuated by MPS on operating parameters that are monitored to provide indications of a LOCA. Some actuation signals have interlocks that are designed to prevent ECCS actuations for expected operational conditions or non-LOCA transients. 3.3.1 Emergency Core Cooling System The ECCS is a two-phase natural circulation system that maintains a liquid water supply to the core during its operation in a LOCA scenario. This results in a collapsed liquid level in the RPV that is above the top of the core. The ECCS consists of two independent divisions of RVVs and RRVs. ECCS is initiated by actuating the RVVs on the top of the RPV in the pressurizer region and the RRVs on the side of the RPV in the downcomer region. The RRVs are designed to provide a low-resistance flow path for coolant to flow from the CNV into the RPV. The RVVs are designed to equalize pressure between the two vessels allowing steam from the reactor to vent to the containment and to provide hydrostatic equalization that allows coolant flow through the RRVs back into the reactor. The ECCS actuation creates a steam flow path from the pressurizer to the containment and an RPV downcomer flow path to and from containment. The RPV depressurizes due to liquid and steam exiting the ECCS valves. Steam entering containment is condensed on the containment wall, which in turn is cooled by the reactor pool. Initially, the containment pressure increases to a peak, and then decreases as flow from the RPV decreases and heat is transferred from the CNV to the reactor pool. The RPV water inventory decreases while the containment level increases due to inventory transferred from the RPV. As the pressure between the two vessels reach a near-equilibrium condition, the collapsed liquid level in the containment rises to a level higher than the RRV elevation, creating enough static head to overcome the pressure difference between the RPV and CNV. At this point, the condensed liquid in containment enters the RPV through the RRVs while steam exits the RPV through the RVVs. This stable process continues maintaining a collapsed water level above the top of the active fuel. Some ECCS valve designs are equipped with an inadvertent actuation block (IAB). This feature prevents spurious opening of the ECCS valves at full operating pressure and can allow for delayed opening of some ECCS valves. The IAB prevents the valves from opening when the differential pressure between the RPV and CNV is greater than the IAB threshold pressure setpoint. After the IAB has blocked a spurious opening of the ECCS valve, it allows the valve to open only after the differential pressure between the RPV and CNV decreases below the IAB release pressure setpoint. The ECCS valves open on low differential pressure between the RPV and CNV, independent of an ECCS actuation signal. This action is a function of the mechanical design of the valves, where the valve spring causes the valves to open if the pressure difference across the main chamber drops below approximately 15 psid. The spring © Copyright 2022 by NuScale Power, LLC 53

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 force opening of the ECCS valves is expected during normal shutdown operations. The ECCS valves could open by this function during a rapid depressurization event where the RPV and CNV pressures equalize. This function could open ECCS valves earlier in some depressurization scenarios but does not significantly impact accident progression or results. Some NPM designs include a supplemental boron feature as part of the ECCS. The ECCS supplemental boron (ESB) provides additional boron in containment to be recirculated into the RPV after ECCS actuation. This feature does not affect the LOCA EM because the short-term LOCA or valve opening event progression ends shortly after recirculation is established. 3.3.2 Decay Heat Removal System The DHRS is a passive safety-related system that relies on natural circulation to remove heat from the RCS through the SG and reject heat to the reactor pool through the DHRS condenser. The DHRS is composed of two DHRS trains associated with one of the two NPM SGs. Each DHRS train is capable of independently removing 100 percent of decay heat. The DHRS piping connects to the main steam and feedwater lines specific to the associated SG. During normal operation, the DHRS condenser and piping are isolated by valves on the steam side of the SG. The condensate side of the DHRS is open to the feedwater piping supplying the associated SG. Upon actuation of the DHRS, the feedwater and main steam isolation valves, feedwater regulating valves, and secondary main steam isolation valves close and the DHRS actuation valves open, creating a closed loop between the SG and DHRS condenser. Both liquid and vapor are contained in the DHRS on system actuation. Because the DHRS is a closed system, the total water mass remains constant during the system operation. For successful operation, liquid water enters the SG through the feedwater line and is boiled by heat from the RCS. The vapor exits the SG through the steam line and is directed to the DHRS condenser where it condenses back to liquid before return to the SG. Thus, the loop transfers heat from the RCS to the DHRS fluid and then from the DHRS to the reactor pool water. The bottom of the DHRS condenser is located above the bottom of the SG providing the static head to drive natural circulation. The DHRS provides additional capacity to remove decay heat during the initial blowdown period of a LOCA. DHRS cooling is most significant for smaller breaks, prior to ECCS valve opening, where RPV level drops slowly allowing more heat to be transfered through the SGs to the DHRS heat exchangers. © Copyright 2022 by NuScale Power, LLC 54

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 4.0 Phenomena Identification and Ranking 4.1 Phenomena Identification and Ranking Process The purpose of the NuScale LOCA PIRT is to provide an assessment of the relative importance of phenomena and processes that may occur in the NPM during LOCA conditions in relation to specified FOMs. The PIRT assessment is part of the EMDAP process prescribed by RG 1.203. The initial NuScale LOCA PIRT was developed in 2008. This PIRT was updated in 2013 and 2015 to address design changes. The initial PIRT and these PIRT updates have been developed by a panel of recognized industry experts and NuScale subject matter experts, and are built upon the state-of-knowledge at the time of their development. In 2022, the PIRT phenomena and ranking were re-assessed to address changes in analysis methodology and to accommodate design modifications. The panel members of the initial LOCA PIRT were: Dr. Brent Boyack (Los Alamos National Laboratory, retired) Dr. Larry Hochreiter (Pennsylvania State University) Dr. Mujid Kazimi (Massachusetts Institute of Technology) Dr. Jose Reyes (NuScale Power, Inc.) Dr. Kord Smith (Studsvik Scandpower, Inc.) Dr. Graham Wallis, Chair (Darthmouth University, Creare, Inc.) The panel members for the 2013 NuScale LOCA PIRT were: Mr. Steve Congdon (GE Nuclear Energy, retired) Dr. Tom George (Zachry Nuclear Engineering) Mr. Craig Peterson (Computer Simulation and Analysis) Dr. Jose Reyes (NuScale Power, Inc.) Mr. Gregg Swindlehurst (GS Nuclear Consulting, LLC) Dr. Graham Wallis, Chair (Darthmouth University, Creare, Inc.) The panel members for the 2015 NuScale LOCA PIRT were: Mr. Steve Congdon (GE Nuclear Energy, retired) Dr. Tom George (Zachry Nuclear Engineering) Dr. Jose Reyes (NuScale Power, Inc.) Mr. Gregg Swindlehurst (Chair, GS Nuclear Consulting, LLC) Dr. Graham Wallis, Chair (Darthmouth University, Creare, Inc.) © Copyright 2022 by NuScale Power, LLC 55

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The 2015 NuScale LOCA PIRT incorporates lessons learned from testing and insights gained from computer code simulations of many LOCA scenarios. The PIRT panel received an in-depth briefing on the NPM design, LOCA sequence of events, and computer code predictions of the response of the NPM to LOCA scenarios. The panel then followed the PIRT process by first identifying the structures, systems, and components (SSC) of the NPM that were associated with the LOCA scenario. The LOCA scenario was then separated into phases with each phase representing a distinct process-dominated time period. Then FOMs were selected for each phase. Specifically, the FOMs were chosen to be quantifiable measures of the systems potential to meet regulatory safety limits. Phenomena were identified for each SSC for each phase, and the phenomena were ranked considering their level of importance relative to the FOMs. The panel also established a knowledge ranking for each of the phenomena. The PIRT panel was reconvened for the 2013 and 2015 PIRT updates and evaluated the changes in NPM design and their impact on progression of LOCA. The biographical information for each PIRT panel member is included with each PIRT release. The 2015 NuScale LOCA PIRT was used for the development of LOCA EM. The 2022 PIRT applicability update assessed the ranking and knowledge level for all PIRT phenomena, and revised the ranking and/or knowledge level for a subset of the phenomena based on the NPM design and assessment information that was new or revised since the 2015 PIRT. The following section provides a brief description of the LOCA scenarios and the accident phases considered for the PIRT developed. The definitions of the selected FOMs and the importance and knowledge ranking categories are summarized. Finally, the list of phenomena that were ranked as high importance by the PIRT panel in at least one of the phases of the NuScale LOCA scenarios is provided along with the brief description of the rationale for assigned importance and knowledge level rankings. The rankings for the identified phenomena and detailed description of the rationale are available in the 2015 NuScale LOCA PIRT report and the 2022 LOCA PIRT applicability update. 4.2 Loss-of-Coolant Accident Scenarios Loss-of-coolant accidents are postulated breaks in the reactor coolant pressure boundary that result in leakage of reactor coolant at a rate exceeding the capability of the normal reactor coolant makeup system, as defined in 10 CFR 50.46(c)(1). Breaks of various sizes, types, and orientations are postulated to occur in piping connected to the RPV. With the elimination of most primary coolant piping in the NPM design, breaks are limited to RCS injection and discharge lines, pressurizer spray supply line, and pressurizer high point vent line. Two types of LOCA scenarios were addressed in the PIRT development process. The first type of LOCA scenarios were ((

                                                                                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                  }}2(a),(c) Section 9.1 provides further description of the progression of each LOCA scenario.

The PIRT panel divided the NPM LOCA scenarios into two phases for the phenomena identification: LOCA blowdown (Phase 1a) Phase 1a begins with a postulated breach in the RCS pressure boundary that initiates the blowdown of the RCS into the CNV and ends when the MPS actuates ECCS to open the RVVs and RRVs. Phase 1a generally extends longer if the ECCS valves have the IAB feature installed because either ECCS actuation is designed to occur later in the event progression to avoid the IAB block threshold, or the RCS must depressurize below the IAB release threshold before the valves open. ECCS actuation (Phase 1b) Phase 1b begins after MPS actuates ECCS and the ECCS valves open and ends when the recirculation flow through the RRVs is established. The pressures and levels in containment and RPV approach a stable condition (i.e., initiation of long-term cooling), Phase 2. Subsequent to the PIRT identification of LOCA phases, an additional evaluation methodolgy was developed to determine MCHFR in parallel to the existing phases. The parallel phase was termed Phase 0 and runs concurrently with Phase 1a and Phase 1b. ((

                                      }}2(a),(c) Phase 0 analysis is defined separately from LOCA Phase 1a due to the unique methodology considerations for the specific short-term MCHFR analysis compared to the longer-term ECCS performance focused evaluation. Figure 4-1 contains a graphical representation of the relative timings.

Example calculations are performed for both steam space and liquid space LOCAs. The progression of the steam and liquid space LOCA events are similar, with the exception of different timing of the sequence of events and the liquid/steam composition of the break flow. The example calculations in this report are done for NPM designs of ECCS valves and ECCS actuation, and cover all LOCA phenomena. Therefore, the example calculations in this report are representative. © Copyright 2022 by NuScale Power, LLC 57

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 4-1 Timing of LOCA Phases ((

                                                                                                    }}2(a),(c) 4.3      Figures of Merit The safe operation of the NPM was considered in the primary design phase. This produced a reactor system that protects the fuel using simple passive safety features.

The NPM retains sufficient water in the RPV that the core does not uncover. For such a system, the LOCA PCT occurs at time zero (normal operating temperature). There is no heatup due to CHF or uncovering the core after event initiation; PCT remains below the 10 CFR 50.46 acceptance criterion of 2,200 degrees Fahrenheit (1,204 degrees Celsius) throughout the event. Hence, PCT is not an FOM for the NuScale PIRT process. The critical heat flux ratio (CHFR) is an important FOM as it demonstrates there is no significant heatup of the cladding. One of the primary design fundamentals of the NPM is to protect the fuel from a CHF event. Therefore, an assessment of CHF becomes important. Collapsed liquid level above the core is an additional FOM as it demonstrates there is an adequate supply of liquid water available to the core. Heatup of the fuel does not occur under LOCA conditions as long as the core is covered with coolant and CHF conditions do not exist. To ensure ECCS performance, the containment must be intact and remain below pressure and temperature design limits. Consequently, peak containment pressure and temperature are evaluated to ensure compliance with 50.46 criteria. The peak containment pressure and temperature for containment performance are calculated with different assumptions and initial conditions than those used in LOCA analysis. The containment peak pressure and temperature analyses are described in this report and sample calculations are presented in Section 9.7. 4.4 Definitions of Importance and Knowledge Level Rankings Each phenomenon identified in the PIRT was assigned an importance ranking and knowledge level ranking. Table 4-1 and Table 4-2 describe the importance rankings and knowledge level rankings developed by the PIRT panel. © Copyright 2022 by NuScale Power, LLC 58

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 4-1 Importance Rankings Importance Ranking Definition High (H) Significant influence on FOM Medium (M) Moderate influence on FOM Low (L) Small influence on FOM Inactive (I) Phenomenon not present or negligible Table 4-2 Knowledge Levels Knowledge Level Definition 4 Well known/small uncertainty 3 Known/moderate uncertainty 2 Partially known/large uncertainty 1 Very limited knowledge/uncertainty cannot be characterized 4.5 Systems, Structures, and Components To aid in the identification of phenomena, the PIRT panel divided the NPM into the SSC presented in Table 4-3. Phenomena were then identified in each SSC and for two of the LOCA phases. Table 4-3 Systems, Structures, and Components Reactor Pressure Decay Heat Removal Containment Vessel Reactor Building Pool Vessel System

RPV heat source
Reactor core
Decay heat removal Reactor Building pool
Containment vessel
Hot leg riser heat exchanger heat sink
Pressurizer
DHRS isolation valves
RRVs
RRVs
SG
RVVs
RVVs
Break
Break
Upper plenum
Downcomer
SG shell side (primary)
SG tube side (secondary)
LP 4.6 High-Ranked Phenomena Separate PIRTs were developed for the two types of LOCA scenarios defined in Section 4.2. More than 80 phenomena were identified and ranked in each PIRT. Only a few differences were identified between the two LOCA scenarios with respect to the phenomena that might occur and their associated ranking. Table 4-4 summarizes the phenomena that were ranked as high importance by the PIRT panel in at least one of the two phases of the LOCA Scenarios 1 and 2. The knowledge level assigned by the PIRT panel is also included. These high-ranked phenomena are addressed in the development of NuScale LOCA EM. These phenomena and their ranking are described in Table 4-4 below. Changes between the 2015 PIRT and the 2022 PIRT update are briefly noted in

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 4-4. Additional discussion of the high-ranked phenomena is provided in Section 4.6.1. A PIRT was also performed in 2021 for the inadvertent opening of a reactor valve (IORV) applying the same ranking approach identified in Section 4.4 where the FOM was CHFR. This PIRT is applicable to the initial portion of the IORV event and is termed Phase 0 as defined in Section 4.2. A listing of the newly identified high-ranked phenomena, relative to the LOCA PIRT, and their associated knowledge levels, is given in Table 4-5. The high-ranked phenomena are discussed further in Section 4.6.2. These phenomena are related to other phenomena that were assessed for LOCA. Table 4-5 lists the associated LOCA phenomena that address the IORV PIRT phenomena. Table 4-4 LOCA High-Ranked Phenomena ((

                                                                                                       }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 4-4 LOCA High-Ranked Phenomena (Continued) ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 4-4 LOCA High-Ranked Phenomena (Continued) ((

                                                                                             }}2(a),(c)

Table 4-5 Inadvertent Opening of a Reactor Valve Newly High-Ranked Phenomena Relative to LOCA PIRT ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 4.6.1 Discussion of LOCA Phenomena Ranked High Importance ((

                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                 }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                   }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                                }}2(a),(c) 4.6.2         Discussion of IORV Phenomena Ranked High Importance

((

                                      }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                    }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                     }}2(a),(c) 4.6.3         Treatment of Boron Transport Phenomena Originating in Phase 1a, 1b

((

                                    }}2(a),(c) 4.7      Phenomena Identification and Ranking Table Summary Some of the high-ranked phenomena identified in the NuScale PIRT are also important for existing reactors and have been the subject of considerable model development, testing, and analysis. Other phenomena are more unique to the NPM design due to the natural circulation coolant flow, integral RCS design, helical coil SG, unique passive ECCS and DHRS, reactor pool as the ultimate heat sink, and high-pressure containment.

Phenomena associated with the helical coil SG and the DHRS are ranked high importance due to their effect on the response for smaller breaks in the LOCA spectrum. ((

                                                                          }}2(a),(c) The LOCA break analysis in Section 9.3 shows that crediting the DHRS removes a large amount of energy from the primary fluid and significantly reduces the energy released to the CNV for these smaller breaks, or can significantly affect the timing of ECCS valve opening and collapsed liquid level for designs with IABs. Since DHRS is assumed available the phenomena are ranked high importance. Phenomena identification and ranking associated with the steam

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 generator and DHRS are further detailed in the Non-LOCA Topical Report (Reference 114). Some of the unique phenomena have a more developed knowledge base due to occurrence of the phenomena in other designs with different geometries, e.g., natural circulation. The PIRT identified the phenomena within the specified components as the high-importance phenomena that have a low-knowledge level. These high importance, low knowledge phenomena are given the greatest focus in the development of the LOCA EM. ((

                                                           }}2(a),(c)

High importance, low knowledge phenomena for IORV include: ((

                                                                    }}2(a),(c) 4.8      LOCA Containment Response PIRT Summary The results of the LOCA scenario PIRT are directly applicable to the primary system M&E release and resultant CNV pressure and temperature response that are the focus of the containment response methodology. The basis for this statement is that ((
                                                                               }}2(a),(c) Therefore, the LOCA scenario PIRT is also considered to be the LOCA containment response analysis methodology PIRT.

4.9 Non-Loss-of-Coolant Accident Event Phenomena Identification and Ranking Table Results NuScale has performed and documented a PIRT for the non-LOCA events. The results of the non-LOCA PIRT are summarized in the non-LOCA evaluation model (Reference 114). The results of the non-LOCA PIRT are directly applicable to the secondary system M&E release and CNV pressure and temperature response that are the focus of the containment response analysis methodology. The basis for this statement is that ((

                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 5.0 Evaluation Model Description This section provides a detailed description of the NPM LOCA model. The nodalization and modeling options selected for each NPM component are discussed along with the rationale for each choice. Justification is provided for the boundary and initial conditions selected for the model. A description of a break spectrum consistent with the requirements of 10 CFR 50 Appendix K is also provided. The NPM LOCA model is consistent with the SET and IET assessments used to validate NRELAP5 (Section 7.0). The model follows the recommended best practices for the preparation of a RELAP5-3D© input (Reference 8) that are applicable to the NRELAP5 LOCA model, as well as the NuScale-specific LOCA guidelines summarized in this report. The model conforms to the applicable requirements of 10 CFR 50 Appendix K, as described in Section 2.0. The results of the break spectrum calculations and the sensitivity studies (i.e., nodalization, time step, initial and boundary conditions, and selected model parameters) that supported the development of the LOCA EM are summarized in Section 9.0. 5.1 NRELAP5 Loss-of-Coolant Accident Model for the NuScale Power Module The unique design features of the NPM permit a simple and reliable approach to evaluate and mitigate the consequences of postulated LOCAs by: ensuring that all LOCAs are contained within the containment pressure vessel by designing the NPM such that the isolation of the CNV is a safety-related system. actuating the ECCS valves, which depressurizes the RPV into the CNV to establish pressure equalization to allow return of discharged fluid back into the RPV to cool the core. maintaining stable natural circulation flow through the ECCS valves with the reactor pool acting as the ultimate heat sink. In the event of a LOCA, these design features result in a simple, predictable transient progression, that can be explained by a standard mass and energy balance over the RPV and CNV considering: choked or unchoked flow through the break and ECCS valves between RPV and CNV. core decay heat generation and RCS stored energy release, heat transfer between CNV and reactor pool that is characterized by steam condensation at the CNV inside surface and free convection at the CNV outside surface to reactor pool. 5.1.1 General Model Nodalization The NRELAP5 model for analyzing a NPM LOCA is developed by reviewing the postulated scenarios and the key phenomena described in the NuScale LOCA PIRT, © Copyright 2022 by NuScale Power, LLC 71

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 summarized in Section 4.0. The model describes the key components of the NPM participating in a LOCA, as follows: RPV with internals

                  -   LP
                  -   reactor core
                  -   riser including the riser upper plenum
                  -   upper and lower downcomer
                  -   pressurizer CNV SG secondary side with DHRS condensers reactor pool ECCS valves postulated break locations RPV internal heat structures and heat structures between components (i.e., RPV to CNV to reactor pool) riser holes - this feature provides one or more sets of flowpaths between the riser and the downcomer promoting boron mixing.

The nodalization diagram of these key components is shown in Figure 5-1 for the NPM-160 and Figure 5-2 for the NPM-20. The nodalization diagrams have minor differences in the number of nodes for some components, such as the downcomer and riser, that are functionally equivalent. For components where differences reflect design changes or modeling changes, additional detail is provided. The details of the NRELAP5 NPM model are described in the following sections. © Copyright 2022 by NuScale Power, LLC 72

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 5-1 Noding Diagram of NRELAP5 Loss-of-Coolant Accident Input Model for the NPM-160 ((

                                                                                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 5-2 Noding Diagram of NRELAP5 Loss-of-Coolant Input Model for the NPM-20 ((

                                                                                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 5.1.2 Reactor Coolant System The RCS model is composed of the LP, reactor core, riser (lower, transition, and upper sections), riser plenum, downcomer (upper section containing the helical coil SGs and lower section), and pressurizer. 5.1.2.1 Lower Plenum ((

                                            }}2(a),(c) 5.1.2.2            Reactor Core 5.1.2.2.1              General Model The reactor core assembly is modeled with ((
                                     }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                           }}2(a),(c)

Various passive heat structures inside the reactor core are also considered to increase the release of the stored energy accumulated in non-heat-generating structures with appreciable metal mass. These structures include: core support assembly including core barrel, reflector, upper support blocks, and lower core plate, additional mass in fuel assemblies including top and bottom nozzles, upper and lower end caps, spacer grids, control rod assembly, instrument guide tubes, and springs. 5.1.2.2.2 Initial Power In accordance with 10 CFR 50 Appendix K, Section I.A, the initial reactor power level is set to 102 percent of the rated thermal power to account for 2 percent measurement uncertainty. 5.1.2.2.3 Core Power Distribution The power distribution ((

                                                              }}2(a),(c) The sensitivity calculations presented in Section 9.6.6 show that axial power shape has negligible impact on Phase 1 LOCA FOMs. The use of single power shape for all of the assemblies is also consistent with the point kinetics model inputs. Additional sensitivity studies for axial power distributions are performed for Phase 0 MCHFR calculations to ensure a limiting power distribution is used for that analysis.

5.1.2.2.4 Fuel Stored Energy The fuel rods are initialized at the maximum initial stored energy condition as required by 10 CFR 50 Appendix K, Section I.A.1. ((

                                                                                                    }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 (( }}2(a),(c) For the longer term fuel stored energy analysis, the UO2 fuel thermal conductivity, volumetric heat capacity, and fuel-cladding gap conductance are considered to be the key thermo-physical properties defining the stored energy in the fuel, as the amount of stored energy is inversely proportional to the thermal diffusivity of the fuel and fuel-cladding gap conductance. Based on the burnup-dependent fuel performance analysis provided by the fuel vendor, choosing ((

                                      }}2(a),(c)

Direct moderator heating is considered in the Phase 0 MCHFR calculation. For Phase 1a and Phase 1b analysis, neglecting direct moderator heating and assuming all reactor core power is deposited in the fuel is conservative. 5.1.2.2.5 Point Kinetics Model and Decay Heat The reactor kinetics model accounts for fission power due to prompt and delayed neutrons, decay power due to fission products, and actinides. The reactivity equations are solved using a point kinetics model with the

                      'separable' option (Section 6.4) to calculate the fission power. This model simulates reactivity changes due to reactor trip and feedback reactivity due to Doppler and moderator density effects. The reactivity change due to the insertion of control rods is modeled using a scram reactivity table as a function of time after the reactor trip. The table reflects conservative representations for the onset of rod motion, the rod position as a function of time after trip and the inserted reactivity as a function of rod position. The scram rod total worth considers that the most reactive rod remains stuck and does not insert into the core following reactor trip. Sensing signal delays are accounted for based on the instrument type (e.g., pressure and temperature) to determine if a reactor trip should be initiated and an additional delay is added to account for the initiation of rod insertion.

The Doppler and moderator density feedback input parameters depend on the average fuel burn-up. The weighted average fuel temperature and moderator density for the feedback reactivity calculations consider the given core power (flux) distribution to be consistent with the point kinetics model. ((

                                         }}2(a),(c)

Six groups of delayed neutron precursors are considered based on the SIMULATE analysis for different cycles, beginning-of-cycle, middle-of-cycle, and end-of-cycle. Beginning-of-cycle kinetic parameters are used as bounding © Copyright 2022 by NuScale Power, LLC 77

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 for point kinetic input with the smallest prompt neutron lifetime to maximize initial energy inventory by prolonging the fission power transient. The variation in precursor decay constants is insensitive to the time in cycle. Additional biasing is introduced to kinetic parameters to account for uncertainty in calculated values in such a way that both the prompt lifetime and the decay constants for each precursor group are decreased. The objective is to maximize the delayed neutron contribution to the total fission power. The 'gamma-ac' option in the NRELAP5 point kinetics model activates the models for the transient effects of decay heat and actinides. Use of the actinides model complies with 10 CFR 50 Appendix K, Section I.A.3. The ANS71 option represents an explicit implementation of four time-dependent exponentials detailed in the draft ANS73 decay heat standard and includes a built-in 1.2 multiplier. Activation of a trip associated with this option initiates evaluation of the power decay as a function of time. However, this time-evaluation of power does not account for post-trip prompt and delayed fissions that can add additional decay heat precursors and increase the integral heat release. The ANS73 model represents an 11 group exponential fit of production and decay constants of the decay heat defined in the standard. User input of a 1.2 multiplier addresses 10 CFR 50 Appendix K requirements for addressing prediction uncertainties (10 CFR 50 Appendix K, Section I.A.4). Decay heat is predicted by the behavior of the 11 precursor groups, and no explicit reactor trip is applied. Instead the model predicts precursor concentrations from the prompt and delayed fissions rate, and so naturally follows fission power. Sensitivities have determined that the ANS71 option requires careful selection of an additional delay time beyond a reactor trip before activating the power decay to account for post-trip prompt and delayed fissions. This delay time, which depends on control rod insertion speed and reactivity, can be more than (( }}2(a),(c) seconds. Once the delay is accounted for, both options are consistent for short term LOCA through the first 1000 seconds. Considering the correct response of the ANS73 option without any special evaluation of delay times, the ANS73 option was chosen as the standard choice in the LOCA guideline. The best-estimate ANS79 actinides model is used to account for heat deposition from actinide decay. The actinide model includes the decay energies from the production and decay of 239U, 239Np, and 239Pu. 5.1.2.3 Riser ((

                                                                                                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                 }}2(a),(c) 5.1.2.4            Downcomer

((

                                      }}2(a),(c) 5.1.2.5            Pressurizer

((

                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                      }}2(a),(c) 5.1.2.6            Collapsed Liquid Level Calculation The collapsed liquid levels in the riser, downcomer, and containment are calculated based on the total liquid volume calculated in each part of the NPM and volume elevation table for these regions. The riser volume includes the lower plenum, core and bypass region, and riser section up to the pressurizer baffle plate. Similarly, the downcomer volume includes the lower plenum, lower and upper downcomer section, and upper riser plenum up to the bottom of the pressurizer or the pressurizer baffle plate.

This approach to collapsed liquid level allows for there to be substantial flashing and momentary voiding in the core, such as that seen at near stagnant flow conditions for small liquid space breaks with delayed ECCS valve opening. This is discussed in Section 9.2 where assurance of no fuel CHF, and hence no fuel heat-up, is shown with transient MCHFR remaining above the steady-state CHFR value for all cases. ((

                                                                                   }}2(a),(c) Evaluation methodology for the collapsed liquid level for the 72-hour period following and event is provided in Reference 13.

5.1.3 Helical Coil Steam Generators Two helical coil SGs are represented using an NRELAP5-specific helical SG component that models the component-specific internal pressure drop and heat transfer effects, as described in Section 6.7. The two independent SGs are thermally connected to the upper downcomer to transfer heat to the steam turbine during normal operation. During off-normal operations, each SG transfers energy to an independent safety-related DHRS (Section 5.1.7) to discharge energy to the reactor pool. © Copyright 2022 by NuScale Power, LLC 80

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The tube-to-coil diameter ratio is specific to the SG geometry. ((

                                                                    }}2(a),(c) 5.1.4         Containment Vessel and Reactor Pool

((

                       }}2(a),(c)

The reactor pool is the ultimate heat sink in the NPM design. The reactor pool volume corresponding to an individual NPM is represented by a ((

                                                                        }}2(a),(c) A wide range of initial

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 reactor pool temperatures is exercised to show the effect of the pool conditions on the LOCA behavior in Section 9.6.5. ((

                                                   }}2(a),(c) 5.1.5         Chemical and Volume Control System The entirety of the CVCS is not necessary to be included in the LOCA model. The CVCS, a nonsafety-related system, is not automatically actuated. ((
                                                                                                     }}2(a),(c)

Continued operation of the CVCS through operator action would add cold water that is non-conservative. The only CVCS piping that is necessary to be represented in the model is the injection line from the RPV wall, through the downcomer and into the riser. It connects the injection line break to the containment vessel at the correct elevation and accounts for a small loss through the line. The discharge line connection at the downcomer, and the two spray supply line and high point vent line connections at the top of the pressurizer are used as break locations. The volumes corresponding to CVCS piping constitute a small fraction of the total RPV and CNV volume; therefore, it has negligible impact on the progression of NPM LOCA. ((

                                                              }}2(a),(c) 5.1.6         Secondary System The model represents the secondary feedwater and steam lines with two helical coil SGs, described in Section 5.1.3. ((
                                                                          }}2(a),(c) The secondary side includes the DHRS with two trains of heat exchangers with feed and steam line piping, described in Section 5.1.7.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                             }}2(a),(c) 5.1.7         Decay Heat Removal System Both DHRS trains are included in the NPM LOCA model. The two independent trains of the DHRS are safety-related systems. The DHRS system removes heat from the RCS and transfers it to the reactor pool through the steam generators and the DHRS heat exchangers. The function of the DHRS is most significant in smaller breaks when there is significant time for heat transfer prior to ECCS actuation. The reduction of energy in the RCS is significant in the CNV peak pressure analyses.

5.1.8 NRELAP5 Modeling Options The NPM LOCA analysis is performed with the latest released version of NRELAP5. ((

                                }}2(a),(c) 5.1.8.1            Junction Options

((

                                                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 5-1 Default Junction Options for the NRELAP5 Loss-of-Coolant Accident Model ((

                                                                                                      }}2(a),(c)

((

                                                                  }}2(a),(c)
  © Copyright 2022 by NuScale Power, LLC                                                           84

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                                   }}2(a),(c) 5.1.8.2            Volume Options

((

                                         }}2(a),(c) This format is described by Table 5-2.

Table 5-2 Default Volume Options for the NRELAP5 Loss-of-Coolant Accident Model ((

                                                                                                     }}2(a),(c)

((

                                            }}2(a),(c) 5.1.8.3            Heat Structure Options

((

                                                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                                                 }}2(a),(c) 5.1.9         Time Step Size Control The NuScale LOCA EM uses the NRELAP5 semi-implicit scheme for the solution of the hydrodynamics. The heat structure solution is implicitly coupled to the hydrodynamic solution. With given user-specified minimum and maximum time step sizes, the code determines the appropriate time step in such a way that the current time step cannot be larger than the courant-time step size determined based on the limiting volume.

no significant mass error accumulation occurs during the solution and halving of the current time step when it is deemed necessary. NRELAP5 provides the capability of providing the user-defined maximum time step size through the definition of control variable. A control variable that defines the fraction of the current courant time-step size during the solution is used to set the user-defined maximum time step size. This approach has the advantage of taking larger time steps when larger courant time step sizes exist during the solution; therefore, the code takes larger time steps when the solution indicates smooth transient progression. (( }}2(a),(c) © Copyright 2022 by NuScale Power, LLC 86

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                        }}2(a),(c)

A sensitivity study is performed on the fraction specified to demonstrate that the selected maximum time-step size has no or insignificant impact on the LOCA figures of merit such as peak containment pressure and collapsed liquid level in the RPV riser above the TAF. 5.2 Analysis Setpoints and Trips A number of safety-related measurements exist in the NPM to detect off-normal conditions. Table 5-3 shows the typical safety-related measurements relevant to LOCA, IORV, and CNV pressure or temperature analysis along with their functions. The safety analysis analytical limits specify the setpoints, or range of setpoints, and the sensing delay for each safety-related signal. These parameters can be monitored in the NRELAP5 model to credit safety-related signal actuations. A single parameter may be used for multiple actuations. For example, a high containment pressure signal could result in reactor trip, containment isolation, and DHRS actuation. The MPS actuation signals and limits vary by NPM design and may be incorporated into the NRELAP5 model for the design. Table 5-3 Typical NuScale Power Module Safety-Related System Measurement Parameters ((

                                                                                                       }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 5-3 Typical NuScale Power Module Safety-Related System Measurement Parameters (Continued) ((

                                                                                                        }}2(a),(c)

((

                                                                                                    }}2(a),(c)

The mixture level detection typically uses a simple approximation of the mixture level such as based on ((

                         }}2(a),(c) 5.3      Initial Plant Conditions Table 5-4 provides the basis for conservatively biasing the initial conditions for LOCA analysis. These ranges are intended to account for both the normal control system deadband and the system/sensor measurement uncertainty without specifically quantifying the portion of the range applied to either uncertainty. For Phase 0 MCHFR analysis, the conservative bias directions are consistent, or a range of initial conditions are evaluated to confirm limiting bias directions.

Table 5-4 Plant Initial Conditions ((

                                                                                                        }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 5-4 Plant Initial Conditions (Continued) ((

                                                                                                       }}2(a),(c) 5.4      Loss-of-Coolant Accident Break Spectrum 10 CFR 50 Appendix K describes the break spectrum as a set of LOCA scenarios that are uniquely defined based on location, configuration and size. Additional analyses were performed on availability of DHRS, availability of power, and postulated single failures.

The break spectrum for the NuScale LOCA EM is summarized in this section. 5.4.1 Break Location The postulated break locations in the NPM design are the RCS injection and discharge lines, the pressurizer spray supply line, and high point vent lines. These break locations establish a flow path between RPV and CNV leading to CNV pressurization during the early phase of LOCA (i.e., Phase 1a): The injection line enters the RPV through a shell penetration and piping internal to the RPV that passes through the downcomer and terminates at an upwardly-oriented nozzle in the riser. The discharge line is connected to a RPV penetration to the downcomer. The pressurizer spray supply line is connected to the top of the pressurizer at two separate penetrations. At each penetration there is a nozzle within the RPV wall. Outside the RPV wall (but within the CNV), the pressurizer spray supply lines connect to a tee which in turn connects to isolation valves on the CNV wall. The high point vent line connects to the top of the pressurizer. All of the connections to the RPV are normally open, except for the high point vent. Each connection can be isolated by two independent safety-related isolation valves in series that close on the containment isolation signal. As a result, discharged break fluid is retained within the CNV for eventual return to the RPV when ECCS actuates. 5.4.2 Break Configuration and Size Table 5-5 summarizes the size and location of the breaks considered as part of the break spectrum of the NuScale LOCA EM. Break area dimensions in Table 5-5 © Copyright 2022 by NuScale Power, LLC 89

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 correspond to 2-inch Schedule 160 piping. ((

                    }}2(a),(c),ECI Table 5-5 Summary of Analyzed Break Sizes

((

                                                                                                  }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 5.4.3 Single Failures 10 CFR 50 Appendix K requires that single failures be considered within the break spectrum. This includes analyzing a system/component classified as nonsafety related if the inclusion of that system/component would introduce a more limiting condition. The following scenarios are considered: no single failure failure of a single RVV to open failure of a single RRV to open failure of one ECCS division (i.e., one RVV and one RRV fail to open) The ECCS valves are held closed with direct current (DC) power and operate on two independent divisions. In the event of an inadvertent actuation of a division (removal of DC power from that division or inadvertent MPS signal), one division of ECCS valves is available to open immediately. For ECCS valves with an IAB installed, the IAB setpoint prevents the opening of those valves that include that feature until the differential pressure between RCS and CNV falls below the IAB release pressure. The other division still actuates on the ECCS actuation signal creating a staggered release. This is a non-limiting case as a staggered release has a smaller impact on system pressures, levels, and core coolability relative to the actuation of both ECCS trains. For ECCS valves that do not have the IAB feature, the valves open immediately on the inadvertent signal. The remaining ECCS valves open when an ECCS signal is actuated (and IAB release pressure is reached for those valves so equipped). 5.4.4 Loss of Power Coincident with a postulated LOCA, two scenarios for loss of power are considered within the LOCA methodology: complete loss of normal alternating current (AC) and DC (e.g., EDSS for NPM-160, EDAS for NPM-20) power and complete loss of only AC power with DC power availability. The loss of DC power can impact the LOCA progression by immediately triggering valves to go to their fail-safe position. Table 5-6 presents the valves along with their fail-safe state. © Copyright 2022 by NuScale Power, LLC 91

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 5-6 NuScale Power Module Valve Fail-Safe Positions with Loss of DC Power ((

                                                                                                       }}2(a),(c)

For the ECCS valves with an IAB, after loss of DC power, the IAB arming valves close because the valve differential pressure is greater than the threshold setpoint, thereby preventing the immediate opening of these ECCS valves. As the RPV pressure decreases and the CNV pressure increases, the valves open as soon as the differential pressure drops below the IAB release pressure setpoint. When normal AC power is lost, the feedwater pumps coast down and the turbine trip is initiated. Upon loss of normal AC power (with a time delay) the reactor trip, containment isolation, and DHRS actuation signals are generated. In addition, on a reactor trip, in some NPM designs, a timer is initiated to actuate ECCS that can be manually bypassed by the operator if subcriticality under cold conditions is maintained. This operator action is not credited. If ECCS does not actuate on the post-trip timer and AC power is not restored, the ECCS actuates after 24 hours. The ECCS is still available to actuate based on its normal actuation signals at all times, even if the post-trip timer is bypassed. 5.4.5 Decay Heat Removal System Availability When the SG tubes are uncovered within the RPV, operation of the DHRS results in condensation of steam on the external surface of the helical coil SGs and retains liquid inventory in the RPV instead of releasing it to the CNV through vaporization. DHRS reduces RPV pressure and maintains a higher minimum inventory in the RPV. There is no single failure that can prevent a single DHRS train from actuating. Therefore, both trains of DHRS are available as long as secondary piping remains intact. © Copyright 2022 by NuScale Power, LLC 92

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 5.5 Containment Response Analysis Model The LOCA models in NRELAP5 are listed in Table 5-7 along with the application in the containment response analysis methodology. Table 5-7 LOCA Versus Containment Response Models Application in the Containment Response LOCA Analysis Model Analysis Methodology Condensation Heat Transfer Used for condensation heat transfer on the CNV

(( inside diameter and inside the DHRS heat exchanger tubes.

                       }}2(a),(c)

Critical flow Used for two-phase saturated critical flow

Moody critical flow model for two-phase flow conditions Helical Coil SG component Used for modeling the helical coil SGs
Heat transfer correlation
Friction correlation Pool heat transfer Churchill-Chu is used for modeling the CNV outside
Churchill-Chu natural convection correlation diameter, the reactor pressure vessel outside correction to use bulk fluid properties diameter and outside the DHRS heat exchanger tubes (vertical surfaces only)

Interfacial drag multiplier Used in containment response analysis

Input multiplier added to allow minimizing liquid methodology to evaluate the effect of liquid entrainment in break and valve flow entrainment on break and valve flow Void drift velocity
Kataoka-Ishii alternative formulation set to default Used for two phase flow Critical heat flux options, Section 5.1.8 The 2006 Groenveld tables are used in the containment response analysis methodology. CHF does not occur for all LOCA and non-LOCA scenarios in the containment response analysis methodology.

Decay heat Not used in the containment response analysis

1971 ANS Standard including actinides methodology 5.6 Containment Response Analysis M&E Release Model 5.6.1 NRELAP5 Primary Release Event Analysis Model Overview The NRELAP5 model used to model NPM performance for primary system (LOCA and ECCS valve opening) release event analyses is similar to the model used in the LOCA evaluation model described in this chapter. The NPM geometry inputs and conservative fuel inputs in the containment response analysis model are consistent

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 with those used by the LOCA Evaluation Model, except for those parameters biased to maximize CNV temperatures and pressure as specified in Table 5-10. The following substantive differences are related to the objective of determining the maximum containment peak pressure and peak temperature scenarios. This is accomplished by conservatively maximizing the M&E release and minimizing containment heat removal. Figure 3-1 is an illustration of the NPM, during power operation that shows the main design features. Figure 3-2 illustrates the ECCS mode of operation and shows the RVVs and RRVs along with the DHRS, CNV and reactor pool that provide containment heat removal and ultimate heat sink. The nodalization diagram in Figure 5-1 and Figure 5-2 plus the assumptions described in this section constitute the NRELAP model used to simulate primary release scenarios resulting from bounding breaks and valve opening events. The following modification is included in the primary release event containment response analysis model: LOCA Pipe Break and Valve Opening Modeling ((

                                        }}2(a),(c)

Conservative modeling of the LOCA pipe break spectrum and the valve opening events to ensure a bounding M&E release includes the following elements: all break locations are considered maximum credible break size at each location critical flow with discharge coefficient of 1.0 saturated liquid - Moody critical flow subcooled liquid - (( }}2(a),(c) modified pressure volume work term maximum RRV and RVV flow areas liquid entrainment evaluated by use of interfacial drag multiplier in upper riser, riser upper plenum, pressurizer baffle, pressurizer, and downcomer © Copyright 2022 by NuScale Power, LLC 94

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Containment Vessel and Reactor Pool Models The CNV nodalization in the NRELAP5 loss-of-coolant accident and valve opening event containment response analysis model (Figure 5-1 and Figure 5-2, Component 500) is consistent with the LOCA evaluation model. The CNV is maintained at a partial vacuum with an assumed high initial pressure (e.g., 3.0 psia), and with the maximum total mass of noncondensables that could exist within the CNV during operation, in order to capture the effects of CNV non-condensable gases. Also, during a LOCA or valve opening event, the maximum total mass of noncondensables that could exist within the RPV during operation are released to the CNV model, in order to capture the effects of RPV non-condensable gases. The LOCA or valve opening event M&E release into the CNV results in a rapid heating and pressurization of the CNV. The steam is condensed on the CNV inside diameter and the condensate film flows downward and forms a pool in the bottom of the CNV. As the CNV pool level rises boiling occurs on the RPV surface. Heat transfer from the CNV outside diameter to the reactor pool initially maintains the vessel at a low temperature except for the upper section of the vessel that is above the pool surface elevation. Following the LOCA or valve opening event, the condensing of steam and convection from the CNV pool increases the vessel temperature, and heat transfer from the CNV outside diameter to the reactor pool increases. Heat transfer on the CNV outside diameter is by pool convection and pool nucleate boiling, except for the upper section that is not submerged in the reactor pool. In the upper section heat transfer is neglected. The initial CNV wall temperature is conservatively addressed based on the assumed reactor pool level. At normal operation conditions, there is very little heat loss from the RPV to CNV, such that the CNV metal below the pool surface is very close to the pool temperature. This region of the CNV heat structure is initialized within NRELAP5 based on explicit treatment of the pool heat transfer on the CNV outer surface. The region of the CNV heat structure above the pool level is conservatively given an adiabatic outer surface boundary condition. The CNV wall temperatures ((

                                                                                                     }}2(a),(c)

Due to the 1-D modeling of the CNV volume and modeled heat transfer from the RPV wall heat structures to the CNV volume, NRELAP5 calculates a large spread in the steady state CNV volume temperatures above the maximum CNV normal operation temperature, particularly in those volumes connected to the PZR wall heat structure where the RPV wall heat transfer area is large. These temperature variations exceed the maximum CNV normal operation temperature, which is conservative with respect to peak CNV pressure and temperature. Ultimately, the initial CNV wall temperatures are the critical inputs to calculating the peak CNV wall temperature figure of merit in © Copyright 2022 by NuScale Power, LLC 95

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 this analysis, ((

                                                                                                }}2(a),(c)

Ultimately, these steady state CNV volume temperature variations are acceptable because they are conservatively high and because the CNV heat structures (used for calculating the peak CNV temperature figure of merit) are appropriately initialized at their analytical limit. Conservative modeling of the heat transfer to and from the CNV inside diameter, and from the CNV outside diameter to the reactor pool, to ensure a bounding peak CNV pressure and temperature response following a LOCA or valve opening event, includes the following elements: ((

                                                                  }}2(a),(c)

Table 5-8 shows the heat transfer correlations and models for all of the processes that could impact the CNV peak pressure and temperature response. Table 5-8 Containment Vessel and Reactor Pool Heat Transfer Modeling Heat Transfer Process Correlation/Model Radiation enclosure model is considered in analysis. Radiant heating from RPV outside diameter to However, inclusion of the radiation enclosure model has CNV inside diameter a negligible impact on CNV peak pressure and temperature results. Convection from RPV outside diameter to CNV Vertical Surfaces pool ((

                                                                                     }}2(a),(c)

Non-Vertical Surfaces ((

                                                                }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 5-8 Containment Vessel and Reactor Pool Heat Transfer Modeling (Continued) Heat Transfer Process Correlation/Model Convection from CNV outside diameter to Vertical Surfaces reactor pool ((

                                                                                          }}2(a),(c)

Non-Vertical Surfaces ((

                                                                     }}2(a),(c)

Reactor pool mixing No mixing is modeled Reactor pool cooling to ambient Assumed adiabatic Reactor pool mixing with other modules No mixing with other modules is modeled 5.6.1.1 Primary System Release Event Initial Conditions Initial conditions for the spectrum of primary system release containment response analyses are selected to ensure a conservative CNV peak pressure and peak temperature result. The process of selecting the initial conditions is consistent with the guidance in DSRS Section 6.2.1.3. The selection process ensures that energy sources are maximized and energy sinks are minimized. Table 5-9 presents the primary system initial conditions for the primary system release containment response analyses. Table 5-9 Primary System Initial Conditions ((

                                                                                                           }}2(a),(c)

The initial conditions in the secondary system, in particular ((

                                                                                                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                      }}2(a),(c) The SG initial conditions result from the NRELAP5 initialization process and are consistent with the conservative primary system initial conditions.

The basis for the initial conditions for the CNV and the reactor pool are shown in Table 5-10. These initial conditions ensure that the CNV heat sink is minimized so that the peak containment pressure and temperature are modeled conservatively. Table 5-10 Containment Vessel and Reactor Pool Initial Conditions ((

                                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 5-10 Containment Vessel and Reactor Pool Initial Conditions (Continued) ((

                                                                                                      }}2(a),(c) 5.6.1.2            Primary System Release Event Boundary Conditions Boundary conditions for the spectrum of primary system M&E release analyses are selected to ensure a conservative CNV peak pressure and peak temperature result. The process of selecting the boundary conditions is consistent with the guidance in DSRS Section 6.2.1.3. The selection process ensures that energy sources are maximized and energy sinks are minimized. Due to the simplicity of the NPM design there are few postulated single failures for the primary system M&E release scenarios. Failure of ECCS valves to open are analyzed as sensitivity studies, and failure of MSIVs or FWIVs to close are considered, but they have minimal effect on the CNV pressure and temperature response as the secondary system is immediately isolated for the primary side events. Table 5-11 presents the boundary conditions for the LOCA containment response analyses.

Table 5-11 Containment Vessel Peak Pressure and Temperature Calculation Assumptions ((

                                                                                                    }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 5-11 Containment Vessel Peak Pressure and Temperature Calculation Assumptions (Continued) ((

                                                                                      }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 5-11 Containment Vessel Peak Pressure and Temperature Calculation Assumptions (Continued) ((

                                                                                               }}2(a),(c) 5.6.2         NRELAP5 Secondary System Mass and Energy Release Analysis Model 5.6.2.1            Secondary Side Pipe Breaks The NRELAP5 non-LOCA model is the starting point for developing the MSLBs and FWLBs models in the containment response analysis methodology.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 5-3 Typical Non-LOCA Nodalization Diagram MSLB and FWLB scenarios start with the blowdown of the secondary inventory through the pipe break and into the CNV. The reactor trips on high CNV pressure or low steam line pressure. The reactor trip causes a turbine trip along with main steam isolation and feedwater isolation. One SG depressurizes as the CNV pressurizes, and an equilibrium is approached. The DHRS actuates, subsequent to feedwater isolation, and transfers decay heat to the reactor pool. Steam released into the CNV condenses on the CNV inner surface that is cooled by conduction and convection to the reactor pool. The MSLB or FWLB inside containment analysis includes the following modeling considerations: break modeling with (( }}2(a),(c) reactor trip on low steam line pressure, high CNV pressure, or appropriate MPS signal © Copyright 2022 by NuScale Power, LLC 102

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 MSIVs actuation FWIVs and FWRVs actuation feedwater pump dynamic response feedwater pipe inventory flashing DHRS actuation with or without loss of normal AC and DC electrical power limiting single failure The NRELAP5 MSLB or FWLB modeling for the containment response analysis methodology focuses on a conservative analysis of the CNV peak pressure and temperature response. Overview The NRELAP5 model used for secondary system pipe break analysis in the containment response analysis methodology is similar to the NRELAP5 model used in the non-LOCA accident methodology. The differences are related to the objective of determining the maximum containment peak pressure and peak temperature scenarios. This is accomplished by conservatively maximizing the M&E release, and minimizing containment heat removal. The models for analyzing the M&E for secondary line breaks are described in the following sections. Feedwater System Model The feedwater system is an important boundary condition for the secondary system M&E release analyses. The initial secondary inventory in the helical coil SG is small and does not by itself cause a significant CNV pressurization following a secondary line break. The main source of mass is the feedwater system due to an assumed single failure of the FWIV on the affected helical coil SG. Also, the feedwater pump is assumed to respond to the decrease in helical coil SG pressure by a corresponding increase in feedwater flow. Feedwater flow continues to supply the affected helical coil SG until the FWRV automatically closes to back up the FWIV. Secondary Pipe Break Model The secondary pipe break spectrum modeling in the containment response analysis methodology is the same as in the non-LOCA methodology, with the limiting break size being the double-ended break which are typically modeled with a break flow path on either end of the break pipe and a normal flow path. At the time of the break the normal flow path is closed while the other two are opened at the approximate break location elevation in the CNV. Main steam isolation valve closure isolates the unaffected SG from the affected SG. A single failure of one © Copyright 2022 by NuScale Power, LLC 103

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 MSIV to close is addressed by automatic closure of the secondary MSIV on each steam line. Conservative modeling of the secondary pipe breaks to ensure a bounding M&E release includes the following elements: ((

                                                         }}2(a),(c) 5.6.2.2            Main Steam Line Break Initial Conditions Initial conditions for the MSLB containment response analyses are selected to ensure a conservative CNV peak pressure and peak temperature result. The process of selecting the initial conditions is consistent with the guidance in DSRS Section 6.2.1.4. The selection process ensures that energy sources are maximized and energy sinks are minimized. ((
                                          }}2(a),(c) Table 5-12 presents the secondary system initial conditions used by the MSLB containment response analyses.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 5-12 Secondary System Initial Conditions ((

                                                                                                     }}2(a),(c) 5.6.2.3            Main Steam Line Break Boundary Conditions Boundary conditions for the MSLB mass and energy release analyses are selected to ensure a conservative CNV peak pressure and peak temperature result. The process of selecting the boundary conditions is consistent with the guidance in DSRS Section 6.2, and specifically DSRS Section 6.2.1.4. The selection process ensures that energy sources are maximized and energy sinks are minimized.

The largest break size is assumed to maximize the secondary system M&E release rate into the CNV and thereby maximize the resulting CNV pressurization and temperature increase. However, a subsequent primary system M&E release © Copyright 2022 by NuScale Power, LLC 105

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 following ECCS actuation and opening of the RVVs may result in the peak CNV pressure and temperature response for some scenarios. As the DHRS cools the primary system, a delayed M&E release through the RVVs is smaller, and the second CNV pressurization when the RRVs open is lower. Furthermore, the steam line break CNV pressure and temperature response remains bounded by the LOCA. Therefore, the maximum MSLB size is bounding and a break spectrum analysis is not necessary. Due to the simplicity of the NPM design, there are few postulated single failures for the secondary system M&E release scenarios. Failure of ECCS valves to open is considered for the scenarios in which ECCS actuation occurs. Failures of MSIVs or FWIVs to close are analyzed as sensitivity studies. Table 5-11 presents the boundary conditions for the primary system containment response analysis methodology, and they are the same for the MSLB containment response analysis methodology except for those presented in Table 5-13. Table 5-13 Boundary Conditions for the Main Steam Line Break Containment Response Analysis Methodology ((

                                                                                                      }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 5-13 Boundary Conditions for the Main Steam Line Break Containment Response Analysis Methodology (Continued) ((

                                                                                                         }}2(a),(c) 5.6.2.4            Feedwater Line Break Initial Conditions Initial conditions for the FWLB mass and energy release analyses are selected to ensure a conservative CNV peak pressure and peak temperature result. The process of selecting the initial conditions is consistent with the guidance in DSRS Section 6.2, and DSRS Section 6.2.1.4 specifically. The selection process ensures that energy sources are maximized and energy sinks are minimized.

((

                                                                    }}2(a),(c) 5.6.2.5            Feedwater Line Break Boundary Conditions Boundary conditions for the FWLB mass and energy release analyses are selected to ensure a conservative CNV peak pressure and peak temperature result. The process of selecting the boundary conditions is consistent with the guidance in DSRS Section 6.2, and specifically DSRS Section 6.2.1.4. The selection process ensures that energy sources are maximized and energy sinks are minimized. Section 5.6.2.3 and Table 5-13 present the boundary conditions

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 used by the MSLB containment response analyses, these boundary conditions are also used by the FWLB containment response analyses. The largest break size is assumed to maximize the initial M&E release into the CNV. However, it is the subsequent second M&E release following ECCS actuation and opening of the RVVs that results in the peak CNV pressure and temperature response. Therefore, the initial break size is unimportant as the secondary M&E release is similar, and the sequence of events leading to the opening of the RVVs is similar. Furthermore, the feedwater line break CNV pressure and temperature response is bounded by the LOCA. Therefore, a break spectrum analysis is not necessary. Due to the simplicity of the NPM design, there are few postulated single failures for the secondary system M&E release scenarios. Failure of ECCS valves to open is considered for the scenarios in which ECCS actuation occurs. Failures of a MSIV or a FWIV to close are analyzed as sensitivity studies 5.7 Sensitivity Studies The sensitivity calculations described in Section 9.6 are performed in the following categories: sensitivies required by 10 CFR 50 Appendix K, (e.g., nodalization and time-step size to demonstrate the stability and consistency of the numerical scheme used by the NRELAP5 code), sensitivities related to key phenomena and design input parameters considered to be important to the LOCA progression and LOCA FOMs (e.g., CCFL at pressurizer baffle plate, ECCS parameters). sensitivities to determine input parameters to ensure conservativism (e.g., reactor pool initial temperature, core power distribution). sensitivity of reduced DHRS capacity on LOCA event progression. © Copyright 2022 by NuScale Power, LLC 108

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 6.0 NRELAP5 Code Description The NuScale LOCA EM is based on the NRELAP5 system thermal-hydraulics code. The NRELAP5 code includes models for characterization of hydrodynamics, heat transfer between structures and fluids, modeling of fuel, reactor kinetics models, and control systems. NRELAP5 uses a two-fluid, non-equilibrium, non-homogenous model to simulate system thermal-hydraulic responses. This section provides a general overview of the code structure, models, and correlations. This section also addresses the LOCA-specific code models and improvements implemented to address unique design features and phenomena for the NPM. The adequacy of code models and correlations essential for modeling all high-ranked PIRT phenomena is discussed in Section 8.0. The full details of the models and correlations that makeup NRELAP5 can be found in the NRELAP5 Theory Manual (Reference 9). RELAP5-3D©, version 4.1.3, was used as the baseline development platform for the NRELAP5 code. RELAP5-3D© was procured and as part of the procurement process commercial grade dedication was performed by NuScale to establish the baseline NRELAP5 code. Subsequently, features were added and changes were made to NRELAP5 to address the unique aspects of the NPM design and licensing methodology. Those aspects of NRELAP5 that are new or revised specifically for the NPM application include: ((

                                                  }}2(a),(c)

The previous RELAP5 series of codes were developed at the INL under sponsorship of the DOE, the NRC, members of the International Code Assessment and Applications Program, members of the Code Applications and Maintenance Program, and members of the International RELAP5 Users Group. Specific applications of the code have included simulations of transients in light water reactor systems, such as LOCAs, anticipated transients without scram, and anticipated operational occurrences, such as loss of feedwater, loss of offsite power, station blackout, and turbine trip. The RELAP5 code, including the RELAP5-3D© version that was used as the development platform for NRELAP5, has an extensive record of usage and acceptable performance for nuclear safety analysis. RELAP5-3D© is the latest version of the RELAP5 code that has been under continuous development since 1975, first under NRC sponsorship and then with additional DOE sponsorship beginning in the early 1980s. While NRC sponsorship ended in 1997, the DOE continued sponsorship of RELAP5-3D© to meet its own reactor safety assessment needs. The RELAP5 code was chosen by DOE as the thermal-hydraulic analysis tool because of its widespread acceptance. © Copyright 2022 by NuScale Power, LLC 109

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Systematic safety analyses were carried out for the DOE that included the N reactor at Hanford, the K and L reactors at Savannah River, the Advanced Test Reactor at the Idaho National Engineering and Environmental Laboratory, the High Flux Isotope Reactor and Advanced Neutron Source at Oak Ridge, and the High Flux Beam Reactor at Brookhaven. The DOE also chose RELAP5 for the independent safety analysis of the New Production Reactor proposed for Savannah River . RELAP5-3D© has worldwide usage for nuclear safety analysis. Users participate in the International RELAP5 Users Group (IRUG) which provides a forum for code users to share their RELAP5 development and analysis experiences. Meeting participants also communicate new features and applications that have been developed for RELAP5-3D©. Code users include reactor vendors, nuclear industry suppliers, a naval nuclear propulsion laboratory, universities, and international organizations. NuScale is a participant in IRUG. RELAP5-3D© has been chosen as a code development platform for small break LOCA analysis by Mitsubishi Heavy Industries for APWR (Reference 21). Furthermore, the NRC has performed a detailed adequacy evaluation of RELAP5/MOD3 Version 3.2.1.2 for analysis of design-basis small break LOCAs in the Westinghouse AP600 reactor (Reference 74). This usage of RELAP5-3D© over a long period of time has produced a large amount of user feedback. Submission of code error reports and the follow up code development has resulted in a robust code which can be used with a high level of confidence that significant code problems have been identified and corrected. The more than 20 year history of code assessment and successful application of the RELAP5-3D© code, and codes based on the RELAP5-3D© platform, by the worldwide user community has successfully exercised the fundamental capabilities of RELAP5-3D© that are the critical characteristics required of NRELAP5 for NuScale's application. The NRELAP5 code is developed following the requirements of the NuScale QAPD (Reference 4). The NuScale corporate Software Configuration Management Plan provides a framework for NRELAP5 configuration management and change control in conformance with the requirements outlined in the NuScale Software Program Plan. Review and approval of the NuScale corporate Software Configuration Management Plan is not within the scope of this report. 6.1 Quality Assurance Requirements The NuScale QAPD complies with the requirements of 10 CFR 50 Appendix B, Quality Assurance Criteria for Nuclear Power Plants and Fuel Reprocessing Plants (Reference 10) and American Society of Mechanical Engineers (ASME) NQA-1-2008 and NQA-1a-2009 Addenda, "Quality Assurance Program Requirements for Nuclear Facility Applications," (Reference 12). © Copyright 2022 by NuScale Power, LLC 110

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 6.2 NRELAP5 Hydrodynamic Model The NRELAP5 hydrodynamic model is a transient, two-fluid model for flow of a two phase vapor/gas-liquid mixture that can contain non-condensable components in the vapor/gas phase as well as a soluble component (i.e., boron) in the liquid phase. The two-fluid equations of motion that are used as the basis for the NRELAP5 hydrodynamic model are formulated in terms of volume and time-averaged parameters of the flow. Phenomena that depend upon transverse gradients, such as friction and heat transfer, are formulated in terms of the bulk properties using empirical transfer coefficient formulations. In situations where transverse gradients cannot be represented within the framework of empirical transfer coefficients, such as subcooled boiling, additional models specially developed for the particular situation are employed. The system model is solved numerically using a semi-implicit, finite-difference technique. 6.2.1 Field Equations The NRELAP5 thermal-hydraulic model solves eight field equations for eight primary dependent variables. The primary dependent variables are pressure, phase-specific internal energies, vapor or gas volume fraction, phasic velocities, non-condensable quality, and boron density. For the one-dimensional equations, the independent variables are time and distance. Non-condensable quality is defined as the ratio of the non-condensable gas mass to the total vapor or gas phase mass. The secondary dependent variables used in the equations are phasic densities, phasic temperatures, saturation temperature, and non-condensable mass fraction in the non-condensable gas phase for the ith non-condensable species. The basic field equations for the two-fluid, non-equilibrium model consist of two phasic continuity equations, two phasic momentum equations, and two phasic energy equations. The equations are time averaged and one-dimensional. The phasic continuity equations are shown in Equation 6-1 and Equation 6-2. 1 ( ) + --- ( v A ) = g Equation 6-1 t g g Ax g g g 1 ( f f ) + --- ( f f v f A ) = f Equation 6-2 t A x Continuity consideration yields the interfacial condition of Equation 6-3. f = - g Equation 6-3 The interfacial mass transfer model assumes that total mass transfer can be partitioned into mass transfer at the vapor/liquid interface in the bulk fluid ( ig ) and © Copyright 2022 by NuScale Power, LLC 111

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 mass transfer at the vapor/liquid interface in the thermal boundary layer near the walls ( w ) as defined by Equation 6-4. g = ig + w Equation 6-4 The phasic momentum equations are in the form of Equation 6-5 and Equation 6-6. 2 v 1 v P g g A --------g- + --- g g A --------g = - g A ------ + g g B x A - ( g g A )FWG v g t 2 x x

                                            + g A ( v gI - v g ) - (  g  g A )FIG  ( v g - v f )                                      Equation 6-5

( vg - vf ) v g v f

                                             -C  g  f  m A -----------------------
                                                                                    - + v f --------- - v g ---------

t x x 2 v 1 v P f f A --------f- + --- f f A --------f = - f A ------ + f f B x A - ( f f A )FWF v f t 2 x x

                                            + g A ( v f I - v f ) - (  f  f A )FIF  ( v f - v g )                                     Equation 6-6

( vf - vg ) v f v g

                                            -C  g  f  m A ------------------------
                                                                                    - + v g --------- - v f ---------

t x x The force terms on the right sides of Equation 6-5 and Equation 6-6 are, respectively, the pressure gradient, the body force (i.e., gravity and pump head), wall friction, momentum transfer due to interface mass transfer, interface frictional drag, and force due to virtual mass. The terms FWG and FWF are part of the wall frictional drag, which are linear in velocity, and are products of the friction coefficient, the frictional reference area per unit volume, and the magnitude of the fluid bulk velocity. The coefficients FIG and FIF are part of the interface frictional drag; two different models (drift flux and drag coefficient) are used for the interface friction drag, depending on the flow regime. Conservation of momentum at the interface requires that the force terms associated with interface mass and momentum exchange sum to zero as shown by Equation 6-7. © Copyright 2022 by NuScale Power, LLC 112

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ( vg - vf ) g Av gI ( g g A )FIG ( v g - v f ) - C g f m A ------------------------- t Equation 6-7 ( vf - vg )

                      - g Av f I (  f  f A )FIF  ( v f - v g ) - C  f  g  m A ------------------------- = 0 t

The phasic thermal energy equations are defined by the following two equations: 1 g P

                          ---- (  g  g U g ) + --- ----- (  g  g U g v g A ) = -P --------- - --- ----- (  g v g A )

t A x t A x Equation 6-8

                                             +Q wg + Q ig -  ig h g -  w h + DISS g g

1 P

                       ---- (  f  f U f ) + ---- ----- (  f  f U f v f A ) = -P ---------f - ---- ----- (  f v f A )

t A x t A x Equation 6-9 Q wf + Q if - ig h - w h + DISS f f f In the phasic energy equations, Q wg and Q wf are the phasic wall heat transfer rates per unit volume. These phasic wall heat transfer rates satisfy Equation 6-10 where Q is the total wall heat transfer rate to the fluid per unit volume. Q = Q wg + Q wf Equation 6-10 The vapor generation (or condensation) consists of two parts, vapor generation that results from energy exchange in the bulk fluid ( ig ) and energy exchange in the thermal boundary layer near the wall ( w ) (Equation 6-4). Each of the vapor generation (or condensation) processes involves interface heat transfer effects. The interface heat transfer terms ( Q ig and Q if ) appearing in Equation 6-8 and Equation 6-9 include heat transfer from the fluid states to the interface due to interface energy exchange in the bulk and in the thermal boundary layer near the wall. The vapor generation (or condensation) rates are established from energy balance considerations at the interface. © Copyright 2022 by NuScale Power, LLC 113

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The phasic energy dissipation terms, DISS g and DISS f , are the sums of wall friction, pump, and turbine effects. The dissipation effects due to interface mass transfer, interface friction, and virtual mass are neglected. 6.2.2 State Relations The six-equation model uses five independent state variables with an additional equation for the non-condensable gas component. The independent state variables are chosen to be P , g , U g , U f , and X n . All the remaining thermodynamic fluid variables (temperatures, densities, partial pressures, qualities, etc.) are expressed as functions of these five independent state variables. In addition to these variables, several state derivatives are needed for some of the linearizations used in the numerical scheme. g g f f

                          --------g-             , ----------         ,  ---------         ,  --------- ,  ----------

P U g Xn U Equation 6-11 g PX n X n PU g P f U f U P The interphase mass and heat transfer models use an implicit (linearized) evaluation of the temperature potentials T I - T f and T I - T g . The quantity T I is the temperature that exists at the phase interface. The implicit (linearized) evaluation of the temperature potentials in the numerical scheme requires the derivatives of the phasic and interface temperatures defined by Equation 6-12. T

                                      --------g-               T g                T g                  T f P  g n  U                                                          P  U f g P ,X n  X n P ,U g U   , X Equation 6-12 T f                     s                T     s              T      s T
                                          ---------- ,  ---------- U ,X ,  ----------             ,  ----------

U f P P g n U g P ,X X n P ,U n g 6.2.2.1 Water Property Tables The set of basic properties for light water is used for all LOCA calculations. Implementation is activated by the user. These thermodynamic tables tabulate saturation properties as a function of temperature, saturation properties as a function of pressure, and single-phase properties as a function of pressure and temperature. The tables are based on the 1995 Steam Tables from the International Association for the Properties of Water and Steam (IAPWS) and are known as IAPWS-95. The temperature and pressure range covered in the property table is 273.16 K (32.018 degrees F) to 5000 K (8540.33 degrees F) and 611.6 Pa (0.0887 psia) to 100 MPa (14,504 psia). The properties and derivatives in the tables are saturation pressure, saturation temperature, specific volume ( ), © Copyright 2022 by NuScale Power, LLC 114

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 specific internal energy, specific entropy, and three derivatives: the isobaric thermal expansion coefficient ( ), the isothermal compressibility ( ), and the specific heat at constant pressure ( C p ). 6.2.2.2 Single-Component, Two-Phase Mixture Liquid properties are obtained from the thermodynamic tables, given P and U f . All the desired density and temperature derivatives can then be obtained from the derivatives of f , f , and C pf . In the case of the vapor being subcooled or the liquid being superheated, (i.e., metastable states) the calculation of , T, , , and C p incorporates a constant pressure extrapolation from the saturation state for the temperature and specific volume. 6.2.3 Flow Regime Maps The one-dimensional nature of the field equations for the two-fluid model used in NRELAP5 precludes direct simulation of effects that depend upon transverse gradients of any physical parameter, such as velocity or energy. Consequently, such effects must be accounted for through algebraic terms added to the conservation equations. The mapping for flow conditions to a specific flow regime is required to provide closure to the two-fluid equations. The selected flow regime determines the constitutive relationships that are applied for interphase friction, the coefficient of virtual mass, wall friction, wall heat transfer, and interphase heat and mass transfer. The flow regime maps are based on the work of Taitel and Dukler (Reference 14 and Reference 15) and Ishii (Reference 16, Reference 17, and Reference 18). Taitel and Dukler have simplified flow regime classifications and developed semi-empirical relations to describe flow regime transitions. However, some of their transition criteria are complex, and further simplification was carried out in order to efficiently apply these criteria in NRELAP5. The flow regime maps for the volumes and junctions are identical but used differently as a result of the finite difference scheme and staggered mesh used in the numerical scheme. The volume map is based on volume quantities. It is used for interphase heat and mass transfer, wall friction, and wall heat transfer. Meanwhile, the junction map is based on junction quantities and is used to calculate the interfacial friction coefficient. Three flow-regime maps in both volumes and junctions for two-phase flow are used in the NRELAP5 code: (a) a horizontal map for flow in pipes; (b) a vertical map for flow in pipes, annuli, and bundles; and (c) a high mixing map for flow through pumps. Wall heat transfer depends on the volume flow regime maps in a less direct way. Generally, void fraction and mass flux are used to incorporate the effects of the flow © Copyright 2022 by NuScale Power, LLC 115

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 regime. Since the wall heat transfer is calculated before the hydrodynamics, the flow information is taken from the previous time step. 6.2.3.1 Vertical Volume Flow Regime Maps The vertical volume flow regime map is for upflow, downflow, and counter current flow in volumes whose inclination (vertical) angle is such that 60 < 90 degrees. An interpolation region between vertical and horizontal flow regimes is used for volumes whose absolute value of the inclination (vertical) angle is between 30 and 60 degrees. This map is modeled as nine regimes: four regimes for pre-CHF heat transfer - bubbly, slug, annular-mist, and dispersed (droplet or mist) four regimes for post-CHF heat transfer - inverted annular, inverted slug, mist, and dispersed (droplet or mist) one regime for vertical stratification A schematic of the vertical flow regime map as coded in NRELAP5 is shown in Figure 6-1. The schematic is three-dimensional to illustrate flow-regime transitions as functions of void fraction ( g ), average mixture velocity ( v m ), and boiling. © Copyright 2022 by NuScale Power, LLC 116

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 6-1 Schematic of Vertical Flow-Regime Map Indicating Transitions 6.2.3.2 Junction Flow Regime Maps The junction map is based on both junction and volume quantities. It is used for the interphase drag and shear, as well as the coefficient of virtual mass. The flow regime maps used for junctions are the same as used for the volumes and are based on the work of Taitel and Dukler (Reference 14 and Reference 15), Ishii (Reference 16), and Tandon, et. al. (Reference 19) As with the volumes, three junction flow regime maps are used: horizontal map for flow in pipes vertical map for flow in pipes/bundles high mixing map for flow in pumps The vertical flow regime map is for junctions whose junction inclination (vertical) angle j is such that 60 j 90 degrees. The horizontal flow regime map is for junctions whose junction inclination (vertical) angle j is such that 0 j 30 degrees. An interpolation region between vertical and horizontal flow regimes is used for junctions whose junction inclination (vertical) angle j is such that © Copyright 2022 by NuScale Power, LLC 117

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 30 < j < 60 degrees. This interpolation region is used to smoothly change between vertical and horizontal flow regimes. Junction quantities used in the map decisions are junction phasic velocities, donored (based on phasic velocities) phasic densities, and donored (based on superficial mixture velocity) surface tension. The junction void fraction ( g, j ) is calculated from either of the volume void fractions of the neighboring volumes, g,k or g,L , using a donor direction based on the mixture superficial velocity ( j m ). 6.2.4 Momentum Closure Relations NRELAP5 uses two different models for the phasic interfacial friction force computation, the drift flux method and the drag coefficient method. The choice of which model to use depends upon the flow regime. 6.2.4.1 Drift Flux Model The drift flux approach is used only in the bubbly and slug-flow regimes for vertical flow. The drift flux model specifies the distribution coefficient and the vapor/gas drift velocity. These two quantities must be converted into a constitutive relation for the interfacial frictional force per unit volume. Such a relation can be found by assuming that the interfacial friction force per unit volume is given by Equation 6-13. F i = C i v R v R = f g ( f - g )g Equation 6-13 where the interfacial frictional force per unit volume is balanced by the buoyancy force per unit volume where C i is an unknown coefficient and v R is the relative velocity between the phases. Within the context of the drift flux model, the relative velocity between the phases is not the difference between the phasic velocities but is a weighted difference between the phase velocities given by Equation 6-14. vR = C1 vg - C0 vf Equation 6-14 where C 0 is given by the drift flux correlations and C 1 is given by Equation 6-15. 1 - g C0 C 1 = ---------------------

                                                                           -                           Equation 6-15 1 - g

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Substituting these relations into Equation 6-13 gives the interfacial friction force per unit volume in terms of the phasic velocities, given by Equation 6-16. Fi = Ci C1 vg - C0 vf ( C1 vg - C0 vf ) Equation 6-16 Here the coefficient C i is yet undetermined. The drift flux model also specifies that the relative velocity ( v R ) can be written as the ratio of the vapor/gas drift velocity and the liquid volume fraction, and is given by Equation 6-17. vg j v R = ------- - Equation 6-17 f where the vapor/gas drift velocity ( v g j ) is given by the drift flux correlations. Substituting this value of the relative velocity into Equation 6-13 allows the coefficient C i to be determined from Equation 6-18. 3 g ( f - g )g f C i = ---------------------------------------- Equation 6-18 v g2 j 6.2.4.2 Drag Coefficient Model The drag coefficient approach is used in all flow regimes other than vertical bubbly and slug-flow. The model uses correlations for drag coefficients and for the computation of the interfacial area density. The constitutive relation for the frictional force on a body moving relative to a fluid is given by Equation 6-19. 1 2 F = --- C D A Equation 6-19 2 where, F = drag force

                   = fluid density,
                   = velocity of body relative to the fluid, C D = drag coefficient, and

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 A = projected area of the body. Expressing the frictional force for a group of bodies moving relative to a fluid (e.g., bubbles moving through liquid or droplets moving through vapor/gas) in terms of the frictional force for each body leads to the constitutive relation of Equation 6-20 for the interfacial frictional force per unit volume: 1 F i = --- c g - f ( g - f )C D S F a gf = C i g - f ( g - f ) Equation 6-20 8 where, F i = interfacial friction force per unit volume, 1 C i = --- c C D S F a gf 8 c = density of continuous phase a gf = interfacial area per unit volume, and S F = shape factor. The additional factor of 1/4 comes from the conversion of the projected area of 2 2 spherical particles (i.e., r ) into the interfacial area (i.e., 4r ) and the shape factor is included to account for non-spherical particles. The drag coefficient model for the global interfacial friction coefficient was reduced to the specification of the continuous density, drag coefficient, interfacial area density, and shape factor for the flow regimes. Once these quantities have been computed, the interfacial friction force per unit volume ( F i ) is computed from Equation 6-20 from which the global interfacial friction coefficient can be computed. 6.2.4.3 Wall Friction The wall friction is determined based on the volume flow regime map. The wall friction force terms include only wall shear effects. Losses due to abrupt area change are calculated using mechanistic form-loss models. Other losses due to elbows or complicated flow passage geometry are modeled using energy-loss coefficients that must be input by the user. The semi-implicit scheme, one-dimensional, finite difference equations for the sum momentum equation and the difference momentum equation contain the terms of Equation 6-21 that represent the phasic wall frictional pressure drop. © Copyright 2022 by NuScale Power, LLC 120

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 n n+1 n n+1 FWG ( g ) x j t and FWF ( f ) x j t Equation 6-21 j j j j These terms represent the pressure loss due to wall shear from cell center to cell center of the cell volumes adjoining the particular junction that the momentum equation is considering. The wall drag or friction depends not only on the phase of the fluid, but also on the flow regime characteristics. The wall friction model is based on a two-phase multiplier approach in which the two-phase multiplier is calculated from the heat transfer and fluid flow service (HTFS) modified Baroczy correlation. The individual phasic wall friction components are calculated by apportioning the two-phase friction between the phases using a technique derived from the Lockhart-Martinelli model (Reference 20). The model is based on the assumption that the frictional pressure drop may be calculated using a quasi-steady form of the momentum equation, as used by Chisholm. This wall friction partitioning model is used with the drag coefficient method of the interphase friction model. The Lockhart-Martinelli model computes the overall two-phase friction pressure drop in terms of the liquid-alone and vapor/gas-alone wall friction pressure drop as shown in Equation 6-22. dP 2 dP 2 dP

                                                           =   ------- =   -------                               Equation 6-22 dx  2             f  dx  f        g  dx  g Here  f and  g are the liquid-alone and vapor/gas-alone two-phase Darcy-Weisbach friction multipliers, respectively. The phasic wall friction pressure gradients are expressed by Equation 6-23 for the liquid and vapor/gas alone.

f Re f M f2 g Re g M f2 dP dP

                                  ------- = ------------------------ and ------- = -------------------------

dx f 2 dx g Equation 6-23 2 2D f A 2D g A Here the prime indicates the liquid and vapor/gas-alone Darcy-Weisbach friction factors, respectively, calculated at the respective Reynolds numbers given by Equation 6-24. Mf D Mg D Re = ------------ and Re = ----------- - Equation 6-24 f f A g g A The liquid and vapor/gas mass flow rates, respectively, are defined by Equation 6-25. M f = f f f A and M g = g g g A Equation 6-25 © Copyright 2022 by NuScale Power, LLC 121

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The overall two-phase friction pressure gradient is calculated using two-phase friction multiplier correlations. The multipliers are interrelated using Equation 6-22 and Equation 6-23 and the Lockhart-Martinelli ratio defined by Equation 6-26. dP

                                                               -------               2 dx                 
                                                        = --------------f- = -----g-2 Equation 6-26 2

dP dx g f The HTFS correlation is used to calculate the two-phase friction multipliers. This correlation was chosen because it is correlated to empirical data over broad ranges of phasic volume fractions, phasic flow rates and phasic flow regimes. The correlation was also shown to give good agreement with empirical data. The HTFS correlation for the two-phase friction multiplier is expressed with Equation 6-27. 2

                                         = 1 + - + -----

1 2 2 and = + C + 1 Equation 6-27 f 2 g C is the correlation coefficient and is the Lockhart-Martinelli ratio given by Equation 6-26. If the HTFS correlation is combined with the wall friction formulations by combining Equation 6-22 and Equation 6-23, Equation 6-25 and Equation 6-26, and Equation 6-27, then the combined two-friction pressure drop is expressed by Equation 6-28. dP 2 dP 2 dP 1

                                        =   ------- =   ------- = -------    f (  f  f ) +

2 dx 2 g dx f g dx f 2D f Equation 6-28 1--- 2 2 2 2 C f ( f f ) g g ( g g ) + g g ( g g ) f The phasic wall friction coefficients are defined by Equation 6-29 and Equation 6-30. 2 dP Z FWF ( f f f )A = f p f = f ------ - -------------------------A Equation 6-29 dx 2 + Z 2 g f dP 1 FWG ( G G G )A = g p g = g ------- ------------------------- A Equation 6-30 dx 2 + Z 2 g f © Copyright 2022 by NuScale Power, LLC 122

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Here Z is defined by Equation 6-31. 2 fw f Re f f ------- - 2 f f Z = -------------------------------------- Equation 6-31 2 gw g Re g g -------- - g g Taking the sum of these two equations gives the overall quasi-static, two-phase wall friction pressure gradient as shown by Equation 6-32. dP FWF ( f f f )A + FWG ( g g g )A = ------- A Equation 6-32 dx 2 The phasic friction factors used in the wall friction model are computed from correlations for laminar and turbulent flows with interpolation in the transition regime. The friction factor model is simply an interpolation scheme linking the laminar, laminar-turbulent transition, and turbulent flow regimes. The laminar friction factor is calculated by Equation 6-33. 64 L = ------------ for 0 Re 2,200 Equation 6-33 Re s Here s is a user-input shape factor for non-circular flow channels ( s is 1.0 for circular channels). The friction factor in the transition region between laminar and turbulent flows is computed by reciprocal interpolation with Equation 6-34. L,T = 3.75 - 8,250

                                                  ------------- (  T,300 -  L,2200 ) +  L,2200 Re Equation 6-34 for 2,200  Re  3,000 Here  L,2200 is the laminar factor at a Reynolds number of 2,200,  T,3000 is the turbulent friction factor at a Reynolds number of 3,000, and the interpolation factor is defined to lie between zero and one.

The turbulent friction factor is given by the Zigrang-Sylvester approximation (Reference 22) to the Colebrook-White correlation (Reference 23) with Equation 6-35, where is the surface roughness. © Copyright 2022 by NuScale Power, LLC 123

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 1 - = -2log ----------- 2.51

                                                          ----------                      21.25- 
                                                                                  - + ------------
                         ---------          10  3.7D- + Re 1.14 - 2log 10  ---   D Re 0.9 T

Equation 6-35 for Re 3,000 6.2.5 Heat Transfer The liquid and vapor/gas energy solutions include the wall heat flux to liquid or vapor/gas. During boiling, the saturation temperature based on the total pressure is the reference temperature, and during condensation the saturation temperature based on the partial pressure is the reference temperature. The general expression for the total wall heat flux is defined by Equation 6-36: q = h wgg ( T W - T g ) + h wgspt ( T W - T spt ) + h wgspp ( T W - T spp ) total Equation 6-36

                                         +h wff ( T W - T f ) + h wfspt ( T W - T spt )

where, h wgg = heat transfer coefficient to vapor/gas, with the vapor/gas temperature as the reference temperature (W/m2 K), h wgspt = heat transfer coefficient to vapor/gas, with the saturation temperature based on the total pressure as the reference temperature (W/m2 K), h wgspp = heat transfer coefficient to vapor/gas, with the saturation temperature based on the vapor partial pressure as the reference temperature (W/m2 K), h wff = heat transfer coefficient to liquid, with the liquid temperature as the reference temperature (W/m2 K), h wfspt = heat transfer coefficient to liquid, with the saturation temperature based on the total pressure as the reference temperature (W/m2 K), T W = wall surface temperature (K), T g = vapor/gas temperature (K), T f = liquid temperature (K), © Copyright 2022 by NuScale Power, LLC 124

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 T spt = saturation temperature based on the total pressure (K), and T spp = saturation temperature based on the partial pressure of vapor in the bulk (K). A boiling curve is used in NRELAP5 to govern the selection of the wall heat transfer correlations when the wall surface temperature is above the saturation temperature (superheated relative to the saturation temperature based on total pressure). When a hydraulic volume is voided and the adjacent surface temperature is subcooled, vapor condensation on the surface is predicted. If non-condensable gases are present, the phenomena are more complex because condensation is based on the partial pressure of vapors present in the region. When the wall temperature is less than the saturation temperature based on total pressure, but greater than the saturation temperature based on vapor partial pressure, a convection condition exists. Figure 6-2 illustrates these three regions of the curve. Figure 6-2 NRELAP5 Boiling and Condensing Curves

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                >7VSS 7Z @                           >7Z7VSW @
    &RQGHQVLQJUHJLRQ
                                           &RQYHFWLRQUHJLRQ The boiling curve uses the Chen boiling correlation (Reference 24) up to the CHF point.

NRELAP5 issues a message and stops running if CHFR reduces below one for heat conductors that are in the core. Post-CHF heat transfer is allowed to occur on surfaces outside the core. For instance, post-CHF heat transfer can occur on the © Copyright 2022 by NuScale Power, LLC 125

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 outside of the RPV where boiling occurs in the pool of liquid that accumulates in the CNV. Post-CHF heat transfer could also occur on the SG tube surfaces, depending on local conditions. 6.3 Heat Structure Models Heat structures provided in NRELAP5 permit calculation of the heat transfer across solid boundaries of hydrodynamic volumes. Modeling capabilities of heat structures are general and include fuel pins or plates with nuclear or electrical heating, heat transfer across SG tubes, and heat transfer from pipe and vessel walls. Temperatures and heat transfer rates are computed from the one-dimensional form of the transient heat conduction equation. Heat structures are represented using rectangular, cylindrical, or spherical geometry. Surface multipliers are used to convert the unit surface of the one-dimensional calculation to the actual surface of the heat structure. Temperature-dependent and space-dependent thermal conductivities and volumetric heat capacities are provided in tabular or functional form either from built-in or user-supplied data. Finite differences are used to advance the heat conduction solutions. Each mesh interval may contain different mesh spacing, a different material, or both. The spatial dependence of the internal heat source, if any, may vary over each mesh interval. The time-dependence of the heat source can be obtained from reactor kinetics, one of several tables of power versus time, or a control system variable. Boundary conditions include symmetry or insulated conditions; a heat transfer correlation package; and tables of surface temperature versus time, heat flux versus time, heat transfer coefficient versus time, and heat transfer coefficient versus surface temperature. The heat transfer correlation package can be used for heat structure surfaces connected to hydrodynamic volumes. The heat transfer correlation package contains correlations for convective, nucleate boiling, transition boiling, and film boiling heat transfer from the wall to the fluid, and it contains reverse heat transfer from the fluid to the wall including correlations for condensation (Section 6.2.5 and Section 6.8). The heat conduction model also includes a gap conduction model and a radiation enclosure model. The integral form of the heat conduction equation is defined by Equation 6-37. T Cp ( T,x ) ----- t

                                          - ( x,t ) dV = k ( t,x )T ( t,x )  ds +

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 S = internal volumetric heat source, t = time, T = temperature, V = volume, x = space coordinates, and C p = volumetric heat capacity. The boundary conditions applied to the exterior surface have the form of Equation 6-38. T ( t ) A ( T )T ( t ) + B ( T ) ------------- = D ( T,t ) Equation 6-38 n The n denotes the unit normal vector away from the boundary surface. Thus, if the desired boundary condition is that the heat transferred out of the surface equals a heat transfer coefficient ( h ) times the difference between the surface temperature ( T ) and the sink temperature ( T sk ) as shown by Equation 6-39. T ( t )

                                                  -k ------------- = h ( T - T sk )                            Equation 6-39 n

then the correspondence between the above expression and Equation 6-38 yields A = h , B = k , and D = hT sk . One-dimensional heat conduction in rectangular, cylindrical, and spherical geometry can be used to represent the heat structures in any of the components in NRELAP5. The equations governing one-dimensional heat conduction are defined by Equation 6-40, Equation 6-41, and Equation 6-42. T T C P ------ = ----- k ------ + S for rectangular geometry Equation 6-40 t x x T 1 T C P ------ = --- ----- rk ------ + S for cylindrical geometry Equation 6-41 t r r r T 1 2 T C P ------ = --- ----- r k ------ + S for spherical geometry Equation 6-42 t r r r © Copyright 2022 by NuScale Power, LLC 127

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Heat may flow across the external heat structure boundaries to either the environment or to the working fluid. For heat structure surfaces connected to hydrodynamic volumes containing the working fluid, a heat transfer package is provided containing correlations for heat transfer from wall-to-liquid and reverse heat transfer from liquid-to-wall. Any number of heat structures may be connected to each hydrodynamic volume, or heat transfer coefficient versus surface temperature can be used to simulate the boundary conditions. The heat conduction equation can be solved by various numerical techniques. NRELAP5 uses the Crank-Nicolson method (Reference 26) for solving this equation. 6.4 Point Reactor Kinetics Model NRELAP5 allows the user to model the power generated in the reactor core as specified from a table or as determined by point-reactor kinetics with reactivity feedback. This power is modeled as an internal heat source in user-defined heat structures and can be partitioned by inputting weighting factors to distribute the energy to the various portions of the core as the user desires. The power is computed using the space-independent or point kinetics approximation, which assumes that power can be separated into space and time functions. The point reactor kinetics model computes both the immediate (prompt and delayed neutrons) fission power and the power from decay of fission products. The immediate (prompt and delayed neutrons) power is that released at the time of fission and includes power from kinetic energy of the fission products and neutron moderation. Decay power is generated as the fission products undergo radioactive decay. The LOCA methodology uses the ANS 1973 decay heat standard (Section 6.10). The point kinetics equations are (Glasstone and Sesonske, Reference 27) defined by Nd dn ( t ) [(t) - ]

                                 ------------ = ------------------------ n ( t ) +

dt i Ci ( t ) + S Equation 6-43 i=1 dC i ( t ) f i

                              --------------- = -------- n ( t ) -  i C i ( t ) i=1, 2, ..., N d                          Equation 6-44 dt (t) = n(t)                                                     Equation 6-45

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 where, t = time (s), n = neutron density (neutrons/m3),

         = neutron flux (neutrons/m2*s),
         = neutron velocity (m/s),

C i = delayed neutron precursor concentration in group i (nuclei/m3), Nd

         = effective delayed neutron fraction =   i ,

i=1

         = prompt neutron generation time (s)
         = reactivity (only the time-dependence has been indicated; however, the reactivity is dependent on other variables),

f i = fraction of delayed neutrons of group i = i i = effective delayed neutron precursor yield of group i , i = decay constant of group i (1/s), S = source rate density (neutrons/m3*s),

         = fission rate (fissions/s),

f = macroscopic fission cross-section (1/m), P f = immediate (prompt and delayed neutron) fission power (MeV/s), Q f = immediate (prompt and delayed neutron) fission energy per fission (MeV/fission), V = volume (m3), and N d = number of delayed neutron precursor groups. © Copyright 2022 by NuScale Power, LLC 129

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 After some modifications and variable substitutions, these equations are solved by the modified Runge-Kutta method. Reactivity feedback can be input into NRELAP5 in one of two models: a separable model and a tabular model. The separable model is so defined that it assumes that each effect is independent of the other effects. This model also assumes nonlinear feedback effects from moderator density and fuel temperature changes and linear feedback from moderator temperature changes. 6.5 Trips and Control System Models The control system provides the capability to evaluate simultaneous algebraic and ordinary differential equations. The capability is primarily intended to simulate control systems typically used in hydrodynamic systems, but it can also model other phenomena described by algebraic and ordinary differential equations. Another use is to define auxiliary output quantities, such as differential pressures, so they can be printed in major and minor edits and be plotted. The control system consists of several types of control components. Each component defines a control variable as a specific function of time-advanced quantities. The time-advanced quantities include hydrodynamic volume, junction, pump, valve, heat structure, reactor kinetics, trip quantities, and the control variables themselves (including the control variable being defined). This permits control variables to be developed from components that perform simple, basic operations. The trip system consists of the evaluation of logical statements. Each trip statement is a simple logical statement that has a true or false result and an associated variable. Two types of trip statements are provided (variable and logical trips). 6.6 Special Solution Techniques Certain models in NRELAP5 have been developed to simulate special processes. Special process models are used in NRELAP5 to model those processes, which are sufficiently complex that they must be modeled by empirical models. The following sections summarize those models. 6.6.1 Choked Flow 6.6.1.1 Moody Critical Flow Model Because the Moody model (Reference 28) is required by 10 CFR 50 Appendix K when the break flow is calculated to be two-phase, a critical flow model that complies with the 10 CFR 50 Appendix K requirements was incorporated in NRELAP5. Two options are available in NRELAP5 for use of the Moody model. ((

                                                                                                      }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                      }}2(a),(c)

Moody developed his critical flow model from theory to predict the maximum flow rate of a single component, two-phase mixture. The model assumes that the liquid phase is incompressible and that the flow is isentropic so that the stagnation enthalpy is constant throughout the system. The flow is maximized with respect to local slip ratio and static pressure for known stagnation conditions. The specific volume ( v ) and specific enthalpy (h) of water can be calculated from two state variables, entropy (s) and pressure (P), i.e., h = h ( s 0 ,P ) and v = v ( s 0 ,P ) , where the subscript 0 denotes break entrance conditions. Because entropy is constant, h and v are functions of P, the stagnation pressure. From the continuity and energy equations for homogeneous flow entering and leaving an ideal nozzle the mass flux, G, satisfies: 2 12 G = [ 2 ( h0 - h ) v ] Equation 6-48 The maximum flow rate occurs at the throat, where dG

                                                   ------- = 0                                        Equation 6-49 dP t 13 Moody showed that the maximum flow occurs when the slip ratio K = ( v g  v f )                       .

With this value of the slip ratio, Moody derived a complex equation for the critical flow rate that was used to create Moody lookup tables for the flow rate as a function of stagnation pressure and stagnation enthalpy. The range of the tables that are used in NRELAP5 covers local static pressure from ((

                                                          }}2(a),(c) with local quality from 0.0 to 1.0 and local stagnation pressures and enthalpies covering the range of saturation states.

((

                                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                     }}2(a),(c) 6.6.1.2            Henry-Fauske Critical Flow Model The principle assumption used in the Henry-Fauske model is that, for most applications, the amount of thermal non-equilibrium at the throat is more important in determining the critical flow rate than the amount of mechanical non-equilibrium. Thus, it is assumed that the phase velocities are equal. Henry and Fauske then argued that for normal nozzle configurations, there is little time for mass transfer to take place, and it is reasonable to assume that the amount of mass transferred during the expansion is negligible and also that the amount of heat transferred between the phases during the expansion is insignificant, so that the liquid temperature is essentially constant. Interfacial viscous terms were neglected. Based on these assumptions Henry and Fauske derived an equation for the mass flux at the throat. The mass flow rate exhibits a maximum with respect to the throat pressure at critical flow, which yields a complex relationship for the critical mass flux that includes dependency on the throat pressure.

The Henry-Fauske model requires only knowledge of the upstream stagnation conditions and, unlike earlier critical flow models, it accounts for the non-equilibrium nature of the flow. Henry and Fauske noted that the critical flow rates are in reasonable agreement with the homogeneous equilibrium model for stagnation qualities greater than 0.10, and that for qualities less than this value, the homogeneous equilibrium model underestimates the data. Therefore, they required that the model input use only stagnation conditions, and yet at the same time account for the non-equilibrium nature of the flow. To address this issue Henry and Fauske correlated the effect of thermal non-equilibrium on the mass transfer rate at the throat as: dX dX eq

                                                  ------- = N ------------                            Equation 6-50 dP t             dP t where N is a thermal non-equilibrium factor defined in terms of the equilibrium quality at the throat ( X eq,t ).

X eq,t N = ----------

                                                                      -                               Equation 6-51 0.14 The final remaining unknown is the value of the pressure at the throat. To determine the throat pressure, the two-phase momentum equation was integrated

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 between the stagnation and the throat locations to give an equation for the critical pressure ratio, i.e., the ratio of the throat pressure to the upstream stagnation pressure when the flow is choked. The use of this equation for the throat pressure in the equation for the critical mass flux results in a transcendental equation for the critical mass flux. The solution of the transcendental equation implicitly involves the critical flow rate and hence its solution yields predictions of the critical pressure ratio and the critical flow rate as functions of the upstream stagnation pressure and quality. The critical pressure ratio determines the transition to non-choked flow. If the mass flux predicted by the critical flow model is less than that resulting from the normal solution of the momentum equations, then the junction is choked. Assessment of the Henry-Fauske model shows excellent agreement against the Marviken 22 and 24 tests (Section 7.2.11). 6.6.1.3 Choked Flow for Orifices, Nozzles and Valves To provide the user with the ability to better characterize the orifice, nozzle or valve behavior, the form of the Henry-Fauske model was retained in RELAP5-3D© and carried over into NRELAP5. The constant in the thermal non-equilibrium factor is included as an adjustable parameter. X eq,t N = MIN 1, ----------- Equation 6-52 C ne where the thermal non-equilibrium constant, C ne , is user input with a default value of 0.14. A user input discharge coefficient, default value of 1.0, can also be applied to the critical mass flux. The ability to input these parameters allows the user to adjust the critical flow model to account for the different amount of thermal non-equilibrium at the throat. While the model development was based on a converging nozzle, the authors of the model extended the results to orifices and short tubes by comparison to experimental data for these geometries. The Henry-Fauske model can be applied to cases where the upstream condition is subcooled liquid or single-phase vapor. While the Henry-Fauske model can handle non-condensable gas, the total amount of non-condensable gas in the NPM is negligible so this capability is not addressed in the following discussion. During the development of the RELAP5 codes, modifications were made to the original model to ensure continuity at phase transitions to better characterize nozzles and orifices. Specifically, the phase transition modifications provide a smooth transition of the critical flow at the subcooled liquid to two-phase mixture interface. two adjustable coefficients, a discharge coefficient and a thermal non-equilibrium constant are provided in order to better characterize nozzles and orifices. The discharge coefficient is a multiplier on the flow area. The © Copyright 2022 by NuScale Power, LLC 133

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 non-equilibrium constant is an assumed throat equilibrium quality that was assigned an average value of 0.14 by Henry and Fauske, but can be specified by the code user. With these modifications, the Henry-Fauske model is applicable to two-phase and single-phase superheated and subcooled critical flow. The two adjustable coefficients allow the code user to more closely match test data from the valve vendor and calibration data from orifices and nozzles used in experimental facilities. 6.6.2 Abrupt Area Change The general reactor system contains piping networks with many sudden area changes and orifices. To apply the NRELAP5 hydrodynamic model to such systems, analytical models for these components are included in the code. The basic hydrodynamic model is formulated for slowly varying (continuous) flow area variations; therefore, special models are not required for this case. The abrupt area change model is based on the Borda-Carnot formulation (Reference 30) for a sudden (i.e., sharp, blunt) enlargement and standard pipe flow relations, including the vena-contracta effect for a sudden (i.e., sharp, blunt) contraction or sharp-edge orifice or both. This is referred to as the full abrupt area change model. It does not include the case where an enlargement, contraction, or orifice is rounded or beveled. Quasi-steady continuity and momentum balances are employed at points of an abrupt area change. The numerical implementation of these balances is such that hydrodynamic losses are independent of upstream and downstream nodalization. In effect, the quasi-steady balances are employed as jump conditions that couple fluid components having abrupt changes in cross-sectional area. This coupling process is achieved without change to the basic numerical time-advancement schemes. The basic assumption used for the transient calculation of two-phase flow in flow passages with points of abrupt area change is that the transient flow process can be approximated as a quasi-steady flow process that is instantaneously satisfied by the upstream and downstream conditions (that is, transient inertia, mass, and energy storage are neglected at abrupt area changes). However, the upstream and downstream flows are treated as fully transient flows. The volume of fluid and associated mass, energy, and inertia at points of abrupt area change is generally small compared with the volume of upstream and downstream fluid components. The transient mass, energy, and inertia effects are approximated by lumping them into upstream and downstream flow volumes. Finally, the quasi-steady approach is consistent with modeling of other important phenomena in transient codes (i.e., heat transfer, pumps, and valves). Activation of the full abrupt area change model in NRELAP5 results in the code internally calculating the form and interfacial losses across a junction. Utilization of the © Copyright 2022 by NuScale Power, LLC 134

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 partial area change model allows the user to specify the form loss while allowing the code to internally calculate the interfacial loss. Activation of the smooth area change model allows the user to specify the form loss with no internal calculation of the interfacial losses. More detailed discussion concerning this model can be found in the NRELAP5 theory manual (Reference 9). 6.6.3 Counter Current Flow Limitation A general CCFL model is implemented in a form proposed by Bankoff (Reference 31), which has the structure 12 12 H + mH = c Equation 6-53 g f where, H g = dimensionless vapor/gas flux, H f = dimensionless liquid flux, 12 c = vapor/gas intercept (value of H g when H f =0, i.e., complete flooding), and m = "slope", that is the vapor or gas intercept divided by the liquid intercept (the value 12 of H when H g =0). f The dimensionless fluxes have the form as defined by Equation 6-54 and Equation 6-55. 12 g H g = j g ------------------------------ Equation 6-54 gw ( f - g ) 12 f H f = j f ------------------------------ Equation 6-55 gw ( f - g ) In these equations j g is the vapor/gas superficial velocity ( g g ), j f is the liquid superficial velocity ( f f ), g is the vapor/gas density, f is the liquid density, g is the vapor/gas volume fraction, f is the liquid volume fraction, g is the gravitational © Copyright 2022 by NuScale Power, LLC 135

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 acceleration. In Equation 6-54 and Equation 6-55, w is the length scale and is given by Equation 6-56. 1-w = D L Equation 6-56 j Where is a user-input constant, D j is the junction hydraulic diameter and L is the Laplace capillary length constant given by Equation 6-57. 12 L = -------------------------- Equation 6-57 g ( f - g ) Bankoff recommends a formula for computing the value of : A 2

                                           = tanh  -----h-  ------  D j                                        Equation 6-58 At tp where A h is the total area of the holes through the plate, A t is the total area of the plate, including the holes, and t p is the thickness of the plate.

The Bankoff correlation specifies that the vapor/gas intercept (c) is of the form: B c = 1.07 + 0.00433 L Equation 6-59 B B when the dimensionless Bond number, L , is less than 200 and c=2 for all L greater than or equal to 200. The bond number is: B g ( f - g ) 1 2 nD L = nD -------------------------- = ------------j Equation 6-60 L where n is the number of holes in the plate at the CCFL junction, D j is the hydraulic diameter, and L is the Laplace capillary length constant, previously defined. More detailed discussion concerning this model can be found in the NRELAP5 theory manual (Reference 9). Assessment of the CCFL model demonstrates excellent agreement against Bankoff perforated plate test data (Section 7.2.10). A sensitivity study of the effects of the CCFL model is presented in Section 9.6.3 for its application at the NPM pressurizer baffle plate. © Copyright 2022 by NuScale Power, LLC 136

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 6.7 Helical Coil Steam Generator Component A new hydrodynamic component and heat transfer package have been added to NRELAP5 to model flow and heat transfer inside a helical coil SG. These are developed based on helical coil geometry-specific heat transfer and wall friction correlations. The need for improved models is based on inadequate agreement with pressure drop and heat transfer performance with the baseline RELAP5-3D© code results against prototypic helical coil SG testing performed at SIET. Improvements and adequacy of the implemented models in NRELAP5 are demonstrated through prototypic assessments of the NuScale helical coil SG using SIET test data (Section 7.4.1 and Section 7.4.2). These tests assessed heat transfer and pressure drop on both the secondary side (within tubes) and primary side (external to tubes) of the helical coil SG. A wide range of pressure drop and heat transfer correlations were investigated for analyzing the inside of the helical coils. A down selection was performed of these investigated models for implementation into the NRELAP5 code based on the applicability of the models to the NPM helical coil SG. 6.7.1 Helical Coil Tube Friction 6.7.1.1 Helical Coil Single-Phase Tube Wall Friction The ((

                                      }}2(a),(c) provided the best global coverage and as such have been implemented into NRELAP5. ((
                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                                                        }}2(a),(c) 6.7.1.2            Helical Coil Two-Phase Tube Wall Friction The two-phase inner wall friction for a helical coil is computed in a similar fashion to the Lockhart-Martinelli model implemented in the RELAP5 code series. A modification is made to the two-phase friction multiplier for the helical coil component as presented in its final form by Equation 6-62.

((

                                                                                                        }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                   }}2(a),(c) 6.7.2         Helical Coil Tube Heat Transfer A new heat transfer package has been added to NRELAP5 and differs from that of the standard RELAP5 pipe geometry in ((
                                              }}2(a),(c) A new geometry type represents the ((
                       }}2(a),(c) 6.7.2.1            Helical Coil Single-Phase Heat Transfer The laminar heat transfer correlation ((
                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                     }}2(a),(c) 6.7.2.2            Helical Coil Two-Phase Subcooled and Saturated Flow Boiling Heat Transfer The saturated flow boiling heat transfer correlation is used for ((
                                                                                         }}2(a),(c) 6.8      Wall Heat Transfer and Condensation Due to the significance of CNV wall heat transfer in reactor core cooling and decay heat removal during a postulated NPM LOCA (Section 8.2.8), a detailed discussion is presented in this section on NRELAP5 wall heat transfer and condensation models.

((

                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 As described below, the ((

                             }}2(a),(c)

Section 6.8.1 below provides further discussion on NRELAP5 evaluation of wall heat transfer with film condensation. The discussion includes the definition of the liquid (film) Reynolds number, partitioning of the total wall heat flux between liquid and vapor phases, and handling the effect of non-condensable gases that may be present in the hydrodynamic volume. Section 6.8.2 summarizes the extended Shah correlation used in NRELAP5 for wall condensation. 6.8.1 NRELAP5 Solution Approach for Wall Condensation Heat Transfer NRELAP5 solves ((

                                                                           }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 6-3 (( }}2(a),(c) ((

                                                                       }}2(a),(c)

((

                                                                   }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                      }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                   }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                              }}2(a),(c) 6.8.2         Wall Condensation Correlation

((

                                                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                      }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                         }}2(a),(c)

Table 6-1 Extended Shah Dimensionless Vapor Velocity Transition Criteria ((

                                                                                                                     }}2(a),(c)

Table 6-2 Extended Shah Condensation Heat Transfer Coefficients Dependent on Regime ((

                                                                                                                     }}2(a),(c) 6.9      Interfacial Drag in Large Diameter Pipes RELAP5-3D© contains the Kataoka-Ishii (Reference 40) formulation of the drift-flux model for use in the bubbly flow case in intermediate ( 0.018 < D  0.08m ) and large pipes

( D > 0.08m ). This same dimensional formulation is maintained within NRELAP5. RELAP5-3D© originally implemented the modified Rouhani distribution coefficient (Reference 42) as shown by Equation 6-87. f gD h C = 1 + 0.2 --------------------------

                                                                             -                            Equation 6-93 G + 0.001

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                                                    }}2(a),(c) 6.10     Fission Decay Heat and Actinide Models The ANS 1973 fission decay heat standard (Reference 46) is presented in terms of the Shure curve (Reference 47) and tabular data. The NRELAP5 implementation of the ANS 1973 standard applies the Shure curve, which is a fit to differential equations for one isotope and 11 groups. Assuming infinite operating time, the fission product decay power is calculated with Equation 6-89. Table 6-3 provides the 11-group constants derived from the Shure curve as implemented into NRELAP5. Figure 6-3 provides the comparison of the ANS 1973 standard to the as implemented curve.

11 P = Po An exp ( -an ts ) Equation 6-95 n=1 where, P = fission decay power, P 0 = infinite operating time fission power prior to shutdown,

         = fission product yield factor, A n , a n = fit coefficients, and t s = time after shutdown.

Table 6-3 ANS 1973 11-Group Fission Decay Constants ((

                                                                                                     }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 6-3 ANS 1973 11-Group Fission Decay Constants (Continued) ((

                                                                                                  }}2(a),(c)

Figure 6-4 NRELAP5 ANS 1973 Implemented Fission Decay Heat Curve The actinide model describes the production of 239U, 239Np, and 239Pu from neutron capture by 238U based on the decay equations of Equation 6-90. © Copyright 2022 by NuScale Power, LLC 149

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 d U ( t )

                                                         - = FU  ( t ) - U U dt d N ( t )
                                              ---------------- =  U  U -  N  N                       Equation 6-96 dt P ( t ) = U U U ( t ) + N N N ( t )

The quantity F U is user-specified and is the number of atoms of 239U produced by neutron capture in 238U per fission from all isotopes. The and values can be user-specified, or default values equal to those stated in the 1979 ANS standard (Table 6-4), the 1994 Standard, or the 2005 Standard can be used. The first equation describes the rate of change of atoms of 239U. The first term on the right represents the production of 239U; the last term is the loss of 239U due to beta decay. The second equation describes the rate of change of 239Np. The production of 239Np is from the beta decay of 239U, and 239Pu is formed from the decay of 239Np. ( t ) is the solution from the NRELAP5 fission source. The implemented model yields the result quoted in the 1979 Standard (Reference 48), the 1994 Standard (Reference 49), and the 2005 Standard (Reference 50) as demonstrated by Figure 6-4. Table 6-4 ANS-79 Actinide Model Constants Isotope (s-1) (MeV) 239 1.772 0.00299 U 239 0.5774 0.00825 Np © Copyright 2022 by NuScale Power, LLC 150

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 6-5 NRELAP5 ANS-79 Implemented Actinide Heat Curve 6.11 Critical Heat Flux Models The CHF is calculated using a combination of the ((

                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                           }}2(a),(c) 6.11.1        ((                       }}2(a),(c)

((

                                                                                      }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 6-6 (( }}2(a),(c) ((

                                                                                  }}2(a),(c)

((

                                      }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                             }}2(a),(c) 6.11.2        Implementation of Critical Heat Flux correlations

(( }}2(a),(c) are implemented in NRELAP5 as follows: ((

                                                  }}2(a),(c) 6.11.3        ((                                    }}2(a),(c)

((

                                                                                        }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                      }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                              }}2(a),(c)

Table 6-5 Coefficient of Revised Pressure Correction Term in Equation 6-108 ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 6-7 ((

                                      }}2(a),(c)

((

                                                                                  }}2(a),(c)

((

                                                        }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 6-6 (( }}2(a),(c) Critical Heat Flux Correlation Application Range ((

                                                                                                     }}2(a),(c)

((

                              }}2(a),(c) 6.11.4        ((                                                 }}2(a),(c)

((

                                                                                     }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                  }}2(a),(c) 6.11.5        NSPN-1 Critical Heat Flux Correlation As discussed in Section 6.11, the DSM and HBM are two approaches that have been traditionally used when evaluating a critical heat flux condition. For Phase 0 conditions where there is potential for a significant axial variation in properties (e.g.,

mass flux, quality) a local condition DSM approach is more appropriate. This is because a HBM model typically utilizes global conditions making assumptions on the flow field in defining an implementable solution (e.g., 1-Dimensional quasi-steady state). Because of these differences and to remove such assumptions a CHF correlation is developed for NRELAP5 based on local conditions utilizing the © Copyright 2022 by NuScale Power, LLC 159

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 approach defined in NuScale Power Critical Heat Flux Correlations, TR-0116-21012-P-A (Reference 115). The NSPN-1 CHF correlation is developed from the design specific CHF testing for the NuFuel-HTP2' fuel design performed at AREVA's Karlstein, Germany, thermal-hydraulic (KATHY) test facility (Reference 115). This section describes the tests, test facilities, statistical methods, and final validation for the NSPN-1 CHF correlation. The NSPN-1 CHF correlation is an empirical based correlation that accounts for local fluid conditions, fuel characteristics, and heat flux. The model includes a non-uniform flux factor (F-factor) that accommodates variability in axial power profiles. The NSPN-1 CHF correlation conservatively predicts CHF for the NuFuel-HTP2' application when used in conjunction with the applicable statistically derived correlation limit. Application of a CHFR safety limit determined with the NSPN-1 CHF correlation ensures with a 95 percent probability at the 95 percent confidence level (95/95 level), that the hot fuel rod in the core does not experience CHF. The NSPN-1 CHF correlation is used in safety analysis evaluations of designs that utilize the NuFuel-HTP2' fuel design with local conditions calculated by the NRELAP5 system thermal-hydraulic code. 6.11.5.1 NSPN-1 Critical Heat Flux Correlation Model The CHF correlation form for determining the uniform critical heat flux applied here (Equation 6-114) is consistent with the form utilized for the NSP4 subchannel CHF correlation developed in Reference 115. In the development of the final expression, the following independent variables were considered: pressure (term A) local mass flux (term B) local equilibrium quality (term C) cold wall factor (term D) boiling length (term E) The empirical expression defined by Equation 6-114 is initially used to correlate the local conditions to determine the local uniform CHF value. The statistically insignificant correlation coefficients are identified by an iterative process and removed to derive the final form of the uniform-CHF model. Using the additive method for linear least squares (LLS), the terms in Equation 6-114 are evaluated for statistical significance as done for the various versions of the NSP correlations (including NSP4) as described in Reference 115. Parameters that have a p-value greater than 0.05, which indicates a lack of sensitivity to that parameter and that the parameter does not describe the data well, are removed and the correlating procedure is repeated until all parameters have a p-value less than 0.05. The final © Copyright 2022 by NuScale Power, LLC 160

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 form of the uniform CHF model with all statistically significant terms is presented in Equation 6-115. The final correlation coefficients of the reduced CHF expression are determined using a 3-fold cross-validation in conjunction with LLS regression. Cross-validation is a technique for assessing how accurately a predictive model performs in its actual application. This process minimizes "over-fitting" of the correlation to data along with undesirable biases. The complete population (all local condition data) is randomly partitioned into three relatively equal size sub-populations, or subsets. Of the three subsets, one is held aside for validation testing, while the other two are combined and used to train the correlation. This process is repeated three times (3-fold) until all three subsets have been used for validation testing. The correlation coefficients derived from the three results are (( }}2(a),(c) correlation coefficients. The coefficients and statistical results of validation testing are compared between the three folds to assure that each predicts similar values. The resulting coefficients are identified in Table 6-7. The modified Tong F-factor discussed in Section 6.11.5.2 converts the uniform CHF value calculated with Equation 6-116 to a non-uniform CHF value. ((

                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                       }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                                                     }}2(a),(c)

Table 6-7 NSPN-1 CHF Correlation Coefficients ((

                                                                                                    }}2(a),(c) 6.11.5.2           NSPN-1 Non-Uniform Flux Factor The Tong non-uniform flux factor (F-factor) from Reference 115 is used to accommodate non-uniform axial power shapes when using a CHF correlation that predicts uniform CHF. The uniform and non-uniform conditions are related by:

((

                                                                                                     }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                                                          }}2(a),(c)

Two axial power shapes were tested for the NuFuel-HTP2' design: ((

                                                    }}2(a),(c) Coefficients for the C term in the F-factor are determined by comparison between cosine data and baseline uniform data.

For the NuFuel-HTP2' design, (( }}2(a),(c) expressed in Equation 6-118, was chosen for the NSP4 CHF correlation because it closely matches the NuFuel-HTP2' form and figure of merit, and is based on a robust data set (Reference 115). ((

                                                                                                          }}2(a),(c)

Although this formulation of the Tong factor was derived in the development of the VIPRE-01 subchannel CHF model, its exact form is applicable for the NRELAP5 application. This is because regardless of the code model and method, the same underlying test data and parameters that define the Tong F factor are used. The only slight differences between model predictions would be the resulting local quality and local mass flux at the CHF location. For a given test point the local parameters predicted between VIPRE-01 and NRELAP5 are consistent (Figure 6-8) thus the use of the NSP4 derived non-uniform model is applicable. Any issues that may arise with the Tong factor are captured by the predictability of the CHF model to data and the resulting design limit. © Copyright 2022 by NuScale Power, LLC 164

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 6-8 Comparison of VIPRE-01 and NRELAP5 Quality and Mass Flux at the Measured CHF Location ((

                                                                                                          }}2(a),(c) 6.11.5.3           Local Conditions Local conditions for CHF test data are calculated with the NRELAP5 system thermal hydraulics code. The NRELAP5 model (Section 6.11.5.4) is used to determine local conditions for each test point in the KATHY database. The term local conditions here refers to the axial varying conditions. The NRELAP5 model for which the correlation is being developed for ((
                                                                                                     }}2(a),(c)

The local conditions are taken at the axial location of where CHF is determined to occur based on thermocouple response. The location of the thermocouples in the KATHY test bundle information are provided in Reference 115. ((

                                                }}2(a),(c) The range of local conditions considered are identified in Table 6-8. The local conditions are plotted in Figure 6-9 on a 3D plot

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 where the contour values represent the local limiting CHF value. As expected, and consistent with NSP4 in Reference 115, CHF generally increases with increasing mass flux with a moderate improvement in the intermediate pressure range due to favorable latent heat behavior in this region. Table 6-8 NSPN-1 Range of Conditions ((

                                                                                                      }}2(a),(c)

Figure 6-9 Pressure, Mass Flux, and Quality Domain with Local CHF Contours ((

                                                                                                      }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 6.11.5.4 NSPN-1 NRELAP5 Model The KATHY NuFuel HTP2' test section is modeled in NRELAP5 using (( }}2(a),(c). The heater rods are modeled with a heat structure that has (( }}2(a),(c). Each of the hydraulic nodes is ((

                              }}2(a),(c) 6.11.5.5           NSPN-1 Design Limit Determination The design limit is determined with consistent methods as applied in Reference 115 and illustrated in Figure 6-10.

Normality of samples is tested with either Shapiro-Wilk (50 or fewer data) or D'Agostino D' (greater than 50 data) tests. These tests are described in detail in Reference 117, Sections 11.9 and 11.10, respectively. Variance between samples is tested with either Bartlett (parametric) or k-Sample Squared Ranks (non-parametric) tests. These tests are described in detail in Reference 117, Sections 14.7 and 25.14, respectively. The null hypothesis is that the variance between samples are equal. The Bartlett test is only used when each of the samples can be shown to be from a normal distribution. In order to determine whether the samples can be combined a test of medians, either Kruskal-Wallis (equal variance) or the median test is used. The null hypothesis is that the medians are equal (and can be combined). These tests are described in detail in Reference 117, Sections 25.10 and 25.9, respectively. Once samples are combined into larger sets, the correlation limit is determined using either parametric (Reference 117, Section 9.12) or non-parametric methods. In the non-parametric method, the n data are ordered ascendingly and the (n-k+1)th value is selected as the tolerance limit. The value of k is maximized such that: I 1 - ( k,n - k + 1 ) Equation 6-119 where, I x = (A,B) is the incomplete Beta function © Copyright 2022 by NuScale Power, LLC 167

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 is the probability that coverage is at least % n is the total number of data The maximum tolerance limit for all of the sub-groups is adopted as the correlation limit, because this limit conservatively bounds all of the groups. Figure 6-10 Design Limit Statistical Methods Flow Chart ((

                                                                                                          }}2(a),(c)

Statistics for the three test sets from the 3-fold cross-validation are tabulated in Table 6-9. The P/M values for each of the 3 folds have means ((

                                              }}2(a),(c). The predicted versus measured CHF values are illustrated in Figure 6-11; the bulk of the data falls along the ideal 45-degrees line.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 6-9 Validation Set Statistics for the 3-Fold Cross-Validation ((

                                                                                                            }}2(a),(c)

A correlation limit is determined using the statistical methods discussed above. Sub-regions are created based on test series, inlet mass flux, pressure, inlet subcooling, local mass flux, and local quality, as tabulated in Table 6-10. Statistical tests are used to determine which sub-regions can be combined and a one-sided tolerance limit for each of the combined sets is calculated. The maximum tolerance limits for each of the sub-regions are tabulated in Table 6-11. The limiting subregion 95/95 design limit is (( }}2(a),(c) as the recommended design limit for the NSPN-1 correlation. Bias plots (P/M ratio versus correlating parameter) for pressure, mass flux, quality, cold wall factor, boiling length, and inlet enthalpy are illustrated in Figure 6-12 through Figure 6-17, respectively. For all of the parameters except for inlet enthalpy, a linear fit of the data falls along P/M of 1.0 with little or no slope indicating very little or no bias is present. A linear fit of the data in the inlet enthalpy bias plot (Figure 6-17) demonstrates a slightly larger slope, but no significant bias. This slope is anticipated because inlet enthalpy is not a correlating parameter for the NSPN-1 CHF correlation. The overall performance of the correlation is depicted in Figure 6-18 showing where the lower and upper 5 percent P/M values reside. From this figure it can be observed that the ((

                                                                      }}2(a),(c)

The CHF correlation defined by Equation 6-115 and Table 6-7 does not result in any significant bias with the correlated terms thus the correlation form is determined to be appropriate to use as the NSPN-1 CHF correlation. © Copyright 2022 by NuScale Power, LLC 169

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 6-10 Subsets of NuFuel-HTP2' Data for NSPN-1 ((

                                                                                               }}2(a),(c)

Table 6-11 Correlation Limit by Sub-Region for NSPN-1 CHF Correlation ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 6-11 NSPN-1 Predicted vs. Measured CHF Binned by Test ID ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 6-12 Pressure Bias Plot for NSPN-1 CHF Correlation ((

                                                                                                 }}2(a),(c)

Figure 6-13 Mass Flux Bias Plot for NSPN-1 CHF Correlation ((

                                                                                                 }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 6-14 Quality Bias Plot for NSPN-1 CHF Correlation ((

                                                                                                 }}2(a),(c)

Figure 6-15 Cold Wall Factor Bias Plot for NSPN-1 CHF Correlation ((

                                                                                                 }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 6-16 Boiling Length Bias Plot for NSPN-1 CHF Correlation ((

                                                                                                  }}2(a),(c)

Figure 6-17 Inlet Enthalpy Bias Plot for NSPN-1 CHF Correlation ((

                                                                                                  }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 6-18 Pressure, Mass Flux, and Quality Domain Sorted by Correlation Performance ((

                                                                                                          }}2(a),(c) 6.11.5.6           NSPN-1 Critical Heat Flux Model and Applicability The NSPN-1 CHF correlation is developed using NuFuel-HTP2' CHF test data.

The NSPN-1 CHF correlation has the form defined by Equation 6-115 with identified coefficient values in Table 6-7. Evaluation of the NSPN-1 CHF correlation indicates that the correlation provides an accurate prediction of NuFuel-HTP2' test data. A correlation design limit of (( }}2(a),(c) ensures at the 95/95 level that CHF is not be experienced on a rod demonstrating a limiting value, which meets the requirements of 10 CFR 50, General Design Requirement 10. The NSPN-1 CHF correlation must be used in conjunction with local condition calculations from the NRELAP5 system code. The ranges of applicability for the NSPN-1 CHF correlation are tabulated in Table 6-8. © Copyright 2022 by NuScale Power, LLC 175

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The NSPN-1 correlation 95/95 design limit ((

                                                                   }}2(a),(c) CHF limit for LOCA and IORV events is 1.20.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.0 NRELAP5 Assessments The following section provides a summary of the SET and IET assessments that have been completed for NRELAP5. The results of these assessments are considered in Section 8.0 to justify the adequacy on NRELAP5 for modeling of high-ranked phenomena in the NuScale LOCA PIRT. To assess the adequacy of NRELAP5, code simulations are compared to measured experimental data. Acceptance criterion from Table 1-2 are applied in rating NRELAP5 performance in terms of "excellent", "reasonable" or "minimal" agreement. These ratings take into consideration the ability to predict overall data trends as well as the magnitude of the data itself. 7.1 Assessment Methodology Various experimental tests, inclusive of SETs, IETs, and analytic problems have been used to assess the performance of NRELAP5 using the process identified in Element 2 of RG 1.203. The database employed to assess the adequacy of the NRELAP5 code was chosen to be consistent with the requirements to adequately model the high-ranked phenomena derived in the NuScale LOCA PIRT. The high-ranked phenomena selected in Section 4.0 are mapped onto an assessment matrix of experiments, and are listed in Table 7-1. The analytic problems (fundamental tests) used to assess NRELAP5 are not shown in Table 7-1. Table 7-1 NRELAP5 Loss-of-Coolant Accident Assessment Matrix ((

                                                                                                       }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 7-1 NRELAP5 Loss-of-Coolant Accident Assessment Matrix (Continued) ((

                                                                                                    }}2(a),(c)

Summarized within this section for each assessment are the following: a brief description and purpose of the experimental facility, a summary of the phenomenon addressed, the experiment procedure, important NRELAP5 modeling techniques, and performance of NRELAP5 against the data. Assessment cases are divided into two categories: Legacy Assessments - these are assessments performed against data collected from historical test programs not encompassed within the NuScale test programs NuScale Test Assessments - these are assessments performed against data collected as part of the NuScale testing program The following sections document the various assessments completed with NRELAP5. These assessments were completed with the current version of NRELAP5 described in Reference 9. © Copyright 2022 by NuScale Power, LLC 178

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.2 Legacy Test Data This section describes those test programs which have produced data that were not performed under the NuScale QAPD (Reference 4). With the exception of Marviken JIT-11 data, these tests are qualified for use by applying non-mandatory guidance provided by NQA-1 2008/1a-2009 Addenda (Reference 10). Use of Marviken JIT-11 data is based on published literature data. 7.2.1 Ferrell-McGee The Ferrell-McGee tests were performed in vertical pipes over a wide range of single-phase and two-phase flow conditions with uniform, contraction, and expansion flow areas. The data assessed includes single- and two-phase pressure drop and void fraction under different pressures, flow rates, and inlet quality. 7.2.1.1 Facility Description The report for the Ferrell-McGee experiments (Reference 58) describes the test facility. Figure 7-1 shows the schematic of the test section. The test apparatus consists of a heated section that controls the degree of sub-cooling of the liquid entering into an adiabatic test section, where pressure drops and void fractions were measured. The lower test section is 40.5 in. (1.0287 m) in height and the upper test section is 49.5 in. (1.2573 m) for a total of 90.0 in. (2.286 m). The two test sections were connected by mating flanges. The tests were organized into seven test groups. Each test group had a different combination of pipe diameters for the lower and upper sections. The tests of Group 1 and Group 4 used pipes of uniform diameter of 0.46 in. (0.0117 m) and 0.34 in. (0.00864 m) and are designed to assess two-phase frictional pressure drop. Tests of Groups 2, 5 and 6 are tests with abrupt area expansion with area ratios of 0.608, 0.332 and 0.546, respectively. The tests of Groups 3 and 7 are tests with abrupt contraction with area ratios of 0.546 and 0.608, respectively. In this section the area ratios are defined as small area or large area. Tests with abrupt area expansion and contractions are designed to assess frictional and form losses. Multiple sets of tests were run with different combinations of pressure, flow, and inlet quality. © Copyright 2022 by NuScale Power, LLC 179

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.2.1.2 Phenomena Addressed The phenomenon addressed with the Ferrell-McGee assessment cases is the ability of NRELAP5 to predict ((

                                                                                                   }}2(a),(c) 7.2.1.3            Experimental Procedure The part of the stainless steel flow loop of particular interest is the vertical adiabatic test section in which an upward flowing steam-water mixture entered at a controlled pressure, mass flow rate, and quality. Pressure drops and steam volume fractions (  ) were measured along the channel at the locations shown in Figure 7-1.

The two-phase mixture exited from the test section into a 0.460-in. inside diameter glass section through which the mixture could be photographed. The vapor-liquid mixture was partially separated in a surge tank and sub-cooled in a bank of six parallel concentric-tube heat exchangers. The sub-cooled liquid passed through a pump, a manual flow control valve, a volumetric flowmeter, and a preheater which controlled the quality of two-phase flow entering the heated section. In the heated section, a 0.462-in. inside diameter by 0.083-in. wall tube heated by an alternating current flowing in the tube wall, the water was brought to the desired quality before injection into the adiabatic test section. System pressure was maintained by a hydraulic accumulator. Loop fluid, cooling water, heated channel wall, and manometer line temperatures were also recorded. The summary of ranges of recorded data for the 201 runs is provided in Table 7-2. The initial boundary conditions covered a range of three different mass flow rates of 460, 920 and 1,150 lbm/hr (209, 417, and 522 kg/hr), a range of three different pressures of 60, 120 and 240 psi (0.414, 0.827, and 1.65 MPa) and void fractions from 0.0 to 1.0. Table 7-2 Summary of Ferrell-McGee Experimental Test Data Range Range Parameter Min Max Units Pressure 60 (0.414) 240 (1.65) psia (MPa) Inlet flow rate 460 (209) 1,150 (522) lbm/hr (kg/hr) Inlet void fraction range1 -0.110 1.038 n/a Expansion area ratio 0.332 0.608 n/a Contraction area ratio 0.546 0.608 n/a 1 Negative void fractions refer to sub-cooling estimates which are calculated. © Copyright 2022 by NuScale Power, LLC 181

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Although the measured data includes void distributions at nine different locations, shown in Figure 7-1, only the void at measurement locations 1, 2 and 3 were considered. These locations have the void measurements near the center of NRELAP5 nodes. The rest of the void fraction measurement taps were not taken into account because placing all void measurement locations near the center of nodes would result in small nodes with a length-to-diameter (L/D) ratio less than 1.0. The total pressure drop measurement uncertainty was estimated to be +/-0.45 psi (0.0031 MPa). The average void fraction measurement uncertainty was estimated to be +/-3 percent. It is noted that test cases 1A6, 1A7, 1A11, 4A5, 4A9, 5A5, 5A10 and 6A9 with void fractions at Void Tap 1 larger than 0.97 were excluded from comparative results. The NRELAP5 total pressure drop predictions for these test cases showed a high deviation from the measurement, which subsequent analysis revealed to result from uncertainty of the inlet void fraction measurement. 7.2.1.4 Special Analysis Techniques For test groups 2, 3, 5, 6 and 7, at the position of expansion and contraction ((

                                              }}2(a),(c) 7.2.1.5            Assessment Results Figure 7-2 shows the predicted versus measured pressure drop for uniform, expansion, and contraction tests. NRELAP5 predicted the experimental data with reasonable-to-excellent agreement. These results validate the ability of NRELAP5 to predict ((
                                                                                                  }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-2 Predicted Versus Measured Pressure Drop for Selected Contraction Tests 7.2.2 GE Level Swell (1 ft) During various phases of a blowdown event in an NPM, such as a LOCA, the fluid within the RCS experiences flashing, vapor generation, level swell, and conditions representative of rapid depressurization. Reference 59 provides a suitable experimental database extending across a large range of pressures and fluid conditions that is used to assess the ability of NRELAP5 to predict ((

                                                                                        }}2(a),(c) The assessment of NRELAP5 against the 1 ft. diameter GE level swell test is summarized in this section, while the assessment of NRELAP5 against the 4 ft. GE level swell test is provided in Section 7.2.3.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.2.2.1 Facility Description The experimental facility is fully described in Reference 58 and summarized in this section. The experiment facility shown in Figure 7-3 consists of a pressure vessel made of carbon steel with a volume of 10 ft3 (0.283 m3), a diameter of approximately 12 in. (0.305 m) and a 14-ft (4.2672 m) length. The small vessel experiments (1 ft.) include a blowdown line with orifice plates that are interchanged to control the blowdown flow rate and depressurization rate. The effluent from the vessel blowdown is discharged into a suppression tank. Figure 7-3 Schematic of the GE 1 ft. Blowdown Vessel

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Three basic types of measurements were obtained during each experiment: static pressures, differential pressures, and temperatures. Figure 7-3 shows the location of the instrumentation. There are six measurement sections between the adjacent differential pressure taps, numbered sequentially. The two-phase density (or mixture density) in each measurement section during blowdown experiments is derived from the axial differential pressure measurements. The fluid mass inventory is obtained from the density and known volume of the measurement section. The average void fraction in each measurement section is determined from the measured mixture density and thermodynamic properties of the liquid and vapor phases at the system pressure as shown in Equation 7-1 (Reference 59). i = ( i - f ) ( g - f ) Equation 7-1 i is the average void fraction in the i-th measurement section, i is the average mixture density in the i-th measurement section, and f , g are the liquid and vapor densities as a function of the measurement section pressure. 7.2.2.2 Phenomena Addressed The phenomena addressed with the GE level swell (1 ft. diameter) test include ((

                                                                                           }}2(a),(c)

Specifically, the GE level swell test assesses the ability of NRELAP5 to predict key in-vessel thermal-hydraulic phenomena associated with a rapid depressurization event. 7.2.2.3 Experimental Procedure The experiment consisted of filling the vessel with demineralized water and boiling the inventory at atmospheric pressure to remove any dissolved gas. The top vent was then closed and the fluid was heated to the specified initial conditions, which was typically saturated conditions at the desired pressure. The initial water level was dependent upon the experiment of interest as listed in Table 7-3. Once conditions were reached a blowdown was initiated from a discharge valve located at the top of the vessel and measurements were recorded. Table 7-3 summarizes the experiment conditions for the GE 1 ft. level swell test selected for the assessment. The parameters listed are used as boundary conditions in the NRELAP5 inputs or for comparisons to the NRELAP5 predictions. The blowdown experiment is initiated at a pressure of 1,011 psia (6.97 MPa) with saturated fluid conditions. The blowdown and fluid response is © Copyright 2022 by NuScale Power, LLC 185

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 measured for approximately 300 seconds. Void fractions ranging from 0.0 to 1.0 are present during the test. Table 7-3 Summary of GE 1 ft. Vessel Level Swell Experiments Orifice Size Restriction Plate Initial Pressure Initial Liquid Level Test Number inches (mm) Configuration psia (MPa) ft (m) 1004-3 3/8 (9.525) No plate 1,011 (6.971) 10.4 (3.17) 7.2.2.4 Special Analysis Techniques Based on sensitivity studies it is determined that applying the ((

                                                                        }}2(a),(c) improved the depressurization comparisons from reasonable-to-excellent agreement. The discussion here, however, provides a summary with default discharge coefficient input of ((                                     }}2(a),(c) 7.2.2.5            Assessment Results Figure 7-4 through Figure 7-8 present the vessel pressure and axial void fraction comparisons between NRELAP5 using the ((                             }}2(a),(c) choking model and the measured data for experiment 1004-3. The initial 100 seconds of the simulations are to confirm a steady state condition. The blowdown event is initiated at 100 seconds and is therefore the initial time for all figures.

Figure 7-4 presents the calculated vessel pressure versus the measured data. The comparisons are in reasonable-to-excellent agreement. The predicted depressurization rate is slightly higher compared to the data. Figure 7-5 through Figure 7-8 present the calculated axial void fraction versus the measured data for several points in time during the transient. The comparisons are presented for 10, 40, 100, and 160 seconds into the transient. The results show reasonable-to-excellent agreement. The trend of increased void fraction along the height of the vessel is predicted rather well. Deviations are observed at the lower elevations. The results of mixture level in the vessel, not presented here, also show reasonable-to-excellent agreement between the calculated NRELAP5 results and the experimental data. The results validate the ability of NRELAP5 to predict key in-vessel thermal-hydraulic phenomena associated with a rapid depressurization event. © Copyright 2022 by NuScale Power, LLC 186

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-8 GE Level Swell 1 ft. Vessel Void Fraction Versus Elevation at 160 Seconds 7.2.3 GE Level Swell (4 ft) 7.2.3.1 Facility Description The 4 ft. GE Level Swell test facility is fully described in Reference 59 and summarized in this section. The experimental facility shown in Figure 7-9 consists of a pressure vessel made of carbon steel with a volume of 160 ft3 (4.5306 m3), 47 in. (1.1938 m) in diameter and 14 ft (4.2672 m) in length. The test facility includes a 10 in. (0.254 m) diameter vertical blowdown dip tube to simulate top break locations and a horizontal blowdown line to simulate bottom break locations. The effluent from the vessel blowdown is discharged to a suppression tank. © Copyright 2022 by NuScale Power, LLC 191

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-9 Schematic of the GE 4 ft. Blowdown Vessel Three basic types of measurements were obtained during each experiment: pressures, differential pressures, and temperatures. As shown in Figure 7-9 the pressure drop is measured at seven sections between the adjacent differential pressure taps, numbered sequentially. Similar to the 1 ft. GE level swell test, the mixture density and void fraction in each measurement section were calculated from the measured pressure drop (Section 7.2.2.1). © Copyright 2022 by NuScale Power, LLC 192

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.2.3.2 Phenomena Addressed The phenomena addressed with the 4 ft. GE level swell are the same as in the 1 ft. GE level test (Section 7.2.2.2). 7.2.3.3 Experimental Procedure The experiment procedure consisted of filling the vessel with demineralized water and boiling the inventory at atmospheric pressure to remove any dissolved gas. The top vent was then closed and the fluid was heated to the starting conditions, which was typically saturated conditions at 1,060 psia (7.308 MPa) for the large blowdown vessel experiments. The initial water level was dependent upon the experiment of interest. Top and bottom break blowdown events were conducted utilizing rupture discs. A test 5801-15 with top break and initial liquid level of 5.5 ft (1.676 m) is selected for the assessment. 7.2.3.4 Special Analysis Techniques The (( }}2(a),(c) improves the depressurization for the sensitivity modeling the discharge and choking into an atmospheric blowdown tank. The discussion here, however, provides a summary with ((

                            }}2(a),(c) 7.2.3.5            Assessment Results The results of the GE level swell (4 ft. vessel) from NRELAP5, using the

(( }}2(a),(c) choking model and the measured data are compared. Key parameters are plotted together with the test data in Figure 7-10 through Figure 7-13. The results show reasonable-to-excellent agreement based on the comparison of the pressure and void fractions in the vessel. These results validate the ability of NRELAP5 to predict key in-vessel thermal-hydraulic phenomena associated with a rapid depressurization event. © Copyright 2022 by NuScale Power, LLC 193

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-13 GE Level Swell 4-ft Vessel Void Fraction Versus Elevation at 20 Seconds 7.2.4 KAIST In the NuScale design, the DHRS is a passive safety-related system that relies on film condensation and natural circulation to remove heat from the RCS through the SG and reject heat to the reactor pool through the DHRS condenser. Reference 60 provides a suitable high pressure steam condensation experimental database which is used to assess the condensation model in NRELAP5. The KAIST test varied the pressure and non-condensable gas fraction of the steam entering the test section (mockup of a condenser tube). ((

                                      }}2(a),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.2.4.1 Facility Description A schematic of the KAIST test facility is shown in Figure 7-14. Figure 7-15 shows the schematic of the test section. The maximum design pressure and temperature of the test facility are 7.5 MPa (1.088 psia) and 300 degrees C (752 degrees F), respectively. The major components of the test facility include: SG which supplied steam (maximum power 200 kWe), test section tube, cooling pool (cools the test section), steam line (transports steam from SG to the test section inlet), condensate drain line, LP (or condensate collection tank), and air supply system. The test section was immersed in the cooling pool and was cooled by boiling and single-phase convective heat transfer on the outside surface of the test section. The test section was a vertical tube with an inside diameter of 4.62 cm (1.82 in.) and an effective heat transfer length of 1.8 m (71 in.). The thickness of the tube wall was 2.3 mm (0.09 in.). To reduce the entrance effect, the top 0.5 m (20 in.) length of the test section was insulated. The test section was submerged in a cooling pool of width 1.2 m x 1.2 m (47 in. x 47 in.) and 2.5 m (98 in.) height. A steam line with an inside diameter 2.34 cm (0.92 in.) was connected from the top of the SG to the top of the test section. The condensate from the test section was drained to the LP (or condensate collection tank) by gravity and then pumped back to the SG. © Copyright 2022 by NuScale Power, LLC 198

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.2.4.2 Phenomena Addressed The phenomena addressed with the KAIST assessment include ((

                                      }}2(a),(c) 7.2.4.3            Experimental Procedure The experiments are started by purging non-condensable gas (i.e., air) from the test loop. This is done by supplying steam to the test loop and venting it to the atmosphere through the vent valve located below the test section. After non-condensable gas is purged, the vent valve is closed and the test section is allowed to fill with the condensate by keeping the condensate drain valve closed.

After the test section is completely filled, the SG pressure is increased to the test pressure. As soon as the test pressure is reached, the condensate drain valve is opened and the condensate recirculation pump is started. A constant water level in the LP is maintained by control of the recirculation pump flow rate. Data acquisition is started after the process reaches a steady state. Parameter ranges for the KAIST tests are summarized in Table 7-4. Table 7-4 Range of KAIST Test Data Parameter Value Pressure (MPa) 0.794 to 7.457 (115 to 1082 psia) 0.036 to 0.34 Reduced pressure (Pr) (using critical pressure of 220.64 bar (3,200 psia) Inlet steam mass flow (kg/s) 0.01 to 0.1 (0.022 to 0.22 lb/s) Inlet air concentration (percent) 0.0 to 30.0 Prandtl number (Prf) 0.84 to 2.63 Liquid Reynolds number (ReLT) 2,300 to 3,2000 Inlet gas Reynolds number (ReGS) 16,400 to 15,0000 7.2.4.4 Special Analysis Techniques Based on sensitivity studies, ((

                                                         }}2(a),(c) 7.2.4.5            Assessment Results The results show reasonable-to-excellent agreement between the NRELAP5 calculations and the KAIST measured experimental data, on the comparison of

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 condensed liquid flows, heat transfer coefficients, and inner wall temperatures. This is a result of implementation of the ((

                                }}2(a),(c) in NRELAP5 (Section 6.8), which is intended to improve the predicted high pressure condensation response.

Figure 7-16 presents the measured versus calculated heat transfer coefficient for the KAIST steam condensation experiments. The majority of the predictions lie within the experimental uncertainty (28 percent for heat transfer coefficient). Figure 7-16 Measured versus predicted heat transfer coefficient Figure 7-17 through Figure 7-19 present heat transfer coefficient, temperature, and mass flow rate versus test section elevation. The majority of the predicted values (all but one) lie within the uncertainty range of the data. Overall, the results show that NRELAP5 calculations are in excellent agreement with the KAIST measured experimental data. These results validate NRELAP5 for prediction of key thermal-hydraulic phenomena associated ((

                                                       }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-19 KAIST and NRELAP5 Axial Liquid Mass Flow Rate 7.2.5 FRIGG The FRIGG loop tests for the Marviken boiling heavy water reactor project were executed in four phases by ASEA-ATOM during the years 1967-1970 (Reference 61). These experiments included measurements of axial and radial void distribution, single-phase and two-phase pressure drop, natural circulation mass velocity, stability limits as well as detailed dynamic characteristics, and burnout in natural and forced circulation. The axial and radial void distribution data as a function of mass flow, inlet sub-cooling, system pressure, and thermal power provide an excellent data set for evaluating the NRELAP5 interphase drag and heat transfer models under two-phase flow conditions. The FRIGG phase 4 (FRIGG-4) tests applied both a non-uniform radial and axial thermal power profile on the heated rod bundle best simulating the power © Copyright 2022 by NuScale Power, LLC 205

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 profiles associated with a typical operating reactor core. As such the FRIGG-4 tests are used to assess NRELAP5 performance. 7.2.5.1 Facility Description The FRIGG-4 test facility consisted of a vertical circular test section containing 36 electrically heated rods, a riser, a steam separator, a downcomer, a condenser, a pump, and connecting pipes. The power supply for the FRIGG loop was capable of providing a maximum of 8 MW of direct current power to the heated rods in the test section. A schematic of the test loop is shown in Figure 7-20. Figure 7-21 shows the locations of the void and pressure sensors used in the test section. Reference 61 provides detailed information on the characteristics of the facility. The rod bundle simulated a full-scale boiling heavy water reactor fuel element. Each rod had a 4.365 m (172 in.) heated length and a 13.8 mm (0.543 in.) outside diameter. The bundle also included a 20 mm (0.787 in) outside diameter unheated center rod that supported the prototype reactor core grid spacers. The heated rods were arranged in equal intervals in three rings, the inner ring having six rods, the middle ring twelve rods, and the outer ring eighteen rods. The rod bundles were contained within a 159.5 mm (6.28 in.) ID shroud. The heated rod bundle had axial and radial thermal power peaking factors typical of an equilibrium reactor core. The FRIGG-4 tests have no thermal power variation in the azimuthal (circumferential) direction. The average heat flux for each test is determined by dividing the total thermal power by the total heated surface. The local heat flux at any given radial or axial zone can be determined by multiplying the measured average heat flux by the radial and axial coordinate scale factors. © Copyright 2022 by NuScale Power, LLC 206

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-22 FRIGG-4 Zones for Evaluation of Radial Void Distribution 7.2.5.2 Phenomena Addressed The phenomena addressed with the FRIGG-4 assessment include ((

                                                                                                     }}2(a),(c)

Specifically, the FRIGG-4 tests assess the ability of NRELAP5 to predict the void distribution data in a rod bundle geometry as a function of mass flow, inlet sub-cooling, system pressure, and thermal power for evaluating interphase drag and heat transfer models under two-phase flow conditions in the core. 7.2.5.3 Experimental Procedure Test points were obtained by specifying the core electric power, inlet flow rate, inlet sub-cooling, and system pressure. Measurements of axial void fractions were collected for each radial zone of the rod bundle. © Copyright 2022 by NuScale Power, LLC 209

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.2.5.4 Special Analysis Techniques There were no special analysis techniques utilized. 7.2.5.5 Assessment Results One-dimensional NRELAP5 model of the test section is used to analyze this test. Figure 7-23 to Figure 7-26 below show the area-weighted average void fractions in axial zones G1 through G7 for tests 613123, 613130, 613010 and 613118. NRELAP5 predicts the experimental void fraction data with reasonable agreement justifying use of ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-24 FRIGG Mean Void Data of NRELAP5 Versus Test 613130 Data Figure 7-25 FRIGG Mean Void Data of NRELAP5 Versus Test 613010 data © Copyright 2022 by NuScale Power, LLC 211

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-26 FRIGG Mean Void Data of NRELAP5 Versus Test 613118 Data 7.2.6 FLECHT-SEASET The FLECHT-SEASET tests (Reference 62 and Reference 63) consisted of forced and gravity reflood experiments using electrical heater rods to simulate fuel bundles similar to the Westinghouse 17 x 17 design. The test program was originally designed to study large-break LOCA events. Following the Three Mile Island accident, it was re-oriented to obtain data relevant to small break LOCA events. Because the NuScale core remains covered with coolant for all design basis LOCA events, reflood phenomena does not occur. However, the test campaign included bundle boil-off tests which are relevant for the NuScale design because the NPM uses boiling in the core to remove heat following a number of accident scenarios that result in actuation of the ECCS. Following ECCS initiation and pressure equalization in the NPM, the RPV and CNV are essentially a pool boiler system with coolant boiled off in the RPV being replaced by an inflow of coolant from the CNV. 7.2.6.1 Facility Description The facility loop with test section is shown in Figure 7-27. The heater rods were manufactured with a prototypical PWR axial cosine power shape. © Copyright 2022 by NuScale Power, LLC 212

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-27 FLECHT-SEASET Experimental Facility

     
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                                                    Ž                                     W    ŽŽ d&

dt d& dt

d&

d&

                               Ž                                                      ŽŽ

d&

                                                                                     Ž
                                                                             
 Ž                                                              
                                                        ŽŽ 7.2.6.2            Experimental Procedure The FLECHT-SEASET boil-off tests were conducted by filling the 12 ft. tall vessel with approximately 10 ft. of (slightly sub-cooled) water. The power to the heater rods was turned on, and the water was allowed to boil. The test was terminated and reflood initiated when a rod thermocouple registered a temperature greater than or equal to 2,000 degrees F. Three separate boil-off tests are used to assess NRELAP5. These tests were conducted with initial system pressures of 20, 40, and 60 psia.

7.2.6.3 Phenomena Addressed The phenomenon addressed with the FLECHT-SEASET assessment include (( }}2(a),(c) Specifically, the FLECHT-SEASET boil off tests assess the ability of NRELAP5 to predict the axial void profile, mixture level (interfacial drag), and cladding temperature response during boil-off of a PWR core. 7.2.6.4 Special Analysis Techniques Based on sensitivity studies using one-dimensional components, it is concluded that ((

                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.2.6.5 Assessment Results The results for Test 35557 performed at 60 psia are shown in Figure 7-28 through Figure 7-35. The predictions for the void fraction at different elevation are shown in Figure 7-28 through Figure 7-31. The comparisons for the collapsed water levels for all sections are provided in Figure 7-32 through Figure 7-35. While the model and data show reasonable agreement, NRELAP5 over-predicts void fractions as a function of time in most of the core region (Figure 7-28 through Figure 7-31) resulting in a conservative earlier prediction of core uncovery when compared to test data. Similar comparisons are obtained for the test runs at 20 psia and 40 psia. Figure 7-28 FLECHT-SEASET Level 1 Void Fraction Versus Time - Test 35557 © Copyright 2022 by NuScale Power, LLC 214

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-29 FLECHT-SEASET Level 2 Void Fraction Versus Time - Test 35557 Figure 7-30 FLECHT-SEASET Level 3 Void Fraction Versus Time - Test 35557 © Copyright 2022 by NuScale Power, LLC 215

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-31 FLECHT-SEASET Level 4 Void Fraction Versus Time - Test 35557 Figure 7-32 FLECHT-SEASET Level 1 Collapsed Water Level Versus Time - Test 35557 © Copyright 2022 by NuScale Power, LLC 216

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-33 FLECHT-SEASET Level 2 Collapsed Water Level Versus Time - Test 35557 Figure 7-34 FLECHT-SEASET Level 3 Collapsed Water Level Versus Time - Test 35557 © Copyright 2022 by NuScale Power, LLC 217

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-35 FLECHT-SEASET Level 4 Collapsed Water Level Versus Time - Test 35557 7.2.7 SemiScale (S-NC-02 and S-NC-10) The Semiscale test loop modeled a typical PWR. The goal of the Semiscale S-NC-2 and S-NC-10 tests was to obtain experimental data on the natural circulation single-phase and two-phase flow conditions at various system inventories for differing system powers. Three powers were investigated: 30 kW, 60 kW, and 100 kW. At each power level the mass inventory was reduced from conditions at or near 100 percent. With the reduction of primary inventory two-phase flow developed resulting in an enhancement of the total system flow rate. Further reduction in system inventory resulted in a degradation of the total system flow rate. 7.2.7.1 Facility Description The Semiscale Mod-2A test facility is a full-height 1/1,705 power-to-volume scaled model of a typical four-loop PWR. Only one loop was used for the S-NC-2 and S-NC-10 tests discussed here. The single-loop configuration is shown in Figure 7-36. The reactor coolant pump was removed and replaced by an orifice to model the loss of a seized pump. © Copyright 2022 by NuScale Power, LLC 218

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-36 Semiscale Mod-2A Single (Intact) Loop Test Facility Configuration 7.2.7.2 Phenomenon Addressed The phenomenon addressed with the Semiscale assessment includes (( }}2(a),(c) © Copyright 2022 by NuScale Power, LLC 219

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Specifically, the Semiscale natural circulation tests assess the ability of NRELAP5 to predict natural circulation during single- and two-phase flow conditions at various system inventories and system powers in a complex geometry. 7.2.7.3 Experimental Procedure Prior to initiation of the tests, the primary system was filled with demineralized water and vented to ensure it was liquid-full. The primary system was heated using core power as a heat source and the SG secondary system as a heat sink. Single-phase natural circulation flow driven by density gradients in the loop was used to thermally condition the system to obtain a specified set of initial conditions. For those steady-state tests in which the primary system was to be drained, the pressurizer was used only to establish initial conditions. The pressurizer was disconnected from the coolant loop prior to draining the primary system. Primary system mass inventory was controlled by draining fluid from the vessel LP in discrete steps. This fluid was condensed and measured using a static pressure transducer. The SG secondary levels were controlled by a feed-and-bleed process combined with secondary system draining. The secondary-system pressure was maintained such that saturation conditions prevailed through the use of a steam control valve. The steady-state natural circulation tests used constant core powers from 30kW to 100kW, representing 1.5 percent to 5.0 percent of the 2,000 kW full Semiscale core power. During the steady-state experiments, the independent variables were controlled in discrete, step-wise manners, allowing steady conditions to be established between the times when changes in the independent variables were made. External heaters were used to offset heat losses from the primary coolant system that would affect loop natural circulation behavior. The heaters were located on the hot leg, pump suction, cold leg and vessel downcomer sections of the experiment system. The external heater powers were adjusted to follow previously-determined system heat loss versus system temperature relations. The effectiveness of the external heaters was verified by ensuring constant temperatures (indicative of no heat losses) across these sections. Three test cases (30kW, 60kW, and 100kW) of S-NC-2 and S-NC-10 were evaluated. 7.2.7.4 Special Analysis Techniques The NRELAP5 assessment of the Semiscale test cases shows that the (( }}2(a),(c) © Copyright 2022 by NuScale Power, LLC 220

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                              }}2(a),(c) 7.2.7.5            Assessment Results The NRELAP5 models for S-NC-02 at 30 kW and 60 kW, and S-NC-10 at 100 kW were run at various system mass inventories. Once steady conditions are obtained for each inventory reduction, the last 200 seconds of the interval are averaged to obtain key FOMs for comparison to data.

The NRELAP5 predictions of loop mass flow rate as a function system inventory are compared to the experimental data in Figure 7-37 to Figure 7-39. In general, NRELAP5 provides reasonable-to-excellent agreement when predicting the trends, the peak two-phase flow rate, and the enhanced flow rate region (region to the right of the peak). Minor discrepancies are noted to exist in the degraded loop flow region (75 percent - 80 percent inventory levels). These results validate the ability of NRELAP5 to predict natural circulation during single- and two-phase flow conditions at various system inventories and system powers in a complex geometry. Figure 7-37 S-NC-2 30 kW Average Mass Flow Rate Versus Percent Inventory © Copyright 2022 by NuScale Power, LLC 221

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-38 S-NC-2 60 kW Average Mass Flow Rate Versus Percent Inventory Figure 7-39 S-NC-10 100 kW Average Mass Flow Rate Versus Percent Inventory © Copyright 2022 by NuScale Power, LLC 222

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.2.8 Wilson Bubble Rise The NPM hot leg riser is a large-diameter pipe. During various phases of a LOCA, a nearly stagnant two-phase mixture is present in the riser. The Wilson bubble rise experimental data are useful to validate NRELAP5 for prediction of void fraction distribution in the hot leg riser. 7.2.8.1 Facility Description The test facility shown in Figure 7-40 includes a steam inlet and exit nozzle as well as an 18-in. (0.457 m) diameter channel inserted vertically within a 36-in. (0.914 m) diameter vessel. A simplified portion of the test section has been modeled with selected boundary conditions. The boundary conditions consist of the inlet steam mass flow rate and exit pressure. Steam enters the test section from the bottom and exits at the top. Figure 7-40 Schematic of Wilson Bubble Rise Test Facility © Copyright 2022 by NuScale Power, LLC 223

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.2.8.2 Experimental Procedure The Wilson bubble rise experiments were executed by the vessel being slowly heated and brought to equilibrium at the desired test pressure. The water level and steam flow rates were adjusted to the desired values. After the system reached equilibrium, the necessary instrument readings were taken. These readings were the vessel pressure, steam flow, and the three radial void fraction readings (outer radial region, median region and central region of channel). After the readings were taken, the next steam flow was set and the process was repeated. The steam flow was varied from 5,000 to 60,000 lb/hr (2,268 to 27,216 kg/hr) and the pressures ranged from 600 to 2,000 psi (4.14 to 13.8 MPa). 7.2.8.3 Phenomena Addressed The phenomenon addressed with the Wilson bubble rise assessment case are (( }}2(a),(c) Specifically, the Wilson bubble rise tests assess the ability of NRELAP5 to predict axial void distribution (dependent on interfacial drag) within a large diameter vertical channel. 7.2.8.4 Special Analysis Techniques Based on sensitivity studies, ((

                                                                       }}2(a),(c) 7.2.8.5            Assessment Results The results in Figure 7-41 through Figure 7-43 show the comparison of predicted and measured void fraction at different pressures. Figure 7-44 shows the data for all cases plotted as predicted versus measured. A reasonable agreement is observed between the calculated NRELAP5 results and the Wilson bubble rise measured experimental data. In general NRELAP5 conservatively predicted higher void fraction (or lower mass inventory).

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-41 NRELAP5 and Wilson Void Fraction Versus Superficial Velocity at 600 psig (4.14 MPa) Figure 7-42 NRELAP5 and Wilson Void Fraction Versus Superficial Velocity 1,000 psig (6.89 MPa) © Copyright 2022 by NuScale Power, LLC 225

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-43 NRELAP5 and Wilson Void Fraction Versus Superficial Velocity 2,000 psig (13.8 MPa) Figure 7-44 Predicted Versus Measured Area Averaged Void Fraction (all cases) © Copyright 2022 by NuScale Power, LLC 226

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.2.9 Marviken Jet Impingement Test (JIT) 11 The Marviken JIT-11 (Reference 66) was chosen to assess the single-phase choked flow model in NRELAP5. 7.2.9.1 Facility Description A schematic of the experimental setup is shown in Figure 7-45. The facility consisted of a pressure vessel of fluid at specified conditions, discharge pipe, ball valve, and discharge nozzle. The facility was constructed with focus on measuring loads due to discharged fluid impingement on a flat plate and full-scale critical flow data. The facility was constructed with a stand-pipe such that only single-phase steam was discharged through the break nozzle. © Copyright 2022 by NuScale Power, LLC 227

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-45 Marviken Jet Impingement Test Facility 7.2.9.2 Experimental Procedure Each test consisted of first obtaining desired initial conditions in the pressure vessel followed by bursting the rupture disk in the discharge pipe. For JIT-11 a stand-pipe was installed in the pressure vessel such that only steam from the upper plenum of the vessel was discharged. The test was conducted at 5 MPa (725 psia) and nearly-saturated liquid in the vessel. The nozzle diameter was 299.0 mm (0.098 ft) with a nozzle length of 1.18 mm (7.4x10-3 in.). © Copyright 2022 by NuScale Power, LLC 228

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.2.9.3 Phenomenon Addressed The Marviken JIT-11 addresses the ability of NRELAP5 to predict single-phase (vapor) choked flow (mass and energy release). 7.2.9.4 Special Analysis Techniques ((

                                                              }}2(a),(c) 7.2.9.5            Assessment Results Figure 7-46 and Figure 7-47 compare the experimental data and the NRELAP5 simulated mass flow rate and density for various values of the discharge coefficient. Excellent agreement is shown with the experimental data for

((

                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-46 Marviken Jet Impingement Test 11 Flowrate ((

                                                                                                }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-47 Marviken Jet Impingement Test 11 Density ((

                                                                                                       }}2(a),(c) 7.2.10        Bankoff Perforated Plate Bankoff, et al. (Reference 67, Reference 68, and Reference 69) conducted air/water and steam/water counter current flow tests in a small scale test apparatus that established counter current flow through a number of different perforated plates. The Bankoff correlation assessment uses the CCFL implementation as described in Section 6.6.3.

7.2.10.1 Facility Description Reference 69 describes the Bankoff CCFL test apparatus. Additional information on the test apparatus and additional tests are reported in Reference 67 and Reference 68. A horizontal perforated plate is located in a vertical test assembly. Steam or air can be introduced below the plate and water can be injected above the plate. A water overflow line is located above the plate to limit the height of the © Copyright 2022 by NuScale Power, LLC 231

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 "bubbly pool" of water above the plate. The perforated plate could be moved so that the height of the "bubbly pool" could be varied. There is a drain at the bottom of the test section to prevent water level from building up below the plate. A beam scale is placed at the drain to measure the flow of water that penetrates through the plate. Air or steam that was not condensed on the injected water exited at the top of the test apparatus. The test simulated for this assessment used a 15-hole plate with a "bubbly pool" height of 267 mm (10.5 in.). The test was conducted at atmospheric pressure. A schematic of the test facility is shown in Figure 7-48. Figure 7-48 Schematic of Bankoff Counter Current Flow Apparatus (from Reference 68) 7.2.10.2 Phenomenon Addressed The phenomenon addressed with the Bankoff assessment case is CCFL at pressurizer baffle plate and upper core plate (UCP) (or top nozzle). 7.2.10.3 Experimental Procedure The test was conducted by establishing a water inlet flow rate and then increasing the air flow rate in a stepwise manner. The rate of water flow through the perforated plate was measured by weighing the flow out of the bottom of the test section. The test was concluded when the air flow was sufficient to prevent water downflow through the perforated plate. © Copyright 2022 by NuScale Power, LLC 232

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.2.10.4 Assessment Results Figure 7-49 compares predicted vapor superficial velocity versus predicted liquid superficial velocity. The comparison shows that the predictions are in excellent agreement with the experimental data. Figure 7-49 Superficial Vapor Velocity Versus Superficial Liquid Velocity 7.2.11 Marviken Critical Flow Test 22 and 24 The Marviken critical flow tests (CFTs) (Reference 70 and Reference 71) were conducted to characterize the conditions of blowdown given ((

                                                                                              }}2(a),(b),(c) 7.2.11.1           Facility Description

((

                                                                                         }}2(a),(b),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                            }}2(a),(b),(c)

Figure 7-50 Schematic of the Marviken Pressure Vessel ((

                                                                                             }}2(a),(b),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-51 Discharge Pipe Dimensions and Instrument Locations ((

                                                                                          }}2(a),(b),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.2.11.2 Experimental Procedure ((

                            }}2(a),(b),(c) 7.2.11.3           Phenomenon Addressed The phenomenon addressed with the Marviken assessment is two-phase and single-phase choked flow.

7.2.11.4 Special Analysis Techniques ((

                                    }}2(a),(b),(c) 7.2.11.5           Assessment Results 7.2.11.5.1             Comparison to Marviken Critical Flow Test-22

((

                                                                                           }}2(a),(b),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                      }}2(a),(b),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-52 Measured Versus Calculated Mass Flow Rate for Marviken Critical Flow Test 22 ((

                                                                                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-53 Marviken Critical Flow Test 22 Comparison to Calculated Mixture Density ((

                                                                                                    }}2(a),(c) 7.2.11.5.2             Comparison to Marviken Critical Flow Test-24

((

                                                    }}2(a),(b),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3

3. ((

                                                }}2(a),(b),(c)

Figure 7-54 Measured Versus Calculated Mass Flow Rate for Marviken Critical Flow Test 24 ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-55 Marviken Critical Flow Test 24 Mixture Density and Calculated Mixture Density ((

                                                                                                       }}2(a),(c)

Analysis shows that NRELAP5 has the capability to perform critical flow calculations with reasonable-to-excellent agreement to test data. 7.3 NuScale Stern Critical Heat Flux Tests The CHF correlations described in Section 6.11 are assessed against steady state CHF experiments performed by NuScale in the Stern facility. This assessment is presented here. The Stern tests were performed on a preliminary prototypical bundle geometrically comparable to the NuFuel HTP2' design, but with ((

                                                                  }}2(a),(b),(c),ECI The Stern preliminary prototypical bundle tests provide data over wide parameter ranges, which encompass the NPM operating parameter values and can be used to assess the capability of NRELAP5 to predict the onset of the CHF. Key FOMs to assess agreement include the critical power and the critical power ratio as a function of mass flux, pressure, and inlet sub-cooling.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.3.1 Facility Description The Stern CHF tests made use of a 5x5 fuel bundle comprised of ((

                      }}2(a),(b),(c),ECI heated length fuel simulators arranged in three configurations including:

(( }}2(a),(b),(c) (U-1 series) (( }}2(a),(b),(c) (U-2 series) (( }}2(a),(b),(c) A prototypical fuel diameter ((

                                                                     }}2(a),(b),(c),ECI Figure 7-56 U1 & C1 (left) versus U2 (right) radial layout

((

                                                                                                      }}2(a),(b),(c),ECI An axial layout of the test section with key instrument locations is shown in Figure 7-57. The test section includes a pressure housing, a channel box (flow channel), fuel simulators, spacer grids, and instrumentation. Four spacers are installed within the heated section of the assembly at prototypical locations with a

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 spacer pitch of (( }}2(a),(b),(c),ECI The resistance temperature detectors are used to measure the average inlet and outlet temperatures of the coolant. In addition to the absolute pressure measurements at the inlet and outlet of the test section, there are nine differential pressure transducers installed within the heated section to measure the pressure drop across various axial sections. Figure 7-57 Stern Test Section Axial Layout ((

                                                                                                 }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.3.2 Experimental Procedure At the Stern test facility, the steady state CHF tests were performed in the following manner: loop conditions were established with the heated assembly at a power below the critical power, loop conditions were maintained steady as much as possible while the power was (( }}2(a),(b),(c) until the critical power was reached, the data acquisition program continuously scans the assembly signals and critical power is considered to occur when the ((

                                                                                       }}2(a),(b),(c),ECI when the occurrence of critical power is confirmed the loop conditions were held steady and the steady-state data were recorded, and once the test point was recorded the power was reduced, as necessary, and loop conditions changed for the next test.

7.3.3 Phenomenon Addressed The Stern CHF benchmark assesses the ability of NRELAP5 to predict CHF and ((

                        }}2(a),(c) 7.3.4         Parameter Ranges Assessed The Stern steady state CHF tests were conducted across a systematic range of mass flows, inlet pressures, and inlet sub-cooling. A total of ((                }}2(a),(b),(c),ECI steady state CHF data were collected for pressures ranging from ((
                                                                                       }}2(a),(b),(c),ECI as described in Table 7-5. A series of repeat tests was also performed to determine the repeatability of the test data. A total of ((         }}2(a),(b),(c),ECI repeat test points are identified. Only the ((      }}2(a),(b),(c),ECI high flow data points with mass fluxes greater than ((                                           }}2(a),(b),(c),ECI are excluded from the assessment presented in this document.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 7-5 Range of Stern Steady State Critical Heat Flux Data ((

                                                                                          }}2(a),(b),(c),ECI 7.3.5         Special Analysis Techniques The NRELAP5 model consists of a ((
                            }}2(a),(c) 7.3.6         Assessment Results

((

                                                                                                 }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                       }}2(a),(c)

Figure 7-58 Predicted Versus Measured Stern Power ((

                                                                                             }}2(a),(b),(c),ECI 7.4      NuScale SIET Steam Generator Tests This section addresses assessments performed against experiments conducted under the NuScale testing program at SIET laboratories, in Piacenza, Italy. Two test programs were conducted as described below.

7.4.1 SIET Tests The SIET experimental program is a two test activity with helical coil SG tubes characterized on an electrically heated test section (TF-1), and on a fluid heated test © Copyright 2022 by NuScale Power, LLC 246

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 section (TF-2). The electrically heated test provides detailed in-tube information for the secondary side, while the fluid heated test allows investigation of the general behavior of the tube bundle heated by the primary side fluid. This section deals with the detailed description of the electrically heated test (TF-1) and NRELAP5 assessment results. The electrically heated test section incorporates three full scale coils of the helical coil SG, providing information focused on the SG secondary side. Direct heating of the test section is provided by passing current through the tubes using three different axial heating zones (subcooled, saturated and superheat). (( }}2(a),(b),(c),ECI coils of the electrically heated test section represent the (( }}2(a),(b),(c),ECI coils of the NuScale SG, in terms of diameter, length and angle of inclination, and they allow investigation into the effects of tube curvature on thermal-hydraulic parameters. The (( }}2(a),(b),(c),ECI coil reproduces the (( }}2(a),(b),(c),ECI coil of the NuScale SG, in terms of diameter and length, ((

                                                         }}2(a),(b),(c),ECI 7.4.1.1            Facility Description The main components and loops of the SIET TF-1 facility in the NuScale helical coil SG test configuration are described here. A pump system drives water from a water storage tank to the pre-heating zone where it is brought to the specified operating conditions and sent to a feedwater header. The header feeds the three coils of the test section that can be activated by valves: singularly or two in parallel. Superheated steam exits the test section toward a header connected to the separation and discharge system. A schematic of the test loop is provided in Figure 7-59.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-59 SIET Electrically-Heated Test Instrumentation Diagram ((

                                                                                                 }}2(a),(b),(c),ECI 7.4.1.2            Phenomena Addressed The SIET TF-1 assessment cases addresses ((
                                                                                 }}2(a),(c) 7.4.1.3            Experimental Procedure For adiabatic tests the inlet flowrate is specified along with the outlet pressure for each test point. For diabatic tests the inlet temperature, flowrate, and tube/zone heat flux (by setting the current) are specified along with the outlet pressure. This section only covers diabatic tests.

((

                                }}2(a),(b),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.4.1.4 Special Analysis Techniques The helical coil component used includes the helical coil friction model and heat transfer packages inside the coil (Section 6.7). 7.4.1.5 Assessment Results In general, NRELAP5 predicts the experimental data with reasonable-to-excellent agreement. The following specific conclusions were drawn from the assessment Calculated axial fluid and wall temperatures are within reasonable-to-excellent agreement of data. Calculated single- and two-phase pressure drops along the coil are in reasonable-to-excellent agreement with the test data. Results from two Coil 2 tests are presented first to illustrate variation between predicted and measured wall temperature along the length of the coil. Subsequently, pressure drop for tests on Coil 1 through 3, and fluid temperature and wall temperature for tests on Coil1 are presented. Wall temperature profile for the three heating zones (subcooled, saturated, and superheat) of coil 2 are depicted in Figure 7-60 and Figure 7-61 for diabatic tests TD0015 and TD0003, respectively. From inspection of the wall temperatures, ((

                                              }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-60 Time Averaged Wall Temperature Profile for Coil 2 Test TD0015 ((

                                                                                     }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-61 Time averaged wall temperature profile for coil 2 test TD0003 ((

                                                                                                    }}2(a),(b),(c),ECI Pressure drops for the five sections along the length of Coil 1 (i.e., axial pressure drop) are given in Figure 7-62. The error bands on these figures represent the uncertainty in measurement of pressure drops. Calculated pressure drops over the first ((
                          }}2(a),(b),(c),ECI are predicted with excellent agreement and within the experimental error, as shown by Figure 7-62. Similar results are shown for pressure drops in Coil 2 and Coil 3 in Figure 7-63 and Figure 7-64. In general, NRELAP5 does a reasonable-to-excellent job of predicting the axial pressure drops taking into account that the standard deviation of the experimental data (not shown on plots) is larger than the reported measurement uncertainty (shown on plot).

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-62 SIET Electrically-Heated Test Differential Pressure for all Coil 1 Diabatic Tests ((

                                                                                       }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-63 SIET Electrically-Heated Test Differential Pressure for all Coil 2 Diabatic Tests ((

                                                                                       }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-64 SIET Electrically-Heated Test Differential Pressure for all Coil 3 Diabatic Tests ((

                                                                                               }}2(a),(b),(c),ECI Fluid temperatures for ((                    }}2(a),(b),(c),ECI of coil 1 are depicted in Figure 7-65. The error bands on these figures represent the uncertainty in measurement of fluid temperature. The calculated values are in reasonable-to-excellent agreement with experimental data.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-65 SIET Electrically-Heated Test Fluid Temperatures for all Coil 1 Diabatic Tests ((

                                                                                              }}2(a),(b),(c),ECI Corresponding wall temperatures at several axial locations of coil 1 are depicted in Figure 7-66. The error bands on these figures represent the uncertainty in measurement of wall temperature. Wall temperature results are similar to the corresponding fluid temperature results ((
                                }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-66 SIET Electrically-Heated Test Wall Temperature for all Coil 1 Diabatic Tests ((

                                                                                                      }}2(a),(b),(c),ECI 7.4.2         SIET Fluid-Heated Test The SIET fluid-heated tests were performed in support of the NuScale design development, with particular emphasis on providing experimental data for validation of NRELAP5 for prediction of helical coil SG primary and secondary heat transfer, primary side pressure drop, and secondary side dryout.

7.4.2.1 Facility Description The SIET TF-2 facility consists of a 252 helical coil tube bundle installed inside a pressure vessel. The tube bundle consists of 5 tube banks, simulating the ((

                          }}2(a),(b),(c),ECI All five tube bundles are placed in an annulus, formed by two cylindrical barrels, installed axially within the pressure vessel. The helical coils are wrapped around the inner barrel and kept in position by four supports, ((
                                                                         }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                         }}2(a),(b),(c),ECI Each tube bank is fed by a feed-water vertical header, inside the vessel, that distributes water to each helical tube. Steam from the exit of each tube bank is collected in a steam vertical header and driven outside the vessel top nozzle by pipes.

The primary-side of the test section consists of an inlet riser barrel, connection bellows, internal barrel, pressure vessel dome, free volume (i.e., unoccupied by the tubes) between the internal and external barrels (i.e., annulus), and free volume of the pressure vessel around the inlet riser and connection bellows. Water on the primary side is circulated by pumps and pre-heated by an electric heater before entering the pressure vessel. The pressure on the primary side is maintained using the electrically heated pressurizer. Primary water, entering the pressure vessel from the bottom nozzle, rises through a vertical channel and enters the central cylindrical part of the test section, representing the riser. After reaching the vessel dome, water turns down into the test section annulus to cross the helical coil tube bundle. Exiting at the bottom, water is driven again to the circulation pumps. Instruments are installed for the measurement of primary side mass flow rate, inlet and exit temperatures, pressures and differential pressures. Instruments are installed to measure the secondary side feed water flow rate, feed water temperature, pressures and differential pressure along the tubes, and exit steam temperatures and flow rates. 7.4.2.2 Phenomenon Addressed Adiabatic tests were performed to characterize the primary side pressure losses in the facility. These tests were run without heat input to the primary flow and there was no secondary flow to the coils. Diabatic tests measured pressure drop and heat transfer on both the primary and secondary, and the thermal crisis (dryout) location on the secondary side during heated operation of the coils. These tests characterize the thermal performance of the coils for a range of primary and secondary side inlet flows and temperatures. 7.4.2.3 Experimental Procedure Target boundary conditions are obtained for the diabatic tests. The duration of the data recording for each test was a minimum 300 seconds for the pre-steady state and 150 seconds for the steady state. © Copyright 2022 by NuScale Power, LLC 257

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.4.2.4 Special Analysis Techniques This benchmark assesses ((

                                                                                                         }}2(a),(c) 7.4.2.5            Assessment Results In general, NRELAP5 predicted the experimental data with reasonable-to-excellent agreement. The following specific conclusions were drawn from the assessment:

((

                                                                                  }}2(a),(b),(c),ECI 7.4.2.5.1              Assessment of Adiabatic Experiment Data Adiabatic experimental data from TF-2 testing is used to assess the modeling of primary side friction and form losses. The primary side pressure drop was measured at ((                     }}2(a),(b),(c),ECI Figure 7-67 shows the comparison of predicted and measured primary side pressure drop at all axial elevations for all adiabatic tests. The error bands represent the uncertainty in measurement of pressure drops. Excellent agreement between NRELAP5 predictions and measured test data exists with primary side pressure drop predicted within the measurement uncertainty. Similar results are obtained for other primary side pressure drop measurement elevations. ((
                                                                                                  }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-67 SIET Fluid-Heated Test Adiabatic Primary Differential Pressure ((

                                                                                       }}2(a),(b),(c),ECI 7.4.2.5.2              Primary Side Pressure Drop and Fluid Temperatures of Diabatic Experiments

((

                                      }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-68 SIET Fluid-Heated Test Diabatic Test Primary Differential Pressure ((

                                                                                      }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-69 SIET Fluid-Heated Test Diabatic Test Primary Temperature ((

                                                                                       }}2(a),(b),(c),ECI 7.4.2.5.3              Steam Generator Tube Wall Temperature of Diabatic Experiments

((

                                   }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-70 Comparison of Wall Temperatures in TD0001 (Case 1A) ((

                                                                                      }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-71 Comparison of Wall Temperatures in TD0005 (Case 1A) ((

                                                                                      }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-72 Comparison of Wall Temperatures in TD0015 (Case 1A) ((

                                                                                               }}2(a),(b),(c),ECI 7.4.2.5.4              Secondary Side Fluid Temperature of Diabatic Experiments Figure 7-73 and Figure 7-74 show the comparison of predicted and measured secondary side fluid temperatures at all elevations in Row 3 for selected tests.

The figures also show the predicted and measured primary fluid temperatures. ((

                                                                                  }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-73 Comparison of Primary and Secondary Side Fluid Temperatures in TD0001 (Case 1A) ((

                                                                                  }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-74 Comparison of Primary and Secondary Side Fluid Temperatures in TD0005 (Case 1A) ((

                                                                                              }}2(a),(b),(c),ECI 7.5      NuScale NIST Test Assessment Cases A scaled facility of the NPM was constructed at Oregon State University, referred to as the NIST facility, to assist in validation of the NRELAP5 system thermal-hydraulic code.

The facility is designed to perform various tests, including LOCA tests. The NIST facility consists of the major components in the NPM. These components include: an RPV, helical coil SG system with DHRS, CNV, and cooling pool vessel (CPV) representing the reactor pool. The NIST ECCS connects the RPV to the CNV and consists of two RVVs and two RRVs, each on separate lines. Breaks can be simulated for the RCS lines that connect the RPV to the CNV to simulate piping breaks within the CNV. This system consists of a RCS discharge line, a RCS injection line, and a pressurizer spray supply line. The CVCS is not functional in the NIST facility and is used only for simulation of CVCS line break LOCAs. Instrumentation is included in the facility to capture the response of the system under steady-state and transient situations. The instrumentation includes pressure, differential pressure, water level, mass flow rate, heat flux, and temperature measurements. © Copyright 2022 by NuScale Power, LLC 266

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.5.1 Test Facility Description Due to the unique nature of the NPM design the number of IET facilities suitable for code assessment is limited. The NIST facility was originally conceived at OSU in 2000 as a proof-of-concept testing platform for development of Small Modular Reactor (SMR) technology. During this period it was referred to as the multi-application small light water reactor facility (Reference 72). Although the NuScale design was based on the Multi-Application Small Light Water Reactor (MASLWR), the concept has evolved considerably since the inception of NuScale in 2008. At the time that NuScale was formed, the facility was renamed the NIST facility. The NIST facility is a scaled, non-nuclear reactor that uses electric heater rods to represent the heat produced from fission. It is designed to produce experimental data in support of verification and validation of thermal-hydraulic codes. In 2014 and 2015, the original NIST facility was modified by NuScale to facilitate accurate simulation and to bring the facility in-line with the current NuScale plant design configuration. Following the upgrade, the NIST facility was renamed NIST-1 facility. A scaling analysis was employed for design of the NIST test facility to ensure that the facility design is capable of capturing important plant phenomena with minimal distortions. Further discussion on the NIST facility scaling and distortions is available in Section 8.3.2 and Section 8.3.4. Updates to the NIST facility included in NIST-1 are: ((

                          }}2(a),(b),(c),ECI The updated NIST-1 facility provides a well-scaled representation of the current NuScale reactor design that minimizes distortions and provides the measurements necessary for safety code and reactor design validation.

The NIST-1 facility was upgraded to the NIST-2 facility in 2018. The primary upgrades were to the secondary and DHRS piping systems to increase the maximum allowable working pressure (MAWP). High level updates are listed below: Main steam system MAWP increased to 900 psig Steam generator and DHRS MAWP increased to 1350 psig © Copyright 2022 by NuScale Power, LLC 267

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Most secondary side piping and valves replaced Improved steam flow control valve installed Instrumentation of subsystems that were not upgraded (i.e. ECCS/break lines, CNV, HTP, and CPV) are unchanged. Upgraded subsystems (i.e. secondary and DHRS) have significant changes to instrumentation. Several small upgrades were made to RPV instrumentation. A set of characterization tests were performed after the NIST-2 upgrade to better define certain facility parameters. Parameters investigated include: Characterization Test 1 (CT-01): Primary Forced Flow

                  -    Primary flow meter performance
                  -    Primary loop flow losses
                  -    Primary loop "bypass" flow caused by leakage between HLR instrument penetrations Characterization Test 2 (CT-02): RPV Thermal Characteristics
                  -    Primary vs secondary side energy balance under SS conditions
                  -    Core stored energy release characteristics
                  -    Core outlet radial temperature distribution via multi-point TCs
                  -    RPV heat loss characteristics Characterization Test 3A (CT-03A): Break and ECCS Unchoked Flow
                  -    Single-phase liquid flow losses from RPV to orifice inlet of each break/ECCS line
                  -    Single-phase liquid flow losses across several orifices
                  -    Single-phase liquid flow losses from orifice outlet to CNV inlet of each break/ECCS line Characterization Test 3B (CT-03B): Break and ECCS Orifice Choked Flow
                  -    Choked flow discharge coefficient for several orifices in applicable break/ECCS lines The findings of each characterization test data assessment were incorporated into the development of the NIST-2 base model, as well as improvements to relevant portions of the NIST-1 base model. A schematic of the NIST facility is shown in Figure 7-75.

Note that the acronym NIST refers generically to the NuScale Integral System Test facility. This facility was modified to test specific NPM configurations. The NIST-1 configuration was utilized from 2015 to 2018 and the NIST-2 configuration was used from 2019 to 2022. © Copyright 2022 by NuScale Power, LLC 268

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-75 Schematic of NuScale Integral Test Facility and NRELAP5 Nodalization ((

                                                                                     }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The NIST facility models the NPM at ((

                        }}2(a),(b),(c),ECI scale. There are three vessels in the NIST facility: the RPV, CNV, and CPV as shown in Figure 7-75. Unlike the NPM, the RPV and CNV are not concentric and the CNV is not immersed in the CPV. Rather the RPV and CNV are connected by piping that contains valves that perform the functions of the RRVs, RVVs and breaks as shown in Figure 7-75. This approach enables flow measurements to be made in this piping during testing. The CNV is connected to the CPV through a HTP that is scaled to allow energy transfer to the pool in the same proportion as in the NPM.

Natural circulation flow in the primary circuit is driven by heat input in the core region and heat removal to the SG tubes. Fluid heated in the core region flows upward through the hot leg riser, and then downward around the outside of the SG tubes, the cold leg and the downcomer. The flow then returns to the core through the LP. The core is composed of a (( }}2(a),(b),(c),ECI electric heater rod bundle with a maximum power of (( }}2(a),(b),(c),ECI kW, a power level scaled to simulate decay heat. System pressure is controlled by the pressurizer component which contains heater rods to bring the pressurizer fluid up to saturation temperature. 7.5.1.1 Reactor Pressure Vessel Major internal components in the RPV are the core, hot leg riser, pressurizer, and SG bundle. The pressurizer is at the top, separated from the lower part of the RPV by a perforated pressurizer baffle plate. The upper plenum occupies the region below the pressurizer baffle plate and above the hot leg riser that extends down to the top of the core. The upper annulus between the hot leg riser and the RPV shell contains the helical coil SG tubes. The lower part of the annulus immediately below the SG tubes is the cold leg. The lower annulus at the core elevation is the downcomer, which is separated from the core by the core shell. The LP occupies the bottom of the RPV and hydraulically connects the downcomer and the core. The RPV shells and flanges are covered by ((

                                       }}2(a),(b),(c),ECI 7.5.1.1.1              Reactor Core The RPV houses the core, which is modeled by a ((
                                                                                                       }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                               }}2(a),(b),(c),ECI 7.5.1.1.2              Hot Leg Riser After leaving the core, the flow enters the chimney of the hot leg riser. The hot leg riser extends from above the core shroud to the upper plenum, creating a riser and downcomer configuration to enable natural circulation. The hot leg riser consists of a lower shell, a conical transition, a middle shell containing the flowmeter for the primary circuit, and an upper shell. Flow exits the riser into the upper plenum, which is the space between the hot leg riser outlet and the bottom of the pressurizer baffle plate.

7.5.1.1.3 Upper Plenum After leaving the top of the hot leg riser, the flow enters the upper plenum and is directed radially outward to flow down in the annulus between the riser and the RPV shell. The pressurizer baffle plate separates the upper plenum from the pressurizer. Hydraulic communication between the pressurizer and the RPV occurs via holes located in the pressurizer baffle plate. 7.5.1.1.4 Pressurizer The pressurizer is located above the upper plenum and is in thermal-hydraulic communication with the upper plenum via the pressurizer baffle plate holes. The pressurizer maintains primary system pressure during normal steady-state and transient conditions through the use of three heater elements. Each element has (( }}2(a),(b),(c),ECI of power and is modulated by the facility control system to maintain system pressure. 7.5.1.1.5 Cold Leg Downcomer After leaving the upper plenum, the flow continues downward through the SG section and into the cold leg downcomer region. The cold leg downcomer is the annular space bounded by the RPV shell ID and the hot leg riser outer diameter. When fluid reaches the hot leg riser conical transition shell, the flow area is reduced. Flow exits the cold leg downcomer into the LP before it recirculates back into the core. 7.5.1.1.6 Steam Generators The SG is a helical-coil, once-through heat exchanger consisting of ((

                                                                                            }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 (( }}2(a),(b),(c),ECI that wrap counter to each other in the annular space between the hot leg riser and the RPV shell inner surface. In the NIST facility, the primary coolant is circulated on the outside of the SG tubes, similar to the NPM. Feedwater supplied from the feedwater storage tank is pumped through the SG coils by a regenerative turbine pump. Pressure in the secondary side is regulated by a pneumatically operated variable position valve located in the steam line portion of the flow loop. 7.5.1.1.7 Lower Plenum The LP is the region bounded between the tubesheet and the lower core flow plate. The LP provides the connection between the downcomer and the core, thus completing the RPV flow loop. 7.5.1.2 Containment The CNV, representing the cavity volume between the RPV outer surface and the containment inside surface, is conjoined to the CPV and thermally separated by a scaled HTP. For the NPM, the RPV is located inside containment. However, with the NIST facility, to maintain both volume and surface area scaling similitude, as well as allow proper instrumentation, the RPV is thermal-hydraulically separated from the CNV. The CNV models the scaled condensation heat transfer surface between the CNV and CPV. Fluid in the CPV, which is at ambient pressures, models the scaled volume in which an NPM CNV is submerged. 7.5.1.3 Cooling Pool Vessel The CPV has a set of four ports allowing for the installation of one of three decay heat removal heat exchangers. The baseline configuration is with a ((

                                                     }}2(a),(b),(c) 7.5.1.4            Emergency Core Cooling System and Chemical and Volume Control System Lines Eight lines connect the RPV to the CNV. Five of these lines belong to the facility ECCS, and the other three are part of the CVCS. As part of the ECCS, there are two independent reactor vent lines near the top of the pressurizer section and two reactor recirculation lines in the lower downcomer of the RPV. The fifth ECCS line is an SET line that also models reactor recirculation. For the CVCS, two lines penetrate the vessel near the bottom of the SG. One of these lines penetrates both the vessel wall and the hot leg riser, simulating the make-up line into the hot

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 leg. The other CVCS line connects to the cold leg and penetrates only the RPV wall. This line represents the facility CVCS discharge break line. A third CVCS line between the RPV and CNV is located at the top of the pressurizer and functions as an analogy for the pressurizer spray supply line. Each line has a pneumatic isolation valve that is actuated through the test facility control system. Any lines that are not installed use a blank flange for isolation. 7.5.1.5 Facility Instrumentation and Control Instrumentation is used throughout the NIST facility to measure the thermal-hydraulic behavior during steady-state and transient operations. The following information is recorded by the DACS: ((

                                                             }}2(a),(b),(c),ECI The data generated and collected by the facility DACS are used to validate applicability of the NRELAP5 thermal-hydraulic code for LOCA analysis.

7.5.1.6 Integral Effects LOCA Test Procedure Prior to startup, a valve and switch lineup is performed to place the facility in the desired configuration for the upcoming test. The break line modeling the break location specified for the test is connected between the RPV and its associated CNV penetration. To prevent an accidental actuation of an incorrect break valve a blind flange is installed in all other break lines. Orifices with specified diameters are installed in the RVV and RRV lines to model the number of valves that are to open when ECCS actuates. © Copyright 2022 by NuScale Power, LLC 273

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Because the NIST facility has a nominal operating pressure of ((

                                   }}2(a),(b),(c),ECI that is less than the NPM pressure of 1,850 psia (12.76 MPa), the test in the NIST facility simulates the NPM transient in progress.

Specifically, the RPV and CNV fluid masses in NIST are scaled such that they are (( }}2(a),(b),(c),ECI that of the RPV and CNV fluid masses in the NPM at a corresponding pressure of (( }}2(a),(b),(c),ECI Thus the initial RCS mass inventory and pressure are preserved on a scaled basis and fluid property similitude is maintained throughout the transient. As part of the NIST LOCA tests, ((

                                }}2(a),(b),(c),ECI 7.5.2         Facility NRELAP5 Model The NRELAP5 model of the NIST facility is constructed to ((
                                                                                                     }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                 }}2(a),(c) These model features are shown in the NRELAP5 nodalization shown in Figure 7-75.

7.5.3 Facility Test Matrix The NIST facility is used to perform design certification IETs and SETs for the purpose of validating NuScale computer codes, model development and assessment, correlation development, verifying compliance with design requirements, demonstrating design features and capabilities, and addressing regulatory concerns. This section briefly describes the test matrix for the NIST facility. Descriptions of tests used for NRELAP5 code validation are provided in Table 7-6. These are the NIST tests that are the essential subset of tests required to validate NRELAP5 for NPM LOCA calculations. Each of the tests indicate whether they were performed under the NIST-1 or NIST-2 configuration. DHRS was not initiated for any of the testing. Table 7-6 Facility High Priority Tests for NRELAP5 Code Validation ((

                                                                                               }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 7-6 Facility High Priority Tests for NRELAP5 Code Validation (Continued) ((

                                                                                    }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 7-6 Facility High Priority Tests for NRELAP5 Code Validation (Continued) ((

                                                                                               }}2(a),(b),(c),ECI Tests NIST-1 HP-06, HP-07 and NIST-2 Runs 1, 2 and 5 are the IETs that are used for validating the NRELAP5 EM for LOCA applications. Tests NIST-1 HP-09, HP-43 and HP-49 and NIST-2 Runs 3, 4, 6 and 7 were performed to support the extension of LOCA EM for the analysis of transients initiated due to inadvertent opening of RPV ECCS valves. NIST-2 IORV Runs 3 and 6 were specifically designed to evaluate Phase 0 phenomena (first 10 seconds of the IORV events). Further discussion on the NRELAP5 validation results against these tests is provided in Section 7.5.4 through Section 7.5.11. These tests also support the containment response analysis methodology.

7.5.4 Separate Effect High Pressure Condensation Tests (NIST-1) The NIST-1 facility HP-02 test is used to assess the capability of NRELAP5 to predict condensation rates at high pressure test conditions by comparing experimental data and NRELAP5 predictions. 7.5.4.1 Facility Description The HP-02 test is an SET performed at the NIST-1 facility. The test involves injecting steam at known conditions into the CNV and measuring the CNV pressure, level, and temperature response. Only the CNV, CPV, and interconnecting HTP are important to this test. During testing, the RPV was pressurized and heated using core heat to supply superheated steam from the SG to the CNV at the desired mass flow rate. The feedwater flowrate was measured with individual Coriolis flowmeters to each of the three SG inlet tube banks. Also, one Coriolis meter measured the total SG © Copyright 2022 by NuScale Power, LLC 277

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 feedwater inlet flow and one vortex flowmeter measured the total steam flow at the SG exit. The Coriolis flowmeter measuring the combined inlet flow was used as a mass flow boundary condition in the NRELAP5 model as it provided the most stable flow measurement with the lowest measurement uncertainty. 7.5.4.2 Phenomenon Addressed The pertinent phenomena validated with the NIST-1 facility HP-02 assessment are ((

                                            }}2(a),(b),(c),ECI 7.5.4.3            Experimental Procedure Initial steam conditions in the CNV were obtained by first operating the NIST-1 facility in its normal mode, heating the RPV with core heaters with heat rejection through the SG to the environment. The SG feedwater flowrate, core power, and steam exit pressure were established to obtain the desired conditions for steam.

Once the desired conditions were established, steam was diverted from the stack (rejected to the environment) to the CNV. Five tests were run to evaluate steady-state condensation at varying CNV pressures. For each test, superheated steam was discharged into the CNV until the CNV target pressure was reached, after which the inlet steam flow was ramped down in an effort to achieve steady state conditions at the target pressure. Steady steam inlet conditions were maintained through the injection period. After steam was injected into the CNV, condensation occurred on the HTP. Condensation energy was then thermally conducted through the HTP and convected into the CPV. 7.5.4.4 Parameter Ranges Assessed Test conditions were selected to obtain condensation data at various CNV pressures. Five tests were conducted at steady CNV pressure varying from ((

                                                               }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.5.4.5 Assessment Results The HP-02 test-data were compared to NRELAP5 predictions designed to simulate the test conditions and test procedures in effect during the experiment. HP-02 test data trends were well predicted by NRELAP5 with reasonable-to-excellent agreement for condensation rates at pressures ranging (( }}2(a),(b),(c),ECI NRELAP5 has demonstrated its capability to predict CNV level, CNV pressure, CNV temperature, and CPV temperature with reasonable-to-excellent agreement. The following subsections provide a brief summary of the results for three HP-02 runs analyzed. 7.5.4.5.1 HP-02 Run 1 Results Both the CNV pressure and level responses for Run 1 depicted in Figure 7-76 and Figure 7-77 are in reasonable-to-excellent agreement with the data. The NRELAP5-simulated pressure peak occurs at the same time as the data; reaching a maximum of ((

                            }}2(a),(b),(c)

The CNV and CPV fluid temperatures predicted by NRELAP5 are in excellent agreement with the data. Figure 7-78 and Figure 7-79 show that the predictions closely following the data trend and magnitude during the earlier transient as well as the steady-state period. © Copyright 2022 by NuScale Power, LLC 279

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-76 HP-02 Run 1 Containment Vessel Pressure Response ((

                                                                                      }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-77 HP-02 Run 1 Containment Vessel Collapsed Level Response ((

                                                                                     }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-78 HP-02 Run 1 Upper Containment Vessel Fluid Temperature Response (in vapor space) ((

                                                                                  }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-79 HP-02 Run 1 Upper Cooling Pool Vessel Temperature Response ((

                                                                                                   }}2(a),(b),(c),ECI 7.5.4.5.2              HP-02 Run 2 Results During run 2 a maximum pressure of ((
                                      }}2(a),(b),(c),ECI was reached. NRELAP5 is in reasonable-to-excellent agreement with the experimental data for CNV pressure, as shown in Figure 7-80. NRELAP5 predicts the general trends for level, with reasonable-to-excellent agreement to data (Figure 7-81), but slightly underpredicts collapsed level.

The NRELAP5 containment vessel and CPV temperatures shown in Figure 7-82 and Figure 7-83 are in excellent agreement with the data, closely following the trend and lying almost entirely within the instrument uncertainty. © Copyright 2022 by NuScale Power, LLC 283

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-80 HP-02 Run 2 Containment Vessel Pressure Response ((

                                                                                      }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-81 HP-02 Run 2 Containment Vessel Collapsed Level Response ((

                                                                                     }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-82 HP-02 Run 2 Upper Containment Vessel Fluid Temperature Response (in vapor space) ((

                                                                                  }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-83 HP-02 Run 2 Upper Cooling Pool Temperature Response ((

                                                                                    }}2(a),(b),(c),ECI 7.5.4.5.3              HP-02 Run 3 Results

((

                                      }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-84 HP-02 Run 3 Containment Vessel Pressure Response ((

                                                                                      }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-85 HP-02 Run 3 Containment Vessel Collapsed Level Response ((

                                                                                     }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-86 HP-02 Run 3 Upper Containment Vessel Fluid Temperature Response (in vapor space) ((

                                                                                  }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-87 HP-02 Run 3 Upper Cooling Pool Temperature Response ((

                                                                                              }}2(a),(b),(c),ECI Based on this assessment, NRELAP5 has demonstrated its capability to predict CNV level, CNV pressure, CNV temperature, and CPV temperature with reasonable-to- excellent agreement for high pressure condensation conditions.

7.5.5 Natural Circulation Test at Power The NIST-1 test HP-05 was used to assess the capability of NRELAP5 to predict natural circulation flow at various core powers and test conditions by comparing experimental data and NRELAP5 predictions. 7.5.5.1 Facility Description The HP-05 test configuration uses the RPV and SG to drive steady-state natural circulation within the RPV at various core heater rod power levels. Core heater rods supply energy to heat the working fluid which, due to buoyancy forces, travels up the riser entering the upper plenum. The fluid then turns 180 degrees and passes over the integrated helical coil SG, exchanging energy to the secondary side. The primary working fluid exits the SG traveling downward © Copyright 2022 by NuScale Power, LLC 291

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 through the downcomer, entering the LP where another 180 degree turn (upward) is made into the entrance of the electrically heated core. Various instruments measure differential pressures, flow, temperatures, pressures, and heater power to assess the loop flowrate and pressure losses. 7.5.5.2 Phenomenon Addressed The pertinent phenomena addressed with the HP-05 assessment case are ((

                                   }}2(a),(c) 7.5.5.3            Experimental Procedure The HP-05 experiment consists of inducing a core power ramp at a constant RPV pressure of approximately ((                                   }}2(a),(b),(c),ECI and a secondary-side pressure of approximately ((                                        }}2(a),(b),(c),ECI Differential pressures around the primary loop were measured to characterize the pressure drops due to form and friction losses. The mass flow rate in the riser and fluid temperatures around the loop are measured. To facilitate comparing to code predictions the core power and temperature rise across the core are used to calculate a theoretical flowrate based on an energy balance.

Test HP-05 initiates from a power of (( }}2(a),(b),(c),ECI, at a pressure of (( }}2(a),(b),(c),ECI, and the steady-state natural circulation flow condition. Once steady-state conditions are achieved, ((

                                                                                      }}2(a),(b),(c),ECI 7.5.5.4            Special Analysis Techniques

((

                              }}2(a),(c) The global response was then confirmed by comparing the experimental loop flow rate to that predicted by NRELAP5.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.5.5.5 Parameter Ranges Assessed ((

                             }}2(a),(b),(c) 7.5.5.6            Assessment Results

((

                                            }}2(a),(b),(c) The NRELAP5 mass flow signal is taken from the same location. The NRELAP5 prediction is closely aligned with the data and shows excellent agreement, with the exception of the behavior demonstrated at the lowest core power level, where reasonable agreement is obtained. At the lower power level, facility constraints on the secondary side made it difficult to obtain steady state conditions.

((

                                                                     }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-88 HP-05 NIST-1 Averaged Mass Flowrate and NRELAP5 Results ((

                                                                                           }}2(a),(b),(c),ECI The core inlet temperature was measured in the LP upstream of where the fluid enters the core. The NRELAP5 signal is taken from the same region.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-89 HP-05 NIST-1 Averaged Core Inlet Temperature and NRELAP5 Results ((

                                                                                               }}2(a),(b),(c),ECI Core outlet temperature was measured in the riser near the core exit. The NRELAP5 signal is located in the same region. Each time the core power is lowered the hot leg temperature first falls and then recovers. In the data, the temperature usually over-shoots the previous steady state value prior to settling down at the next steady state value. Except for these power transients, the data and NRELAP5 predictions are in excellent agreement as demonstrated in Figure 7-90 except the oscillations observed in NRELAP5 at low power/flow conditions.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-90 HP-05 NIST-1 Averaged Core Outlet Temperature and NRELAP5 Results ((

                                                                                              }}2(a),(b),(c),ECI Based on this assessment, NRELAP5 has demonstrated its capability to predict primary flow rate, core inlet temperature, and core outlet temperature with reasonable-to-excellent agreement for natural circulation flow conditions.

7.5.6 RCS Discharge Line Loss-of-Coolant Accident Integral Effects Tests (NIST-1) The HP-06 test was used to assess the capability of NRELAP5 to predict the integral response of the NIST-1 facility for a single-ended discharge line break inside containment. The discharge line and valve connect the downcomer side of the RPV to the CNV. The HP-06b test was similar to the HP-06 test, with the exception of the core power. This test was performed to assess the impact of core power on the progression of the LOCA. ((

                                        }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.5.6.1 Facility Description The entire NIST-1 facility except for the DHRS was used for this IET, including: the SG was active to remove heat from the primary side and drive natural circulation in conjunction with the electrically heated core during the steady state ((

                                                        }}2(a),(b),(c) the CPV was filled to accept rejected heat from the HTP In addition, ((
                                      }}2(a),(b),(c) 7.5.6.2            Phenomenon Addressed The HP-06 and HP-06b tests are IETs modeling a single-ended discharge line break inside containment. The purpose of these IETs is to assess the integral response of the scaled NIST-1 facility. The pertinent phenomena addressed by these tests are:

((

                                       }}2(a),(b),(c),ECI 7.5.6.3            Experimental Procedure The IET test procedure is described in Section 7.5.1.6. When the CNV pressure reached the specified CNV break initiation pressure, the CVCS break valve was opened, initiating the transient.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 open signals are sent to the RRVs and the RVVs (ECCS). The opening of the ECCS valves causes a large amount of mass and energy transfer to occur between the RPV and the CNV over a short period of time. The CNV pressurization and heat-up occurs rapidly, followed by a long depressurization and cooldown profile. Test data were recorded for an extended period of time, well into the long-term cooling phase. 7.5.6.4 Special Analysis Techniques The RCS discharge line orifice has a length of approximately ((

                             }}2(a),(b),(c),ECI and an ID of approximately ((
                                }}2(a),(b),(c),ECI Thus, the orifice has an L/D ratio roughly equal to

(( }}2(a),(b),(c),ECI Analysis indicates that a NRELAP5 discharge coefficient near (( }}2(a),(b),(c),ECI produces reasonable agreement with the break flow test data. The ((

                                                                                                                }}2(a),(c) 7.5.6.5            Assessment Results (HP-06)

The NRELAP5 transient model is designed to simulate initial test conditions and includes logic that follows facility controls and test procedures. The NRELAP5-calculated RCS discharge line break mass flow rate is shown in Figure 7-91 with a peak flowrate of approximately (( }}2(a),(b),(c),ECI lbm/s. For this experiment, the break mass flow rate was not measured. The calculated break flow rate is reasonable because the differential pressure across the RCS discharge line orifice (Figure 7-92), the RPV level response (Figure 7-95), the CNV level response (Figure 7-96), the RPV pressure response (Figure 7-99), and the CNV pressure response (Figure 7-97) are all in excellent agreement. The NIST-1 v-cone flowmeter (measuring primary loop flowrate) is designed for positive single-phase liquid conditions. During the HP-06 test, two-phase conditions occur at the location of the v-cone meter. Figure 7-93 shows that NRELAP5 predicts the RPV primary-flow coast-down after break initiation with reasonable accuracy. The measured RPV mass flow rate after approximately 29 seconds post-test initiation is more uncertain due to potential for two-phase conditions at the v-cone meter. The pressurizer level is compared in Figure 7-94. The comparisons show reasonable-to-excellent agreement. NRELAP5 predicts complete draining of the pressurizer at about (( }}2(a),(b),(c),ECI NRELAP5 provides reasonable-to-excellent agreement for level response in the RPV and CNV as shown in Figure 7-95 and Figure 7-96. The CNV peak pressure © Copyright 2022 by NuScale Power, LLC 298

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 and pressure response are also predicted with excellent agreement to data as shown in Figure 7-97 and Figure 7-98. The timing of ECCS actuation is predicted with reasonable-to-excellent agreement to the test data. Primary pressure response is predicted with reasonable-to-excellent agreement (Figure 7-99). Figure 7-91 NIST-1 HP-06 NRELAP5 Chemical and Volume Control System Discharge Line Break Mass Flow Rate ((

                                                                                             }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-92 NIST-1 HP-06 Break Orifice Differential Pressure ((

                                                                                          }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-93 NIST-1 HP-06 Primary Mass Flow Rate ((

                                                                                         }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-94 NIST-1 HP-06 Pressurizer Level Comparison ((

                                                                                         }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-95 NIST-1 HP-06 Reactor Pressure Vessel Level Comparison ((

                                                                                      }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-96 NIST-1 HP-06 Containment Vessel Level Comparison ((

                                                                                      }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-97 HP-06 Containment Vessel Pressure Comparison ((

                                                                                       }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-98 NIST-1 HP-06 Containment Vessel Pressure Comparison ((

                                                                                      }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-99 NIST-1 HP-06 Primary Pressure Comparison ((

                                                                                            }}2(a),(b),(c),ECI 7.5.6.6            Assessment Results (HP-06b)

((

                                         }}2(a),(b),(c),ECI Other HP-06b initial and boundary conditions were similar to the HP-06 test

((

                                                                                        }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                              }}2(a),(b),(c),ECI Figure 7-100 Comparison of Core Power in HP-06 and HP-06b Tests with the NuScale Power Module Decay Power after Reactor Trip (scaled)

((

                                                                                            }}2(a),(b),(c),ECI Figure 7-101 and Figure 7-102 show the comparisons of predicted and measured RPV and CNV pressures, respectively. Similar comparisons for the RPV and CNV levels are shown in Figure 7-103 and Figure 7-104, respectively. Overall, NRELAP5 predicted the HP-06b data with reasonable-to-excellent agreement.

Figure 7-105 and Figure 7-106 show the comparisons of measured RPV and CNV pressures in HP-06 and HP-06b tests, respectively. Similar comparisons for the measured RPV and CNV levels are shown in Figure 7-107 and Figure 7-108, respectively. ((

                           }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-101 NIST-1 HP-06b Primary Pressure Comparison ((

                                                                                       }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-102 NIST-1 HP-06b Containment Vessel Pressure Comparison ((

                                                                                     }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-103 NIST-1 HP-06b Reactor Pressure Vessel Level Comparison ((

                                                                                     }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-104 NIST-1 HP-06b Containment Vessel Level Comparison ((

                                                                                     }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-105 Comparison of NIST-1 HP-06 and HP-06b Reactor Pressure Vessel Pressure ((

                                                                                }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-106 Comparison of NIST-1 HP-06 and HP-06b Containment Vessel Pressure ((

                                                                                }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-107 Comparison of NIST-1 HP-06 and HP-06b Reactor Pressure Vessel Level ((

                                                                                 }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-108 Comparison of NIST-1 HP-06 and HP-06b Containment Vessel Level ((

                                                                                                 }}2(a),(b),(c),ECI 7.5.7         NIST-1 Pressurizer Spray Supply Line Loss-of-Coolant Accident Integral Effects Test The HP-07 test was used to assess the capability of NRELAP5 to predict the integral response of the NIST-1 facility modeling a single-ended pressurizer spray supply line break inside containment.

7.5.7.1 Facility Description The entire NIST-1 facility, except for the DHRS, was used for this IET. The SG was active to remove heat from the primary side and drive natural circulation in conjunction with the electrically heated core during steady state. ((

                                       }}2(a),(b),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The CPV was filled to accept rejected heat from the HTP. 7.5.7.2 Phenomenon Addressed The phenomena addressed in the test facility HP-07 test are same as in the HP-06 test (Section 7.5.6.2). 7.5.7.3 Experimental Procedure The LOCA test procedure is discussed in Section 7.5.1.6. When the CNV pressure reached the specified CNV break initiation pressure, the pressurizer spray supply line break valve was opened, initiating the transient. Within the NIST-1 facility, the ECCS actuation occurs when the compensated level in the RPV downcomer reaches a pre-specified value. Once this occurs, open signals are sent to the RRVs and the RVVs. The opening of the ECCS valves causes a large amount of mass and energy transfer to occur between the RPV and the CNV over a short period of time. Containment vessel pressurization and heat-up occurs rapidly, followed by a long depressurization and cooldown profile. Test data were recorded for an extended period of time, well into the long-term cooling phase. 7.5.7.4 Special Analysis Techniques The pressurizer spray supply line orifice has a length of approximately ((

                                                                                 }}2(a),(b),(c),ECI. Thus, the orifice has an L/D ratio roughly equal to ((       }}2(a),(b),(c),ECI. Analysis indicates that a NRELAP5 discharge coefficient near ((             }}2(a),(b),(c),ECI produces the best overall agreement with the break flow test data.

The ((

                                                                                                        }}2(a),(c) 7.5.7.5            Assessment Results Figure 7-109 shows the comparison of core power in the HP-07 test to the scaled NPM total power after reactor trip (i.e., fission product decay, actinide decay, and fission power). The HP-07 power is approximately representative ((
                            }}2(a),(b),(c),ECI of the NPM power.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-109 Comparison of Core Power in HP-07 with the NuScale Power Module Power (fission and decay) after Reactor Trip (scaled) ((

                                                                                               }}2(a),(b),(c),ECI The break flowrate predicted by NRELAP5 (Figure 7-110) provided results with excellent agreement to data with the use of the ((                      }}2(a),(c) choking model and a discharge coefficient of ((        }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-110 NIST-1 HP-07 Pressurizer Spray Supply Line Break Discharge Mass Flow Rate ((

                                                                                   }}2(a),(b),(c),ECI

((

                                                                                }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                  }}2(a),(b),(c),ECI Figure 7-111 NIST-1 HP-07 Primary Mass Flow Rate

((

                                                                                         }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-112 NIST-1 HP-07 Reactor Pressure Vessel Level Response Comparison with Data ((

                                                                                   }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-113 NIST-1 HP-07 Containment Vessel Level Response ((

                                                                                       }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-114 NIST-1 HP-07 Containment Vessel Pressure Comparison ((

                                                                                     }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-115 NIST-1 HP-07 Primary Pressure Comparison ((

                                                                                                 }}2(a),(b),(c),ECI 7.5.8         Spurious Reactor Vent Valve Opening Tests The HP-09 test was used to assess the capability of NRELAP5 to predict the integral response of the NIST-1 facility to inadvertent depressurization of the RPV initiated by RVV spurious opening without DHRS. Furthermore, this test also provided bounding depressurization rate for a LOCA initiated by break from pressurizer gas space. This test was repeated with a larger flowpath to simulate three RVVs opening in HP-43.

7.5.8.1 Facility Description The entirety of the NIST-1 facility except for the DHRS was used for this IET. The SG was active to remove heat from the primary side and drive natural circulation in conjunction with the electrically heated core during the steady state. ((

                                       }}2(a),(b),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                                             }}2(a),(b),(c)

The CPV was filled to accept rejected heat from the HTP. 7.5.8.2 Phenomenon Addressed The phenomena addressed in the NIST-1 HP-09 and HP-43 tests are the same as in the HP-06 and HP-07 IETs (Section 7.5.6.2). 7.5.8.3 Experimental Procedure The experimental procedure is consistent with the LOCA test procedure discussed in Section 7.5.1.6. 7.5.8.4 Special Analysis Techniques The ((

                        }}2(a),(c) at the valve orifice. Furthermore, the modified PV term was activated at the valve orifice. The Bankoff CCFL model was applied at the pressurizer baffle plate.

Analysis indicates that a NRELAP5 discharge coefficient near (( }}2(a),(b),(c),ECI produces the best overall agreement with the valve flow test data. 7.5.8.5 Assessment Results Figure 7-116 shows the comparison of core power in the HP-09 test to the scaled NPM total power after reactor trip (i.e., fission product decay, actinide decay, and fission power). The HP-09 core power is ((

                             }}2(a),(b),(c),ECI of the power in the NPM.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-116 Comparison of HP-09 Core Power with the Scaled NuScale Power Module Fission and Decay Power ((

                                                                                              }}2(a),(b),(c),ECI Figure 7-117 through Figure 7-124 compare NIST-1 HP09 test data with the NRELAP5 transient response. The calculated RVV mass flow rate is shown in Figure 7-117 with a peak flowrate of approximately ((        }}2(a),(b),(c),ECI lbm/s.

The mass flow rate is over-predicted during the first (( }}2(a),(b),(c),ECI seconds of the transient. Thereafter, the calculated flow shows excellent agreement with the measured flow rate. © Copyright 2022 by NuScale Power, LLC 326

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-117 NIST-1 HP-09 Valve Mass Flow Rate ((

                                                                                                        }}2(a),(b),(c),ECI The calculated pressurizer pressure is compared to data in Figure 7-118. The calculated pressure shows excellent agreement with the data, including the time of ECCS initiation. An examination of the first ((               }}2(a),(b),(c),ECI seconds of the RPV pressure (Figure 7-119) shows that the NRELAP5-predicted pressure is higher than the measured pressure.

Figure 7-120 compares the NRELAP5-predicted and NIST-1-measured CNV pressure response. Figure 7-121 shows the short-term response. The peak pressure from data and model are (( }}2(a),(b),(c),ECI psia and (( }}2(a),(b),(c),ECI psia, respectively. The comparison shows reasonable-to-excellent agreement with the measured data. As with the RPV pressure response, after about (( }}2(a),(b),(c),ECI seconds, the CNV pressure is under-predicted. The trends of the data are well represented in the calculation. © Copyright 2022 by NuScale Power, LLC 327

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-118 NIST-1 HP-09 Pressurizer Pressure Comparison ((

                                                                                        }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-119 NIST-1 HP-09 Pressurizer Pressure Comparison - 500 Seconds ((

                                                                                   }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-120 NIST-1 HP-09 Containment Vessel Pressure Comparison ((

                                                                                     }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-121 NIST-1 HP-09 Containment Vessel Pressure Comparison - 500 Seconds ((

                                                                                                             }}2(a),(b),(c),ECI The pressurizer and RPV levels are compared in Figure 7-122 and Figure 7-123, respectively. The comparisons show reasonable-to-excellent agreement. Note that the code-to-data comparison presented in Figure 7-122 shows that NIST-1 pressurizer draining fully occurs between ((                                      }}2(a),(b),(c),ECI seconds, i.e., when the data (LDP-1401_calc) reaches a value of approximately

(( }}2(a),(b),(c),ECI inches, the low range of the measurement. When the data are extrapolated out past (( }}2(a),(b),(c),ECI, it appears full draining of the pressurizer occurs at about (( }}2(a),(b),(c),ECI seconds. NRELAP5 predicts pressurizer draining to fully occur at about (( }}2(a),(b),(c),ECI seconds. NRELAP5 predicts ((

                                               }}2(a),(b),(c),ECI As shown in Figure 7-123 the RPV level prediction is in reasonable-to-excellent agreement with the test data.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-122 NIST-1 HP-09 Pressurizer Level Comparison ((

                                                                                         }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-123 NIST-1 HP-09 Reactor Pressure Vessel Level Comparison ((

                                                                                     }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-124 NIST-1 HP-09 Reactor Pressure Vessel Level Comparison - 500 Seconds ((

                                                                                            }}2(a),(b),(c),ECI 7.5.8.6            Assessment Results HP-43 Figure 7-125 through Figure 7-130 present HP-43 transient short-term (0-800 seconds) code-to-data comparisons of selected parameters. The FOM comparisons of pressurizer pressure, RPV level, CNV pressure, and CNV level show reasonable-to-excellent agreement. It should be noted that ECCS actuation occurs at approximately 1190 seconds and does not appear in the short-term period plots.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-125 HP-43 Transient Short-Term Pressurizer Pressure Comparison ((

                                                                                            }}2(a),(b),(c),ECI The code-to-data comparison of pressurizer level presented in Figure 7-126 shows reasonable-to-excellent agreement. Note that the ((
                                      }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-126 HP-43 Transient Short-Term Pressurizer Level Code-to-Data Comparison ((

                                                                                            }}2(a),(b),(c),ECI The code-to-data comparison of short-term RPV level presented in Figure 7-127 shows reasonable-to-excellent agreement. It is notable that the timing of the minimum RPV level occurs ((                               }}2(a),(b),(c),ECI in the NRELAP5 results.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-127 HP-43 Transient Short-Term RPV Level Code-to-Data Comparison ((

                                                                                         }}2(a),(b),(c),ECI The code-to-data comparison of CNV pressure presented in Figure 7-128 shows reasonable-to-excellent agreement with a ((                         }}2(a),(b),(c),ECI by NRELAP5.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-128 HP-43 Transient Short-Term CNV Pressure Code-to-Data Comparison ((

                                                                                                          }}2(a),(b),(c),ECI The code-to-data comparison of the spurious RVV orifice mass flow rate presented in Figure 7-129 shows less than reasonable agreement for ((
                                    }}2(a),(b),(c),ECI of the transient, with reasonable predictions thereafter. However, test data in Figure 7-125 (PZR pressure) and Figure 7-128 (CNV pressure) show ((
                                                                   }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-129 HP-43 Transient Short-Term Spurious RVV Orifice Mass Flow Rate Code-to-Data Comparison ((

                                                                                                  }}2(a),(b),(c),ECI The code-to-data comparison of CNV level presented in Figure 7-130 shows reasonable-to-excellent agreement with a slight under-prediction by NRELAP5

(( }}2(a),(b),(c),ECI The trends in CNV pressure and level presented in Figure 7-128 and Figure 7-130 appear to indicate ((

                         }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-130 HP-43 Transient Short-Term CNV Level Code-to-Data Comparison ((

                                                                                                  }}2(a),(b),(c),ECI 7.5.9         Spurious Reactor Recirculation Valve Opening Integral Effects Test 7.5.9.1            Purpose The HP-49 test was performed at the NIST-1 facility and was used to assess the capability of NRELAP5 to predict the integral response of the NIST-1 facility for a spurious reactor recirculation valve (RRV) opening inside containment. The reactor recirculation line (RRL) and RRV connect the downcomer side of the RPV to the CNV.

7.5.9.2 Facility Description The NIST-1 facility is described in Section 7.5.1. The entire NIST-1 facility except for the CVCS, PZR Spray, and DHRS was used for this IET, including: the SG was active to remove heat from the primary side and drive natural circulation in conjunction with the electrically heated core during the steady state period ((

                                        }}2(a),(b),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                             }}2(a),(b),(c) the CPV was filled to accept rejected heat from the HTP 7.5.9.3            Phenomenon Addressed The HP-49 test is an IET modeling a spurious RRV opening into containment. The pertinent phenomena addressed by this test are:

((

                                          }}2(a),(b),(c),ECI 7.5.9.4            Experimental Procedure The experiment test procedure is consistent with the LOCA test procedure described in Section 7.5.1.6. When the CNV pressure reached the specified CNV transient initiation pressure, the spurious RRV was opened, initiating the transient.

Within the NIST-1 facility, the ECCS actuation occurs when the compensated level in the RPV downcomer reads lower than a specified value. Once this occurs, open signals are sent to the remaining RRV and the RVVs. The opening of the ECCS valves causes a large amount of mass and energy transfer to occur between the RPV and the CNV over a short period of time. The CNV pressurization and heat-up occurs rapidly, followed by a long depressurization and cooldown profile. Test data were recorded for an extended period of time, into the long-term cooling phase. 7.5.9.5 Special Analysis Techniques The RRV discharge line orifice has a length of approximately ((

                             }}2(a),(b),(c),ECI and an ID of approximately ((
                               }}2(a),(b),(c),ECI Thus, the orifice has an L/D ratio roughly equal to

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 (( }}2(a),(b),(c),ECI Analysis indicates that an NRELAP5 discharge coefficient near (( }}2(a),(b),(c),ECI produces reasonable agreement with the spurious RRV flow rate inferred from test data, however, the literature determined value of (( }}2(a),(b),(c),ECI is used for the base case assessment. The ((

                                   }}2(a),(c) 7.5.9.6            Assessment Results (HP-49)

The NRELAP5 transient model is designed to simulate initial test conditions and includes logic that follows facility controls and test procedures. For this experiment, the spurious mass flow rate was not measured. The calculated spurious flow rate is reasonable because the differential pressure across the spurious RRV line orifice (Figure 7-158), the RPV level response (Figure 7-161), the CNV level response (Figure 7-162), the RPV pressure response (Figure 7-165), and the CNV pressure response (Figure 7-163) are all in reasonable agreement for the pre-ECCS opening period of the transient. The NIST-1 v-cone flowmeter (measuring primary loop flowrate) is designed for positive single-phase liquid conditions. During the HP-49 test, two-phase conditions occur at the location of the v-cone meter. As shown in Figure 7-159 NRELAP5 captures the RPV primary-flow coast-down period after transient initiation with reasonable accuracy. Note that after ((

                                                                                 }}2(a),(b),(c),ECI The pressurizer level is compared in Figure 7-160. The comparisons show reasonable agreement. NRELAP5 predicts complete draining of the pressurizer at about ((                           }}2(a),(b),(c),ECI NRELAP5 provides reasonable agreement for level response in the RPV and CNV as shown in Figure 7-161 and Figure 7-162. The CNV peak pressure and pressure response are also predicted with reasonable agreement to data as shown in Figure 7-163 and Figure 7-164. The timing of ECCS actuation is predicted with reasonable agreement to the test data. Primary pressure response is predicted with reasonable agreement (Figure 7-165).

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-131 NIST-1 HP-49 Spurious Orifice Differential Pressure ((

                                                                                        }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-132 NIST-1 HP-49 Primary Mass Flow Rate ((

                                                                                         }}2(a),(b),(c),ECI Figure 7-133 NIST-1 HP-49 Pressurizer Level Comparison

((

                                                                                         }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-134 NIST-1 HP-49 Reactor Pressure Vessel Level Comparison ((

                                                                                      }}2(a),(b),(c),ECI Figure 7-135 NIST-1 HP-49 Containment Vessel Level Comparison

((

                                                                                      }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-136 NIST-1 HP-49 Containment Vessel Peak Pressure Comparison ((

                                                                                     }}2(a),(b),(c),ECI Figure 7-137 NIST-1 HP-49 Containment Vessel Pressure Comparison

((

                                                                                     }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-138 NIST-1 HP-49 Primary Pressure Comparison ((

                                                                                                 }}2(a),(b),(c),ECI 7.5.10        NIST-2 LOCA Integral Effects Test Series This section documents the evaluation of the NIST-2 LOCA test data and the simulated NRELAP5 LOCA transient results. The LOCA test data are intended to provide a better understanding of phenomena related to Phase 1a and 1b for cases of CVCS, RVV, RRV, and high point vent line breaks scenarios.

The NIST-2 LOCA test suite consists of the following characteristics: The following transient scenarios (referred to herein as runs): Run 1 - 100% CVCS discharge line break (similar to NIST-1 HP-06). Run 2 - 100% high point vent line break (similar to NIST-1 HP-07). Run 3 - Inadvertent opening of a single RVV (similar to NIST-1 HP-09). Run 4 - Inadvertent opening of a single RRV (similar to NIST-1 HP-49). Run 5 - 100% High Point Vent line break with ((

                                                               }}2(a),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Run 6 - Inadvertent opening of a single RRV: When ECCS actuates, ((

                            }}2(a),(c),ECI Run 7 - Inadvertent opening of a single RRV ((
                                                                 }}2(a),(c),ECI 7.5.10.1           Facility Description The entirety of the NIST facility except for the DHRS was utilized for this integral effects test, including:

((

                                                                                         }}2(a),(c),ECI 7.5.10.2           Phenomenon Addressed The NIST-2 LOCA test suite addresses integral effects related to Phase 1a and 1b for various break configurations. The pertinent phenomenon addressed with the NIST-2 LOCA assessment are:

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                          }}2(a),(c),ECI Parameters to assess agreement included direct measurements of the CNV pressure, RPV pressure, CNV level, RPV level, primary flowrate, pressurizer level, CPV temperature, CNV temperature, and HTP temperature.

7.5.10.3 Experimental Procedure As part of the NIST-2 LOCA test procedures, ((

                                                                     }}2(a),(c),ECI 7.5.10.4           Parameter Ranges Assessed Table 7-7 through Table 7-9 summarize the characteristic ((
                                                                                            }}2(a),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                             }}2(a),(c),ECI Table 7-7 LOCA Characteristic State-Points Considered for Runs 1 and 3

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                                                                                           }}2(a),(c),ECI Table 7-8 LOCA ECCS Actuation Sequences

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 7-8 LOCA ECCS Actuation Sequences (Continued) ((

                                                                                                }}2(a),(c),ECI Table 7-9 LOCA Characteristic State-Points Considered for Runs 4, 6, and 7

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                                                                                                }}2(a),(c),ECI Table 7-10 LOCA ECCS Actuation Sequences

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                                                                                                }}2(a),(c),ECI 7.5.10.5           Special Analysis Techniques This assessment contains analysis techniques and sensitivities relevant to transient simulation for code-to-data comparisons with primary focus on the

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                    }}2(a),(c),ECI 7.5.10.6           Assessment Results: CVCS Discharge Line Break Run 1 is a liquid space break utilizing the CVCS discharge line. Table 7-11 provides the sequence of events for Run 1. Overall, sequence timings between the data and the simulation match well. For this run, ((
                                                                                  }}2(a),(c),ECI Table 7-11 LOCA Run 1 Sequence of Events

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                                                                                                 }}2(a),(c),ECI Figure 7-139 compares the predicted RPV and CNV pressures with data ((
                                                                                         }}2(a),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                                     }}2(a),(c),ECI Figure 7-139 Run 1 Predicted RPV/CNV Pressure Comparison with Data

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-140 Run 1 Predicted RPV Pressure Comparison with Data for Full HTP Surface Area ((

                                                                                    }}2(a),(c),ECI Figure 7-141 Run 1 Predicted CNV Pressure Comparison with Data for Full HTP Surface Area

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-142 and Figure 7-143 show the RPV and CNV collapsed liquid levels, respectively. The minimum RPV downcomer collapsed level is over predicted by about 5 inches in the simulation. The maximum CNV level is within approximately 2 inches. The minimum/maximum levels are dynamic effects that are mainly effected by the HTP heat transfer area and the fin effect between the CNV/CPV and their connections to the HTP. Figure 7-144 and Figure 7-145 show improvement in the early response of the transient when the full HTP surface area (sensitivity) for heat transfer is available. Figure 7-142 Run 1 Predicted RPV Level Comparison with Data ((

                                                                                                 }}2(a),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-143 Run 1 Predicted CNV Level Comparison with Data ((

                                                                                        }}2(a),(c),ECI Figure 7-144 Run 1 Predicted RPV Level Comparison with Data for Full HTP Surface Area

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-145 Run 1 Predicted CNV Level Comparison with Data for Full HTP Surface Area ((

                                                                                                   }}2(a),(c),ECI Figure 7-146 and Figure 7-147 compare the predicted RPV shell energy and the predicted CNV shell energy with data. In the first 20,000 seconds, the predicted RPV shell energy shows excellent agreement with the data. This suggests that RPV shell heat losses are modeled appropriately with in the model. Likewise, the predicted CNV shell and insulation overall heat losses compare reasonably with the data, indicating that the CNV heat losses are modeled reasonably.

Figure 7-148 compares the predicted HTP energy with data. As observed, RELAP5 is under- predicting the HTP heat loss for the first 8,000 to 10,000 seconds of the transient. Using the reduced HTP surface area reduces the amount of energy transfer through the plate in the early phase of the transient (Phase 1), but tracks well during the second phase of the transient. The effective HTP surface area appears not to have a significant impact in the late transient, but does have significant impact on the early phase, resulting in early transient system pressure and level distortion. The largest impact on using the full HTP surface area, as observed in the sensitivity, is the better response of the HTP energy transfer as observed in Figure 7-149. © Copyright 2022 by NuScale Power, LLC 357

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-146 Run 1 Predicted RPV Shell Energy Comparison ((

                                                                                           }}2(a),(c),ECI Figure 7-147 Run 1 Predicted CNV Shell Energy Comparison

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-148 Run 1 Predicted HTP Energy Transfer Comparison ((

                                                                                               }}2(a),(c),ECI Figure 7-149 Run 1 Predicted HTP Energy Transfer Comparison - Full HTP Surface Area

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                                                                                               }}2(a),(c),ECI With respect to the CPV thermal behavior, Figure 7-150 through Figure 7-152 compares the predicted CPV fluid temperatures with data. Also shown in the

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 figures are the results with the sensitivity (100 percent full HTP area). In the base calculation, the temperatures are under-predicted. However, the sensitivity calculation shows improvement over the base calculation. The under-prediction of the CPV fluid temperatures reflect on the predicted RPV and CNV pressures and temperatures as shown above. Long term cooling behavior (not shown in the figures) shows the predicted temperatures continue to increase and over-predict the CPV fluid temperatures. The over-prediction of the temperatures also impacts the quasi-steady state conditions in the RPV and the CNV, resulting in higher predicted RPV/CNV levels and pressures. Figure 7-153 shows the predicted CPV level against the data. The data show a gradual evaporation of the CPV level, whereas the level increases in the simulation as the fluid thermally expands due to lack of evaporation. At the termination of the transient the CPV level in the experiment declined from 193 inches to about 188 inches, while in the simulation the level increases to about 197 inches. The sensitivity calculation shows the CPV trends with the base case calculation. Figure 7-150 Run 1 Level 1 CPV Fluid Temperature Comparison ((

                                                                                                   }}2(a),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-151 Run 1 Level 5 CPV Fluid Temperature Comparison ((

                                                                                          }}2(a),(c),ECI Figure 7-152 Run 1 Level 6 CPV Fluid Temperature Comparison

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-153 Run 1 CPV Fluid Level Comparison ((

                                                                                                  }}2(a),(c),ECI From Run 1, the following overall observations were made between the NRELAP5 simulation and the data:

Once the RPV and CNV have reached quasi-steady state conditions (at approximately 20,000 seconds and after), the modelled effective HTP surface area has minimal to negligible impact to the RPV, CNV, and HTP metal heat losses. The effective HTP surface area does, however, have considerable impact to the early transient comparisons, namely CNV and RPV pressures and levels. The heat transfer across the HTP showed considerable improvement in the early phase of the transient using the full HTP surface area (sensitivity). The larger HTP surface area resulted in more of the core decay power and stored energy in the RPV and CNV wall to be removed. After about 20,000 seconds, the CPV fluid temperature distortions begin to dominate the transient, resulting in elevated RPV/CNV levels and pressures. These observations apply to all of the LOCA runs simulated. 7.5.10.7 Assessment Results: High Point Vent Line Breaks Runs 2 and 5 are both steam space breaks utilizing the high point vent line. In the test, after the CNV steam bath was performed and the CNV depressurized, another pressurization was performed to 47 psia to bring the CNV to the required © Copyright 2022 by NuScale Power, LLC 362

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 initial conditions. Table 7-12 and Table 7-13 provides the sequence of events for Run 2 and 5, respectively. For Run 2, once the high point vent line initiates the transient, the criterion for actuating the RVV occurs when the RPV downcomer collapsed liquid level decreases below 160.0 inches. The criterion for actuating the RRV occurs when the differential pressure between the RPV and the CNV reaches 450 psid. For Run 5, once the high point vent line initiates the transient, the RVV and RRV lines actuate when the differential pressure between the RPV and the CNV reaches 400 psid. Table 7-12 LOCA Run 2 Sequence of Events ((

                                                                                                    }}2(a),(c),ECI Table 7-13 LOCA Run 5 Sequence of Events

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                                                                                                    }}2(a),(c),ECI Figure 7-154 shows the Run 2 pressure comparisons between the CNV and the PZR during the first 20,000 seconds of the transient. In the earlier phase of the

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 transient there is minimal agreement with the data. However, after about 20,000 seconds of transient time, there is reasonable-to-excellent agreement. Due the HTP distortion, the predicted peak CNV pressure is higher than the measured peak pressure (about 708 psia predicted versus about 648 psia measured). The ability for the HTP to condense the steam in the CNV during the early transient is insufficient compared to the data, resulting in a higher predicted CNV peak pressure. Figure 7-154 Run 2 RPV-CNV Pressure Comparison ((

                                                                                                   }}2(a),(c),ECI Figure 7-155 and Figure 7-156 show the RPV downcomer collapsed liquid level and the CNV collapsed liquid level, respectively. The minimum RPV downcomer collapsed level is over-predicted in the simulation (Table 7-12), whereas the maximum CNV level is within about 2 inches. The minimum/maximum levels are dynamic effects that are mainly effected by the HTP heat transfer area and the fin effect between the CNV/CPV and their connections to the HTP.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-155 Run 2 RPV Level Comparison ((

                                                                                            }}2(a),(c),ECI Figure 7-156 Run 2 CNV Level Comparison

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-157 through Figure 7-159 shows the Run 5 pressure and level comparisons between the CNV and the PZR during the first 20,000 seconds of the transient. The pressure and level response for Run 5 simulation show the same behavior as the Run 2 simulation, i.e. the pressures and levels are over-predicted due to the HTP distortion. With respect to the energy balances in the RPV, CNV and HTP, the predicted results compared to the measured data are similar to the results obtained with the simulation of Run 1. Likewise with the behavior of the CPV for Run 2 and 5. Figure 7-157 Run 5 RPV-CNV Pressure Comparison ((

                                                                                                 }}2(a),(c),ECI Overall, the simulations of Run 2 and Run 5 show that NRELAP5 can reasonably predict the response of the system considering a high point vent line break given different methods for activating the ECCS. Whether it is activating the ECCS by the differential pressure between the RPV and CNV (Run 5) or if it is activating the RVV by RPV level and activation of the RRV by RPV and CNV differential pressure (Run 2), NRELAP5 can reasonably predict the system behavior.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-158 Run 5 RPV Level Comparison ((

                                                                                            }}2(a),(c),ECI Figure 7-159 Run 5 CNV Level Comparison

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.5.10.8 Assessment Results: Inadvertent Opening of an RVV Run 3 is a steam space break utilizing the inadvertent opening of a single RVV. Table 7-14 provides the sequence of events for Run 3 for the experiment and the NRELAP5 simulation. For Run 3, the criterion for actuating the other single RVV occurs when the RPV downcomer liquid level declines below 160.0 inches. The criterion for actuating the RRV occurs when the differential pressure between the RPV and the CNV reaches 450 psid. Table 7-14 LOCA Run 3 Sequence of Events ((

                                                                                                  }}2(a),(c),ECI Figure 7-160 shows the pressure comparisons between the CNV and the PZR during the first 20,000 seconds of the transient. As in Run 1, the earlier phase of the simulation over-predicts the RPV and CNV pressure response. After about 20,000 seconds the predicted pressures lay more in line with the measured data.

The same reason of the over-prediction is due to the ability of the HTP to condense the steam coming from the RPV into the CNV because of the distortion to the HTP heat transfer surface area. Due to the HTP distortion, the predicted peak CNV pressure is about 677 psia compared to the measured peak pressure of about 633 psia). The ability for the HTP to condense the steam in the CNV during the early transient is insufficient compared to the data, resulting in the higher predicted CNV peak pressure. © Copyright 2022 by NuScale Power, LLC 368

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-160 Run 3 RPV-CNV pressure comparison ((

                                                                                                  }}2(a),(c),ECI Figure 7-161 and Figure 7-162 compare the predicted RPV downcomer collapsed liquid level and the CNV collapsed liquid level with data, respectively. The minimum RPV downcomer collapsed level is over predicted, whereas the maximum CNV level is within about an inch of the data. The minimum/maximum levels are dynamic effects that are mainly effected by the HTP heat transfer area and the fin effect between the CNV/CPV and their connections to the HTP.

With respect to the energy balances in the RPV, CNV and HTP, the predicted results compared to the measured data are similar to the results obtained with the simulation of Run 1. Likewise with the behavior of the CPV. The predicted energy balances in the RPV, CNV, and through the HTP generally show reasonable agreement with the data. Early in the transient simulation, the energy transferred through the HTP is under-predicted, which results in the over-prediction of the RPV and CNV pressures and subsequently the RPV and CNV levels. © Copyright 2022 by NuScale Power, LLC 369

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-161 Run 3 RPV level comparison ((

                                                                                              }}2(a),(c),ECI Overall, the simulation of Run 3 shows that NRELAP5 can reasonably predict the response of the system considering a steam space break given different methods for activating the ECCS.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-162 Run 3 CNV Level Comparison ((

                                                                                                  }}2(a),(c),ECI 7.5.10.9           Assessment Results: Inadvertent Opening of an RRV Runs 4, 6, and 7 are variations of an inadvertent opening of an RRV. The variations are in the activation of the ECCS. For Run 4, the RVV line is opened on a downcomer level less than 160 inches and the RRV (unaffected) is opened on a RPV-CNV differential pressure of 450 psid. For Run 6, one of the RVV lines contains an orifice that results in choked flow and is opened when the RPV downcomer level is less than 160 inches. When the RPV/CNV differential pressure is between 50 and 100 psid, the other RVV line (unchoked flow) is opened and the RVV line with the choked flow orifice is closed. The RRV line (unaffected) is opened on a differential pressure between the RPV and CNV of 450 psid. For Run 7, the RVV lines are opened based on a RPV downcomer level less than 160 inches and the RRV line (unaffected) is opened based on a RPV/CNV differential pressure of 450 psid. In addition, the CPV level in Run 7 was reduced by 25 percent. Table 7-15, Table 7-16, and Table 7-17 provide the sequence of events for Runs 4, 6, and 7, respectively.

Table 7-15 LOCA Run 4 Sequence of Events ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 7-15 LOCA Run 4 Sequence of Events ((

                                                                                           }}2(a),(c),ECI Table 7-16 LOCA Run 6 Sequence of Events

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                                                                                           }}2(a),(c),ECI Table 7-17 LOCA Run 7 Sequence of Events

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 7-17 LOCA Run 7 Sequence of Events ((

                                                                                                   }}2(a),(c),ECI The general parameter trends of the simulations of these test runs was from reasonable to excellent.

Figure 7-163, Figure 7-164, and Figure 7-165 show the predicted RPV and CNV pressure comparisons with the measured data for the first 20,000 seconds (time period of interest for LOCAs). As in the other test runs simulated, the RPV and CNV pressures are over-predicted due to the HTP distortion in the plate surface area from the sealant on the CPV side of the plate. As in the sensitivity calculation performed for Run 1, it is expected that using the full HTP surface area produces a better match of the data for the short term. The peak CNV pressures for Runs 4, 6, and 7 are about 61 psia over-predicted. Figure 7-163 Run 4 RPV-CNV pressure comparison ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-164 Run 6 RPV-CNV Pressure Comparison ((

                                                                                          }}2(a),(c),ECI Figure 7-165 Run 7 RPV-CNV Pressure Comparison

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-166 through Figure 7-171 show the predicted comparison of the RPV downcomer level and the CNV level with data for Runs 4, 6, and 7, respectively. The predicted levels for Runs 4, 6, and 7 are consistent with the predicted levels presented above for Runs 1, 2, 3, and 5. As observed, the predicted levels for Runs 4, 6, and 7 show reasonable agreement with the measured data, given the differences in the actuation of the ECCS and with the reduced CPV fluid level (Run 7). Figure 7-166 Run 7 RPV-CNV Pressure Comparison ((

                                                                                                }}2(a),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-167 Run 4 CNV Level Comparison ((

                                                                                            }}2(a),(c),ECI Figure 7-168 Run 6 RPV Level Comparison

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-169 Run 6 CNV Level Comparison ((

                                                                                            }}2(a),(c),ECI Figure 7-170 Run 7 RPV Level Comparison

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-171 Run 7 CNV Level Comparison ((

                                                                                                   }}2(a),(c),ECI The prediction of the energy balance in the RPV and CNV show excellent agreement with the data for Runs 4, 6, and 7. The energy transfer across the HTP for these runs show that for the early part of the simulation, the energy is under-predicted due to the distortion in the modeled HTP surface area. The reduced HTP area from the sealant on the CPV side of the plate resulted in less energy transferred and thus resulted in the over-prediction of the RPV and CNV pressures and subsequently the predicted RPV and CNV levels. Although the comparisons of the energy balances for Runs 4, 6, and 7 are not shown here, the response is similar to those shown for Run 1 above.

The predicted behavior of the CPV for Runs 4, 6, and 7 are similar to those of Run 1 presented above. Evaporation plays a role in removing the energy deposited into the fluid of the CPV through the HTP. Since NRELAP5 lacks an evaporation model, the predicted temperatures in the CPV increase above those measured in the test in the long term. However, for the time frame of interest (0 - 20,000 seconds) the fluid temperature is under-predicted, particularly for the Run 7 simulation. The under-prediction of the fluid temperature reflects the under-prediction of the heat transfer across the HTP. As the CPV temperatures rise in the simulations, the fluid expands and the predicted CPV level increases (CPV fluid temperatures and level is not shown for simulations of Runs 4, 6, and 7, but are typical of the behavior shown above for the Run 1 simulation). In the test data, the CPV level declines due to evaporation. © Copyright 2022 by NuScale Power, LLC 378

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 7.5.11 NIST-2 IORV Integral Effects Test Series In this section, select IORV tests are described, along with the phenomenon addressed, experimental procedure, parameter ranges assessed, and any special analysis techniques. The purpose of these calculations is to document the assessment between the NIST-IORV Test data and the simulated NRELAP5 IORV transient results. The IORV test data are intended to provide a better understanding of phenomena related to Phase 0 in an Inadvertent Opening of a Reactor Valve scenario. Several IORV tests were performed at the NIST-2 facility. Two of the tests' results are presented: IORV Run 3 - Inadvertent opening of all ECCS valves. IORV Run 6 - Inadvertent opening of a single RVV. 7.5.11.1 Facility Description The entirety of the NIST-2 facility except for the DHRS was utilized for these integral effects tests, including: high-frequency sampling rate data system (10 Hz) high-frequency sampling pressure instrumentation and thermocouples High accuracy RTDs ((

                                }}2(a),(c) the SG was active to remove heat from primary side and drive NC in conjunction with the electrically heated core.

a pre-heated CNV was used to limit condensation during the simulated IORV event on the CNV shell walls a pre-pressurization process of the CNV was performed prior to break initiation to obtain a scaled pressure response of the facility relative to the full scale NPM the CPV was filled to accept rejected heat from the HTP 7.5.11.2 Phenomenon Addressed ((

                                   }}2(a),(b),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                                     }}2(a),(b),(c),ECI 7.5.11.3           Experimental Procedure

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                                                        }}2(a),(c) 7.5.11.4           Parameter Ranges Assessed The following conditions summarize the required characteristic state-points of the NIST-2 IORV tests:

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                                                                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                           }}2(a),(c)

Starting from these conditions, each test run is analyzed primarily over Phase 0. Phenomena of interest include break choked and unchoked flow, liquid droplet entrainment, and vapor generation. 7.5.11.5 Special Analysis Techniques Because rapid depressurization scenarios exacerbate the phenomena of interest for IORV, the analysis focuses on Phase 0 of Run 3 and Run 6. This assessment contains analysis techniques and sensitivities relevant to transient simulation for code-to-data comparisons with primary focus on the use of: ((

                                                                                            }}2(a),(c),ECI 7.5.11.6           Assessment Results Figure 7-172 through Figure 7-180 compare NIST-2 IORV test data with the NRELAP5 transient response. The PZR pressure response for Run 3 and Run 6 are shown in Figure 7-172 and Figure 7-173, respectively. ((
                                                                                               }}2(a),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                              }}2(a),(c),ECI Figure 7-174 and Figure 7-175 show the core differential pressure trend for each case. ((
                                      }}2(a),(c),ECI The significant underprediction of the PZR pressure in both cases shown previously indicates that phenomena contributing to the repressurization of the PZR are not well-predicted by NRELAP5. The phenomena involved in the prediction of the PZR depressurization rate are:

Mass/energy lost through the break Mass/energy transferred across the baffle plate Repressurization due to production of steam To investigate which phenomenon or phenomena are responsible for the underprediction, an advanced data analysis technique is employed. Note that this method can only be applied for a scenario with a single break flow path. The technique separates the PZR from the lower RPV (LRPV) at the baffle plate. ((

                                                                            }}2(a),(c)

For the remainder of this section, the focus is on Run 6. The baffle plate is examined first. Figure 7-176 shows the model baffle plate mass flow rate compared to that of the data. The data trend oscillates ((

                                 }}2(a),(c) The simulation shows flow into the PZR (note that negative values indicate flow out of the LRPV) almost immediately, which is due to level swell. Figure 7-177 and Figure 7-178 each also show values at

(( }}2(a),(c) within the same time period. This indicates that, during the period of repressurization in the PZR, there is little to no mass/energy exchange between the LRPV and PZR. This indicates that what causes the underprediction of pressure in the PZR must occur either within the PZR itself, and/or at the break choking plane. © Copyright 2022 by NuScale Power, LLC 382

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The break characteristics are examined next. Figure 7-179 shows the model break total mass flow rate and Figure 7-180 shows the model break total energy flow rate compared to the data. The data trend shows non-physical values within the first second of the transient. This is expected to be due to the inherent uncertainties in the method for calculating break flow rate. As a result, this segment of the data is not used. It is expected that within this time frame NRELAP5 predicts the mass flow rate well, given that it is expected to be pure steam. ((

                              }}2(a),(c)

In summary, the mass and energy entering the PZR from the LRPV and leaving the PZR via the break appear to be reasonably well predicted. ((

                                                         }}2(a),(c),ECI Figure 7-172 IORV Run 3 Pressurizer Pressure

((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-173 IORV Run 6 Pressurizer Pressure ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-174 IORV Run 3 Core Differential Pressure ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-175 IORV Run 6 Core Differential Pressure ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-176 IORV Run 6 Baffle Plate Total Mass Flow Rate Comparison ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-177 IORV Run 6 Baffle Plate Total Energy Flow Rate Comparison ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-178 IORV Run 6 Break Baffle Plate Differential Pressure ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-179 IORV Run 6 Break Total Mass Flow Rate Comparison ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 7-180 IORV Run 6 Break Total Energy Flow Rate Comparison ((

                                                                                              }}2(a),(c),ECI 7.6      Containment Response Methodology Assessment The NRELAP5 code has been qualified or assessed to the separate effects and integral effects tests to demonstrate the capability to simulate LOCAs in the NPM. The results of the NRELAP5 comparisons to data establish the capability of the code to model the NPM design of LOCA analysis. The mode important assessment activities were those comparing to integral LOCA tests conducted in the NIST facility.

The following two key known scaling distortions are relevant to the scope of the containment response analysis methodology: ((

                         }}2(a),(c)

Neither of the above phenomena have an impact on the peak CNV pressure. The first distortion is addressed by the containment response analysis methodology by closure of the MSIVs. The second distortion is addressed by the overall conservative modeling of CNV heat transfer by the containment response analysis methodology, which includes use of conservative initial and boundary conditions. © Copyright 2022 by NuScale Power, LLC 391

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 8.0 Assessment of Evaluation Model Adequacy The adequacy of the NRELAP5 code (Reference 9) for analysis of design-basis LOCAs in the NPM is demonstrated by closure model and correlation reviews, and assessments against relevant experimental data. Establishing the adequacy of the NRELAP5 code as a component of the NuScale LOCA methodology is an essential part of the EMDAP (RG 1.203). 8.1 Adequacy Demonstration Overview The adequacy demonstration process used here is similar to that used for the AP-600 (Reference 74) and is a modified version of the process approved in Revision 2 of this report. As the NuScale PIRT is a primary input to the adequacy evaluation, the findings of the PIRT summarized in Section 4.0 are used to guide the adequacy of the evaluation process. There are no significant changes to PIRT phenomena as part of the 2022 update or the Phase 0 PIRT, (( }}2(a),(c) which is addressed in the Non-LOCA Topical Report (Reference 114). The adequacy of the NuScale LOCA EM is demonstrated through the following steps:

1. Section 8.2 documents the bottom-up assessment of the NRELAP5 models and correlations to determine their adequacy to predict the high (H) ranked phenomena.

The code models used to represent each high (H) ranked phenomena are identified, with emphasis on the phenomena with low-knowledge level. These assessments address the fidelity of the models and correlations to the appropriate fundamental or SET data. Fidelity of the assessments is evaluated using the criteria of excellent, reasonable, minimal and insufficient from RG 1.203. These criteria are defined in Table 1-2. The comparisons to data can identify modeling deficiencies which could impose limitations on the application of the NRELAP5 based LOCA EM.

2. Section 8.3 covers the top-down assessment of the EM including a review of EM governing equations and numerics to determine their applicability to NPM LOCA analysis, and evaluation of the integral code performance based on the assessments of the EM against relevant IETs.

3. Section 8.4 summarizes the adequacy findings. The report shows how each PIRT high (H) ranked phenomenon is covered by the LOCA methodology models and correlations. Models which are marginally adequate, or ranges where PIRT phenomena are not covered, are identified. The manner of addressing code limitations is described.

8.2 Evaluation of Models and Correlations (Bottom-Up Assessment) The adequacy of the models and correlations in NRELAP5 for modeling the high (H) ranked phenomena is examined by considering their pedigree, applicability, and fidelity to appropriate fundamental or SET data (established by assessment of the EM against legacy and NuScale-specific SET data), and scalability to the NuScale LOCA scenario. The following steps are used to perform the evaluations. © Copyright 2022 by NuScale Power, LLC 392

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

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2. ((

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The result of these assessments for each model or correlation is used to identify whether there are any shortcomings in the parametric space and provide information needed for the development effort where additional models, assumptions, or conservatisms may be required. The first three steps described above are addressed in Section 8.2.1. Step 4 is discussed in Section 8.2.2 through Section 8.2.22. 8.2.1 Important Models and Correlations Table 8-1 identifies the dominant code models and correlations for the PIRT, defined as high-ranked phenomena in Section 4.0. Key parameters that are influenced by the dominant models and correlations are listed, along with phenomenological and separate effects cases that are used to assess the model or correlation capabilities. This information is used to establish adequacy of the dominant code models or correlations for NPM LOCA applications. Table 8-1 Dominant NRELAP5 Models and Correlations ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-1 Dominant NRELAP5 Models and Correlations (Continued) ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-1 Dominant NRELAP5 Models and Correlations (Continued) ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-1 Dominant NRELAP5 Models and Correlations (Continued) ((

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Table 8-2 is a summary of the estimated range of key parameters over which each dominant model or correlation should be applicable for the NPM steady-state and design basis LOCA. Parameter ranges obtained are intended to identify the minimum range that needs to be covered; the applicability of models and correlations are not restricted to these ranges. Several sources are used to obtain the values of the ranges. This includes design values, proposed technical specification limits, and limiting initial and boundary conditions. The ranges for some parameters are obtained from NRELAP5 LOCA break spectrum calculations described in Section 9.0. An explanation of how each limiting range was determined is provided in the Comments column of Table 8-2. Table 8-2 Process Parameter Ranges for EM Applicability Evaluation ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-2 Process Parameter Ranges for EM Applicability Evaluation (Continued) ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-2 Process Parameter Ranges for EM Applicability Evaluation (Continued) ((

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Table 8-3 lists the range of geometric parameters that could influence high-ranked phenomena for the NPM LOCA. Values given for each parameter are intended to identify the minimum range geometric parameters that need to be covered by the LOCA EM; applicability of the NuScale LOCA EM is not restricted to these values. These values have been obtained from a compilation of geometric information and plant parameters determined from design drawings. The changes to the NPM-160 geometric parameters have no or little impact on the LOCA EM applicability to NPM-20 because there is no significant NPM geometry change affecting the applicability of the associated NRELAP5 physical models to the NPM-20 design. © Copyright 2022 by NuScale Power, LLC 399

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-3 Geometric Parameters for NPM-160 and NPM-20 ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-3 Geometric Parameters for NPM-160 and NPM-20 (Continued) ((

                                                                                                }}2(a),(c),ECI Each of the NRELAP5 dominant models or correlations listed in Table 8-1 is evaluated with respect to the extent that the model or correlation, as assessed for the NPM LOCA application, covers all or a portion of the NPM range given in Table 8-2.

Where the range provided in the model or correlation development does not cover the full range of the NPM LOCA application, the range is extended by extrapolation of assessments against experimental data, or justification is provided based on legacy RELAP5-3D© assessments and applications. The range covered by models and correlations is discussed for the key parameters that define the response for high-ranked phenomena. 8.2.1.1 Applicability Evaluation To determine adequacy of the models and correlations to simulate the high-ranked phenomena, the results of assessments against phenomenological and SETs are discussed. The assessment results are drawn from the NRELAP5 assessments discussed in Section 7.0 where descriptions of the test facilities, instrumentation, and test procedures are provided. © Copyright 2022 by NuScale Power, LLC 401

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 8.2.1.2 Overview A graded approach is used to address the items described in Section 8.2, Step 4. More emphasis is given to high-ranked phenomena with a low-knowledge level. Less emphasis is placed on phenomena that are well understood with a high-knowledge level. This includes industry standard and handbook models. Each of the following four areas is evaluated to the extent that they are relevant for each high-ranked phenomenon. ((

                                               }}2(a),(c) 8.2.1.2.1              High-Ranked, Low-Knowledge Level Phenomena The PIRT identified some phenomena within specified components as high-importance phenomena that have a low-knowledge level. These high-importance and low knowledge phenomena are given the greatest focus in the development of the LOCA EM. They include:

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                                                                    }}2(a),(c) 8.2.2         ((
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Table 8-4 Marviken Range of Parameters Compared to the NuScale Power Module ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-4 Marviken Range of Parameters Compared to the NuScale Power Module ((

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Table 8-5 Ferrell-McGee Range of Parameters Compared to the NuScale Power Module ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-5 Ferrell-McGee Range of Parameters Compared to the NuScale Power Module ((

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Table 8-6 Dimensions of NuScale Power Module, NIST-1 and Bankoff Pressurizer Plate ((

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                                      }}2(a),(c) 8.2.5         ((                                             }}2(a),(c) 8.2.5.1            Background

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Table 8-7 Range of Riser Interphase Friction - Separate Effects Tests and NuScale Power Module ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-7 Range of Riser Interphase Friction - Separate Effects Tests and NuScale Power Module (Continued) ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 8.2.7 (( }}2(a),(c) 8.2.7.1 Background ((

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                                             }}2(a),(c) 8.2.8.2            Technical Evaluation Figure 8-1 shows a schematic of the major heat transfer modes governing the heat transfer across the CNV wall. ((
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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 8-1 CNV Wall Heat Transfer Modes ((

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Figure 8-2 Thermal Resistance Network Between CNV and UHS ((

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                              }}2(a),(c) 8.2.10        Flashing 8.2.10.1           Background Flashing is the fundamental thermodynamic process of vaporization that occurs when a saturated liquid undergoes a reduction in pressure below its boiling point, resulting in a phase change from liquid to vapor. In the pressurizer, the liquid inventory that is normally at saturated conditions flashes as the RCS depressurizes in response to a LOCA and the actuation of the ECCS. As the RPV continues to depressurize, flashing occurs in the hot leg riser, core, LP, and downcomer. Flashing causes level swell which can affect the quality at the break and at the ECCS valves.

The interphase heat and mass transfer models in NRELAP5 are the dominant models that determine the flashing rate. The vapor generation (or condensation) consists of two parts, vapor generation which results from energy exchange in the bulk fluid (flashing) and energy exchange in the thermal boundary layer near the wall (boiling). Flashing is addressed in this section and boiling in Section 8.2.19. Each of the vapor generation processes involves interfacial heat transfer effects. The interfacial heat transfer area and heat transfer coefficient models used in NRELAP5 are summarized in Table 2.5-1 of Reference 9. The models that govern the flashing phenomenon are those for the superheated liquid fluid state. For bubbly flow bulk interfacial heat transfer between the vapor and liquid phases is handled using the maximum of a correlation derived from the Plesset-Zwick (Reference 90) equation for the growth rate of a bubble and the modified Lee-Ryley (Reference 91) correlation. For droplets in flow regimes such as annular mist and dispersed, NRELAP5 uses a heat transfer coefficient k f D f ( T sf ) where k f is the liquid thermal conductivity, D is hydraulic diameter, and T sf is the difference between the saturation and the liquid temperatures; f ( T sf ) is a flow-type dependent function. A ((

                                          }}2(a),(c) heat transfer coefficient is used for films. In all cases, a large heat transfer coefficient is calculated so that the difference between the superheated liquid temperature and the saturation temperature at any time is small. Hence for the NPM, while the pressure is decreasing, the liquid temperature remains very close to the saturation temperature due to the relatively

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 slow depressurization resulting from the small break sizes. The time constant for the vapor generation process is much smaller than the time constant for the depressurization. A non-equilibrium superheated liquid state can exist for a time period on the order of milliseconds while the depressurization process is taking place over a period of minutes. Hence, high accuracy in the vapor generation rate due to flashing is not required because any model with a small time constant generates the amount of vapor necessary to keep the phases in thermal equilibrium. Implementation of these correlations is described in Reference 9 (Section 2.5.1.1). The pedigree of the model is established by application and validation of RELAP5-3D©. In Table 2.2-2 of Reference 87, it is noted that the flashing model is validated against the Edward pipe and Marviken CFT-22 and CFT-24 tests. Validation of NRELAP5 for the suite of assessment cases confirms that this pedigree is maintained in the NRELAP5 code. 8.2.10.2 Technical Evaluation Important parameters associated with flashing phenomenon are pressure, void fraction/interfacial area, and phasic temperatures that interact to determine the vapor generation rate. The flashing model in NRELAP5 covers the entire range of the water properties tables, which encompasses the NPM LOCA application. The GE level swell (1-ft and 4-ft) tests are the primary assessment cases used to validate the applicability of NRELAP5 to predict the flashing phenomenon. In Section 8.2.6, it is shown that the NRELAP5 predictions of the GE level swell assessments show reasonable-to-excellent agreement with the test data, thus demonstrating that NRELAP5 is applicable for predicting flashing phenomenon that occurs in NPM LOCA events. Furthermore, flashing is an inherent phenomenon in the NIST LOCA IETs. Therefore, NRELAP5 analysis against the NIST IET data provides additional assessment of the flashing model. Because the heat transfer coefficient resulting from the Plesset-Zwick correlation depends only on fluid properties and all of the heat transfer coefficients are very large, there are no scaling restrictions for the vapor generation model in NRELAP5 which could impose limitations on the application of the NRELAP5 model to the configuration and conditions of the NPM hot leg riser in the LOCA transient domain. 8.2.11 (( }}2(a),(c) 8.2.11.1 Background ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 8-4 Transient Void Fraction in Node 4 for the GE 4-ft Level Swell Test Figure 8-5 Transient Void Fraction in Node 6 for the GE 1-ft Level Swell Test © Copyright 2022 by NuScale Power, LLC 432

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 8-6 Transient Void Fraction in Node 5 for the GE 1-ft Level Swell Test ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 8.2.12 (( }}2(a),(c) 8.2.12.1 Background ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 8.2.16 (( }}2(a),(c) 8.2.16.1 Background ((

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                                                                   }}2(a),(c) 8.2.17.2           Technical Evaluation Table 8-1 lists the flow, void fraction, pressure, heat rate, and core geometry as key parameters for the NRELAP5 interfacial drag model, the dominant NRELAP5 model that affects phase slip and flow regimes. Table 8-2 identifies the range of parameters encountered by NPM steady-state and design basis accidents and transients. This is the source of the ranges used for the NPM ((
                                   }}2(a),(c) SETs used to validate the NRELAP5 interphase drag model for the core geometry are FRIGG tests 613130, 613010, 613118 and 613123, and FLECHT-SEASET boil off tests 35557, 35658 and 35759 The FRIGG tests are steady-state runs with set inlet and boundary conditions to a 36-rod electrically-heated bundle. The void fraction profile along the heated channel was measured and compared to NRELAP5 predictions (Section 7.2.5).

The FLECHT-SEASET tests (Section 7.2.6) are boil off transient tests in which a 161-rod bundle was initially filled with saturated water and then boiled to the point at which heater rod temperature reached 2,000 degrees F (1,093 degrees C), and the test was terminated by cutting the rod power and flooding the bundle. © Copyright 2022 by NuScale Power, LLC 441

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-8 shows the ranges of the key variables for the NRELAP5 ((

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Table 8-8 Ranges of Key Parameters for Core Interphase Friction - Separate Effects Tests and Plant ((

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Table 8-9 Range of Key Parameters for Core Flow - Separate Effects Tests and Plant ((

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                                                      }}2(a),(c) 8.2.19        ((                            }}2(a),(c) 8.2.19.1           Background

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Table 8-10 Range of Key Parameters for Core Boiling - Separate effects tests and plant ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 8.2.20.2 Technical Evaluation ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 8.2.21 (( }}2(a),(c) 8.2.21.1 Background ((

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-11 Range of Key Parameters for Subcooling Boiling and Separate Effects Tests and Plant ((

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                            }}2(a),(c) 8.2.22        ((                                           }}2(a),(c) 8.2.22.1           Background

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                                                                                        }}2(a),(c) 8.3      Evaluation of Integral Performance (Top-Down Assessment)

There are three primary areas addressed by the top-down assessment. ((

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To ensure maximum fidelity of the assessments, the NRELAP5 NIST and NPM input models were developed using consistent nodalization and option selection. Code assessments are also performed against SETs to establish code capabilities for predicting local behavior within unique NPM components. Assessments against SETs are addressed in Section 8.2.1. © Copyright 2022 by NuScale Power, LLC 452

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 8.3.1 Review of Code Governing Equations and Numerics The NRELAP5 Theory Manual (Reference 9) describes the NRELAP5 code architecture, field equations, and solution techniques, which are essentially unchanged compared to the RELAP5-3D© code. The descriptions of code modifications/features made to address unique aspects of the NuScale application are included in the NRELAP5 Theory Manual and summarized in Section 6.0. This review is based primarily on the information in Reference 9. The field equations solved by NRELAP5 are discussed in Section 2.1 of Reference 9 and summarized in Section 6.2. Applicability of the field equations to represent the processes and phenomena that can occur in the NPM is evaluated, along with an assessment of the ability of the NRELAP5 numerical solution to approximate the set of governing field equations. This evaluation addresses the mathematical models implemented in NRELAP5 for the NuScale LOCA analysis, and considers the applicability of the assumptions and processes involved in developing the NRELAP5 system of governing equations, and closure relations. The numeric solution evaluation considers convergence, conservation of physical properties, and stability of code calculations performed to solve the set of governing equations for an NRELAP5 NPM model. The objective of this evaluation is to summarize information regarding the domain of applicability of the numerical techniques and user options that may impact the accuracy, stability, and convergence of NRELAP5 calculations. User guidelines for model development and execution were developed based on "lessons learned" during the code reviews and assessments. The guidelines include requirements for assuring convergence of solutions, accounting for uncertainty in results and monitoring code function to ensure that the basic conservation equations are being solved correctly. 8.3.1.1 Conservation of Mass, Momentum and Energy NRELAP5 LOCA applications do not use the three-dimensional modeling capability of RELAP5-3D©. The one-dimensional equations and numerics have been used in versions of the RELAP5 codes for many years so their pedigree has been well established by code assessments and applications. The semi-implicit solution technique used by NuScale has been in the RELAP5 codes as the primary solution technique for the governing conservation equations since the initial development of the code. The solution technique continues to be used in NRELAP5 as discussed in Section 2.1.3 of the NRELAP5 code manual (Reference 9). The basic governing equations for mass, momentum, and energy conservation use area-averaging for vapor and liquid fields. Mass, momentum, and energy conservation equations are written for each field, resulting in what is referred to as a six-equation model. The governing equations are discussed in Section 6.2.1 through Section 6.2.4. The basic governing equations in NRELAP5 are generally © Copyright 2022 by NuScale Power, LLC 453

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 accepted as reasonable representations of the applicable physical laws that govern the steady-state and transient behavior of thermal-hydraulic systems. Energy transfer into and out of the phases from the boundaries is governed by correlations discussed in Section 6.2.5. Heat conduction within structures is modeled by the one-dimensional heat conduction equation discussed in Section 6.3. Models are also included for trips and control systems as discussed in Section 6.5. This feature is used to model the safety-related system actuations, control power, set boundary conditions, determine ranges (minimum and maximum values) of selected variables including the FOMs, and other functions within the NRELAP5 models. NuScale performed acceptance testing and procurement requirements as part of the commercial grade dedication of RELAP5-3D© to serve as the development platform for NRELAP5. The testing and inspection verified that RELAP5-3D© has the necessary critical characteristics to be used as the code development platform for NRELAP5. The critical characteristics include the suitability of the basic governing equations described above for the NuScale application. 8.3.1.2 Numerical Solution Techniques The entire fluid domain of interest is divided into control volumes connected via junctions where the flow velocities are defined. The heat transfer into or out of control volumes are defined through heat structures where the heat conduction equation is solved considering the relevant heat transfer regime in the communicated control volume. The difference equations implement mass and energy conservation by equating accumulation to the rate of mass or energy in through the cell boundaries, minus the rate of mass or energy out through the cell boundaries, plus source and sink terms. This approach necessitates defining mass and energy volume average properties and requiring knowledge of velocities at the volume boundaries. The velocities at the cell edges are defined through the use of momentum control volumes centered on the mass and energy cell boundaries. This approach results in a numerical scheme having a staggered spatial mesh with the momentum control volumes extending from the mass and energy cell centers to the neighboring mass and energy cell centers. The scalar properties of the flow (pressure, specific internal energies, and void fraction) are defined at mass and energy cell centers, while the vector quantities (velocities) are defined on the mass and energy cell boundaries. The governing equations for the system model are solved numerically using a semi-implicit finite-difference technique. A nearly-implicit finite-difference technique which allows violation of the material Courant limit, is also available. However, the LOCA EM and the supporting assessment calculations use only the © Copyright 2022 by NuScale Power, LLC 454

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 semi-implicit numerical scheme. The semi-implicit numerical solution scheme is based on replacing the system of differential equations with a system of finite difference equations partially implicit in time. When generating a solution of finite difference equations, there is a possibility that the solution may not be converged. This could be the result of an ill-posed problem, inappropriate time step size selection, inadequate spatial nodalization, or an instability. Sensitivity studies have proven useful to ensure convergence and stability of the NRELAP5 solutions. Adherence to the modeling requirements of RELAP5 assist in ensuring that the governing equations are well posed. Requirements for nodalization and time step sensitivity studies comply with 10 CFR 50 Appendix K requirements and ensure converged solutions. Solutions are examined to identify unstable or unphysical behavior. 8.3.2 NuScale Facility Scaling The NIST facility is designed to simulate the integral system behavior of a single NPM immersed in a single bay within the reactor pool. The scaling analysis was performed to determine the geometric dimensions and operating conditions for the NIST facility. The purpose of the scaling analysis was to design an IET facility that can be used to obtain quality data for thermal-hydraulic system safety analysis code validation. The hierarchical two-tiered scaling (H2TS) (Reference 99) method was used to perform the RCS natural circulation scaling and the scaling of LOCA and ECCS. The scaling analysis generated the sets of dimensionless groups that needed to be preserved to accurately simulate the high-ranked phenomena identified in the LOCA PIRT. The figures of merit were the peak CNV pressure and the collapsed liquid level above the top of the core. The scaling analysis also documented the scaling distortions between the NIST facility and the NPM design, and evaluated the effects of these distortions. Detailed documentation of the NIST scaling analysis is available in the NIST Facility Scaling Reports. Section 8.3.2.1 summarizes the scaling objectives and methodology. The approaches for RCS scaling natural circulation scaling and the scaling of LOCA and ECCS are briefly presented in Section 8.3.2.2 and Section 8.3.2.3, respectively. 8.3.2.1 Scaling Objectives and Methodology The general objective of the scaling analysis was to obtain the physical dimensions and operating conditions of a reduced-scale test facility capable of simulating the important flow and heat transfer behavior of a NPM under the LOCA conditions. To develop a properly scaled test facility, the following specific objectives were met for each operational mode of interest. The thermal-hydraulic processes that should be modeled were identified. © Copyright 2022 by NuScale Power, LLC 455

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The similarity criteria that should be preserved between the test facility and the full-scale prototype were obtained. The priorities for preserving the similarity criteria were established. Specifications for the test facility design were established. Biases due to scaling distortions were quantified. The critical attributes of the test facility that must be preserved to meet testing requirements were identified. Different similarity criteria were obtained for the different modes of system operation. These criteria depend on the geometry of the components, the scaling level required to address the transport phenomena of interest, and the initial and boundary conditions for each particular mode of operation. To ensure that the scaling objectives were met in an organized and clearly traceable manner, a general design framework (GDF) was established. The model for this framework includes features drawn from the NRC severe accident scaling methodology presented in NUREG/CR-5809 (Reference 99). A flow diagram for the GDF is presented in Figure 8-7. © Copyright 2022 by NuScale Power, LLC 456

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 8-7 General Design Framework for the NuScale Integral System Test Facility Experimental Objectives The first task outlined by the GDF was to specify the experimental objectives. The experimental objectives define the types of tests that will be performed to address specific design or certification needs. These objectives determined the general modes of operation that should be simulated in the test facility. The objective of the NuScale LOCA test program was to obtain qualified data to benchmark the computer codes and models that is used to evaluate the safety of the NPM. This includes: 1) measurements of transient and steady-state, single-phase natural circulation flow in the integrated RPV, and 2) characterization of the thermal-hydraulic phenomena in the RPV, containment, and CPV during the three periods of the LOCA. Loss-of-Coolant Accident Phenomena Identification and Ranking Table The second task outlined by the GDF was the development of a PIRT. The PIRT presented in Section 4.0 was used as the basis. The nature of scaling forbids © Copyright 2022 by NuScale Power, LLC 457

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 exact similitude of all of the parameters of a reduced-scale test facility with those of a full-scale prototype. As a result, the design and operation of the test facility was based on simulating the thermal-hydraulic processes most important to the system operational modes that are explored. The PIRT identified the different phases of a LOCA and most important thermal-hydraulic phenomena within those phases that should be simulated in the test facility. All of the highly ranked integral system phenomena identified in the LOCA PIRT are observed in the NIST facility to some degree. Although majority of the high-ranked PIRT phenomena are fully covered in NIST, the NIST facility is not the primary source of validation data for some phenomena. For example, the NIST facility does not model the details of the core fuel rods or core sub-channels; therefore, CHF data were obtained in a separate full-scale test facility (Section 7.3). Similarly, detailed information regarding the helical coil SG thermal-hydraulic performance is obtained from SIET TF-1 and TF-2 experiments (Section 7.4). Description of the H2TS Method The third step in the GDF was to perform a scaling analysis for each of the hierarchical levels (e.g., systems and subsystems) and their modes of operation defined in the previous section. The H2TS method has been successfully used to develop the similarity criteria necessary to scale the APEX-600 and APEX-1000 systems for LOCA transients. The H2TS method was developed by the NRC and is fully described in Appendix D of NUREG/CR 5809 (Reference 99). Figure 8-8 presents the four basic elements of the H2TS analysis method. The first element consists of subdividing the plant into a hierarchy of systems. Each system is subdivided into interacting subsystems which are further subdivided into interacting modules which are further subdivided into interacting constituents (materials) which are further subdivided into interacting phases (liquid, vapor or solid). Each phase can be characterized by one or more geometrical configurations and each geometrical configuration can be described by three field equations (mass, energy and momentum conservation equations). Each field equation can incorporate several processes. Figure 8-9 presents the breakdown of the NuScale system into hierarchical levels and high level processes to be scaled. It represents a roadmap used to structure the scaling analyses. The RCS and the ECCS were the focus of the scaling study. © Copyright 2022 by NuScale Power, LLC 458

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 8-8 Flow Diagram for the Hierarchical, Two-Tiered Scaling Analysis (NUREG/CR-5809) 6<67(0 6&$/( 723'2:16<67(06&$/,1*

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Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 8-9 NuScale System Breakdown into Hierarchical Levels and Primary Operational Modes to be Scaled ((

                                                                                                          }}2(a),(c)

After identifying and subdividing the system of interest, the next step was to identify the scaling level at which the similarity criteria should be developed. This was determined by examining the phenomena being considered. For example, if the phenomenon being considered involves mass, momentum or energy transport between materials such as water and solid particles, then the scaling analysis would be performed at the constituent level. If the phenomenon of interest © Copyright 2022 by NuScale Power, LLC 460

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 involves mass, momentum, or energy transport between vapor and liquid, then the scaling analysis would be performed at the phase level. Therefore, identifying the scaling level depends on the phenomenon being addressed. Thermal-hydraulic phenomena involving integral RCS interactions, such as primary system depressurization or loop natural circulation, would be examined at the "system" level. Thermal-hydraulic phenomena, such as SG heat transfer, would be examined at the "subsystem" level. Specific interactions between the steam-liquid mixture and the stainless steel structure would be examined at the "constituent" level. The H2TS method required performing a "top-down" (system) scaling analysis. The top-down scaling analysis examines the synergistic effects on the system caused by complex interactions between the constituents deemed important by the PIRT. Its purpose is to use the conservation equations at a given scaling level to obtain characteristic time ratios and similarity criteria. It also identifies the important processes to be addressed in the bottom-up scaling analysis. The H2TS method also required performing a "bottom-up" (process) scaling analysis. This analysis provides similarity criteria for specific processes such as flow pattern transitions and flow dependent heat transfer. The focus of the bottom-up scaling analysis is to develop similarity criteria to scale individual processes of importance to system behavior as identified by the PIRT. Test Facility Specifications and Scaling Ratios The fourth step of the GDF was to document all of the test facility design and operation specifications. All of the essential geometric features and operating parameters that must be carefully measured and documented to ensure accurate code simulations of the important thermal-hydraulic phenomena were identified and designated as critical attributes. The NIST facility was developed by modifying the existing MASLWR test facility at Oregon State University. This was accomplished by establishing a fixed set of scale factors for component lengths, flow areas, and volumes. These scale factors were obtained through an iterative process that included a practical assessment of component costs, ease of operation, material availability, and instrumentation accuracy for the scale selected. Having fixed the length, volume, and flow area scale factors for the test facility, and assuming fluid property similitude, the scaling ratios obtained using the governing equations for loop natural circulation were used to define the scale factors for the adjustable parameters. That is, the core power, component heat transfer areas, SG heat removal rate, and loop resistance were adjusted to preserve the requirement of isochronicity. Table 8-12 lists the required temporal and geometric scale factors for the NIST facility under the requirement of isochronicity and fluid property similitude. © Copyright 2022 by NuScale Power, LLC 461

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-12 Scaling Factors for NIST Facility ((

                                                                                                            }}2(a),(c)

The NIST tests start from steady-state natural circulation conditions. ((

                                           }}2(a),(c). The test facility operates near prototypic pressures and temperatures and operates with the same working fluid: water.

Therefore, fluid property similitude is invoked. This means that the fluid property ratios are near to unity in all of the scale ratios, thereby simplifying the analysis. 8.3.2.2 Reactor Coolant System Natural Circulation Scaling Figure 8-10 provides a flow diagram that describes the scaling analysis process for the RCS natural circulation operational mode. First, a top-down scaling analysis was performed. This included an analysis at the system level (integrated loop behavior) for normal operating conditions. ((

                                      }}2(a),(c) Further details are available in the NIST facility scaling reports.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 8-10 Scaling Analysis Flow Diagram for Single-Phase Primary Loop Natural Circulation ((

                                                                                                  }}2(a),(c) 8.3.2.3            Loss-of-Coolant Accident and Emergency Core Cooling System Scaling The scaling analysis approach for LOCA and actuation of the ECCS includes the following four related scaling analyses:

RCS depressurization containment vessel pressurization long-term recirculation cooling Reactor Building pool heat-up During ECCS operation, the RPV transports energy and mass to the containment. The mass and energy leaving the RPV is captured by the CNV. The CNV transports energy to the reactor pool. © Copyright 2022 by NuScale Power, LLC 463

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The objectives of the LOCA/ECCS scaling analyses were to scale the RCS depressurization and containment pressurization behavior during the blowdown and venting phases of the LOCA. RCS, containment cooling, and the reactor pool heat-up during the long-term recirculation cooling phase of the LOCA. Reactor Coolant System Depressurization Scaling Figure 8-11 shows top-down and bottom-up scaling analyses performed for RCS depressurization scaling. This included an analysis at the system level (integrated loop behavior) for RCS initial conditions at 50 percent power. ((

                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 8-11 Scaling Analysis Flow Diagram for Reactor Coolant System Depressurization ((

                                                                                                    }}2(a),(c)

Containment Pressurization Scaling Top-down and bottom-up scaling analyses performed for the scaling of containment pressurization are shown in Figure 8-12. ((

                                                                                                  }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                             }}2(a),(c)

Figure 8-12 Scaling Analysis Flow Diagram for Containment Pressurization ((

                                                                                                    }}2(a),(c)

Long-Term-Cooling Phase Scaling Long-term recirculation occurs after the CNV and RPV pressures have become nearly equalized and the flow through the RRVs is from containment to RPV. As shown in Figure 8-13, top-down and bottom-up scaling analyses were performed for long-term phase scaling. ((

                                                                                                }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                                      }}2(a),(c)

Figure 8-13 Scaling Analysis Flow Diagram for Long-Term Recirculation Cooling Mode ((

                                                                                                          }}2(a),(c)

Reactor Pool Heatup Scaling Heat transferred from the CNV exterior surface creates a heated plume of fluid that rises to the top of the pool. The heated plume mixes with the liquid at the top of the pool to create a thermally stratified layer. The thermally stratified layer consists of two regions; a well-mixed layer with uniform temperature at the surface of the pool and a partially mixed thermocline that extends downward and serves as a transition layer to the colder liquid in the pool. The thermal stratification layer grows over time. During the heat up of the pool, heat and mass are lost from the pool to the air at the pool interface due to evaporation. Heat is also transferred to the Reactor Building pool steel liner and concrete structures. © Copyright 2022 by NuScale Power, LLC 467

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Top-down and bottom-up scaling analyses performed for reactor pool heatup scaling are shown in Figure 8-14. ((

                                                                                   }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 8-14 Scaling Analysis Flow Diagram for Reactor Building Pool Heat-Up ((

                                                                                                         }}2(a),(c) 8.3.2.4            As-Built NuScale Facility Scaling Summary Comparison of plant behavior and the NIST test facility behavior for various events is presented in this section and in Section 8.3.4. Distortions exist between the plant and NIST IET as it does for any other scaled test facility. The purpose of the IET is to provide a scaled facility simulating the phenomena important to plant behavior with relative magnitudes that are similar to the plant for use in validating the models and integral behavior of the analysis code. The distortion analyses presented in this section for as-built NIST facility and in Section 8.3.4 for as-performed NIST tests show that the NIST facility meets these criteria. The top-down portion of the NIST scaling analysis presented in Section 8.3.2.3 was expanded to perform an additional quantitative evaluation of the distortions in the NIST facility. The mass/energy balance equations were re-defined to include additional terms that better quantify the distortion in various phenomena seen in the RCS and CNV during a typical LOCA. The control volume balance equations derived for the RCS and CNV include

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                                             }}2(a),(c)

For quantifying the distortions, the following terms in the energy balance equations were explicitly accounted for in the top-down scaling analysis ((

                                                            }}2(a),(c)

The dimensionless forms of the mass/energy balance equations were derived by identifying the characteristic scales appearing in the balance equations. groups characterizing the ratio of characteristic times for each process were defined based on the dimensionless equations. The NIST-2 LOCA and IORV IETs address some of the key high-ranked phenomena identified in the LOCA PIRT. The other high-ranked LOCA PIRT phenomena are addressed through the NRELAP5 code assessment using other IETs or SETs. The top-down LOCA scaling analysis with the group comparison, confirms that the high-ranked LOCA PIRT phenomena addressed in the NIST-2 LOCA tests are consistent with the dominant groups. The system transient behavior during the LOCA/IORV event progression is governed mainly by blowdown (break/IORV/ECCS), phase change (boiling/flashing/condensation), and heat transfer through solid structures. Top-down scaling analysis comparing the NPM-160/NIST-1 design and NPM-20/NIST-2 design demonstrated that the dimensionless PI groups quantifying phenomena importance are similar between the NPM-160, NPM-20 and NIST-2. Similarly, top-down scaling for the IORV tests demonstrated similarity between the NPM and NIST-2 groups. Table 8-13, Table 8-14, and Table 8-15 summarize the mass flow paths and heat flow paths for the RCS and CNV considered in the top-down scaling analysis. The heat and mass flows into the control volume have a positive sign; whereas the negative sign represents heat and mass flow out of the control volume. Three mass flow rates are identified for both NPM and NIST and are symmetric between the RCS and CNV. The same number of heat flow paths are identified for the RCS in NPM and NIST. ((

                                                                             }}2(a),(c) Two major heat transfer paths are identified for the CNV: the first path is the heat transfer on the

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 inner surface of the containment wall, the second is the heat transfer from the outer surface of the reactor pressure vessel. ((

                                                  }}2(a),(c)

A summary of groups in the dimensionless mass/energy balance equations is given in Table 8-16 and Table 8-17. As depicted in Table 8-16 and Table 8-17, ((

                                                                                   }}2(a),(c)

Table 8-13 Mass Flow Paths for NPM and NIST (RCS and CNV) ((

                                                                                                      }}2(a),(c)

Table 8-14 Heat Flow Paths for RCS in NPM and NIST ((

                                                                                                      }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-15 Heat Flow Paths for Containment in NPM and NIST ((

                                                                                                }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-16 Description of Groups for the RCS Mass/Energy Balance (( Group Description

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-17 Description of Groups for the Containment Mass/Energy Balance (( Group Description

                                                                                                    }}2(a),(c)

These groups were evaluated based on NRELAP5 simulations of the NPM and as built NIST facility for the following events: 100 percent discharge line break on the CVCS line (similar to NIST-1 HP-06 and NIST-2 Run 1) © Copyright 2022 by NuScale Power, LLC 474

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 100 percent high point vent line break (Similar to NIST-1 HP-07 and NIST-2 Run 2) Inadvertent opening of a single RVV (Similar to NIST-1 HP-09 and NIST-2 Run 3) ((

                                                            }}2(a),(c) The key conclusions of this analysis are summarized below.

1. ((

                                                       }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3

4. ((

                                      }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3

10. ((

                                                                                              }}2(a),(c) The scaling and distortion analysis methodology presented above is used to analyze the impact of biases in initial and boundary conditions and differences in operating procedures of the final NIST IET data used in Section 8.3.4.

8.3.3 Assessment of NuScale Facility Integral Effects Test Data The NIST IET data that supports the validation of NRELAP5 for NPM LOCA and IORV analysis includes the following tests. ((

                                                                                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                    }}2(a),(c)

In addition, the following IETs were performed at the NIST-2 facility. The purpose of these tests were to evaluate differences in system response for design changes in the NPM-20. Additional information about NIST-2 tests is in Section 7.5.10. ((

                                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Additional IETs were performed at NIST-2 to evaluate the phenomena related to ((

                                                                      }}2(a),(c)

As shown in Section 7.5.5 to Section 7.5.10, in general, NRELAP5 predicted the NIST IET data with excellent agreement. This shows that NRELAP5 is capable of predicting the phenomena and process occurring in the NIST facility including system interactions. Further, evaluations of these assessments for each high-ranked PIRT phenomenon are summarized in Table 8-19. 8.3.4 Evaluation of NuScale Integral Effects Tests Distortions and NRELAP5 Scalability The scaling and distortion analysis summarized in Section 8.3.2.4 identified and quantified scaling distortions in the NIST facility ((

                                     }}2(a),(c)

The NuScale NRELAP5 LOCA EM is updated to simulate NIST-1 and NIST-2 IETs in the NPM. ((

                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The results showed that the biases, differences, and distortions between the NPM design and the NIST facility can be accounted for using NRELAP5, and NRELAP5 is scalable to model phenomena and process in the NPM during LOCA events. 8.3.4.1 NuScale Facility Powered Natural Circulation Test (HP-05) The NPM relies on natural circulation flow as the primary mechanism to remove energy produced in the core and to deposit that energy in the SG tubes. The core power provides the driving force, with resistance to the flow caused by form and friction losses along the primary coolant path. The NIST-1 test facility is a scaled model of the NPM that uses the same natural circulation mechanism to move energy from the core heater rods to the model SG tubes. Scaling factors for the NIST-1 facility are as shown in Table 8-12. Test NIST-1 HP-05 was conducted to characterize the natural circulation flow rate and pressure drop in the NIST-1 test facility at various core power levels. As shown in Section 7.5.5 NRELAP5 predicted the HP-05 test data for the primary loop flow rate, core inlet temperature, and upper riser inlet temperature with reasonable-to-excellent agreement (Figure 7-88 to Figure 7-90). These results demonstrate the applicability of NRELAP5 to predict the natural circulation flow in the NIST-1 facility over a range of power levels. ((

                                               }}2(a),(c)

Figure 8-15 shows that the scaled NPM feedwater flow compares well with the test data. ((

                                                                 }}2(a),(c) Figure 8-16 shows comparison of scaled NPM natural circulation flow to the test data. ((
                                                                                                       }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                       }}2(a),(c)

Figure 8-15 Comparison of HP-05 Feedwater Flow to Test Data ((

                                                                                          }}2(a),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 8-16 Comparison of HP-05 Reactor Pressure Vessel Flow to Test Data ((

                                                                                                }}2(a),(c),ECI The upper riser inlet temperature comparison to test data is shown in Figure 8-17.

((

                                      }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 In summary, the comparisons of scaled NRELAP5 calculations of the NPM RPV flow and fluid temperatures to NIST-1 NRELAP5 calculations and the test data indicate that the NIST facility is well scaled. ((

                                                               }}2(a),(c)

Figure 8-17 Comparison of HP-05 Upper Riser Inlet Temperature to Test Data ((

                                                                                                      }}2(a),(c),ECI

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 8-18 Comparison of HP-05 Core Inlet Temperature to Test Data ((

                                                                                                  }}2(a),(c),ECI 8.3.4.2            NuScale Facility Loss-of-Coolant Accident and Inadvertent Reactor Vent Valve Opening Integral Effects Tests (HP-06, HP-07, and HP-09)

This section summarizes the results of the distortion analysis performed for the NIST-1 LOCA and inadvertent RVV opening IETs. The following initial/boundary condition biases, differences and scaling distortions between NPM and the as performed NIST-1 IETs have been identified to have a noticeable impact on the important LOCA parameters (i.e., RPV/CNV pressures and levels): Initial Conditions: The NIST-1 tests start from the steady-state natural circulation conditions. ((

                                    }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                                                         }}2(a),(c)

The impact of bias in some of the initial conditions is summarized below: Initial core power: ((

                                                           }}2(a),(c)

Initial RCS temperature/subcooling distribution: ((

                                                                                                     }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                  }}2(a),(c)

Initial reactor pool temperature: ((

                                                             }}2(a),(c)

Reactor Core Power following Reactor Trip: ((

                                       }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 CNV Wall Thickness and Material: ((

                         }}2(a),(c)

Steam Generator Secondary Side Operation and Quantity of Steam Generators: ((

                                      }}2(a),(c)

NIST-1 CNV Shell: ((

                                       }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                               }}2(a),(c)

RPV Outside Surface Heat Transfer: ((

                                                                                                   }}2(a),(c)

RPV stored energy: ((

                                                    }}2(a),(c) 8.3.5         Calculation of Peak CNV pressure Since containment is an integral part of the NPM ECCS, Section 4.3 identifies peak containment pressure as one of the LOCA EM FOMs. However, as identified earlier in Section 4.3, the peak containment pressure and temperature for containment performance are calculated with a different methodology (Containment Response Analysis Methodology - Reference 109). The top-down scaling analysis of  groups representing the inventory and energy balance equations (Section 8.3.2) can be used to provide more insights on the processes/phenomena governing the peak containment pressure. It is observed that the CNV pressurization during Phase 1a of the liquid space break is governed by ((
                                   }}2(a),(c). As described in Section 9.0 of this report, the peak CNV pressure occurs following ECCS actuation in liquid space breaks. It is observed that the major processes that contribute to CNV pressurization during the early part of

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Phase 1b are ((

                                                                                 }}2(a),(c) 8.4      Summary of Adequacy Findings 8.4.1         Findings from Bottom-Up Evaluation The bottom-up evaluation focused on determining the pedigree, applicability, fidelity to SET data, and scalability of the NRELAP5 closure relations and correlations that model the high-ranked phenomena as determined by the PIRT panel.

The pedigree of the identified closure relations and correlations was first established based on their historical development and subsequent assessment in the literature. Assessment cases were then identified to demonstrate the capability of NRELAP5 to predict the experimental data responses with reasonable-to-excellent agreement. Applicability of NRELAP5 to model the subject phenomena is established by demonstrating that the assessment cases cover the range of parameters that approximates the NPM range. The scalability evaluation was limited to whether the specific model or correlation is applicable for the NPM configuration over the range of conditions encountered in LOCA events. Based on the results from the NRELAP5 assessments with the NIST-2 IORV tests, the GE level swell tests, and the NPM core IORV analysis model evaluation, the applicability of NRELAP5 to the IORV analysis is evaluated. ((

                                                                                                    }}2(a),(c) is conservative in terms of the CHFR for the first few seconds of primary interest after the IORV initiation for the NPM IORV analysis.

((

                                                                                                    }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                    }}2(a),(c) to obtain reasonably conservative CHFR calculation results for the NPM IORV analysis with NRELAP5.

The IORV PIRT identified several phenomena that had some distinction from those identified in LOCA PIRT. These phenomena are identified in Table 4-5. The table also includes related LOCA phenomena that adequately assess these phenomena. Therefore, no new assessments of these phenomena is provided. The assessment of (( }}2(a),(c) was retained due to its relevance to some of these phenomena. Results of the bottom up evaluation are summarized in Table 8-18. © Copyright 2022 by NuScale Power, LLC 490

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3

                                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3

                                                                                                }}2(a),(c)

Table 8-18 Summary of Bottom-Up Evaluation of NRELAP5 Models and Correlations (Continued) ((

                  © Copyright 2022 by NuScale Power, LLC                                                                                          492

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3

                                                                                                               }}2(a),(c)

Table 8-18 Summary of Bottom-Up Evaluation of NRELAP5 Models and Correlations (Continued) ((

                  © Copyright 2022 by NuScale Power, LLC                                                                                  493

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3

                                                                                                     }}2(a),(c)

Table 8-18 Summary of Bottom-Up Evaluation of NRELAP5 Models and Correlations (Continued) ((

                  © Copyright 2022 by NuScale Power, LLC                                                                             494

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3

                                                                                                                      }}2(a),(c)

Table 8-18 Summary of Bottom-Up Evaluation of NRELAP5 Models and Correlations (Continued) ((

                  © Copyright 2022 by NuScale Power, LLC                                                                             495

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3

                                                                                                                                    }}2(a),(c)

Table 8-18 Summary of Bottom-Up Evaluation of NRELAP5 Models and Correlations (Continued) ((

                  © Copyright 2022 by NuScale Power, LLC                                                                               496

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                                                                                                                                    }}2(a),(c)

Table 8-18 Summary of Bottom-Up Evaluation of NRELAP5 Models and Correlations (Continued) ((

                  © Copyright 2022 by NuScale Power, LLC                                                                               497

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                                                                                                                   }}2(a),(c)

Table 8-18 Summary of Bottom-Up Evaluation of NRELAP5 Models and Correlations (Continued) ((

                  © Copyright 2022 by NuScale Power, LLC                                                                                 498

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                                                                                                }}2(a),(c)

Table 8-18 Summary of Bottom-Up Evaluation of NRELAP5 Models and Correlations (Continued) ((

                  © Copyright 2022 by NuScale Power, LLC                                                                                499

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 8.4.2 Findings from Top-Down Evaluation Results of the adequacy evaluation based on the NIST-1 IETs are summarized in Table 8-19 below. All high-ranked phenomena are included in the table. Where the NIST-1 IETs do not provide information, or provide limited information, regarding NRELAP5 applicability to model the phenomenon an explanation is provided. Areas not covered, or partly covered, by the IETs are addressed by SETs or other means, e.g., sensitivity studies, bounding assumptions, component test data. The NIST-2 IET results are consistent with the NIST-1 IETS. The NIST-2 testing did not uniquely address any high ranked phenomena, nor did the scaling analyses show any necessary re-working of the LOCA PIRT. High-ranked phenomena for the NPM-20 are consistent with those previously identified and assessed against NIST-1 tests. Although some design changes affect the loss-of-coolant event progression, there are no newly identified phenomena for the NPM-20. High-ranked phenomena were added from some SSCs or phases, but were previously evaluated by validation of NIST-1 testing. There are no significant changes to PIRT phenomena, (( }}2(a),(c) which is addressed in the Non-LOCA Topical Report (Reference 114). Table 8-19 Applicability Summary for High-Ranked Phenomena ((

                                                                                                    }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-19 Applicability Summary for High-Ranked Phenomena (Continued) ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-19 Applicability Summary for High-Ranked Phenomena (Continued) ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-19 Applicability Summary for High-Ranked Phenomena (Continued) ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-19 Applicability Summary for High-Ranked Phenomena (Continued) ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-19 Applicability Summary for High-Ranked Phenomena (Continued) ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 8-19 Applicability Summary for High-Ranked Phenomena (Continued) ((

                                                                                                    }}2(a),(c) 8.4.3         Summary of Biases and Uncertainties The NRELAP5 based LOCA EM was evaluated for applicability to analyzing LOCA events in the NPM. The applicability evaluation confirmed that the models and correlations in the NuScale LOCA EM are acceptable for simulating the important, i.e., high ranked, phenomena that determine the NPM response. Results of the LOCA EM applicability evaluation based on the bottom-up approach are summarized in Table 8-18. The overall evaluation of NRELAP5 applicability based on the top down approach is summarized in Table 8-19. The summaries in these tables show that the code is applicable for predicting LOCA response for the high-ranked phenomena that govern LOCA response in the NPM. A key element of the applicability confirmation is provided by SET and IET assessments that demonstrate reasonable-to-excellent agreement between NRELAP5 predictions and relevant experimental data.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 9.0 Loss-of-Coolant Accident Calculations The primary purpose of the representative break spectrum calculations and sensitivity studies presented in this section is to support the development of the LOCA EM and to demonstrate its application for the evaluation of the NPM ECCS performance during postulated LOCAs. The specific objectives of this section are to: describe the progression of typical LOCA scenarios in the NPM with regard to the key phenomena and processes during different phases of the LOCA identified by the PIRT (Section 4.0), present the results of the representative LOCA break spectrum calculations and other sensitivity calculations required by 10 CFR 50 Appendix K, and present the results of additional sensitivity calculations that address the uncertainties in modeling of key phenomena affecting the LOCA progression. 9.1 Loss-of-Coolant Accident Progression in the NuScale Power Module The example calculations for the LOCA EM are done for the NPM-20 design. The LOCA EM in this report is applicable to the NPM-160 design but the sample calculations presented in Revision 2 of this report are still applicable for the NPM-160 approved design. The LOCA progression for both a liquid and steam space break is presented in this section. A detailed discussion is provided for a 100 percent break of the RCS injection line and the high point vent line. As described in Section 4.2, the NPM LOCA has the following phases:

1. An initiation phase (Phase 0) that begins at blowdown initiation (RPV valve opening or LOCA) and runs in parallel with Phase 1a and 1b. ((

                                               }}2(a),(c)

2. A LOCA blowdown phase ( Phase 1a) begins with a postulated break in the RCS pressure boundary initiating a blowdown into the CNV and ends with opening the ECCS valves.

3. Phase 1b begins with the opening of ECCS valves resulting in pressure equalization between the RPV and CNV and the return of discharged fluid from the CNV to the RPV.

4. The long-term cooling phase begins when the pressure and level between the RPV and CNV stabilizes, and a stable natural recirculation flow pattern is established

(( }}2(a),(c) The LOCA calculations are extended to (( }}2(a),(c) following the flow reversal on the RRVs to ensure that the stable equilibrium collapsed levels are achieved in the riser. The LOCA scenarios described in the following sections assume full-break area, no loss of AC or DC power, and do not assume a single failure. These conditions were chosen to represent a typical application of the conservative 10 CFR 50 Appendix K LOCA EM. © Copyright 2022 by NuScale Power, LLC 507

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The following discussion uses the NPM-20 MPS logic and plant response. The overall methodology is also applicable to NPM-160 design, although timing and event sequences will vary due to MPS logic, ECCS valve and other design differences. Some examples include both NPM-160 and NPM-20 results for comparison. 9.1.1 Liquid Space Break Phase 0 Analysis (NPM-20) The RCS injection line connects from the CVCS system through the CNV and RPV into the riser above the core. The discharge line connects from the RPV downcomer above the RRV and exits through the RPV and CNV. Breaks in either of these lines between the RPV and the CNV discharges RCS fluid into the CNV. The Phase 0 LOCA progression for liquid space breaks (discharge line and injection line) is presented in this section. Results for the discharge line break spectrum for MCHFR with electric power (AC and EDAS) available are given in Table 9-1 and results for reactor power, pressurizer pressure, RCS flow and CHFR are plotted on Figure 9-1 through Figure 9-4. After break initiation, blowdown flow from the RPV to containment depressurizes the RPV. ((

                                                                               }}2(a),(c) Minimum CHFR occurs near the time of minimum flow before scram occurs. As the transient progresses, increasing containment pressure results in scram on the high containment pressure signal. Following scram, decreasing core power terminates the CHF excursion. MCHFR is not limiting at the end of Phase 0 since reactor power is significantly decreased while flow remains relatively high. The results show that larger breaks are limiting for MCHFR. Larger breaks result in faster depressurization, more core voiding, and a more severe flow transient.

Results for the injection line break spectrum MCHFR with electric power (AC and EDAS) available are presented in Table 9-2 and results for reactor power, pressurizer pressure, RCS flow and CHFR are plotted on Figure 9-5 through Figure 9-8. The overall system level progression of the injection line break is similar to that of the discharge line break. However, because the injection line connects to the RPV riser section above the core, ((

                                       }}2(a),(c) The discharge line break is more limiting compared to the injection line break due to the RPV break location.

When normal AC power is assumed lost at the time of the LOCA, the primary impact on the transient progression is the loss of secondary cooling and subsequent module © Copyright 2022 by NuScale Power, LLC 508

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 heat up. ((

                                                                                          }}2(a),(c) Again the discharge line break bounds the injection line break results when loss of AC power is assumed. Results for the discharge line break MCHFR with a loss of AC power are given in Table 9-1 and results for reactor power, pressurizer pressure, RCS flow and CHFR are plotted on Figure 9-9 through Figure 9-12. Results for the injection line break MCHFR with a loss of AC power are given in Table 9-2 and the results for reactor power, pressurizer pressure, RCS flow and CHFR are plotted in Figure 9-13 through Figure 9-16.

If EDAS power is lost at the time of the LOCA, safety-related valves go to their fail safe position. Most important for Phase 0 MCHFR is both RVVs open, which significantly increases the total break area. The RRVs are prevented from opening by the IAB since the vessel differential pressure is greater than the threshold pressure at the time EDAS power is lost. With both RVVs open, the RPV rapidly depressurizes and CHFR quickly reaches a minimum value in less than one second. Since the large total break area of both RVVs dominate the module response, the transient appears similar across all breaks sizes. However the influence of the discharge line break is still apparent, and the 100% break size is limiting. The injection line break has an insignificant impact on the transient compared to both RVVs being opened, and MCHFR is approximately equal across the break spectrum. As with the previous power scenarios, the discharge line break results bound the injection line break. Given the impact of both RVVs being open, the loss of DC power is the limiting electric power scenario for the liquid region LOCA breaks. Results for the discharge line break MCHFR with a loss of EDAS power are given in Table 9-1 and the results for reactor power, pressurizer pressure, RCS flow and CHFR are plotted on Figure 9-17 through Figure 9-20. Results for the injection line break MCHFR with a loss of EDAS power are given in Table 9-2 and results for reactor power, pressurizer pressure, RCS flow and CHFR are plotted on Figure 9-21 through Figure 9-24. © Copyright 2022 by NuScale Power, LLC 509

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 9-1 Discharge Line Break MCHFR Results ((

                                                                                                  }}2(a),(c)

Table 9-2 Injection Line Break MCHFR Results ((

                                                                                                  }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-1 Reactor Power for Discharge Line Break Spectrum - AC and EDAS Power Available ((

                                                                                            }}2(a),(c)

Figure 9-2 Pressurizer Pressure for Discharge Line Break Spectrum - AC and EDAS Power Available ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-3 RCS Flow for Discharge Line Break Spectrum - AC and EDAS Power Available ((

                                                                                          }}2(a),(c)

Figure 9-4 CHFR for Discharge Line Break Spectrum - AC and EDAS Power Available ((

                                                                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-5 Reactor Power for Injection Line Break Spectrum - AC and EDAS Power Available ((

                                                                                             }}2(a),(c)

Figure 9-6 Pressurizer Pressure for Injection Line Break Spectrum - AC and EDAS Power Available ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-7 RCS Flow for Injection Line Break Spectrum - AC and EDAS Power Available ((

                                                                                            }}2(a),(c)

Figure 9-8 CHFR for Injection Line Break Spectrum - AC and EDAS Power Available ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-9 Reactor Power for Discharge Line Break Spectrum - Loss of AC Power ((

                                                                                            }}2(a),(c)

Figure 9-10 Pressurizer Pressure for Discharge Line Break Spectrum - Loss of AC Power ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-11 RCS Flow for Discharge Line Break Spectrum - Loss of AC Power ((

                                                                                            }}2(a),(c)

Figure 9-12 CHFR for Discharge Line Break Spectrum - Loss of AC Power ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-13 Reactor Power for Injection Line Break Spectrum - Loss of AC Power ((

                                                                                             }}2(a),(c)

Figure 9-14 Pressurizer Pressure for Injection Line Break Spectrum - Loss of AC Power ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-15 RCS Flow for Injection Line Break Spectrum - Loss of AC Power ((

                                                                                              }}2(a),(c)

Figure 9-16 CHFR for Injection Line Break Spectrum - Loss of AC Power ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-17 Reactor Power for Discharge Line Break Spectrum - Loss of EDAS Power ((

                                                                                            }}2(a),(c)

Figure 9-18 Pressurizer Pressure for Discharge Line Break Spectrum - Loss of EDAS Power ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-19 RCS Flow for Discharge Line Break Spectrum - Loss of EDAS Power ((

                                                                                           }}2(a),(c)

Figure 9-20 CHFR for Discharge Line Break Spectrum - Loss of EDAS Power ((

                                                                                           }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-21 Reactor Power for Injection Line Break Spectrum - Loss of EDAS Power ((

                                                                                             }}2(a),(c)

Figure 9-22 Pressurizer Pressure for Injection Line Break Spectrum - Loss of EDAS Power ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-23 RCS Flow for Injection Line Break Spectrum - Loss of EDAS Power ((

                                                                                              }}2(a),(c)

Figure 9-24 CHFR for Injection Line Break Spectrum - Loss of EDAS Power ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 9.1.2 Steam Space Break Phase 0 Analysis The pressurizer spray line and the high point vent line are located in the steam space of the pressurizer. Breaks in either of these lines between the RPV and the CNV releases steam into the CNV. The Phase 0 LOCA progression for vapor space breaks is presented in this section. MCHFR results are summarized in Table 9-3. Results for these breaks with electric power available, a loss of AC power, and a loss of EDAS power are plotted in Figure 9-25 through Figure 9-36 for each of the following parameters: reactor power, pressurizer pressure, RCS flow and CHFR. The overall transient progression is very similar to that described for the liquid region breaks in Section 9.1.1. The key difference is the break location on the RPV. ((

                                                                 }}2(a),(c) Overall, the vapor space breaks are less limiting than the discharge line break, but more limiting than the injection line break in terms of MCHFR.

Similar to the liquid space breaks, the larger breaks yield a more limiting MCHFR when all electric power is available. When AC power is lost, ((

                                 }}2(a),(c) If EDAS power is lost, the larger breaks are again more limiting, but the overall transient is very similar for all breaks sizes due to the large RVV flow area.

For all power scenarios, the high point vent line breaks bounds the pressurizer spray line break. When AC power is assumed lost, the high point vent line cases are more limiting than the liquid LOCA breaks (( }}2(a),(c) However, under the most limiting loss of EDAS power scenario, the discharge line break bounds all vapor region breaks. Table 9-3 High Point Vent and Pressurizer Spray Line Break MCHFR Results ((

                                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 9-3 High Point Vent and Pressurizer Spray Line Break MCHFR Results (Continued) ((

                                                                                            }}2(a),(c)

Figure 9-25 Reactor Power High Point Vent and Pressurizer Spray Spectrum - AC and EDAS Power Available ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-26 Pressurizer Pressure High Point Vent and Pressurizer Spray Spectrum - AC and EDAS Power Available ((

                                                                                                }}2(a),(c)

Figure 9-27 RCS Flow High Point Vent and Pressurizer Spray Spectrum - AC and EDAS Power Available ((

                                                                                                }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-28 CHFR High Point Vent and Pressurizer Spray Spectrum - AC and EDAS Power Available ((

                                                                                          }}2(a),(c)

Figure 9-29 Reactor Power High Point Vent and Pressurizer Spray Spectrum - Loss of AC Power ((

                                                                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-30 Pressurizer Pressure High Point Vent and Pressurizer Spray Spectrum - Loss of AC Power ((

                                                                                          }}2(a),(c)

Figure 9-31 RCS Flow High Point Vent and Pressurizer Spray Spectrum - Loss of AC Power ((

                                                                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-32 CHFR High Point Vent and Pressurizer Spray Spectrum - Loss of AC Power ((

                                                                                           }}2(a),(c)

Figure 9-33 Reactor Power High Point Vent and Pressurizer Spray Spectrum - Loss of EDAS Power ((

                                                                                           }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-34 Pressurizer Pressure High Point Vent and Pressurizer Spray Spectrum - Loss of EDAS Power ((

                                                                                           }}2(a),(c)

Figure 9-35 RCS Flow High Point Vent and Pressurizer Spray Spectrum - Loss of EDAS Power ((

                                                                                           }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-36 CHFR High Point Vent and Pressurizer Spray Spectrum - Loss of EDAS Power ((

                                                                                                  }}2(a),(c) 9.1.3         Phase 1 LOCA Results Summary The following conclusions are reached based on the LOCA Phase 1 calculations and sensitivity studies:

((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                     }}2(a),(c)

It is shown that the 100 percent injection line break with DHRS operation, no loss of power, and failure of one ECCS division (one RRV/RVV failed to open) is the most limiting case for minimum collapsed level above TAF. Table 9-4 provides a summary of the limiting cases for each of the figures of merit considered in NPM LOCA evaluation model. The minimum Phase 1 MCHFR values do not challenge the 95/95 CHFR design limit and shows significant margin to the limit among all event types, break sizes, power scenarios, and single failures. Table 9-4 Summary of Phase 1 Limiting LOCA Parameters by Break Type ((

                                                                                                       }}2(a),(c) 9.1.3.1            Liquid Space Break Phase 1 Analysis The RCS injection line is the limiting event for CLL above TAF. The sequence of events is described in Table 9-5. The break occurs at time zero and choking at the break location occurs immediately. The mass and energy release into the CNV through the break results in rapid pressurization of the CNV and depressurization of the RPV. The MPS generates the reactor trip signal based on the high CNV pressure (greater than 9.5 psia) followed by CNV isolation. The reactor trip system includes a conservative delay before the scram reactivity insertion begins in the NPM core. The control rods drop to insert large negative reactivity and the control rod insertion is complete at approximately 7 seconds after break initiation.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 The Phase 0 LOCA MCHFR analysis provides the short-term limiting MCHFR analysis compared to the longer-term ECCS performance focused evaluation for Phase 1. Therefore, the MCHFR values reported in this section correspond to the minimum CHFR values calculated in Phase 1. The Phase 1 MCHFR is calculated from the end of Phase 0 to the end of transient (Phase 2). Phase 1a of the NPM LOCA includes the mass and energy release from the break location into the CNV and is terminated by the opening of the ECCS valves (Phase 1b). For the 100 percent injection line break scenario, the RVVs are actuated on the low riser level signal. Figure 9-37 compares the break flow with the net ECCS valve flow during the transient. The two-peak profile of the total ECCS flow is the result of staggered timing of RVV and RRV opening, with the RVVs opening first upon the low riser level signal (large peak) and the RRVs opening shortly thereafter upon reaching the IAB release pressure (small peak, followed by extended flow). The RPV and CNV pressure responses shown in Figure 9-38 are a result of the behavior of each component of the energy balance. The energy release to the CNV through the break and ECCS valve flow is significantly larger than the energy release to the RPV by core heat transfer. Due to a relatively early ECCS actuation timing, ECCS energy is the dominant heat transfer mechanism rather than break energy. Heat transfer from the CNV wall and DHRS to the reactor cooling pool or Ultimate Heat Sink (UHS) causes a continuous depressurization of both RPV and CNV after the initial pressurization of the CNV. As shown in Figure 9-38, the peak containment pressure occurs after the ECCS valve opening. As the RCS loses inventory, first through the break and later through the ECCS valves, the collapsed level continuously drops in the RPV riser and downcomer, as shown in Figure 9-39. After approximately 208 seconds, the pressurizer is completely emptied and the ECCS valves open when the collapsed liquid level is approximately 34 feet above TAF. The collapsed liquid level drops an additional 20 feet after the ECCS actuation. After the pressure equalization occurs between the RPV and CNV, an equilibrium level is achieved at approximately 10 feet above the TAF. Following the reactor trip and secondary isolation, the RCS flow is significantly reduced. ((

                                                }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                                                    }}2(a),(c) Since the core remains covered and CHF violation is not calculated, the peak cladding and fuel centerline temperature, shown in Figure 9-44, follow the core decay power transient (see Figure 9-42);

both the maximum cladding temperature (720 degrees F) and maximum fuel centerline temperature (3732 degrees F) across the NPM-20 core occur at the beginning of the transient similar to the transient minimum core MCHFR (Figure 9-43). Therefore, the 10 CFR 50.46 requirement of maximum allowed clad temperature of 2200 degrees F is not challenged. The Phase 1b of the NPM LOCA begins with the opening of the ECCS valves, which produces the flow reversal at the RRVs and has relatively short duration. The stable circulation flow across the ECCS valves is established following the RRVs opening, i.e. the steam flowing into the CNV is condensed inside the CNV and condensate flow enters the RPV through the RRVs. The depressurization of both RPV and CNV continues as the heat transfer into the CNV wall is significantly larger than the core power. It is important to note that due to large thermal inertia of the thick CNV wall, the major heat transfer occurs via heating up the wall before the heat eventually reaches the ultimate heat sink (UHS). The negative values of the heat transfer into the CNV wall and eventually the UHS denote the heat transfer out of the module. The LOCA progression described based on the full-size break of the RCS injection line is also valid for the discharge line as demonstrated in Figure 9-45 and Figure 9-46. Both RPV and CNV pressure transient and the collapsed liquid level in the RPV follow very similar trends except shifted timing of events. Initially the fluid in the downcomer section is approximately 132 degrees more subcooled than the fluid in the lower riser section. This results in substantially higher break flow due to flow being un-choked as shown in Figure 9-47; therefore, faster depressurization and quicker reduction of the liquid inventory inside the RPV yields early ECCS actuation. Figure 9-48 and Figure 9-49 provides a plot comparison of system pressure and system level for the NPM-160 and NPM 20 designs for the base case injection line break. Although the removal of IABs on the RVVs and the addition of the low riser trip for NPM-20 results in an earlier ECCS actuation timing, Figure 9-50 (NPM-160) and Figure 9-51 (NPM-20) have similar peak CNV pressure and level profiles for base case injection line break. The opening of the RVVs for NPM-20 immediately upon the lower riser level ECCS signal (rather than high CNV level) © Copyright 2022 by NuScale Power, LLC 533

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 results in an abrupt drop in collapsed level, rather than a gradual decline in collapsed level until reaching the high CNV level as observed for NPM-160. The dip in collapsed level below equilibrium level (i.e. <10ft above TAF) at the time of ECCS actuation does not occur for NPM-20, since ECCS actuation occurs at a point when collapsed level is relatively high. Table 9-5 RCS Injection Line Sequence of Events (Limiting Collapsed Level Case) ((

                                                                                                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-37 Break Flow and Net ECCS Valve Flow for 100 percent RCS Injection Line Break ((

                                                                                            }}2(a),(c)

Figure 9-38 RPV and CNV Pressure for 100 percent RCS Injection Line Break ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-39 Collapsed Liquid Levels for 100 percent RCS Injection Line Break ((

                                                                                             }}2(a),(c)

Figure 9-40 Recirculation Flow Rate and Core Flow Distribution for 100 percent RCS Injection Line Break ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-41 Core Flow Transient for 100 percent RCS Injection Line Break ((

                                                                                              }}2(a),(c)

Figure 9-42 Components of Reactor Power for 100 percent RCS Injection Line Break ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-43 Post-Scram MCHFR during 100 percent RCS Injection Line Break ((

                                                                                          }}2(a),(c)

Figure 9-44 Peak Clad and Fuel Centerline Temperature during 100 percent RCS Injection Line Break ((

                                                                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-45 Comparison of RPV and CNV Pressure Response 100 percent RCS Injection and Discharge Line Breaks ((

                                                                                                 }}2(a),(c)

Figure 9-46 Comparison of Collapsed Riser Level 100 percent RCS Injection and Discharge Line Breaks ((

                                                                                                 }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-47 Comparison of Break Flows between 100 percent RCS Injection and Discharge Line Breaks ((

                                                                                             }}2(a),(c)

Figure 9-48 NPM-160 RPV and CNV Pressure for Base Case RCS Injection Line Break ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-49 NPM-20 RPV and CNV Pressure for Base Case RCS Injection Line Break ((

                                                                                           }}2(a),(c)

Figure 9-50 NPM-160 Collapsed Liquid Levels for Base Case RCS Injection Line Break ((

                                                                                           }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-51 NPM-20 Collapsed Liquid Levels for Base Case RCS Injection Line Break ((

                                                                                                       }}2(a),(c) 9.1.3.2            Steam Space Break Phase 1 Analysis The largest steam space break occurs on the pressurizer high point vent line. The sequence of events is described in Table 9-6. The RPV and CNV pressure response for the high point vent line break is similar to the RCS injection line break, as shown in Figure 9-52. The 100 percent break on the high point vent line causes a reactor trip signal based on high containment pressure. The full insertion of the control rods is complete within approximately seven seconds of the break initiation. Containment isolation, secondary system isolation and DHRS actuation occur after the reactor trip signal.

Approximately 130 seconds following the LOCA, the collapsed liquid level drops below the top of the upper riser and the upper plenum becomes empty, disrupting natural circulation. ((

                                      }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 depressurization of the RPV and pressurization of the CNV, which immediately causes the reactor trip as shown in Table 9-6. The initial choked flow for the vent and the injection line breaks (Figure 9-53 and Figure 9-57) is very similar. With same break size, the pressurization rate of the CNV or the depressurization rate of the RPV is faster due to high enthalpy steam flow through the vent line break as shown in Figure 9-52. The collapsed level above the TAF decreases following the LOCA initiation as shown in Figure 9-54. As the RCS inventory loss continues due to the break, the collapsed level continues to decrease. The low riser level ECCS signal is reached at approximately 130 seconds. The actuation of the ECCS valves further reduces the RPV collapsed level and establishes the equilibrium level at approximately 9 feet, which is very similar to that of the injection line break. The post scram MCHFR does not challenge acceptance criteria as demonstrated in Figure 9-55. The post-scram MCHFR occurs immediately following reactor trip plus the Phase 0 window. ((

                          }}2(a),(c)

Figure 9-56 and Figure 9-57 provide a plot comparison of RPV and CNV pressure for the NPM-160 and NPM-20 designs for the base case high point vent line break. Figure 9-58 and Figure 9-59 compare collapsed liquid levels for the NPM-160 and NPM-20 designs for the base case high point vent line break. Although the removal of IABs on the RVVs and the addition of the low riser trip for NPM-20 results in an earlier ECCS actuation timing, NPM-160 and NPM-20 have similar pressure and level profiles for the base case high point vent line break, with respect to system depressurization rate and characteristic trends in collapsed level above TAF (i.e. decrease during initial blowdown, increase during draining of liquid from pressurizer to upper riser, decrease once pressurizer is empty). It is noted that in the NPM-160 case, ECCS actuates later on high CNV level, while peak CNV pressure occurs in LOCA Phase 1a. © Copyright 2022 by NuScale Power, LLC 543

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 9-6 RCS High Point Vent Line Sequence of Events - Limiting Collapsed Level Case ((

                                                                                          }}2(a),(c)

Figure 9-52 NPM-20 RPV and CNV Pressure Comparing Injection Line Break to High Point Vent Line Break ((

                                                                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-53 Break and ECCS Flow for 100 percent High Point Vent Line Break ((

                                                                                             }}2(a),(c)

Figure 9-54 Collapsed Liquid Levels for 100 percent High Point Vent Line Break ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-55 Post-Scram MCHFR for 100 percent High Point Vent Line Break ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-56 NPM-160 RPV and CNV Pressure Base Case High Point Vent Line Break ((

                                                                                        }}2(a),(c)

Figure 9-57 NPM-20 RPV and CNV Pressure Base Case High Point Vent Line Break ((

                                                                                        }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-58 NPM-160 Collapsed Liquid Levels Base Case High Point Vent Line Break ((

                                                                                          }}2(a),(c)

Figure 9-59 NPM-20 Collapsed Liquid Levels Base Case High Point Vent Line Break ((

                                                                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 9.2 Break Size The previous section discussed the LOCA progression based on two unique break locations, the RCS injection line and pressurizer high point vent line. The purpose of this section is to discuss the effects of break size and locations on the NPM-20 LOCA FOMs. The spectrum of break areas for different break locations are summarized in Table 5-5. The justification of the selected matrix is discussed in Section 5.4. The full-size breaks for RCS injection and high point vent line breaks are covered in detail the previous section. The maximum break area for the pressurizer spray line is determined by the orifice size on the connecting piping that has a diameter of 1 inch. The minimum break area for the liquid and steam space breaks is determined by examining wide range of break areas such that limiting values for the NPM LOCA FOMs are obtained within the analyzed range. The timing of events are directly affected by the break area via the choking flow rate at the break location. The break flow rate is proportional to the break area for similar upstream conditions. The smaller break size yields slower depressurization and lower mass/energy inventory redistribution between the RPV and CNV. Therefore, the transient times are longer compared to the larger break sizes. The CNV pressure as a function of time with different break areas are presented for the RCS injection line (labeled CL or IL on figures), RCS discharge line (DL), and high point vent (HPV) lines in Figure 9-61, Figure 9-64, and Figure 9-67, respectively. ((

                                                    }}2(a),(c)

The peak CNV pressures ((

                   }}2(a),(c)

Figure 9-62 and Figure 9-68 show the collapsed liquid level above TAF in the RPV riser section as a function of break size for both RCS injection and high point vent line breaks. The final equilibrium levels established after the pressure equalization are independent of break size and locations. The equilibrium collapsed level is directly related to the geometry of the RPV/CNV as well as the value of the equilibrium pressure between two pressure vessels. As shown in Figure 9-68 and Figure 9-74, the minimum collapsed level is typically the equilibrium level of approximately 10 feet for all break sizes. © Copyright 2022 by NuScale Power, LLC 549

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                      }}2(a),(c)

Section 9.1 discusses the LOCA progression based on the full size liquid and steam space breaks and evaluates the limiting Phase 0 MCHFR. The post-scram MCHFR as a function of break size is plotted in Figure 9-69 and Figure 9-70 for both RCS injection and high point vent line breaks. The post-scram CHFR margin rapidly increases following the reactor trip ((

                                                                                                     }}2(a),(c)

Figure 9-73 and Figure 9-74 summarize two other important LOCA FOMs; namely peak CNV pressure and minimum collapsed level above TAF in the RPV as a function of break size and location. Figure 9-74 includes all the break size and locations with DHRS operation, without loss of power, and no single failures. ((

                                         }}2(a),(c) The minimum collapsed levels of approximately 9 feet above TAF are observed with all of the steam-space breaks and liquid-space breaks.

Figure 9-72 shows the timing of the ECCS actuation as a function of the break size at different break locations and Figure 9-71 shows the pressure differential across the RVVs when the ECCS actuates. For the smallest break size, the ECCS actuation time ranges from about 50 to 250 minutes after the break initiation. The post-scram MCHFR is observed following the reactor trip and the Phase 0 window. © Copyright 2022 by NuScale Power, LLC 550

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-60 RPV Pressure Transient for RCS Injection Line Breaks ((

                                                                                               }}2(a),(c)

Figure 9-61 CNV Pressure Transient for RCS Injection Line Breaks ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-62 Collapse Liquid Level above TAF for RCS Injection Line Breaks ((

                                                                                              }}2(a),(c)

Figure 9-63 RPV Pressure Transient for RCS Discharge Line Breaks ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-64 CNV Pressure Transient for RCS Discharge Line Breaks ((

                                                                                              }}2(a),(c)

Figure 9-65 Collapse Liquid Level above TAF for RCS Discharge Line Breaks ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-66 RPV Pressure Transient for High Point Vent Line Breaks ((

                                                                                               }}2(a),(c)

Figure 9-67 CNV Pressure Transient for High Point Vent Line Breaks ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-68 Collapse Liquid Level above TAF for High Point Vent Line Breaks ((

                                                                                               }}2(a),(c)

Figure 9-69 Post-Scram MCHFR for RCS Injection Line Breaks ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-70 Post-Scram MCHFR for High Point Vent Line Breaks ((

                                                                                               }}2(a),(c)

Figure 9-71 Pressure Differential across the RVV at Actuation by Break Size and Location ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-72 ECCS Valve Actuation Time by Break Size and Location ((

                                                                                                }}2(a),(c)

Figure 9-73 CNV Peak Pressure by Break Size and Location ((

                                                                                                }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-74 Minimum Collapsed Liquid Level by Break Size and Location ((

                                                                                                      }}2(a),(c) 9.3      Decay Heat Removal System The DHRS is a safety-related system and is assumed to function per design in all LOCA scenarios. The DHRS adds an additional heat sink capacity during the NPM LOCA that impacts primarily the smaller break sizes. ((
                                                      }}2(a),(c) However, when the DHRS operation is taken into account, all break sizes behave similarly and minimum collapsed liquid levels are the same as the final equilibrium level for most all of the break sizes. Sensitivities for DHRS performance are included in Section 9.6.

9.4 Power Availability The discussion presented previously assumes that both AC and DC power are available during the NPM-20 LOCA. Loss of power is considered by assuming loss of only AC and loss of both AC and DC (EDAS for NPM-20). Figure 9-75 demonstrates that the loss of both AC and EDAS power has significant impact on peak containment pressure for the high point vent line break. The loss of all power causes immediate reactor trip and de-energizing the solenoids of the ECCS valves. Because the IABs were removed from the RVVs for the NPM-20 design, the RVV valve opening is immediate and not governed by the IAB release pressure as it is in the NPM-160 design. The immediate RVV opening caused by the loss of AC and EDAS power results in higher CNV peak pressures across the break sizes for the HPV break compared to the all power available scenario. Also, the © Copyright 2022 by NuScale Power, LLC 558

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 CNV peak pressure values across the break sizes for the high point vent and injection line breaks are all around ((

                                             }}2(a),(c)

Figure 9-75 Effect of Power Availability on Peak CNV Pressure for RCS Injection Line Break ((

                                                                                                 }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-76 Effect of Power Availability on Peak CNV Pressure for High Point Vent Line Break ((

                                                                                             }}2(a),(c)

Figure 9-77 Effect of Power Availability on Collapsed Liquid Level for RCS Injection Line Break ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-78 Effect of Power Availability on Collapsed Liquid Level for High Point Vent Line Break ((

                                                                                                      }}2(a),(c) 9.5      Single Failure In the previous discussion, no single failure is assumed. As discussed in Section 5.4.3, the following single failures are considered in this section:

failure of a single RVV to open, failure of a single RRV to open, and failure of one ECCS division (i.e., one RVV and one RRV) Figure 9-79 and Figure 9-80 demonstrate that the single failures listed above have ((

                                                                                                 }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                          }}2(a),(c)

Figure 9-79 Effect of Single Failures on Peak CNV Pressure for RCS Injection Line Break ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-80 Effect of Single Failures on Peak CNV Pressure for High Point Vent Line Break ((

                                                                                              }}2(a),(c)

Figure 9-81 Effect of Single Failures on Collapsed Liquid Level for RCS Injection Line Break ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-82 Effect of Single Failures on Collapsed Liquid Level for High Point Vent Line Break ((

                                                                                              }}2(a),(c)

Figure 9-83 Effect of Single Failures on Collapsed Liquid Level for RCS Injection Line Break ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-84 Effect of Single Failures on RRV Flow for RCS Injection Line Break ((

                                                                                              }}2(a),(c)

Figure 9-85 Effect of Single Failures on Collapsed Liquid Level for RCS Discharge Line Break ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-86 Effect of Single Failures on RRV Flow for RCS Discharge Line Break ((

                                                                                              }}2(a),(c)

Figure 9-87 Effect of Single Failures on Collapsed Liquid Level for High Point Vent Line Break ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-88 Effect of Single Failures on RRV Flow for High Point Vent Line Break ((

                                                                                                      }}2(a),(c)

Figure 9-89 Peak CNV Pressure with DHRS Operation for Break Locations with Different Power Availability and Single Failures ((

                                                                                                      }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-90 Collapsed Liquid Level with DHRS Operation for Break Locations with Different Power Availability and Single Failures ((

                                                                                                      }}2(a),(c) 9.6      Sensitivity Studies Several sensitivity studies are performed to establish the basis for the NuScale LOCA EM. Some of the sensitivity studies in this section (Sections 9.6.1 through 9.6.6) were done for the NPM-160, but the generic nature of the evaluation makes them applicable to the NPM-20 as well. Similarly, some sensitivity studies were done with NPM-20 example data but also apply to the NPM-160. For cases that are design specific, notations are made indicating module applicability. The sensitivity calculations are performed to address the effects of the modeling parameters such as nodalization, time-step size selection, CCFL behavior at the pressurizer baffle plate, ECCS valve parameters (such as IAB release pressure differential threshold, size/capacity, as well as valve stroke time).

An additional sensitivity study is performed on core power distribution addressing the effects of core axial power shape and radial peaking assigned to the hot assembly. The sensitivity calculations are also performed to determine the impact of initial reactor cooling pool temperature. Justifications for other initial and boundary conditions selected for the conservative LOCA analysis are provided in Section 5.3. 9.6.1 Model Nodalization Performing a nodalization sensitivity study is important to determine its impact on the key LOCA FOMs such as peak containment pressure and collapsed liquid level above TAF in the RPV riser. As described in Section 5.1, the NRELAP5 model uses one-dimensional components. In order to address the impact of nodalization on the © Copyright 2022 by NuScale Power, LLC 568

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 NPM LOCA behavior, three nodalization schemes that conform to general NRELAP5 modeling guidelines are selected as shown in Table 9-7. Table 9-7 Number of Volumes in Reactor Pressure Vessel and Containment Vessel Nodalization ((

                                                                                                       }}2(a),(c)

The full range of break sizes for both RCS injection line and high point vent line breaks with three nodalization schemes are investigated. Both break locations are considered without DHRS operation, no loss of power, and no single failure. Figure 9-91 shows the RPV and CNV pressures and collapsed liquid level above TAF in RPV riser for the RCS injection break with 100 percent break area without DHRS, no loss of power, and no single failure. The same parameters are plotted in Figure 9-92 for the RCS injection break with 10 percent break area without DHRS, no loss of power, and no single failure. With three different nodalization schemes, two key LOCA FOMs are shown to be similar including timing of event during the transient, ((

                      }}2(a),(c)

The results shown in Figure 9-93 for the 100 percent high point vent line break cases show the similarities. ((

                                                                    }}2(a),(c) three different nodalization schemes provide similar LOCA response in RPV and CNV pressures and collapsed levels for the high point vent line break scenario.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-91 Reactor Pressure Vessel and Containment Vessel Pressure (Left) and Collapsed Level Above Top of Active Fuel (Right) for 100 Percent Reactor Coolant System Injection Line Break ((

                                                                                                   }}2(a),(c)

Figure 9-92 Reactor Pressure Vessel and Containment Vessel Pressure (Left) and Collapsed Level above Top of Active Fuel (Right) for 10 Percent Reactor Coolant System Injection Line Break ((

                                                                                                   }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-93 Reactor Pressure Vessel and Containment Vessel Pressure (Left) and Collapsed Level above Top of Active Fuel (Right) for 100 Percent High Point Vent Line Break ((

                                                                                                           }}2(a),(c)

The hot assembly mass flux and core-wide MCHFR during RCS injection line break are shown in Figure 9-94 for the three nodalization schemes. The initial core MCHFR at the beginning of the transient is not affected by the number of hydrodynamic volumes in an NPM core, ((

                                                                                      }}2(a),(c) Furthermore, very similar MCHFRs are calculated for the ((
                                                                            }}2(a),(c) correlation. When the hot channel assembly flow goes ((
                    }}2(a),(c) correlation that includes the ((                               }}2(a),(c) is used.

((

                                         }}2(a),(c) Therefore, the core nodalization has no material impact on predicting the CHF margin during a postulated NPM LOCA. However, a small difference is observed in the initial hot assembly flows with coarse and coarser nodalization. As described in Table 9-7, both coarse and coarser nodalization schemes use a coarse representation of the core and steam generators. As a result, the steady state natural circulation flow rate is slightly different when compared to the finer nodalization due to relatively small shift in natural circulation loop thermal center.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-94 Hot Channel Core Flow (Left) and Core Critical Heat Flux Ratio (right) During 100 Percent Reactor Coolant System Injection Line Break ((

                                                                                                          }}2(a),(c) 9.6.2         Time-Step Size Selection The NRELAP5 NuScale LOCA EM uses a semi-implicit numerical scheme with implicit coupling of the hydrodynamic and heat conduction solutions. The time-step size is restricted by the courant time-step size and the accumulation of the mass-error during the time integration. In general, the NPM LOCA simulations have a courant time-step size at approximately ((                  }}2(a),(c) In order to address the effect of time-step size selection on the key NPM LOCA FOMs, various fractions of the problem courant time-step size are examined as shown in Figure 9-95 through Figure 9-98 for full size injection line and high point vent line breaks. For multipliers above approximately ((                                      }}2(a),(c) the max time-step size allowed for the calculations is mainly determined by the mass-error management. The figures show that the containment and RPV pressures, minimum collapsed level above the TAF in the RPV riser, hot channel mass flux, and hot channel MCHFR are all independent of the time-step sizes selected for the simulation.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-95 Time-Step Size Sensitivity on Reactor and Containment Vessel Pressures and Reactor Pressure Vessel Collapsed Liquid Level for 100 Percent Reactor Coolant System Injection Line Break. ((

                                                                                                    }}2(a),(c)

Figure 9-96 Time-Step Size Sensitivity on Hot Assembly Flow and Minimum Critical Heat Flux Ratio for 100 Percent Reactor Coolant System Injection Line Break ((

                                                                                                    }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-97 Time-Step Size Sensitivity on Reactor and Containment Vessel Pressures and Reactor Pressure Vessel Collapsed Liquid Level for 100 Percent High Point Vent Line Break ((

                                                                                                    }}2(a),(c)

Figure 9-98 Time-Step Size Sensitivity on Hot Assembly Flow and Minimum Critical Heat Flux Ratio for 100 Percent High Point Vent Break ((

                                                                                                    }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 9.6.3 Counter Current Flow Limitation Behavior on Pressurizer Baffle Plate ((

                      }}2(a),(c) A few of the break spectrum cases activated the CCFL flag at the pressurizer baffle plate, which did not allow liquid to readily drain from the pressurizer to the downcomer in the presence of upward steam flow. These break cases were limited to the larger pressurizer spray and vent line breaks. A study was performed

((

                             }}2(a),(c)

Figure 9-99 Effect of Counter Current Flow Limitation Line Slope on Levels for 100 Percent High Point Vent Line Break ((

                                                                                                        }}2(a),(c) 9.6.4         Emergency Core Cooling System Valve Parameters (NPM-160 only)

Operation of the ECCS valves varies based on the valve characteristics. The NPM ECCS valve specification provides minimum and maximum valve sizes and a range of differential pressures at which the IAB arming valve closes (locks) and opens © Copyright 2022 by NuScale Power, LLC 575

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 (releases). A study was performed with liquid and steam breaks to evaluate separate and combined effects of the range of these valve characteristics on the LOCA FOMs. Figure 9-100 shows the effect of IAB release pressure on peak CNV pressure and minimum collapsed liquid level as function of break size for injection line break. Since the large break size results in relatively rapid RCS depressurization, ((

                                              }}2(a),(c)

Figure 9-101 and Figure 9-102 show the effect of RRV and RVV sizes on peak CNV pressure and minimum collapsed liquid level as function of break size. Overall the impact on ECCS valve size on peak CNV pressure and collapsed liquid level is (( }}2(a),(c). Figure 9-101 shows only (( }}2(a),(c) in minimum collapsed liquid level with (( }}2(a),(c). Conclusions of this study show that the ((

                                                                                                     }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-100 Effect of Inadvertent Actuation Block Release Pressure on Peak Containment Vessel Pressure and Minimum Collapsed Liquid Level above Top of Active Fuel as a Function of Break Size for Reactor Coolant System Injection Line Break ((

                                                                                               }}2(a),(c)

Figure 9-101 Effect of Reactor Recirculation Valve Size on Peak Containment Vessel Pressure and Minimum Collapsed Liquid Level for Reactor Coolant System Injection Line Break ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-102 Effect of Reactor Vent Valve Size on Peak Containment Vessel Pressure and Minimum Collapsed Liquid Level for Reactor Coolant System Injection Line Break ((

                                                                                                              }}2(a),(c) 9.6.5         Initial Reactor Pool Temperature The maximum initial reactor cooling pool temperature of ((                                   }}2(a),(c) is used in the LOCA break spectrum calculations as discussed in the previous sections.

Sensitivity studies covering the range of initial pool temperatures are performed to investigate the impact on the NuScale LOCA EM FOMs. Reactor pool temperatures ranging from (( }}2(a),(c) to (( }}2(a),(c) are considered. The RCS injection line break with sizes down to (( }}2(a),(c) of the full-break size break area are analyzed. The peak CNV pressure and the minimum collapsed liquid level above TAF as a function of break size are plotted for the pool temperatures of (( }}2(a),(c) in Figure 9-103. The effect of the initial pool temperature on the peak CNV pressure is more pronounced at (( }}2(a),(c). Figure 9-104 compares the various components of the RPV and CNV energy balance for 100 percent (left figure) and 10 percent (right figure) injection line breaks. ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                             }}2(a),(c) For all the initial pool temperatures investigated in the sensitivity calculation, no CHF violation is observed; therefore, the minimum MCHFR is defined by a value close to the steady state value.

Figure 9-103 Effect of Initial Reactor Pool Temperature on Peak Containment Vessel Pressure and Minimum Collapsed Liquid Level above Top of Active Fuel for Reactor Coolant System Injection Line Break ((

                                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-104 Containment Vessel to Pool Energy Transfer at Different Initial Pool Temperatures for 100 Percent (Left) and 10 Percent (Right) Reactor Coolant System Injection Line Break ((

                                                                                                         }}2(a),(c) 9.6.6         Core Power Distribution The sensitivity study is performed based on a full-range of break sizes for the RCS injection line break for the core power distribution considering:

Generic axial power shapes to bound the axial peakings (( }}2(a),(c) core channel Generic axial power shapes as shown in Figure 9-105 are used to investigate the effect on the key LOCA behavior and FOMs. The axial power shapes are chosen to represents a typical (( }}2(a),(c). The (( }}2(a),(c) shown in Figure 9-105 is used for all the LOCA calculations performed in this report. The axial peaking is determined to bound the values observed in the NPM core design. ((

                                                                                                  }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-105 Generic Normalized Axial Power Shapes ((

                                                                                                     }}2(a),(c)

The RCS injection line break with full break area spectrum is analyzed without DHRS operation, no power loss, and no single failure. Figure 9-106 compares the RPV and CNV pressures and collapsed liquid level above TAF for different axial power shapes for RCS injection line break. Figure 9-107 shows the impact of axial power shapes on peak CNV pressure and minimum collapsed liquid level as function of different injection line break sizes. ((

                        }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-108 shows the impact of axial power shapes on the hot assembly mass flux and MCHFR ((

                                        }}2(a),(c) However, in all of the cases analyzed, the minimum core MCHFR transient value is close to the value determined at the initiation of the event.

Figure 9-106 Effect of Axial Power Shape on Reactor Pressure Vessel and Containment Pressures and Collapsed Liquid Level above Top of Active Fuel for Reactor Coolant System Injection Line Break ((

                                                                                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-107 Effect of Axial Power Shape on Peak Containment Vessel Pressure and Minimum Collapsed Liquid Level above Top of Active Fuel for Reactor Coolant System Injection Line Break ((

                                                                                               }}2(a),(c)

Figure 9-108 Effect of Axial Power Shape on Hot Assembly Flow and Minimum Critical Heat Flux Ratio during Reactor Coolant System Injection Line Break ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 9.6.7 DHRS Capacity The following sensitivity was performed to identify the sensitivity of the LOCA event progression and figures of merit to the DHRS capacity assumed upon DHRS actuation. The base case LOCA events were re-run with reduced DHRS capacity ((

                       }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-109 RPV Pressure for Injection Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) ((

                                                                                              }}2(a),(c)

Figure 9-110 CNV Pressure for Injection Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-111 Collapsed Liquid Level for Injection Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) ((

                                                                                                 }}2(a),(c)

Figure 9-112 CNV Pressure for Injection Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) ((

                                                                                                 }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-113 RPV Pressure for High Point Vent Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) ((

                                                                                               }}2(a),(c)

Figure 9-114 CNV Pressure for High Point Vent Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-115 Collapsed Liquid Level for High Point Vent Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) ((

                                                                                               }}2(a),(c)

Figure 9-116 CNV Pressure for High Point Vent Line Break with DHRS 100 percent Capacity (solid line) vs. (( }}2(a),(c) Capacity (dashed) ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 9.7 Primary and Secondary System Release Scenario Containment Response Analysis This section presents sample results of the NRELAP5 limiting analyses of the spectrum of primary system break scenarios for the NPM-20, listed in the following Table 9-8, that are determined using the LOCA analysis methodology presented earlier in this report. Input parameters and assumptions are chosen to maximize peak CNV temperature and pressure. Example results for the NPM-160 are presented in technical report TR-0916-51299 (Reference 11). The primary side events include the CVCS discharge line break (case prefix DL), CVCS charging or injection line break (CL), HPV line break (HPV), inadvertent RVV opening (RVV), inadvertent RRV opening (RRV), and inadvertent actuation of ECCS signal (ECC). Following transient initiation, mass and energy released into the CNV through the break or opening cause CNV pressurization and RPV depressurization. Depending on the evaluated scenario, the high containment pressure signal or loss of EDAS power causes reactor trip, containment isolation, and DHRS actuation. Reactor trip, containment isolation, and DHRS actuation occur immediately when all electric power is assumed lost. As the transient progresses, coolant is continuously lost to the CNV through the break or valve opening until all ECCS valves open. When EDAS power is available, the ECCS valve opening occurs once an ECCS actuation signal is generated. For the loss of normal AC and EDAS power, the ECCS valve opening occurs immediately following the loss of power. While the RVVs open once an ECCS actuation signal is generated, IABs prevent the RRVs from opening until the IAB release pressure is reached. After the ECCS valves open, coolant lost to CNV is free to flow back into the RPV downcomer through the RRVs. A two-phase natural circulation cooling loop is established and the module transitions to long-term cooling. Table 9-8 Primary System Mass and Energy Release Scenarios Subsequent RVV and RRV Initiating Event Analysis Case Actuations on ECCS LOCA in RCS discharge line (DL) from Two RVVs and two RRVs actuate 1 downcomer LOCA in RCS injection line from riser Two RVVs and two RRVs actuate 2 LOCA in RPV high point vent line near top of Two RVVs and two RRVs actuate 3 vessel RVV opens due to a mechanical failure One RVV and two RRVs actuate 4 RRV opens due to a mechanical failure Two RVVs and one RRV actuate 5 Both RVVs open due to inadvertent ECCS Two RRVs actuate at IAB release 6 signal point 9.7.1 Analysis Approach The approach to determine the limiting peak CNV pressure event from the spectrum of primary mass and energy release scenarios for an NPM, and the limiting peak CNV temperature for each primary release event was as follows: ((

                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 ((

                       }}2(a),(c) 9.7.2         Base Case Analysis and Sensitivity Results The following insights are obtained from the results of the NRELAP5 analyses of the five primary system M&E release cases and associated sensitivity studies.

The peak CNV pressure scenario is the DL break (Case 1) with a loss of normal AC power. The DL break mass and energy release causes an initial heatup and pressurization of the CNV, and then ECCS actuation results in a second M&E release with all RVVs and RRVs opening, which pressurize the CNV to the highest peak pressure. The peak CNV wall temperature scenario is also the DL break (Case 1) with a loss of normal AC power. The break in this location combines a high temperature liquid initial M&E release followed by a high temperature M&E release through both RVVs following an ECCS actuation signal. The sensitivity parameters have only a small effect on the peak CNV pressure and temperature results of the limiting cases. No single failures had a significant © Copyright 2022 by NuScale Power, LLC 590

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 impact on the results for the limiting cases. The loss of power sensitivity that results in early ECCS actuation was the most important sensitivity parameter. The value of the RPV riser level ECCS actuation setpoint did not impact limiting CNV pressure or temperature. The RPV riser level setpoint did have a minor impact on some non-limiting cases. For the non-limiting all power available case, the low bias ECCS RPV riser level setpoint (540 inches) resulted in higher CNV wall temperature than the high bias setpoint (552 inches). 9.7.3 Primary Release Scenario Pressure and Temperature Results The initial conditions used by NRELAP5 analyses for each of the six cases in Table 9-8 are shown in Table 9-9. The initial condition values in the second column of Table 9-9 are the nominal values plus the uncertainty or conservative allowance in parentheses. The assumed parameter values are consistent with the LOCA methodology and maximize heat sources while minimizing heat sinks. The decay heat conservatively used by these analyses is 120 percent of the 1979 ANS standard rather than the methodology assumption (1979 ANS standard plus 2-sigma uncertainty). The 120 percent assumption bounds the required 2-sigma uncertainty. Table 9-9 Initial Conditions for Mass and Energy Release Event Analyses ((

                                                                                                       }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 9-9 Initial Conditions for Mass and Energy Release Event Analyses (Continued) ((

                                                                                                          }}2(a),(c)

The results of each of the five primary containment response analysis release analysis cases are summarized in the following sections, with more detailed results and discussion provided for the limiting CNV peak pressure scenario, and for the limiting CNV peak temperature scenario (Case 1 - RCS DL break LOCA). 9.7.3.1 Case 1: Reactor Coolant System Discharge Line Break Loss-of-Coolant Accident The LOCA in the RCS discharge line initiates an M&E release from the downcomer into the CNV. A summary of the spectrum of DL line breaks and results is included in Table 9-11. The sequence of events for the limiting case (Table 9-11, DL-2) is shown in Table 9-10. The CNV pressure response and temperature response are shown in Figure 9-117 and Figure 9-118. The CNV peak pressure is 910 psia for the base case (Table 9-11, Case DL-1), and 937 psia with the combined effect of the adverse sensitivity parameters (loss of normal AC power, Table 9-11, Case DL-2). The peak CNV wall temperature is 531 degrees F for the base case, and 533 degrees F with the combined effect of the adverse sensitivity parameters (loss of normal AC power). For the DL break, a loss of normal AC power at event initiation is more limiting than all power available with respect to peak CNV pressure and wall temperature, since it results in an immediate turbine trip (i.e., termination of SG heat removal) while the reactor trip is delayed until the High CNV Pressure signal is reached (Table 9-11, Case DL-2). Also, a loss of normal AC and EDSS power is less limiting than all power available with respect to peak CNV pressure and wall temperature, since it results in an immediate ECCS actuation, at which point no significant liquid level has accumulated within the CNV to occupy CNV free volume prior to ECCS valve opening (Table 9-11, Case DL-3). For the DL break, a low-biased Low RCS Level setpoint (i.e. 540") is more limiting with respect to peak CNV wall temperature for the all power available scenario, since it results in a later ECCS actuation and longer duration of CNV wall heatup prior to RVV and RRV opening and rapid system depressurization (Table 9-11, © Copyright 2022 by NuScale Power, LLC 592

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Case DL-4). Assuming single failure of an RVV, RRV, or RVV/RRV train results in a slight decrease in peak CNV pressure for the all power available scenario, since it limits the release rate through the RVVs/RRVs upon ECCS actuation. Figures showing collapsed liquid level, CNV level and CNV wall temperature are shown in Figure 9-119, Figure 9-120 and Figure 9-121, respectively. Table 9-10 Case 1 Sequence of Events - Reactor Coolant System Discharge Line Break Loss-of-Coolant Accident Peak CNV Peak CNV Pressure Event Temperature Case Case Time (sec) Time (sec) LOCA in RCS discharge line For peak pressure case:

Loss of normal AC power 0
Turbine trip Same
Loss of feedwater For peak temperature case:
Same High CNV pressure signal resulting in:

For peak pressure case:

Containment isolation
Reactor trip 1
DHRS actuation Same
Secondary system isolation (SSI)
CVCS isolation For peak temperature case:
Same ECCS actuation on low RPV riser level (RVVs open after 87 Same 60 second delay, RRVs armed)

RVVs open, RRVs open shortly afterward on IAB release 147 Same pressure Peak CNV temperature reached: 168 Same For peak pressure case: 533 °F For peak temperature case: Same Peak CNV pressure is reached: 170 Same For peak pressure case: 937 psia For peak temperature case: Same ~3190 CNV pressure decreases to <50% of peak pressure Same © Copyright 2022 by NuScale Power, LLC 593

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 9-11 Spectrum of Discharge Line Breaks Evaluated for Peak CNV Temperature and Pressure ((

                                                                                                         }}2(a),(c) 9.7.3.2            Case 2: Limiting Loss-of-Coolant Event - Reactor Coolant System Injection Line Break Loss-of-Coolant Accident The LOCA in the RCS injection line (CL) initiates an M&E release from the riser into the CNV. The results of the primary release event M&E release break spectrum analysis and sensitivity analyses have determined that Case 2 is not limiting for CNV peak pressure and CNV wall temperature events. A summary of the spectrum of CL line breaks and results is included in Table 9-12. The detailed results for key parameters are shown in Figure 9-122 through Figure 9-126. The CNV peak pressure is 890 psia for the base case, and 909 psia with the combined effect of the adverse sensitivity parameters (loss of normal AC). The peak CNV wall temperature is 528.5 degrees F for the base case, and 531 degrees F with the combined effect of the adverse sensitivity parameters. The primary sensitivity parameter that contributes to the temperature increase is the loss of normal AC power.

The following observations are made from the CL event results presented: Electric Power Availability - For the CL break, a loss of normal AC power at event initiation (Table 9-12, Case CL-2) is more limiting than all power available (Table 9-12, Case CL-1) with respect to peak CNV pressure and wall temperature, since it results in an immediate turbine trip (i.e. termination of SG heat removal) while the reactor trip is delayed until the High CNV Pressure signal is reached. Also, a loss of normal AC and EDSS power (Table 9-12, Case CL-3) is less limiting than all power available with respect to peak CNV pressure and wall temperature, since it results in a much earlier ECCS actuation. Low RCS Level Signal - For the CL break, a low-biased Low RCS Level setpoint (i.e. 540") is more limiting with respect to peak CNV wall temperature for the all power available scenario, since it results in a later ECCS actuation and longer © Copyright 2022 by NuScale Power, LLC 594

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 duration of CNV wall heatup prior to RVV and RRV opening and rapid system depressurization. Single Failures - For the CL break, assuming single failure of an RVV, RRV, or RVV/RRV train results in a slight decrease in peak CNV pressure for the all power available scenario, since it limits the release rate through the RVVs/RRVs upon ECCS actuation. Table 9-12 Spectrum of Injection Line Breaks Evaluated for Peak CNV Temperature and Pressure ((

                                                                                                         }}2(a),(c) 9.7.3.3            Case 3: Reactor Pressure Vessel High Point Vent Line Loss-of- Coolant Accident The LOCA in the RPV high point vent line initiates an M&E release from the top of the pressurizer into the CNV. A summary of the spectrum of HPV line breaks and results is included in Table 9-13. The CNV pressure response and temperature response are shown in Figure 9-117 and Figure 9-118. The CNV peak pressure is 807 psia for the base case, and 877 psia with the combined effect of the adverse sensitivity parameters (loss of normal AC and EDAS). The peak CNV wall temperature is 517 degrees F for the base case, and 525 degrees F with the combined effect of the adverse sensitivity parameters (loss of normal AC and DC power).

The following observations are made from the HPV event results presented: Electric Power Availability - For the HPV break, a loss of normal AC power at event initiation (Table 9-13, Case HPV- 2) is more limiting than all power available (Table 9-13, Case HPV- 1) with respect to peak CNV pressure and wall temperature, since it results in an immediate turbine trip (i.e. termination of SG heat removal) while the reactor trip is delayed until the High CNV Pressure signal is reached. Also, a loss of normal AC and EDAS power (Table 9-13, Case HPV-3) is more limiting than all power available and loss of AC power alone with respect to peak CNV pressure, since it results in an earlier ECCS actuation. Earlier ECCS © Copyright 2022 by NuScale Power, LLC 595

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 actuation is more limiting for vapor space breaks, since they result in much less liquid level accumulation within the CNV (compared to the liquid space breaks) to occupy CNV free volume prior to ECCS valve opening. Therefore, for vapor space breaks, peak CNV pressure is driven less by CNV liquid level accumulation occupying CNV free volume prior to ECCS actuation and more by RCS conditions at the time of ECCS actuation. The earlier the ECCS actuation, the higher the energy of the RCS fluid at the time of ECCS valve opening. Low RCS Level Signal - For the HPV break, assuming a low-biased Low RCS Level setpoint (i.e. 540") is less limiting with respect to peak CNV pressure and wall temperature for the all power available scenario, since it results in a much later ECCS actuation (i.e. almost 1 hour later) that does not occur until after the CNV pressure profile has already peaked and steadily decreased. Single Failures - For the HPV break, assuming single failure of an RVV, RRV, or RVV/RRV train results in a slight decrease in peak CNV pressure for the all power available scenario, since it limits the release rate through the RVVs/RRVs upon ECCS actuation. Table 9-13 Spectrum of High Point Vent Line Breaks Evaluated for Peak CNV Temperature and Pressure ((

                                                                                                         }}2(a),(c) 9.7.3.4            Case 4: Inadvertent Reactor Vent Valve Opening Anticipated Operational Occurrence The inadvertent RVV actuation anticipated operational occurrence (AOO) initiates an M&E release from the top of the pressurizer into the CNV. A summary of the spectrum of RVV opening events and results is included in Table 9-14. The CNV pressure response and temperature response are shown in Figure 9-119 and Figure 9-120. The CNV peak pressure is 822 psia for the base case, and 876 psia with the combined effect of the adverse sensitivity parameters (loss of normal AC and DC power). The peak CNV temperature is 519 degrees F for the base case, and 525 degrees F for the case with the combined effect of the adverse sensitivity

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 parameters (loss of normal AC and EDAS power). Case 4 is non-limiting and was confirmed to be non-limiting in comparison to the DL break limiting event. The following observations are made from the RVV event results presented: Electric Power Availability - For the inadvertent RVV opening, a loss of normal AC power at event initiation (Table 9-14, Case RRV- 2) is more limiting than all power available (Table 9-14, Case RRV- 1) with respect to peak CNV pressure and wall temperature, since it results in an immediate turbine trip (i.e. termination of SG heat removal) while the reactor trip is delayed until the High CNV Pressure signal is reached. Also, a loss of normal AC and EDSS power (Table 9-14, Case RRV-3) is more limiting than all power available and loss of AC power alone with respect to peak CNV pressure, since it results in an earlier ECCS actuation. Low RCS Riser Level Signal - For the inadvertent RVV opening, assuming a low-or high-biased low RCS riser level setpoint (i.e. 552" or 540") does not impact the peak CNV pressure and temperature for the all power available scenario. In this case, ECCS actuation occurs on the low differential pressure condition between the RPV and CNV (i.e., less than 15 psid) and the peak values are reached prior to the time of the Low RCS riser level signal. To clarify, the low RCS level setpoint is actually reached first, but the 15 psid condition is reached before the 60-second delay on the low RCS riser level trip is finished. Single Failures - For the inadvertent RVV opening, assuming single failure of an RVV, RRV, or RVV/RRV train results in a slight decrease in peak CNV pressure for the all power available scenario, since it limits the release rate through the RVVs/RRVs upon ECCS actuation. Table 9-14 Spectrum of Reactor Vent Valve Opening Events Evaluated for Peak CNV Temperature and Pressure ((

                                                                                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 9.7.3.5 Case 5: Inadvertent Reactor Recirculation Valve Opening Anticipated Operational Occurrence The inadvertent RRV actuation initiates an M&E release from the downcomer into the CNV. The results of the primary release event M&E release break spectrum analysis and sensitivity analyses are presented in Table 9-15 (Case 5). The following observations are made from the RRV event results presented: Electric Power Availability - For the inadvertent RRV opening, a loss of normal AC power at event initiation (Table 9-15, Case RRV- 2) is more limiting than all power available (Table 9-15, Case RRV- 1) with respect to peak CNV pressure and wall temperature, since it results in an immediate turbine trip (i.e. termination of SG heat removal) while the reactor trip is delayed until the High CNV Pressure signal is reached. Also, a loss of normal AC and EDSS power (Table 9-15, Case RRV-3) is less limiting than all power available with respect to peak CNV pressure and wall temperature, since it results in an earlier ECCS actuation. Low RCS Riser Level Signal - For the inadvertent RRV opening, a low-biased low RCS riser level setpoint (i.e. 540") is more limiting with respect to peak CNV wall temperature for the all power available scenario, since it results in a later ECCS actuation and longer duration of CNV wall heatup prior to RVV and RRV opening and rapid system depressurization. Single Failures - For the inadvertent RRV opening, assuming single failure of an RVV, RRV, or RVV/RRV train results in a slight decrease in peak CNV pressure for the all power available scenario, since it limits the release rate through the RVVs/RRVs upon ECCS actuation. Table 9-15 Spectrum of Reactor Recirculation Valve Opening Events Evaluated for Peak CNV Temperature and Pressure ((

                                                                                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 9.7.3.6 Case 6: Inadvertent Actuation of ECCS Anticipated Operational Occurrence The inadvertent RRV actuation initiates an M&E release from the downcomer into the CNV. The results of the primary release event M&E release break spectrum analysis and sensitivity analyses are presented in Table 9-16. The following observations are made from the inadvertent ECCS event results presented: Electric Power Availability - For the inadvertent ECCS actuation signal event, a loss of normal AC power at event initiation (Table 9-16, Case ECC-2) is more limiting than all power available (Table 9-16, Case ECC-1) with respect to peak CNV pressure and wall temperature, since it results in an immediate turbine trip (i.e. termination of SG heat removal) while the reactor trip is delayed until the High CNV Pressure signal is reached. Also, a loss of normal AC and EDAS power (Table 9-16, Case ECC-3) is more limiting than all power available with respect to peak CNV pressure and temperature, since it results in an earlier ECCS actuation. Low RCS Level Signal - For the inadvertent ECCS actuation signal event, assuming a low- or high biased. Low RCS Level setpoint (i.e. 552" or 540") does not impact the peak CNV pressure and temperature. Single Failures - For the inadvertent ECCS actuation signal event, assuming single failure of an RVV, RRV, or RVV/RRV train results in a slight decrease in peak CNV pressure for the all power available scenario, since it limits the release rate through the RVVs/RRVs upon ECCS actuation. Table 9-16 Spectrum of Inadvertent ECCS Actuation Events Evaluated for Peak CNV Temperature and Pressure ((

                                                                                                       }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-117 Discharge Line Break - RPV Pressure Response ((

                                                                                                }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-118 Discharge Line Break - CNV Pressure Response ((

                                                                                                }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-119 Discharge Line Break - RPV Level above Top of Active Fuel ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-120 Discharge Line Break - CNV Level ((

                                                                                                   }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-121 Discharge Line Break - CNV Wall Temperature Response ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-122 Injection Line Break - RPV Pressure Response ((

                                                                                                  }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-123 Injection Line Break - CNV Pressure Response ((

                                                                                                  }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-124 Injection Line Break - RPV Level above Top of Active Fuel ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-125 Injection Line Break - CNV Level ((

                                                                                                    }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-126 Injection Line Break - CNV Wall Temperature Response ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-127 High Point Vent Line Break - RPV Pressure Response ((

                                                                                                }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-128 High Point Vent Line Break - CNV Pressure Response ((

                                                                                                }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-129 High Point Vent Line Break - RPV Level above Top of Active Fuel ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-130 High Point Vent Line Break - CNV Level ((

                                                                                                   }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-131 High Point Vent Line Break - CNV Wall Temperature Response ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-132 Inadvertent RVV Opening - RPV Pressure Response ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-133 Inadvertent RVV Opening - CNV Pressure Response ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-134 Inadvertent RVV Opening - RPV Level above Top of Active Fuel ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-135 Inadvertent RVV Opening - CNV Level ((

                                                                                                 }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-136 Inadvertent RVV Opening - CNV Wall Temperature Response ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-137 Inadvertent RRV Opening - RPV Pressure Response ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-138 Inadvertent RRV Opening - CNV Pressure Response ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-139 Inadvertent RRV Opening - RPV Level above Top of Active Fuel ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-140 Inadvertent RRV Opening - CNV Level ((

                                                                                                 }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-141 Inadvertent RRV Opening - CNV Wall Temperature Response ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-142 Inadvertent ECCS Actuation Signal - RPV Pressure Response ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-143 Inadvertent ECCS Actuation Signal - CNV Pressure Response ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-144 Inadvertent ECCS Actuation Signal - RPV Level above Top of Active Fuel ((

                                                                                           }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-145 Inadvertent ECCS Actuation Signal - CNV Level ((

                                                                                                }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-146 Inadvertent ECCS Actuation Signal - CNV Wall Temperature Response ((

                                                                                                     }}2(a),(c) 9.7.4         Main Steamline Break Containment Pressure and Temperature Results This section evaluates the main steam line break (MSLB) inside the CNV. The analyses in this section used the same input parameters that are used in the primary system release analysis given in Table 9-8.

Following transient initiation, mass and energy released into the CNV through the break cause CNV pressurization and secondary side depressurization. Depending on the evaluated scenario, the high containment pressure signal or loss of EDAS power causes reactor trip, containment isolation, and DHRS actuation. Reactor trip, containment isolation, and DHRS actuation occur immediately when all electric power is assumed lost. As the transient progresses, secondary side inventory is © Copyright 2022 by NuScale Power, LLC 629

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 continuously lost to the CNV through the break until secondary side isolation valves are fully closed, which isolates the feedwater source from the broken line. The spectrum of MSLB events evaluated, single failures and power loss scenarios considered and CNV temperature and pressure results are provided in Table 9-17. The sequence of events for the limiting secondary system piping break (Case SLB-3) is given in Table 9-18. The following observations are made from the MSLB event results presented: Electric Power Availability - For the MSLB, a loss of normal AC power (Table 9-17, Case SLB-2) is less limiting than all power available (Table 9-17, Case SLB-1) with respect to peak CNV pressure, since it results in an immediate feedwater pump trip (i.e., termination of source of secondary side M&E release). Also, a loss of normal AC and EDAS power (Table 9-17, Case SLB-3) is more limiting than all power available with respect to peak CNV pressure and wall temperature, since it results in an immediate ECCS actuation. Low RPV Riser Level ECCS Signal - For the MSLB, the secondary side M&E release is isolated prior to the RCS experiencing significant level shrink due to the cooldown. ECCS actuation will not occur on the low RPV riser level signal, even with the setpoint biased high (i.e., 552 inches). Therefore, the low RPV riser level setpoint bias does not impact the MSLB peak CNV pressure and wall temperature. Single Failures - For the MSLB, assuming single failure of an FWIV (Table 9-17, Case SLB 9) is more limiting for the all power available scenario (Table 9-17, Case SLB-1) with respect to peak CNV pressure and temperature because failing to isolate the SG allows more mass to be fed and discharged out of the break into the CNV. Failure of an MSIV (Table 9-17, Case SLB-8) has little impact on CNV temperature or pressure. Also, assuming single failure of an RVV, RRV, or RVV/RRV train (Table 9-18, Cases SLB-5, 6 and 7) results in an insignificant change or a decrease in peak CNV pressure and temperature for the loss of AC and EDAS power scenario (Table 9-18, Case MSLB-3), since it limits the release rate through the RVVs/RRVs upon ECCS actuation. The MSLB system responses for Cases SLB-1 through SLB-7 are presented in the following figures. Figure 9-147 shows the RPV pressure response and Figure 9-148 shows the CNV pressure response. SLB-3 results in the peak CNV pressure of 900 psia. This peak pressure is driven primarily due to the loss of EDAS power, which results in ECCS actuation, opening both RVVs immediately and the RRVs when the release pressure is reached. The MSLB peak CNV pressure is less than that of the limiting primary break, DL-2, discussed in Section 9.7.3.1. Figure 9-149 shows the collapsed liquid level, which does not drop below 10 feet above the top of active fuel. Figure 9-150 shows the CNV level above the bottom of the RPV. Figure 9-151 show the CNV wall temperature response. The limiting CNV wall temperature cases are SLB-3 and SLB-6 at just under 530 degrees F. These cases are less than the limiting primary release CNV wall temperature (Case DL-2). Case SLB-3 is the limiting secondary system break © Copyright 2022 by NuScale Power, LLC 630

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 for containment temperature and pressure (bounds all feedwater and steam line break events) but is bounded by the primary side discharge line break (DL-2). Table 9-17 Spectrum of Main Steam Line Break Events Evaluated for Peak CNV Temperature and Pressure ((

                                                                                                    }}2(a),(c)

Table 9-18 Limiting Secondary Break Containment Peak Pressure and Temperature (Case SLB-3) - Sequence of Events ((

                                                                                                    }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-147 Main Steam Line Break - RPV Pressure Response ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-148 Main Steam Line Break - CNV Pressure Response ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-149 Main Steam Line Break - RPV Level above Top of Fuel ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-150 Main Steam Line Break - CNV Level ((

                                                                                                  }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-151 Main Steam Line Break - CNV Wall Temperature ((

                                                                                                     }}2(a),(c) 9.7.5         Feedwater Line Break Containment Pressure and Temperature Results This section evaluates the main feedwater line break (FWLB) inside the CNV. The analyses in this section used the same input parameters that are used in the primary system release analysis given in Table 9-8.

Following transient initiation, mass and energy released into the CNV through the break cause CNV pressurization and secondary side depressurization. Depending on the evaluated scenario, the high containment pressure signal or loss of EDAS power causes reactor trip, containment isolation, and DHRS actuation. Reactor trip, containment isolation, and DHRS actuation occur immediately when all electric power is assumed lost. As the transient progresses, secondary side inventory is © Copyright 2022 by NuScale Power, LLC 636

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 continuously lost to the CNV through the break until secondary side isolation valves are fully closed, which isolates the feedwater source from the broken line. The spectrum of FWLB events evaluated, single failures and power loss scenarios considered and CNV temperature and pressure results are provided in Table 9-19. The following observations are made from the FWLB event results presented: Electric Power Availability - For the FWLB, a loss of normal AC power is more limiting than all power available with respect to peak CNV pressure and wall temperature, since it results in an immediate turbine trip, faster pressurization of the unbroken FW train, correspondingly faster heatup and pressurization of the RCS, and higher enthalpy flow through the broken SG train. Also, a loss of normal AC and EDAS power is more limiting than all power available with respect to peak CNV pressure and wall temperature, since it results in an immediate ECCS actuation. Low RPV Riser Level ECCS Signal - For the FWLB, the secondary side M&E release is not significant enough prior to secondary side isolation for the RCS level to shrink to the Low RCS Level signal, even with the low RPV riser level setpoint biased high (i.e., 552 inches). Therefore, the low RPV riser level setpoint bias does not impact the FWLB peak CNV pressure and wall temperature. Single Failures - For the FWLB, assuming single failure of an MSIV (Table 9-19, Case FLB-8) or FWIV (Table 9-19, Case FLB-9) has very little impact on peak CNV pressure and wall temperature compared to the all power available scenario (Table 9-19, Case FLB-1). MSIV failure allows more high-enthalpy steam to be trapped within the system (up to the secondary MSIV) and discharged out of the break after secondary side isolation. Assuming single failure of an RVV, RRV, or RVV/RRV train (Table 9-19, Cases FLB-5, 6 and 7) results in an insignificant change or a decrease in peak CNV pressure and temperature for the loss of AC and EDAS power scenario (Table 9-19, Case FLB-3), since it limits the release rate through the RVVs/RRVs upon ECCS actuation. The FWLB system responses for Cases FLB-1 through FLB-7 are presented in the following figures. Figure 9-152 shows the RPV pressure response and Figure 9-153 shows the CNV pressure response. Case FLB-3 results in the peak CNV pressure of 886 psia. This peak pressure is driven primarily due to the loss of EDAS power, which results in ECCS actuation, opening both RVVs immediately and the RRVs when the release pressure is reached. The FWLB peak CNV pressure is less than that of the limiting steam line break (Case SLB-3, Section 9.7.4) and the limiting primary break, DL 2, discussed in Section 9.7.3.1. Figure 9-154 shows the collapsed liquid level, which does not drop below 10 feet above the top of active fuel. Figure 9-155 shows the CNV level above the bottom of the RPV. Figure 9-156 shows the CNV wall temperature response. The limiting CNV wall temperature cases are FLB-3 and FLB-6 at 526 degrees F. These cases are less than the limiting MSLB (Case SLB-3) and the primary release CNV wall temperature (Case DL-2). © Copyright 2022 by NuScale Power, LLC 637

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 9-19 Spectrum of Feedwater Line Break Events Evaluated for Peak CNV Temperature and Pressure ((

                                                                                                }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-152 Main Feedwater Line Break - RPV Pressure Response ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-153 Main Feedwater Line Break - CNV Pressure Response ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-154 Main Feedwater Line Break - RPV Level above Top of Fuel ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-155 Main Feedwater Line Break - CNV Level ((

                                                                                                  }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-156 Main Feedwater Line Break - CNV Wall Temperature ((

                                                                                                    }}2(a),(c) 9.8      Inadvertent Opening of Reactor Pressure Vessel Valves (IORV)

The purpose of this section is to describe representative results of the evaluation model and methodology applied to analyze NRC Standard Review Plan (Reference 101) event 15.6.1, Inadvertent Opening of a PWR Pressurizer Pressure Relief Valve, as well as the Inadvertent Operation of the Emergency Core Cooling System (ECCS) event as defined in Section 15.6.6 of the Design-Specific Review Standard (DSRS) (Reference 100) for an NPM. The methodology and EM are presented in the main body of this report, and allow for demonstration via analysis that the acceptance criteria for Anticipated Operational Occurrences (AOOs) listed in DSRS Section 15.0 (Reference 107) are met. © Copyright 2022 by NuScale Power, LLC 643

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Although it is very unlikely due to the mechanical design features, spurious opening of an ECCS valve equipped with an IAB is evaluated. In addition to mechanical failures, for ECCS valves that do not include the IAB, a spurious ECCS actuation or loss of power can also result in a valve opening. Thus, for the NPM-20, the spurious actuations of one or more RVVs and a single RRV are considered separately. The spurious opening of reactor valves is not included in the LOCA break spectrum definition; however the spurious opening of the ECCS valves is considered here with the evaluation model developed for the NPM LOCA. 9.8.1 Phase 0 IORV Analysis The Phase 0 IORV event progression is presented in this section. Representative MCHFR results are summarized in Table 9-20. Representative results for electric power available, loss of AC power, and loss of DC power are respectively plotted in Figure 9-157 through Figure 9-168. The overall Phase 0 transient progression for the RRV is similar to that described for the LOCA liquid region breaks in Section 9.1.1, and the RVV and RSV progression is similar to that described for the LOCA vapor region breaks in Section 9.1.2. While the IORV events are modeled as valves rather than instantaneous breaks, the stroke time is still short (0.1 seconds) with little sensitivity to faster opening rates. The primary difference compared to the LOCA break spectrum is the modeled break area. The RRV venturi area is bounded by the maximum LOCA liquid break. However, the RSV area is larger than the maximum LOCA vapor break, and the RVV venturi area bounds the RSV area. When all electric power is available, the RRV event gives approximately equal MCHFR as the RSV event, despite having only (( }}2(a),(c) percent of the RSV break flow area. As with the LOCA break spectrum, this indicates the downcomer liquid region is the limiting break location. The inadvertent ECCS actuation is the most limiting IORV event, however this is driven by the significantly larger total break area. Similar to the large LOCA breaks, when AC power is lost, the module heat up tends to limit the depressurization which increases MCHFR. The impact is more pronounced for the liquid space RRV opening compared to the vapor space RVV and RSV openings. When loss of DC power is assumed, only one additional RVV opens for the RVV initiating event and MCHFR is less than the RRV and RSV events. Again the difference between the RRV and RSV MCHFR is small, but the RRV is slightly more limiting when loss of DC power is considered due to the larger decrease in RCS flow. The limiting IORV event is the inadvertent ECCS actuation with electric power available. Both RVVs opening without immediate scram is more limiting than the larger break area of the inadvertent RRV or RSV with loss of EDAS power and immediate scram. However, the difference between the limiting MCHFR for the inadvertent ECCS and RRV with loss of EDAS power is small. © Copyright 2022 by NuScale Power, LLC 644

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 9-20 Representative Results of IORV Event Analyses ((

                                                                                                 }}2(a),(c)

Figure 9-157 Reactor Power for IORV Event Spectrum - Electric Power Available ((

                                                                                                 }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-158 Pressurizer Pressure for IORV Event Spectrum - Electric Power Available ((

                                                                                           }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-159 Reactor Coolant System Flow for IORV Event Spectrum - Electric Power Available ((

                                                                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-160 Critical Heat Flux Ratio for IORV Event Spectrum - Electric Power Available ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-161 Reactor Power for IORV Event Spectrum - Loss of AC Power ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-162 Pressurizer Pressure for IORV Event Spectrum - Loss of AC Power ((

                                                                                           }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-163 Reactor Coolant System Flow for IORV Event Spectrum - Loss of AC Power ((

                                                                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-164 Critical Heat Flux Ratio for IORV Event Spectrum - Loss of AC Power ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-165 Reactor Power for IORV Event Spectrum - Loss of EDAS Power ((

                                                                                           }}2(a),(c) 2

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-166 Pressurizer Pressure for IORV Event Spectrum - Loss of EDAS Power ((

                                                                                           }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-167 Reactor Coolant System Flow for IORV Event Spectrum - Loss of EDAS Power ((

                                                                                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-168 Critical Heat Flux Ratio for IORV Event Spectrum - Loss of EDAS Power ((

                                                                                                        }}2(a),(c) 9.8.2         Phase 1 IORV Analysis 9.8.2.1            Phase 1 IORV Analysis with Power Available For IORV events, Phase 0 and Phase 1 initiate at the start of the event. Phase 0 ends ((                                          }}2(a),(c) Phase 1 ends when ECCS recirculation through the RRVs is established and Phase 2 begins. The limiting MCHFR for IORV events occurs within the first two seconds and is calculated using the methodology described in the Phase 0 analysis. The Phase 1 analysis does not calculate MCHFR during the Phase 0 time period. The MCHFR values reported for Phase 1 are calculated for the period between Phase 0 and Phase 2.

IORV events are similar to RCS discharge line and vent line breaks, the reactor trip is initiated by the high containment pressure followed by containment isolation. Representative RPV and CNV pressure transients for the spurious opening of a single RVV and a single RRV is shown in Figure 9-169. The figure clearly demonstrates the pressure response similarities between a single RVV opening and high point vent line break and between a single RRV opening and © Copyright 2022 by NuScale Power, LLC 656

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 RCS discharge line break (see Section 9.1.3 for Phase 1 LOCA figures). The larger flow area associated with a single RVV compared to the 100 percent high point vent line break causes much faster depressurization of the RPV and leads to opening of the remaining ECCS valves at a much smaller differential pressure between RPV and CNV. With a single RRV opening, the remaining ECCS valves open later than a single RVV opening, since it takes longer to reach the low riser level condition. Figure 9-170 compares the collapsed liquid level above TAF between the spurious opening of a single RRV and RVV. The RRV opening causes rapid accumulation of the liquid level inside the CNV and establishing equilibrium levels in both vessels. The final equilibrium level for both cases is approximately 10 feet. The similarities to the liquid and steam space breaks discussed in previous sections can also be seen from the behavior of the collapsed level. The U-shaped increase in collapsed level following ECCS actuation (i.e., 100-300 sec for RVV opening, 300-400 sec for RRV opening) is due to the draining of liquid from the PZR to the upper riser, as shown in Figure 9-170 and Figure 9-171. This is because the calculated collapsed level value shown in Figure 9-170 only spans from TAF to the upper plenum (excludes the PZR). The hot assembly flow at the inlet is shown in Figure 9-172. The flow behavior in both cases shows similar trends with both cases reaching a long-term hot channel flow of 0.1 Mlb/ft2.hr, which is similar to the long term behavior observed in both RCS injection and high point vent line break scenarios. As shown in Figure 9-173, the Phase 1 MCHFR as a function of time in both cases shows acceptable values, and the margin for CHF increases almost exponentially following the reactor trip due to power/flow mismatch throughout the transient. An inadvertent ECCS signal results in both RVVs opening. This event is very similar to the single RVV opening event. Figure 9-174 shown the RPV depressurization and Figure 9-175 shows the CNV pressurization. Figure 9-176 show the collapse liquid level above TAF. Table 9-23 provides a summary of the Phase 1 limiting IORV impacts on MCHFR, collapse liquid level, CNV pressure and CNV temperature. The RRV opening event has a more significant impact on CNV temperature and pressure than the opening of one or two RVVs. The collapsed liquid level above TAF is essentially the same for all three events. The most significant impact of two RVVs opening is on Phase 0 MCHFR discussed in Section 9.8.1. For comparison purposes, Figure 9-177 through Figure 9-185 provide a plot comparison of system pressure and system level for the NPM-160 and NPM-20 designs for the base case inadvertent RVV opening and base case inadvertent RRV opening. RVV and RRV flow capacity is reduced by venturis for the NPM-20, resulting in slightly slower depressurization for the inadvertent RVV and RRV opening events. ECCS actuation on RPV riser level results in an earlier RRV actuation and collapsed level drop for the NPM-20 inadvertent RVV opening case. The U shaped increase in collapsed level observed in the NPM-20 plot is due to © Copyright 2022 by NuScale Power, LLC 657

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 the draining of RCS fluid from the PZR to the riser, since collapsed level is based on the RCS mass balance below the baffle plate. For the base case inadvertent RRV opening, ECCS actuates later for the NPM-20 (low riser level trip) compared to NPM-160 (high CNV level trip). The dip in collapsed level below equilibrium level (i.e., less than10 feet above TAF) at the time of ECCS actuation does not occur for NPM 20, since ECCS actuation occurs at a point when collapsed level is relatively high. The slower decrease in collapsed level for NPM-20 is due to the addition of the RRV venturi, which limits the RCS liquid flow rate through the RRV. 9.8.2.2 Phase 1 IORV Analysis Impact of Loss of Normal AC and EDAS Power For inadvertent RVV or RRV opening events, the loss of normal AC and EDAS power has some impact on the event progressions. The most significant impact is addressed in Section 9.8.1 for Phase 0 MCHFR analysis. There is no significant impact in Phase 1 on MCHFR for the loss of power cases (Figure 9-188 and Figure 9-191). The loss of EDAS power opens the RVVs immediately and causes CNV and RPV pressure to equalize more quickly and causes collapsed liquid level to drop earlier in the event. However, the loss of EDAS does not have a significant impact on the peak CNV pressure (Figure 9-186 and Figure 9-189) or minimum collapsed liquid level (Figure 9-187 and Figure 9-190). Loss of AC power causes a delay in the event progression but no significant impact on peak CNV pressure or minimum collapsed liquid level. Table 9-21 Sequence of Events for Inadvertent RVV Opening for NPM-20 ((

                                                                                                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Table 9-22 Sequence of Events for Inadvertent RRV Opening for NPM-20 ((

                                                                                               }}2(a),(c)

Table 9-23 Summary of Phase 1 Limiting IORV Parameters Valve Opening Collapse Liquid CNV Pressure CNV Temperature MCHFR Level (feet above Event (psia) (ºF) TAF) RVV 2.26 8.6 766 476 RRV 2.01 8.6 851 492 Two RVVs 2.26 8.6 776 476 © Copyright 2022 by NuScale Power, LLC 659

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-169 NPM-20 RPV/CNV Pressure Response for Single ECCS Valve Opening ((

                                                                                        }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-170 NPM-20 Collapsed Liquid Level above TAF for Single ECCS Valve Opening ((

                                                                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-171 NPM-20 Pressurizer Level Response for Single ECCS Valve Opening ((

                                                                                           }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-172 NPM-20 Hot Channel Mass Flux for Single ECCS Valve Opening ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-173 NPM-20 Phase 1 MCHFR for Single ECCS Valve Opening ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-174 NPM-20 Phase 1 Reactor Vessel Pressure for Inadvertent ECCS Signal (Two RVV Opening) ((

                                                                                          }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-175 NPM-20 Phase 1 Containment Pressure for Inadvertent ECCS Signal (Two RVV Opening) ((

                                                                                           }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-176 NPM-20 Phase 1 Collapse Liquid Level over TAF for Inadvertent ECCS Signal (Two RVV Opening) ((

                                                                                                }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-177 NPM-20 Phase 1 MCHFR for Inadvertent ECCS Signal (Two RVV Opening) ((

                                                                                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-178 NPM-160 System Pressure - Inadvertent RVV Opening ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-179 NPM-20 System Pressure - Inadvertent RVV Opening ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-180 NPM-160 System Level - Inadvertent RVV Opening ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-181 NPM-20 System Level - Inadvertent RVV Opening ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-182 NPM-160 System Pressure - Inadvertent RRV Opening ((

                                                                                             }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-183 NPM-20 System Pressure - Inadvertent RRV Opening ((

                                                                                              }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-184 NPM-160 System Level - Inadvertent RRV Opening ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-185 NPM-20 System Level - Inadvertent RRV Opening ((

                                                                                                }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-186 Reactor Pressure Vessel and Containment Pressure for Inadvertent RVV Opening - Loss of Power ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-187 Riser Collapsed Level for Inadvertent RVV Opening - Loss of Power ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-188 Core MCHFR for Inadvertent RVV Opening - Loss of Power ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-189 Reactor Pressure Vessel and Containment Pressure for Inadvertent RRV Opening - Loss of Power ((

                                                                                               }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-190 Riser Collapsed Level for Inadvertent RRV Opening - Loss of Power ((

                                                                                            }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Figure 9-191 Core MCHFR for Inadvertent RRV Opening - Loss of Power ((

                                                                                                   }}2(a),(c) 9.9      Loss-of-Coolant Accident and Inadvertent Opening of Reactor Pressure Vessel Valve Calculation Summary The following conclusions are reached based on the beak spectrum calculations and sensitivity studies:

1. The core Phase 1 MCHFR rapidly increases following the initiation of a LOCA due to power/flow mismatch - power decreases faster than flow due to differences in process time constants.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3

2. ((

                                                                                 }}2(a),(c) In conclusion, there is no fuel CHF and hence no fuel heat-up for a NPM LOCA.

3. The most sensitive LOCA analysis parameters are determined to be DHRS capacity,

(( }}2(a),(c)

4. ((

                                                                            }}2(a),(c)

5. For all the break cases and sizes, the (( }}2(a),(c) collapsed RPV level above TAF converges to approximately (( }}2(a),(c) This value is independent of the LOCA progression. This value is directly related to the ((

                                                                                    }}2(a),(c) and the initial mass/energy inventory.

6. Minimum RPV collapsed level during NPM LOCA transient is invariant of the break size ((

                                      }}2(a),(c)

7. ((

                                           }}2(a),(c)

8. ((

                                                                  }}2(a),(c)

9. ((

                                                                                                              }}2(a),(c)

10. For IORV events, MCHFR occurs within the first two seconds and recovers quickly, within 5 seconds of the minimum for the limiting events.

11. Minimum RPV collapsed liquid level during IORV transient does not drop below the

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 10.0 Conclusions The NPM design is unique when compared to any current operating power plant. It is based on an integral PWR design without coolant loops, coolant pumps, or pressurizer surge lines. The primary reactor system is driven by natural circulation with few connecting pipes and a simple safety-related system to mitigate the consequences of postulated accidents. In particular, the NPM design is not significantly challenged by LOCA events as primary system coolant is captured completely by the CNV, cooled, and returned to the RPV using a large reactor pool as the ultimate heat sink, which can provide cooling for many weeks. The LOCA EM uses a conservative bounding approach to analyzing LOCA transients. It adheres to the relevant requirements of 10 CFR 50 Appendix K and follows the EMDAP described in RG 1.203. Multiple layers of conservatism are incorporated in the LOCA EM to ensure that a conservative analysis result is obtained. These conservatisms stem from application of the relevant modeling requirements of 10 CFR 50 Appendix K and through a series of conservative modeling features that have been incorporated. The methodology uses the proprietary NRELAP5 computer code. RELAP5-3D© was procured and commercial grade dedication was performed as part of the procurement process by NuScale to establish the baseline NRELAP5 code for development. Subsequently, features were added and changes made to NRELAP5 to address the unique aspects of the NPM design and licensing methodology. The models and correlations used in the NRELAP5 code have been reviewed and, where appropriate, modified for use within the NuScale LOCA EM. Features added and changes made to address unique aspects of the NPM design and NuScale LOCA EM that applies to 10 CFR 50 Appendix K include the following: helical coil SG heat transfer and pressure drop models core CHF models wall condensation models critical flow models interfacial drag models for large diameter pipes The NRELAP5 code includes all of the necessary models for characterization of the NPM hydrodynamics, heat transfer between structures and fluids, modeling of fuel, reactor kinetics models, and control systems. The geometry of certain NPM components dictates the use of specific correlations, ((

                                                                    }}2(a),(c) The CHF correlations chosen to assess fuel conditions are selected based on full-scale fuel bundle performance tests over the range of conditions (flows, temperatures, and pressures) anticipated in an NPM during a LOCA event.

A number of conservatisms are built into the NuScale LOCA EM to ensure that conservative analysis results are obtained. Not only are applicable 10 CFR 50 Appendix © Copyright 2022 by NuScale Power, LLC 684

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 K conservatisms present, but additional conservatisms above and beyond 10 CFR 50 Appendix K have been incorporated. Conservatisms that are in addition to the 10 CFR 50 Appendix K requirements include: ((

                                                                  }}2(a),(c)

A PIRT was developed that identified all of the important phenomena that could occur during a LOCA event. Phenomena and process ranking was performed in relation to specified FOMs, as described by RG 1.203. The PIRT also established a knowledge ranking for each of the phenomena identified. Using these FOMs, over 20 phenomena are identified as important to correctly capture in the LOCA EM. Extensive NRELAP5 code validation has been performed to ensure that the LOCA EM is applicable for all important phenomena and processes over the range of conditions encountered in the NPM LOCA. The validation suite includes many legacy SETs and IETs, as well as many SETs and IETs developed and run specifically for the NPM application. The SETs run for the NPM application were performed at the SIET facility on a model helical coil SG and at the Stern facility to obtain CHF data on a full-scale rod bundle test section. Integral effects tests were performed at the NIST facility, a scaled representation of the complete NPM primary and secondary systems, as well as the CPV. The EMDAP requires an applicability demonstration of the NRELAP5 code and tests. A unique aspect of the EMDAP applicability demonstration is the comparison of NRELAP5 simulations of LOCA events to NIST test data and NRELAP5 simulation of the same LOCA event in an NPM. ((

                                                                                        }}2(a),(c) The reasonable-to-excellent agreement obtained by these comparisons establishes both the fidelity of the NIST design to the NPM, and the applicability of NRELAP5 to accurately predict LOCA phenomena at both the NIST and NPM scales. Limitations and modeling requirements are determined in this assessment process and are accounted for in the application of the LOCA methodology.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 This topical report provides an example application of the LOCA EM in order to aid the reader's understanding of the context of the application of the NuScale LOCA EM. These calculations are presented for break spectra that cover a range of break locations, break sizes, single failures, equipment unavailability, and initial and boundary conditions. The nodalization and time-step sensitivity required by 10 CFR 50 Appendix K and additional sensitivity calculations that address the uncertainties in modeling of key phenomena are performed. The analyses conducted demonstrate that an NPM retains sufficient water inventory in the primary system such that the core does not uncover or experience a CHF condition during a LOCA such that the minimum CHF ratio is greater than the ((

                                    }}2(a),(c) as described in Section 6.11.5 and the Phase 1 analysis limit of ((         }}2(a),(c) as described in Section 7.3.6, and that containment design pressure is not challenged. The PCT is shown to occur at the beginning of the LOCA event and cladding temperature decreases as the transient evolves. Because no fuel heatup occurs for any design-basis LOCAs, the following regulatory acceptance criteria from 10 CFR 50.46 are met:

PCT remains below 2,200 degrees Fahrenheit (1,204 degrees Celsius). Maximum fuel oxidation is less than 0.17 times total cladding thickness. Maximum hydrogen generation is less than 0.01 times that generated if all cladding were to react. Coolable geometry is retained. The methodology in this report is also used to support other analyses including:

1. events as described in Topical Report TR-0516-49416-P, "Non-Loss of Coolant Accident Methodology,"

2. containment peak pressure analysis as described in NPM-160 Technical Report TR-0516-49084-P, "Containment Response Analysis Methodology,"

3. long term cooling as described in NPM-160 Technical Report, TR-0919-51299-P, "Long-Term Cooling Methodology," and

4. Topical Report, TR-124587-P, "Extended Passive Cooling and Reactivity Control Methodology" (NPM-20).

This report presents the NuScale containment response analysis methodology for determining primary system and secondary system mass and energy releases and the resultant CNV pressure and temperature response for an NPM. The scope of the methodology is the short-term CNV response for comparison to the CNV pressure and temperature design limits. Equipment qualification and the long-term NPM response are not in the scope of this report. The containment response analysis methodology is shown to meet the intent of Section 6.2 of the NuScale DSRS. Based on the systematic application of conservative initial conditions and boundary conditions in the containment response analysis methodology, the margin in the containment response analysis methodology is judged to be sufficient. © Copyright 2022 by NuScale Power, LLC 686

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 Conservative NRELAP5 demonstration analyses of the containment response analysis methodology have been performed for a spectrum of primary system LOCAs and ECCS valve opening events, and for the MSLB and FWLB accident secondary system events. Sensitivity studies have been used to identify the bounding scenarios and trends. The following insights were obtained: The bounding scenarios for both peak CNV pressure and temperature are determined to be primary system release events. The secondary system break events may include ECCS actuation, which essentially combines an initial secondary system M&E release with a subsequent primary system M&E release, but they are non-limiting scenarios. The limiting M&E release scenario is characterized by an initial heatup and pressurization of the CNV due to the LOCA or ECCS valve opening, and then the subsequent opening of the RVVs on following the pressure differential decreasing below the IAB release pressure. It is the second M&E release that drives the CNV to the peak CNV pressure and peak CNV wall temperature results. The heat capacity of the CNV wall, rather than heat transfer to the reactor pool, provides the short-term heat sink to limit the peak CNV pressure and temperature. The results of full-size guillotine breaks bound those of smaller break sizes. For the limiting cases the results of the sensitivity studies, including postulated single failures, showed only a limited impact on the key figures-of-merit. The loss of normal AC and DC power and the timing of ECCS valve opening are the most important sensitivity parameters. © Copyright 2022 by NuScale Power, LLC 687

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 11.0 References 1 U.S. Nuclear Regulatory Commission, "Transient and Accident Analysis Methods," Regulatory Guide 1.203, Rev. 0, December 2005. 2 U.S. Code of Federal Regulations, "Domestic Licensing of Production and Utilization Facilities," Part 50, Title 10, Appendix K, "ECCS Evaluation Models," (10 CFR 50 Appendix K). 3 U.S. Code of Federal Regulations, "Domestic Licensing of Production and Utilization Facilities", Part 50, Title 10, Section 50.46,"Acceptance Criteria for Emergency Core Cooling System for Light-Water Nuclear Power Reactors," (10 CFR 50.46). 4 NuScale Topical Report, Quality Assurance Program Description," MN-122626 Rev. 0. 5 U.S. Nuclear Regulatory Commission, "Clarification of TMI Action Plan Requirements," NUREG-0737, November 1980. 6 U.S. Nuclear Regulatory Commission, "Design-Specific Review Standard for NuScale SMR Design," Chapter 15, Section 15.6.5, Rev.0, June 2016. 7 U.S. Nuclear Regulatory Commission, "Design-Specific Review Standard for NuScale SMR Design, Section 4.4, Thermal and Hydraulic Design," Rev. 0, June 2016. 8 RELAP5-3D© Code Manual Volume V: "User's Guidelines," INEEL-EXT-98-00834 Revision 4.1, September 2013. 9 SwUM-0304-17023, Revision 10, NRELAP5 Version 1.6 Theory Manual, January 31, 2022. 10 U.S. Code of Federal Regulations, "Domestic Licensing of Production and Utilization Facilities," Part 50, Title 10, Appendix B, "Quality Assurance Criteria for Nuclear Power Plants and Fuel Reprocessing Plants", (10 CFR 50 Appendix B). 11 NuScale Technical Report, "Long Term Cooling Methodology", TR-0916-51299-P, Rev. 3. 12 American Society of Mechanical Engineers, Quality Assurance Program Requirements for Nuclear Facility Applications, ASME NQA-1-2008, NQA-1a-2009 Addenda 13 NuScale Topical Report, Extended Passive Cooling and Reactivity Control Methodology, TR-124587-P, Rev. 0. 14 Taitel, Y., and A.E. Dukler, "A Model of Predicting Flow Regime Transitions in Horizontal and Near Horizontal Gas-Liquid Flow," AIChE Journal,(1976): 22:47-55. © Copyright 2022 by NuScale Power, LLC 688

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 15 Taitel, Y., D. Bornea, and A.E. Dukler, "Modeling Flow Pattern Transitions for Steady Upward Gas-Liquid Flow in Vertical Tubes," AIChE Journal, (1980): 26:345-354. 16 Ishii, M., and G. De Jarlais, "Inverted Annular Flow Modeling," Advanced Code Review Group Meeting, Idaho Falls, ID, July 27, 1982. 17 U.S. Nuclear Regulatory Commission, "Local Drag Laws in Dispersed Two-Phase Flow," NUREG/CR-1230, December 1979. 18 U.S. Nuclear Regulatory Commission, "Study of Two-Fluid Model and Interfacial Area," NUREG/CR-1873, December 1980. 19 Tandon, T.N., H.K. Varma, and C.P. Gupta, "A New Flow Regime Map for Condensation Inside Horizontal Tubes," Journal of Heat Transfer, (1982)": 104:763-768. 20 Lockhart, R.W., and R.C. Martinelli, "Proposed Correlation of Data for Isothermal Two-Phase, Two-Component Flow in Pipes," Chemical Engineering Progress, (1949): 45:39-48. 21 Small Break LOCA Methodology for US-APWR, MUAP-07013-NP, October 2010. 22 Zigrang, D.J., and N.D. Sylvester, "A Review of Explicit Friction Factor Equations," Transactions of the ASME, Journal of Energy Resources Technology, (1985): 107:280-283. 23 Colebrook, C.F., "Turbulent Flow in Pipes with Particular Reference to the Transition Region Between Smooth and Rough Pipe Laws," Journal of Institute of Civil Engineers, (1939): 11:133-156. 24 Chen, J.C., "A Correlation for Boiling Heat Transfer to Saturated Fluids in Convective Flow," Process Design and Development, (1966), 5:322-327. 25 Shah, M.M., "A General Correlation for Heat Transfer during Film Condensation Inside Pipes," International Journal of Heat and Mass Transfer, (1979): 22:547-556. 26 Crank, J., and P. Nicolson, "A Practical Method for Numerical Evaluation of Solutions of Partial Differential Equations of the Heat Conduction Type," Proceedings of the Cambridge Philosophical Society, (1947): 43:50-67. 27 Glasstone, S., and A. Sesonske, Nuclear Reactor Engineering, Von Nostrand Reinhold, New York, NY, 1981. 28 Moody, F.J., "Maximum Flow Rate of a Single Component, Two-Phase Mixture," Transactions of the ASME, Journal of Heat Transfer, vol. 87, No. 1, 1965, pp. 134-142. © Copyright 2022 by NuScale Power, LLC 689

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 29 ((

                                                           }}2(a),(c) 30     Crane Co. "Flow of Fluids Through Valves, Fittings and Pipe", Crane Technical Paper No. 410, 1988.

31 Bankoff, S.G., R.S. Tankin, M.C. Yuen, and C.L. Hsieh, "Countercurrent Flow of Air Water and Steam/Water Through a Horizontal Perforated Plate", International Journal of Heat and Mass Transfer, (1981): Vol. 24, No. 8 pp 1381-1395. 32 Ito, H., "Friction factors for turbulent flow in curved pipes," Transactions of the ASME, Journal of Basic Engineering, (1959): 81:123-124. 33 Sreenivasan K.R., and P.J. Strykowski, "Stabilization Effects in Flow Through Helically Coiled Pipes," Experiments in Fluids 1, 1983, pp. 31-36. 34 Seban, R.A., and E.F. McLaughlin, "Heat transfer in tube coils with laminar and turbulent flow," International Journal of Heat and Mass Transfer, (1963): 6:387-395. 35 Prasad, B.V.S.S.S, D.H. Das, and A.K. Prabhakar, "Pressure drop, heat transfer and performance of a helically coiled tubular exchanger," Heat Recovery Systems and CHP, (1989), 9: 249-256. 36 Dittus, F.W., and L.M.K. Boelter, "Heat transfer in automobile radiators of the tubular type," International Communications in Heat and Mass Transfer, (1985), Vol 12, Issue 1, pp. 3-22. Originally published in University of California Publications in Engineering, Vol. 2, No. 13, October 13, 1930, pp. 443-46. 37 Colburn, A.P., and O.A. Hougen, "Design of Cooler Condensers for Mixtures of Vapors with Noncondensing Gases," Industrial and Engineering Chemistry, (1934): 26:1178-1182. 38 Green, D.W. and Perry, R.H., Perry's Chemical Engineers' Handbook, 8th Edition, McGraw-Hill, New York, NY, 2008. 39 ((

                        }}2(a),(c) 40     Kataoka, I., Ishii, M., "Drift Flux Model for Large Diameter Pipe and New Correlation for Pool Void Fraction", International Journal of Heat and Mass Transfer, (1987):

Vol. 30, No. 9, pp. 1927-1939. 41 Churchill, S.W., and H.H.S. Chu, "Correlating Equations for Laminar and Turbulent Free Convection From a Vertical Plate," International Journal of Heat and Mass Transfer, (1975): 18:1323-1329. © Copyright 2022 by NuScale Power, LLC 690

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 42 Rouhani, S.Z., "Modified Correlations for Void and Pressure Drop," AB Atomenergi, Sweden, Internal Report AE-RTC 841, March 1969. 43 Reserved. 44 Draft American Nuclear Society, "Decay Energy Release Rate Following Shutdown of Uranium-Fueled Thermal Reactors," Proposed Standard ANS 5.1, LaGrange Park, IL, October 1973. 45 Ishii, M., "One-Dimensional Drift-Flux Model and Constitutive Equations for Relative Motion between Phases in Various Two-Phase Flow Regimes", Report No. ANL-77-47, Argonne National Laboratory, October 1977. 46 American Nuclear Society, "Decay Energy Release Rate Following Shutdown of Uranium-Fueled Thermal Reactors," Proposed Standard ANS 5.1, LaGrange Park, IL, October 1971 (revised October 1973). 47 Shure, K., "Fission-Product Decay Energy," WAPD-BT-24, Westinghouse Atomic Division, Bettis, December 1961. 48 American National Standards Institute/American Nuclear Society, "Decay Heat Power in Light Water Reactors," ANSI/ANS-5.1-1979, LaGrange Park, IL. 49 American National Standards Institute/American Nuclear Society, "Decay Heat Power in Light Water Reactors," ANSI/ANS-5.1-1994, LaGrange Park, IL. 50 American National Standards Institute/American Nuclear Society, "Decay Heat Power in Light Water Reactors," ANSI/ANS-5.1-2005, LaGrange Park, IL. 51 Healzer, J.M., J.E. Hench, E. Janssen, and S. Levy, "Design Basis for Critical Heat Flux Condition in Boiling Water Reactors," APED-5186, GE Company Private Report, July 1966. 52 General Electric, "General Electric BWR Thermal Analysis Basis (GETAB): Data, Correlation and Design Application," NEDO-10958-A, 1977. 53 Biasi, L., et.al., "Studies on Burnout Part 3- A New Correlation for Round Ducts and Uniform Heating and Its Comparison with World Data," Energy Nucl., (1967), 14:530-7. 54 Electric Power Research Institute, "Parametric Study of CHF Data," Vol. 2, Palo Alto, California, 1983. 55 ((

                       }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 56 Zuber, N., "Hydrodynamic Aspects of Boiling Heat Transfer," Ph.D. thesis, University of California at Los Angeles, 1959. 57 Groeneveld, D.C., et al., "Lookup Tables for Predicting CHF and Film-Boiling Heat Transfer: Past, Present and Future," Nuclear Technology, (2005): 152:87-104. 58 Ferrell, J.K., and J. W. McGee, "Two-Phase Flow Through Abrupt Expansions and Contractions, Final Report, "Study of Convection Boiling Inside Channels", TID-23394 (Vol. 3), June 1966. 59 U.S. Nuclear Regulatory Commission, "BWR Refill-Reflood Program Task 4.8-Model Qualification Task Plan," NUREG/CR-1899, August 1981. 60 Kim, S.J., "Turbulent film condensation of high pressure steam in a vertical tube of Passive Secondary Condensation System," PhD thesis, Korea Advanced Institute of Science and Technology, 2000. 61 Nylund O., et.al, "Hydrodynamic and heat transfer measurements on a full-scale simulated 36-rod BHWR fuel element with non-uniform axial and radial heat flux distribution", 1970. 62 U.S. Nuclear Regulatory Commission, "Analysis of the FLECHT SEAST Unblocked Bundle Steam Cooling and Boiloff Tests," NUREG/CR-1533, January 1981. 63 U.S. Nuclear Regulatory Commission, "PWR FLECHT SEASET Unblocked Bundle, Forced and Gravity Reflood Task Data Report," NUREG/CR-1532, Volume 2, Appendix C, September 1981. 64 U.S. Nuclear Regulatory Commission, "Results of the Semiscale MOD-2A Natural Circulation Experiments," NUREG/CR-2335, September 1982. 65 U.S. Nuclear Regulatory Commission, Staff Requirements - SECY-19-0036 - Application of the Single Failure Criterion to NuScale Power LLCs Inadvertent Actuation Block Valve., July 2, 2019. 66 Electric Power Research Institute, "Final Report, Two-Phase Jet Modeling and Data Comparison," EPRI NP-4362, March 1986. 67 U.S. Nuclear Regulatory Commission, "Countercurrent Flow of Air/Water and Steam/Water Flow above a Perforated Plate," NUREG/CR-1808, November 1980. 68 U.S. Nuclear Regulatory Commission, "Countercurrent Steam/Water Flow Above a Perforated Plate-Vertical Injection of Water," NUREG/CR-2323, September 1981. 69 Bankoff, S.G., R.S. Tankin, M.C. Yuen, and C.L. Hsieh, "Countercurrent Flow of Air Water and Steam/Water Through a Horizontal Perforated Plate", International Journal of Heat and Mass Transfer, (1981): Vol. 24, No. 8 pp 1381-1395. © Copyright 2022 by NuScale Power, LLC 692

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 70 U.S. Nuclear Regulatory Commission, "The Marviken Full Scale Critical Flow Tests, Summary Report, Joint Reactor Safety Experiments in the Marviken Power Station Sweden," NUREG/CR-2671, May 1982. 71 The Marviken Full Scale Critical Flow Tests, Results from Test 22, Joint Reactor Safety Experiments in the Marviken Power Station Sweden, MXC-222, September 1979. 72 Modro, S.M., et. al., "Multi-Application Small Light Water Reactor Final Report," INEEL/EXT-04-01626, Idaho National Engineering and Environmental Laboratory, December 2003. 73 Idelchik, I.E., "Handbook of Hydraulic Resistance," Hemisphere Publishing, New York, NY, 3rd Edition. 74 Fletcher, C.D., et al., "Adequacy Evaluation of RELAP5/MOD3, Version 3.2.1.2 for Simulating AP600 Small-Break Loss-of-Coolant Accidents," INEL-96/0400, April 1997. 75 NuScale Power, LLC, "Topical Report: Subchannel Analysis Methodology," TR-0915-17564-P-A, Rev. 2, February 2019. 76 The RELAP5-3D Code Development Team, "RELAP5-3D Code Manual, Volume III: Developmental Assessment", INEEL-EXT-98-00834, Revision 4.1, October 2013. 77 Jeandey, C., et al., "Auto Vaporisation D'Ecoulements Eau/Vapeur, Departement des Reacteurs a Eau Service des Transferts Thermiques (Centre D'Etudes Nucleaires de Grenoble)," Report T.T. No. 163, July 1981. 78 U.S. Nuclear Regulatroy Commission, "Assessment of Two-Phase Critical Flow Models Performance in RELAP5 and TRACE against Marviken Critical Flow Tests," NUREG/IA-0401, February 2012. 79 Elias, E., and G. S. Lellouche, "Two-Phase Critical Flow," International Journal of Multiphase Flow, (1994): Vol. 20, No. 91-168. 80 US Nuclear Regulatory Commission, "TRACE V5.0 Theory Manual - Volume 1: Field Equations, Solution Methods, and Physical Models," June 2008, Agencywide Document Access and Management System (ADAMS) Accession No. ML120060218. 81 RELAP5 MOD3.3 Code Manual, Volume IV: Models and Correlations, October 2010. 82 ((

                                         }}2(a),(c)

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 83 Lee, K.W., H.C. No, and C.H. Song, "Onset of Water Accumulation in the Upper Plenum with a Perforated Plate," Nuclear Engineering and Design, (2007): 237:1088-1095. 84 Wallis, G.B., One-dimensional Two-Phase Flow, McGraw-Hill, New York, NY, 1969. 85 Ilic, V., S. Banerjee, and S. Behling, "Qualified Database for the Critical Flow of Water, Final Report", EPRI-NP-4556, May 1986. 86 Zuber, N., and J. A. Findlay, "Average Volumetric Concentrations in Two-Phase Flow Systems", Transactions of the ASME, Journal of Heat Transfer, (1965): 87:453-568. 87 The RELAP5-3D Code Development Team, "RELAP5-3D Code Manual, Volume III: Developmental Assessment", INEEL-EXT-98-00834, Revision 4.1, October 2013. 88 McAdams, W.H., Heat Transmission, 3rd Edition, McGraw-Hill, New York, NY, 1954. 89 Minkowycz, W.J., and E. M., Sparrow, "Local Nonsimilar Solutions for Natural Convection on a Vertical Cylinder," Journal of Heat Transfer, (1974): 96(2), 178-183. 90 Plesset, M.S., and S. A. Zwick, "The Growth of Vapor Bubbles in Superheated Liquids", Journal of Applied Physics, (1954): 25, 493. 91 Lee, K., and D. J. Ryley, "The Evaporation of Water Droplets in Superheated Steam", Transactions of the ASME, Journal of Heat Transfer, (1968): pp. 445-451. 92 Aumiller, D.L., "The Effect of Nodalization on the Accuracy of the Finite-Difference Solution of the Transient Conduction Equation", 2000 RELAP5 International Users Seminar, Jackson Hole, Wyoming, September 12-14, 2000. 93 Electric Power Research Institute, "The Chexal-Lellouche Void Fraction Correlation for Generalized Applications," NSAC-139, April 1991. 94 Inayatov, A.Y., "Correlation of Data on Heat Transfer Flow Parallel to Tube Bundles at Relative Pitches of 1.1 < s/d < 1.6," Heat Transfer-Soviet Research, (1975): 7, 3, pp. 84-88. 95 ((

                                    }}2(a),(c) 96     Van den Eynde, G., "Comments on "A Resolution of the Stiffness Problem of Reactor Kinetics"," Nuclear Science and Engineering, (2006): 153:200-202.

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 97 Saha, P., and N. Zuber, "Point of Net Vapor Generation and Vapor Void Fraction in Subcooled Boiling," Proceedings Fifth International Heat Transfer Conference, September 3-7, 1974, Tokyo, Japan: 4:175-179. 98 Lahey, R.T., "A Mechanistic Subcooled Boiling Model," Proceedings Sixth International Heat Transfer Conference, August 7-11, 1978, Toronto, Canada: 1:293

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99 U.S. Nuclear Regulatory Commission, "An Integrated Structure and Scaling Methodology for Severe Accident Technical Issue Resolution," NUREG/CR-5809, Appendix D, November 1991. 100 U.S. Nuclear Regulatory Commission, "Design-Specific Review Standard for NuScale SMR Design," Chapter 15, Section 15.6.6, Rev.0, June 2016. 101 U.S. Nuclear Regulatory Commission, NUREG-0800, Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants: LWR Edition, Section 15.6.1, Revision 2, March 2007. 102 U.S. Nuclear Regulatory Commission, NUREG-0800, Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants: LWR Edition, Section 15.0, Revision 2, March 2007. 103 American Nuclear Society, "Nuclear Safety Criteria for the Design of Stationary Pressurized Water Reactor Plants," ANSI N18.2-1973. 104 U.S. Nuclear Regulatory Commission, NUREG-0800, Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants: LWR Edition, Section 5.2.3, Revision 2, March 2007. 105 GE Nuclear Energy, "ABWR Design Control Document," Revision 4, March 1997. 106 U.S. Nuclear Regulatory Commission, "Applying Statistics," NUREG-1475, Rev. 1, March 2011. 107 U.S. Nuclear Regulatory Commission, "Design-Specific Review Standard for NuScale SMR Design," Chapter 15, Section 15.0, Rev. 0. June 2016. 108 U.S. Nuclear Regulatory Commission, NUREG-1503, "Final Safety Evaluation Report Related to the Certification of the Advanced Boiling Water Reactor Design, Main Report," July 1994 109 U.S. Code of Federal Regulations, 10CFR 50, Appendix A, "General Design Criteria for Nuclear Power Plants." 110 U.S. Nuclear Regulatory Commission, "Design-Specific Review Standard for NuScale SMR Design," Chapter 6, Section 6.2.1, Rev. 0, June 2016. © Copyright 2022 by NuScale Power, LLC 695

Loss-of-Coolant Accident Evaluation Model TR-0516-49422-NP Revision 3 111 U.S. Nuclear Regulatory Commission, "Design-Specific Review Standard for NuScale SMR Design," Chapter 6, Section 6.2.1.1.A, Rev. 0, June 2016. 112 U.S. Nuclear Regulatory Commission, "Design-Specific Review Standard for NuScale SMR Design," Chapter 6, Section 6.2.1.3, Rev. 0, June 2016. 113 U.S. Nuclear Regulatory Commission, "Design-Specific Review Standard for NuScale SMR Design," Chapter 6, Section 6.2.1.4, Rev. 0, June 2016. 114 NuScale Power, LLC, Topical Report "Non-LOCA Transient Analysis Methodology," TR 0516-49416-P, Rev. 3. 115 NuScale Power, LLC, Topical Report, "NuScale Power Critical Heat Flux Correlations," TR-0116-21012-P-A, Revision 1. 116 Tong, L.S., Currin H.B., et. al., "Influence of axially non-uniform heat flux on DNB," AIChE Chemical Engineering Progress Symposium Series, Vol. 62, No. 64. 117 U.S. Nuclear Regulatory Commission, "Applying Statistics," NUREG-1475, Rev. 1. © Copyright 2022 by NuScale Power, LLC 696

LO-133399 : Affidavit of Mark W. Shaver, AF-133400 NuScale Power, LLC 1100 NE Circle Blvd., Suite 200 Corvallis, Oregon 97330 Office 541.360.0500 Fax 541.207.3928 www.nuscalepower.com

NuScale Power, LLC AFFIDAVIT of Mark W. Shaver I, Mark W. Shaver, state as follows: (1) I am the Licensing Manager of NuScale Power, LLC (NuScale), and as such, I have been specifically delegated the function of reviewing the information described in this Affidavit that NuScale seeks to have withheld from public disclosure, and am authorized to apply for its withholding on behalf of NuScale (2) I am knowledgeable of the criteria and procedures used by NuScale in designating information as a trade secret, privileged, or as confidential commercial or financial information. This request to withhold information from public disclosure is driven by one or more of the following: (a) The information requested to be withheld reveals distinguishing aspects of a process (or component, structure, tool, method, etc.) whose use by NuScale competitors, without a license from NuScale, would constitute a competitive economic disadvantage to NuScale. (b) The information requested to be withheld consists of supporting data, including test data, relative to a process (or component, structure, tool, method, etc.), and the application of the data secures a competitive economic advantage, as described more fully in paragraph 3 of this Affidavit. (c) Use by a competitor of the information requested to be withheld would reduce the competitors expenditure of resources, or improve its competitive position, in the design, manufacture, shipment, installation, assurance of quality, or licensing of a similar product. (d) The information requested to be withheld reveals cost or price information, production capabilities, budget levels, or commercial strategies of NuScale. (e) The information requested to be withheld consists of patentable ideas. (3) Public disclosure of the information sought to be withheld is likely to cause substantial harm to NuScales competitive position and foreclose or reduce the availability of profit-making opportunities. The accompanying report reveals distinguishing aspects about the method by which NuScale develops its Loss-of-Coolant Accident Evaluation Model. NuScale has performed significant research and evaluation to develop a basis for this method and has invested significant resources, including the expenditure of a considerable sum of money. The precise financial value of the information is difficult to quantify, but it is a key element of the design basis for a NuScale plant and, therefore, has substantial value to NuScale. If the information were disclosed to the public, NuScale's competitors would have access to the information without purchasing the right to use it or having been required to undertake a similar expenditure of resources. Such disclosure would constitute a misappropriation of NuScale's intellectual property, and would deprive NuScale of the opportunity to exercise its competitive advantage to seek an adequate return on its investment. (4) The information sought to be withheld is in the enclosed report entitled Loss-of-Coolant Accident Evaluation Model. The enclosure contains the designation Proprietary" at the top of each page containing proprietary information. The information considered by NuScale to be proprietary is identified within double braces, "(( }}" in the document. (5) The basis for proposing that the information be withheld is that NuScale treats the information as a trade secret, privileged, or as confidential commercial or financial information. NuScale relies upon the exemption from disclosure set forth in the Freedom of Information Act ("FOIA"), 5 USC § AF-133400 Page 1 of 2

552(b)(4), as well as exemptions applicable to the NRC under 10 CFR §§ 2.390(a)(4) and 9.17(a)(4). (6) Pursuant to the provisions set forth in 10 CFR § 2.390(b)(4), the following is provided for consideration by the Commission in determining whether the information sought to be withheld from public disclosure should be withheld: (a) The information sought to be withheld is owned and has been held in confidence by NuScale. (b) The information is of a sort customarily held in confidence by NuScale and, to the best of my knowledge and belief, consistently has been held in confidence by NuScale. The procedure for approval of external release of such information typically requires review by the staff manager, project manager, chief technology officer or other equivalent authority, or the manager of the cognizant marketing function (or his delegate), for technical content, competitive effect, and determination of the accuracy of the proprietary designation. Disclosures outside NuScale are limited to regulatory bodies, customers and potential customers and their agents, suppliers, licensees, and others with a legitimate need for the information, and then only in accordance with appropriate regulatory provisions or contractual agreements to maintain confidentiality. (c) The information is being transmitted to and received by the NRC in confidence. (d) No public disclosure of the information has been made, and it is not available in public sources. All disclosures to third parties, including any required transmittals to NRC, have been made, or must be made, pursuant to regulatory provisions or contractual agreements that provide for maintenance of the information in confidence. (e) Public disclosure of the information is likely to cause substantial harm to the competitive position of NuScale, taking into account the value of the information to NuScale, the amount of effort and money expended by NuScale in developing the information, and the difficulty others would have in acquiring or duplicating the information. The information sought to be withheld is part of NuScale's technology that provides NuScale with a competitive advantage over other firms in the industry. NuScale has invested significant human and financial capital in developing this technology and NuScale believes it would be difficult for others to duplicate the technology without access to the information sought to be withheld. I declare under penalty of perjury that the foregoing is true and correct. Executed on 01/08/23. Mark W. Shaver AF-133400 Page 2 of 2}}

v t e Letter Sequence Request
EPID:L-2022-TOP-0016, LLC, Submittal of Topical Report Loss-of-Coolant Accident Evaluation Model, TR-0516-49422, Revision 3 (Open)
Initiation Request Acceptance... Administration Questions Answers Results Approval... Other: ML23068A105, ML23068A106, ML23205A008, ML23265A145
MONTHYEAR2023-03-17 [Table View]

ML23008A002

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SUBJECT:

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