State-of-the-Art Reactor 

Consequence Analyses 

Project

Volume 2: 

Surry Integrated Analysis

Office of Nuclear Regulatory Research

NUREG/CR-7110, Vol. 2, 

Rev. 1
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State-of-the-Art Reactor 

Consequence Analyses 

Project

Volume 2: 

Surry Integrated Analysis

Manuscript Completed: May 2012 

Date Published: August 2013

Prepared by:

Sandia National Laboratories

Albuquerque, New Mexico 87185

Operated for the U.S. Department of Energy

NRC Job Code N6306

Office of Nuclear Regulatory Research

NUREG/CR-7110, Vol. 2, 

Rev. 1
Sandia National Laboratories is a multi-program laboratory managed and 

operated by Sandia Corporation, a wholly owned subsidiary of Lockheed 

Martin Corporation, for the U.S. Department of Energy’s National Nuclear 

Security Administration under contract DE-AC04-94AL85000.
iii

ABSTRACT

The evaluation of accident phenomena and the offsite consequences of severe reactor accidents 

has been the subject of considerable research by the U.S. Nuclear Regulatory Commission 

(NRC) over the last several decades. As a consequence of this research focus, analyses of severe 

accidents at nuclear power reactors are more detailed, integrated, and realistic than at any time in 

the past. A desire to leverage this capability to address conservative aspects of previous reactor 

accident analysis efforts was a major motivating factor in the genesis of the State-of-the-Art 

Reactor Consequence Analysis (SOARCA) project. By applying modern analysis tools and 

techniques, the SOARCA project developed a body of knowledge regarding the realistic

outcomes of severe nuclear reactor accidents. To accomplish this objective, the SOARCA 

project used integrated modeling of accident progression and offsite consequences using both 

state-of-the-art computational analysis tools and best modeling practices drawn from the

collective wisdom of the severe accident analysis community. This study has focused on 

providing a realistic evaluation of accident progression, source term, and offsite consequences 

for the Surry Nuclear Power Station. By using the most current emergency preparedness 

practices, plant capabilities, and best available modeling, these analyses are more detailed, 

integrated, and realistic than past analyses. These analyses also consider all mitigative measures, 

contributing to a more realistic evaluation.

PAPERWORK REDUCTION ACT STATEMENT

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to the requirements of the Paperwork Reduction Act of 1995 (44 USC 3501, et seq.).

PUBLIC PROTECTION NOTIFICATION

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information or an information collection requirement unless the requesting document displays a 

currently valid OMB control number.

v

TABLE OF CONTENTS

ABSTRACT.................................................................................................................................. iii

TABLE OF CONTENTS ..............................................................................................................v

LIST OF FIGURES ..................................................................................................................... ix

LIST OF TABLES.....................................................................................................................xvii

EXECUTIVE SUMMARY ....................................................................................................... xxi

ACKNOWLEDGEMENTS ......................................................................................................xxxi

ACRONYMS..............................................................................................................................xxxiii

1. INTRODUCTION................................................................................................................ 1-1

1.1 Background ................................................................................................................. 1-1

1.2 Objective ..................................................................................................................... 1-2

1.3 Outline of the Report................................................................................................... 1-3

2. ACCIDENT SCENARIO DEVELOPMENT.................................................................... 2-1

2.1 Sequences Initiated by Internal Events....................................................................... 2-1

2.2 Sequences Initiated by External Events...................................................................... 2-3

2.3 Mitigative Measures.................................................................................................... 2-3

2.3.1 Mitigation of Sequences Initiated by Internal Events ........................................... 2-4

2.3.2 Mitigation of Sequences Initiated by External Events .......................................... 2-4

3. ACCIDENT SCENARIO DEFINITIONS......................................................................... 3-1

3.1 Long-Term Station Blackout ...................................................................................... 3-1

3.1.1 Initiating Event ...................................................................................................... 3-1

3.1.2 System Availabilities............................................................................................. 3-2

3.1.3 Operator Actions and Mitigative Measures........................................................... 3-2

3.1.4 Scenario Boundary Conditions.............................................................................. 3-3

3.2 Short-Term Station Blackout ...................................................................................... 3-8

3.2.1 Initiating Event ...................................................................................................... 3-8

3.2.2 System Availabilities............................................................................................. 3-9

3.2.3 Mitigative Actions................................................................................................. 3-9

3.2.4 Scenario Boundary Conditions.............................................................................. 3-9

3.3 Spontaneous SGTR................................................................................................... 3-14

3.3.1 Initiating Event .................................................................................................... 3-14

3.3.2 System Availabilities........................................................................................... 3-14

3.3.3 Mitigative Actions............................................................................................... 3-15

3.3.4 Boundary Conditions........................................................................................... 3-16

3.4 Interfacing Systems LOCA....................................................................................... 3-19

3.4.1 Initiating Event .................................................................................................... 3-19

3.4.2 System Availabilities........................................................................................... 3-19

3.4.3 Mitigative Actions............................................................................................... 3-19

3.4.4 Boundary Conditions........................................................................................... 3-21

3.5 Surry Seismic PRA Study ......................................................................................... 3-23
vi

4. MELCOR MODEL OF THE SURRY PLANT................................................................ 4-1

4.1 Vessel and Reactor Coolant System ........................................................................... 4-4

4.2 Primary and Secondary System Relief Valve Modeling........................................... 4-12

4.2.1 Primary System Relief Valves............................................................................. 4-12

4.2.2 Secondary System Relief Valves......................................................................... 4-12

4.2.3 PWR versus BWR Valve Failure Modeling........................................................ 4-13

4.3 Decay Heat Power Modeling .................................................................................... 4-14

4.4 Natural Circulation Modeling ................................................................................... 4-15

4.5 Core Degradation Modeling...................................................................................... 4-18

4.6 Containment.............................................................................................................. 4-20

4.7 Containment Leakage Model .................................................................................... 4-23

4.8 Auxiliary Building .................................................................................................... 4-26

4.9 Best Modeling Practices............................................................................................ 4-27

4.9.1 Approach to Modeling Important Phenomena .................................................... 4-28

4.9.2 Early Containment Failure Phenomena............................................................... 4-30

4.10 Safeguards Area, Containment Spray Pump Area, and MSV House ....................... 4-31

4.10.1 Building Interconnectivity................................................................................... 4-35

4.10.2 Existing Potential Fission Product Release Pathways......................................... 4-36

4.10.3 Potential Building Boundary Over-Pressure Failures ......................................... 4-37

4.11 Safeguards Ventilation System ................................................................................. 4-38

4.12 Low Head Safety Injection Piping ............................................................................ 4-40

4.13 Radionuclide Deposition in LHSI Piping by Turbulent Deposition and 

Impaction Deposition in the LHSI Piping................................................................. 4-45

4.14 Analysis Methodology Involving Two MELCOR Models for LHSI Piping............ 4-46

5. INTEGRATED THERMAL HYDRAULICS, ACCIDENT PROGRESSION, AND 

RADIOLOGICAL RELEASE ANALYSIS ...................................................................... 5-1

5.1 Long-Term Station Blackout ...................................................................................... 5-1

5.1.1 Unmitigated Long-Term Station Blackout............................................................ 5-1

5.1.2 Mitigated Long-Term Station Blackout .............................................................. 5-12

5.1.3 Unmitigated Long-Term Station Blackout with Early RCP Seal Failure ........... 5-18

5.1.4 Mitigated Long-Term Station Blackout with Early RCP Seal Failure................ 5-26

5.2 Short-Term Station Blackout .................................................................................... 5-31

5.2.1 Unmitigated Short-Term Station Blackout.......................................................... 5-31

5.2.2 Mitigated Short-Term Station Blackout .............................................................. 5-42

5.2.3 Uncertainties in the Hydrogen Combustion in the Mitigated Short-Term 

Station Blackout .................................................................................................. 5-51

5.3 Short-Term Station Blackout with Thermally-Induced SGTR................................. 5-63

5.3.1 Unmitigated Short-Term Station Blackout with Thermally-Induced 

Steam Generator Tube Rupture ........................................................................... 5-65

5.3.2 Mitigated Short-Term Station Blackout with Thermally Induced Steam 

Generator Tube Rupture ...................................................................................... 5-80

5.3.3 Uncertainties in the Failure of the Thermally-Induced Steam Generator 

Tube versus the Hot Leg...................................................................................... 5-87

5.4 Spontaneous SGTR................................................................................................... 5-90

5.4.1 Mitigated Spontaneous Steam Generator Tube Rupture ..................................... 5-91

5.4.2 Unmitigated - Spontaneous SGTR with Failed Operator Action........................ 5-98
vii

5.4.3 Unmitigated - Spontaneous SGTR with Failed Operator Action and 

Faulted Steam Generator SORV....................................................................... 5-104

5.5 Interfacing System Loss of Coolant Accident ........................................................ 5-110

5.5.1 Unmitigated ISLOCA........................................................................................ 5-110

5.5.2 ISLOCA Separate-effects Calculation............................................................... 5-131

5.5.3 Mitigation of the ISLOCA................................................................................. 5-139

5.6 Other Sensitivity Studies......................................................................................... 5-158

5.6.1 Chemical Form of Iodine................................................................................... 5-158

5.6.2 Additional Source Term from Iodine Spiking................................................... 5-161

5.6.3 Air Ingression into the Vessel ........................................................................... 5-161

5.6.4 Aerosol Settling Rate in the Containment ......................................................... 5-163

6. EMERGENCY RESPONSE............................................................................................... 6-1

6.1 Population Attributes .................................................................................................. 6-3

6.1.1 Population Distribution.......................................................................................... 6-4

6.1.2 Evacuation Time Estimates................................................................................... 6-5

6.2 WinMACCS................................................................................................................ 6-7

6.2.1 Hotspot and Normal Relocation and Habitability ................................................. 6-7

6.2.2 Shielding Factors................................................................................................... 6-7

6.2.3 Potassium Iodide.................................................................................................... 6-8

6.2.4 Adverse Weather ................................................................................................... 6-8

6.2.5 Modeling using Evacuation Time Estimates......................................................... 6-9

6.2.6 Establishing the Initial Cohort in the Calculation ............................................... 6-10

6.3 Accident Scenarios.................................................................................................... 6-11

6.3.1 Unmitigated LTSBO............................................................................................ 6-11

6.3.2 Unmitigated STSBO............................................................................................ 6-15

6.3.3 Unmitigated STSBO with TI-SGTR ................................................................... 6-19

6.3.4 Mitigated STSBO with TI-SGTR........................................................................ 6-22

6.3.5 Unmitigated ISLOCA.......................................................................................... 6-25

6.4 Sensitivity Studies..................................................................................................... 6-28

6.4.1 Sensitivity 1 ISLOCA.......................................................................................... 6-31

6.4.2 Sensitivity 2 ISLOCA.......................................................................................... 6-33

6.4.3 Sensitivity 3 ISLOCA Delay of Implementation of Protective Actions ............. 6-36

6.5 Analysis of Earthquake Impact ................................................................................. 6-38

6.5.1 Soils Review........................................................................................................ 6-38

6.5.2 Infrastructure Analysis ........................................................................................ 6-39

6.5.3 Electrical and Communications........................................................................... 6-42

6.5.4 Emergency Response........................................................................................... 6-42

6.5.5 Development of WinMACCS parameters........................................................... 6-44

6.5.6 Seismic Analysis STSBO with TI-SGTR............................................................ 6-45

6.6 Accident Response and Mitigation of Source Terms................................................ 6-49

6.6.1 External Resources.............................................................................................. 6-52

6.6.2 Mitigation Strategies............................................................................................ 6-54

6.6.3 Truncation Summary ........................................................................................... 6-59

6.7 Emergency Preparedness Summary and Conclusions.............................................. 6-59
viii

7. OFF-SITE CONSEQUENCES........................................................................................... 7-1

7.1 Introduction................................................................................................................. 7-1

7.2 Surry Source Terms..................................................................................................... 7-1

7.3 Consequence Analyses................................................................................................ 7-3

7.3.1 Unmitigated Long-Term Station Blackout............................................................ 7-3

7.3.2 Unmitigated Short-Term Station Blackout............................................................ 7-7

7.3.3 Unmitigated Short-Term Station Blackout with TI-SGTR ................................... 7-9

7.3.4 Mitigated Short-Term Station Blackout with TI-SGTR Rupture........................ 7-12

7.3.5 Unmitigated ISLOCA.......................................................................................... 7-15

7.3.6 Sensitivity Analyses on Size of the Evacuation Zone and Start Time ................ 7-19

7.3.7 Evaluation of the Effect of the Seismic Activity on Emergency Response ........ 7-20

7.3.8 Evaluation of SST1 Source Term........................................................................ 7-20

7.3.9 Surface Roughness .............................................................................................. 7-24

7.3.10 Importance of Chemical Classes ......................................................................... 7-26

8. REFERENCES..................................................................................................................... 8-1

APPENDIX A SURRY CONTAINMENT PERFORMANCE.............................................A-1

APPENDIX B SURRY RADIONUCLIDE INVENTORY ...................................................B-1

APPENDIX C INPUT PARAMETERS FOR CONSEQUENCE ANALYSIS...................C-1

APPENDIX D ISLOCA MODELING DETAILS .................................................................D-1
ix

LIST OF FIGURES

Figure 4-1 Surry Reactor Coolant System Hydrodynamic Nodalization .............................. 4-7

Figure 4-2 Surry Reactor Vessel Core, Lower Plenum, and Upper Plenum and 

Steam Dome Hydrodynamic and COR Structure Nodalization .......................... 4-8

Figure 4-3 Surry Reactor Core Radial Power Profile and Nodalization ............................... 4-9

Figure 4-4 Surry Steam Generator A Hydrodynamic Nodalization .................................... 4-10

Figure 4-5 Surry Steam Generator B and C Hydrodynamic Nodalization.......................... 4-11

Figure 4-6 Natural Circulation Flow Patterns in a PWR..................................................... 4-18

Figure 4-7 Depiction of the Fuel Rod Degradation............................................................. 4-20

Figure 4-8 Containment Hydrodynamic Nodalization ........................................................ 4-22

Figure 4-9 Containment Hydrodynamic Nodalization, Plan View ..................................... 4-23

Figure 4-10 Nominal Containment Leakage Model.............................................................. 4-25

Figure 4-11 Containment Failure Leakage Model................................................................. 4-25

Figure 4-12 Auxiliary Building Hydrodynamic Nodalization............................................... 4-27

Figure 4-13 Safeguards Buildings MELCOR Nodalization.................................................. 4-33

Figure 4-14 Safeguards Buildings Ventilation MELCOR Nodalization............................... 4-34

Figure 4-15 Low Head Safety Injection Piping MELCOR Nodalization.............................. 4-43

Figure 4-16 Low Head Safety Injection Pipe Sections in Safeguards Area .......................... 4-44

Figure 4-17 Cold Leg Pressure and Temperature Imposed on ISLOCA

Deposition Calculation....................................................................................... 4-48

Figure 4-18 Cold Leg Partial Pressure Imposed on ISLOCA Deposition Calculation ......... 4-49

Figure 4-19 Safeguards Area Pool Level and Temperature Imposed on ISLOCA 

Deposition Calculation....................................................................................... 4-50

Figure 4-20 Radioactive Aerosol Sourcing into ISLOCA Deposition Calculation............... 4-51

Figure 4-21 Masses of Radionuclides Sourced into the ISLOCA 

Deposition Calculation by Size.......................................................................... 4-51

Figure 5-1 Unmitigated LTSBO primary and secondary pressure history............................ 5-6

Figure 5-2 Unmitigated LTSBO vessel two-phase coolant level history.............................. 5-6

Figure 5-3 Unmitigated LTSBO core temperature history.................................................... 5-7

Figure 5-4 Unmitigated LTSBO lower head inner and outer temperature history................ 5-7

Figure 5-5 Unmitigated long-term station blackout containment pressure history ............... 5-8

Figure 5-6 Unmitigated long-term station blackout containment gas concentrations........... 5-8

Figure 5-7 Unmitigated LTSBO iodine fission product distribution history ...................... 5-10

Figure 5-8 Unmitigated LTSBO cesium fission product distribution history..................... 5-11

Figure 5-9 Unmitigated LTSBO environmental release history of fission products........... 5-11

Figure 5-10 Mitigated long-term station blackout primary and secondary

pressure history .................................................................................................. 5-15

Figure 5-11 Mitigated long-term station blackout vessel two-phase coolant level............... 5-15

Figure 5-12 Mitigated long-term station blackout core temperature history......................... 5-16

Figure 5-13 Mitigated long-term station blackout lower head inner and outer 

temperature history ............................................................................................ 5-16

Figure 5-14 Mitigated long-term station blackout containment pressure history.................. 5-17

Figure 5-15 Mitigated LTSBO vessel emergency make up and pump seal leakage 

flows................................................................................................................... 5-17
x

Figure 5-16 Unmitigated long-term station blackout primary and secondary 

pressure history .................................................................................................. 5-22

Figure 5-17 Comparison of the unmitigated long-term station blackout RCP seal 

leakages with and without early RCP seal failures............................................ 5-22

Figure 5-18 Comparison of the unmitigated long-term station blackout vessel level 

responses with and without early RCP seal failures.......................................... 5-23

Figure 5-19 Comparison of the unmitigated long-term station blackout containment 

pressure responses with and without early RCP seal failures............................ 5-23

Figure 5-20 Comparison of the unmitigated long-term station blackout containment 

water pool masses with and without early RCP seal failures ............................ 5-24

Figure 5-21 Comparison of the unmitigated long-term station blackout noble gas 

releases to the environment with and without early RCP seal failures.............. 5-25

Figure 5-22 Comparison of the unmitigated long-term station blackout iodine 

releases to the environment with and without early RCP seal failures.............. 5-25

Figure 5-23 Comparison of the unmitigated long-term station blackout cesium 

releases to the environment with and without early RCP seal failures.............. 5-26

Figure 5-24 Primary and secondary pressure responses for the mitigated long-term 

station blackout with early RCP seal failure...................................................... 5-29

Figure 5-25 Comparison of the mitigated long-term station blackout RCP seal 

leakages with and without early RCP seal failures............................................ 5-30

Figure 5-26 Comparison of the mitigated long-term station blackout vessel level 

with and without early RCP seal failures........................................................... 5-30

Figure 5-27 Comparison of the mitigated long-term station blackout portable pump 

injection rates with and without early RCP seal failures................................... 5-31

Figure 5-28 Unmitigated STSBO primary and secondary pressures history ........................ 5-35

Figure 5-29 Unmitigated short-term station blackout vessel two-phase coolant level.......... 5-35

Figure 5-30 Unmitigated STSBO core temperature history.................................................. 5-36

Figure 5-31 Unmitigated STSBO containment leakage area ................................................ 5-36

Figure 5-32 Unmitigated STSBO lower head inner and outer temperature history.............. 5-37

Figure 5-33 Unmitigated short-term station blackout containment pressure history ............ 5-37

Figure 5-34 Unmitigated short-term station blackout containment leakage area.................. 5-38

Figure 5-35 Unmitigated short-term station blackout containment gas concentrations........ 5-38

Figure 5-36 Unmitigated STSBO iodine fission product distribution history....................... 5-40

Figure 5-37 Unmitigated STSBO cesium fission product distribution history ..................... 5-41

Figure 5-38 Unmitigated STSBO environmental release history of fission products........... 5-41

Figure 5-39 Mitigated STSBO primary and secondary pressure history .............................. 5-45

Figure 5-40 Mitigated short-term station blackout vessel two-phase coolant level .............. 5-46

Figure 5-41 Mitigated short-term station blackout core temperature history........................ 5-46

Figure 5-42 Mitigated STSBO lower head inner and outer temperature history .................. 5-47

Figure 5-43 Mitigated short-term station blackout containment pressure history................. 5-47

Figure 5-44 Mitigated short-term station blackout containment gas concentrations ............ 5-48

Figure 5-45 Mitigated and unmitigated short-term station blackout containment gas 

concentrations.................................................................................................... 5-48

Figure 5-46 Mitigated short-term station blackout iodine fission product 

distribution history............................................................................................. 5-50

Figure 5-47 Mitigated STSBO cesium fission product distribution history.......................... 5-50
xi

Figure 5-48 Mitigated STSBO environmental release history of all fission products .......... 5-51

Figure 5-49 Comparison of the mitigated short-term station blackout containment 

pressure history versus the dome hydrogen mass and total hydrogen 

production .......................................................................................................... 5-52

Figure 5-50 Comparison of mitigated short-term station blackout containment cavity 

gas concentration history and potential ignition source temperatures............... 5-54

Figure 5-51 Comparison of the mitigated short-term station blackout containment 

dome gas concentration history and potential ignition source temperature....... 5-54

Figure 5-52 Mitigated short-term station blackout containment pressure history for

the sensitivity calculation with delayed ignition................................................ 5-56

Figure 5-53 Mitigated short-term station blackout containment dome gas 

concentrations versus stoichiometric ................................................................. 5-57

Figure 5-54 Comparison of the containment pressure responses for the sensitivity 

calculations with delayed ignition ..................................................................... 5-58

Figure 5-55 Comparison of the cesium release to the environment for the sensitivity 

calculations with delayed ignition ..................................................................... 5-59

Figure 5-56 Comparison of the iodine release to the environment for the sensitivity 

calculations with delayed ignition ..................................................................... 5-59

Figure 5-57 Mitigated short-term station blackout containment gas concentration 

history for the sensitivity calculation with delayed ignition.............................. 5-61

Figure 5-58 Mitigated short-term station blackout containment pressure history for 

the sensitivity calculation with delayed ignition................................................ 5-61

Figure 5-59 Mitigated short-term station blackout containment airborne radionuclide 

history versus hydrogen concentrations for the sensitivity calculation 

with delayed ignition.......................................................................................... 5-62

Figure 5-60 Unmitigated 100% and 200% TI-SGTR STSBO primary and secondary 

pressures histories.............................................................................................. 5-69

Figure 5-61 Comparison of the hot leg nozzle and steam generator tube structure 

temperatures and creep rupture indices for the STSBO with TI-SGTR............ 5-69

Figure 5-62 Unmitigated 100% and 200% TI-SGTR STSBO primary and TI-SGTR 

flowrate histories................................................................................................ 5-70

Figure 5-63 Unmitigated 100% and 200% TI-SGTR STSBO vessel convective 

heat removal rate from the fuel.......................................................................... 5-70

Figure 5-64 Unmitigated 100% and 200% TI-SGTR STSBO fuel oxidation power 

before the RCS hot leg failure ........................................................................... 5-71

Figure 5-65 Unmitigated 100% and 200% TI-SGTR short-term station blackout vessel 

two-phase coolant level...................................................................................... 5-71

Figure 5-66 Unmitigated 100% and 200% TI-SGTR short-term station blackout fuel 

oxidation power after the RCS hot leg failure ................................................... 5-72

Figure 5-67 Unmitigated 100% and 200% TI-SGTR short-term station blackout lower 

head inner and outer temperature histories........................................................ 5-72

Figure 5-68 Unmitigated 100% and 200% TI-SGTR short-term station blackout 

containment pressure histories........................................................................... 5-73

Figure 5-69 Unmitigated 100% and 200% TI-SGTR short-term station blackout 

containment and TI-SGTR leakage areas.......................................................... 5-73
xii

Figure 5-70 Unmitigated 100% TI-SGTR STSBO iodine fission product distribution 

history ................................................................................................................ 5-76

Figure 5-71 The unmitigated 200% TI-SGTR short-term station blackout iodine fission 

product distribution history................................................................................ 5-76

Figure 5-72 The unmitigated 100% TI-SGTR short-term station blackout cesium 

fission product distribution history .................................................................... 5-77

Figure 5-73 The unmitigated 200% TI-SGTR short-term station blackout cesium 

fission product distribution history .................................................................... 5-77

Figure 5-74 The unmitigated 100% and 200% TI-SGTR short-term station blackout 

iodine fission product distribution history......................................................... 5-78

Figure 5-75 The unmitigated 100% and 200% TI-SGTR short-term station blackout 

cesium fission product distribution history........................................................ 5-78

Figure 5-76 The unmitigated 100% TI-SGTR short-term station blackout 

environmental release history of all fission products ........................................ 5-79

Figure 5-77 The unmitigated 200% TI-SGTR short-term station blackout 

environmental release history of all fission products ........................................ 5-79

Figure 5-78 The mitigated STSBO primary and secondary pressure history........................ 5-82

Figure 5-79 The mitigated short-term station blackout vessel two-phase coolant level ....... 5-83

Figure 5-80 The mitigated short-term station blackout containment pressure history.......... 5-83

Figure 5-81 The mitigated short-term station blackout containment pressure history.......... 5-84

Figure 5-82 The iodine distribution in the containment for short-term station 

blackout with a 100% thermally-induced SGTR with spray mitigation............ 5-85

Figure 5-83 The cesium distribution in the containment for short-term station 

blackout with a 100% thermally-induced SGTR with spray mitigation............ 5-86

Figure 5-84 The short-term station blackout with a 100% thermally-induced SGTR 

with and without spray mitigation iodine environmental release ...................... 5-86

Figure 5-85 The short-term station blackout with a 100% thermally-induced SGTR 

with and without spray mitigation cesium environmental release..................... 5-87

Figure 5-86 The hot leg creep rupture failure index in the short-term station blackout 

sensitivity case with a 100% thermally-induced SGTR and no hot leg 

failure ................................................................................................................. 5-89

Figure 5-87 The hot leg temperature response in the thermally-induced steam 

generator tube rupture cases with and without hot leg failure ........................... 5-90

Figure 5-88 The hot leg creep rupture failure index and iodine release to environment 

for the thermally-induced SGTR cases with and without hot leg failure .......... 5-90

Figure 5-89 SGTR with Operator Action – System Pressures.............................................. 5-94

Figure 5-90 SGTR with Operator Action – RCS Conditions Relative to RHR Entry 

Conditions.......................................................................................................... 5-95

Figure 5-91 SGTR with Operator Action – RPV Level ........................................................ 5-95

Figure 5-92 SGTR with Operator Action – Maximum Cladding and Lower 

Head Temperatures............................................................................................ 5-96

Figure 5-93 SGTR with Operator Action – RWST and ECST Inventories .......................... 5-96

Figure 5-94 SGTR with Operator Action- Containment Pressure......................................... 5-97

Figure 5-95 SGTR with Operator Action- Steam Generator Level....................................... 5-97

Figure 5-96 SGTR with Failed Operator Action – System Pressures................................. 5-101

Figure 5-97 SGTR with Failed Operator Action – RPV Water Level ................................ 5-102
xiii

Figure 5-98 SGTR with Failed Operator Action – Maximum Cladding and Lower 

Head Temperatures.......................................................................................... 5-102

Figure 5-99 SGTR with Failed Operator Action – RWST and ECST Inventories ............. 5-103

Figure 5-100 SGTR with Failed Operator Action-Containment Pressure............................. 5-103

Figure 5-101 SGTR with Failed Operator Action- Steam Generator Level.......................... 5-104

Figure 5-102 SGTR with Failed Operator Action and SORV – System Pressures............... 5-107

Figure 5-103 SGTR with Failed Operator Action and SORV – RPV Water Level.............. 5-107

Figure 5-104 SGTR with Failed Operator Action and SORV – Max Clad and Lower 

Head Temperature............................................................................................ 5-108

Figure 5-105 SGTR with Failed Operator Action with SORV – RWST and ECST 

Inventories........................................................................................................ 5-108

Figure 5-106 SGTR with Failed Operator Action and SORV- Containment Pressure ......... 5-109

Figure 5-107 SGTR with Failed Operator Action and SORV- Steam Generator Level ....... 5-109

Figure 5-108 ISLOCA Reactor System Pressure .................................................................. 5-113

Figure 5-109 ISLOCA Break Flow (RCS to Safeguards) ..................................................... 5-113

Figure 5-110 ISLOCA ECCS Flow....................................................................................... 5-114

Figure 5-111 ISLOCA RWST Inventory .............................................................................. 5-114

Figure 5-112 ISLOCA Integral Accumulator Flow............................................................... 5-115

Figure 5-113 ISLOCA Reactor Vessel Water Level ............................................................. 5-115

Figure 5-114 ISLOCA In-Vessel Hydrogen Production ....................................................... 5-116

Figure 5-115 ISLOCA Maximum Core Temperature ........................................................... 5-116

Figure 5-116 ISLOCA Reactor Vessel Lower Head Temperature........................................ 5-117

Figure 5-117 ISLOCA Containment Pressure....................................................................... 5-118

Figure 5-118 ISLOCA Containment Temperature ................................................................ 5-118

Figure 5-119 ISLOCA Safeguards Pressure.......................................................................... 5-120

Figure 5-120 ISLOCA Safeguards Temperature................................................................... 5-120

Figure 5-121 ISLOCA Safeguards Level Short Term........................................................... 5-121

Figure 5-122 ISLOCA Safeguards Level Long Term........................................................... 5-122

Figure 5-123 ISLOCA Integral Mass of H2 & CO Burned in Safeguards Area.................... 5-123

Figure 5-124 ISLOCA Energy Release from H2 and CO Burns in Safeguards Area............ 5-123

Figure 5-125 ISLOCA Safeguards Exhaust Ventilation Fan Head....................................... 5-125

Figure 5-126 ISLOCA Safeguards Exhaust Ventilation Flows............................................. 5-125

Figure 5-127 ISLOCA Safeguards Exhaust Ventilation Filter Aerosol Loading.................. 5-126

Figure 5-128 ISLOCA Safeguards Exhaust Ventilation Filter Unit Pressure Diff ............... 5-126

Figure 5-129 ISLOCA Safeguards Exhaust Ventilation Duct and Filter Inlet 

Temperature ..................................................................................................... 5-127

Figure 5-130 ISLOCA Decay Power of Fission Products Captured in Safeguards 

Exhaust Ventilation HEPA Filters................................................................... 5-127

Figure 5-131 ISLOCA Fission Product Release to the Environment.................................... 5-129

Figure 5-132 ISLOCA Cesium Distribution ......................................................................... 5-129

Figure 5-133 ISLOCA Iodine Distribution............................................................................ 5-130

Figure 5-134 ISLOCA CsI Releases by Available Pathways (% Instantaneous Total) ........ 5-130

Figure 5-135 Pressure Along LHSI Piping in ISLOCA Deposition Calculation .................. 5-132

Figure 5-136 Vapor Temperature Along LHSI Piping in ISLOCA Deposition 

Calculation ....................................................................................................... 5-132
xiv

Figure 5-137 Pipe Wall Inside Surface Temperature Along LHSI Piping in ISLOCA 

Deposition Calculation..................................................................................... 5-133

Figure 5-138 Vapor Velocity along LHSI Piping in ISLOCA Deposition Calculation........ 5-133

Figure 5-139 LHSI Piping CsI and Cs2MoO4 Capture in ISLOCA Deposition 

Calculation ....................................................................................................... 5-135

Figure 5-140 LHSI Piping Ba and Te Capture in ISLOCA Deposition Calculation............. 5-136

Figure 5-141 CsI Capture along LHSI Piping in ISLOCA Deposition Calculation 

(Early) .............................................................................................................. 5-136

Figure 5-142 CsI Capture along LHSI Piping in ISLOCA Deposition Calculation.............. 5-137

Figure 5-143 Cs2MoO4 Capture along LHSI Piping in ISLOCA Deposition Calc............... 5-137

Figure 5-144 Captured Particulate Volume versus Pipe Volume in ISLOCA 

Deposition Calculation..................................................................................... 5-138

Figure 5-145 ISLOCA RCS Pressure and Temperature Relative to the RHR Entry 

Pressure and Temperature Criteria for the Unmitigated ISLOCA................... 5-145

Figure 5-146 ISLOCA Vessel Swollen Water Level for the Unmitigated ISLOCA ............ 5-145

Figure 5-147 ISLOCA Break Flow for the Unmitigated ISLOCA ....................................... 5-146

Figure 5-148 RWST Volume for the Unmitigated ISLOCA................................................. 5-146

Figure 5-149 Auxiliary Building Water Volume for the Unmitigated ISLOCA................... 5-147

Figure 5-150 Facility pressure response to PORV actuation at 0.5hrs.................................. 5-149

Figure 5-151 Pressurizer liquid level response to PORV actuation at 0.5hrs ....................... 5-150

Figure 5-152 Flow Rate to Pressurizer Relief Tank – PORVs Open at 0.5 hrs .................... 5-150

Figure 5-153 Flow Rates with CsI Invessel Release – PORVs Open at 0.5 hrs.................... 5-151

Figure 5-154 Cesium Release to the Environment – PORVs Open at 0.5 hrs ...................... 5-151

Figure 5-155 Iodine Release to the Environment – PORVs Open at 0.5 hrs ........................ 5-152

Figure 5-156 Facility pressure response to PORV actuation at Core Degradation ............... 5-152

Figure 5-157 Flow Rates - CsI Invessel Release – PORVs Open at Core Degradation........ 5-153

Figure 5-158 Cesium Release to Environment – PORVs Open at Core Degradation........... 5-153

Figure 5-159 Iodine Release to Environment – PORVs Open at Core Degradation............. 5-154

Figure 5-160 ISLOCA HHSI Throttling Sensitivity - RWST Inventory .............................. 5-155

Figure 5-161 ISLOCA HHSI Throttling Sensitivity - Fission Product Decay Power........... 5-156

Figure 5-162 ISLOCA HHSI Throttling Sensitivity - RCS Pressure .................................... 5-156

Figure 5-163 ISLOCA HHSI Throttling Sensitivity - Core Water Level ............................. 5-157

Figure 5-164 ISLOCA HHSI Throttling Sensitivity - Maximum Core Temperature ........... 5-157

Figure 5-165 The additional gaseous iodine source term using Phebus data compared 

to iodine source term for unmitigated short-term station blackout.................. 5-160

Figure 5-166 The additional gaseous iodine source term using Phebus data is 

compared to the iodine source term for the unmitigated short-term 

station blackout with a thermally-induced steam generator tube rupture ........ 5-160

Figure 5-167 Comparison of the oxygen concentration and UO2 mass in the vessel 

during the unmitigated STSBO........................................................................ 5-162

Figure 5-168 Unmitigated STSBO airborne aerosol mass in the containment 

following vessel failure.................................................................................... 5-163

Figure 6-1 Surry 10 and 20 Mile Areas................................................................................. 6-2

Figure 6-2 Unmitigated LTSBO Emergency Response Timeline....................................... 6-12

Figure 6-3 Duration of Protective Actions - Unmitigated LTSBO ..................................... 6-12

Figure 6-4 Unmitigated STSBO emergency response timeline .......................................... 6-16
xv

Figure 6-5 Duration of Protective Actions - Unmitigated STSBO...................................... 6-16

Figure 6-6 Unmitigated STSBO with TI-SGTR emergency response timeline .................. 6-20

Figure 6-7 Duration of Protective Actions - Unmitigated STSBO with TI-SGTR............. 6-20

Figure 6-8 Mitigated STSBO with TI-SGTR emergency response timeline ...................... 6-22

Figure 6-9 Duration of Protective Actions - Mitigated STSBO with TI-SGTR.................. 6-23

Figure 6-10 Unmitigated ISLOCA Emergency Response Timeline ..................................... 6-26

Figure 6-11 Duration of Protective Actions - Unmitigated ISLOCA.................................... 6-26

Figure 6-12 Evacuation Timeline from Surry for the 10 to 20 Mile Region......................... 6-30

Figure 6-13 Unmitigated ISLOCA Timeline for Sensitivity 1.............................................. 6-31

Figure 6-14 Duration of Protective Actions - Unmitigated ISLOCA Sensitivity 1 .............. 6-31

Figure 6-15 Unmitigated ISLOCA Timeline for Sensitivity 2.............................................. 6-34

Figure 6-16 Protective Action Durations - Unmitigated ISLOCA Sensitivity 2................... 6-34

Figure 6-17 Unmitigated ISLOCA Timing for Sensitivity 3................................................. 6-36

Figure 6-18 Protective Action Durations - Unmitigated ISLOCA Sensitivity 3................... 6-37

Figure 6-19 Highway 199 Over Highway 321 ...................................................................... 6-39

Figure 6-20 Roadway Network Identifying Potentially Affected Roadways and 

Bridges............................................................................................................... 6-40

Figure 6-21 STSBO with TI-SGTR Emergency Response Timeline (Seismic) ................... 6-46

Figure 6-22 Protective Action Durations - STSBO with TI-SGTR (Seismic)...................... 6-46

Figure 6-23 Surry and Rushmere Fire Departments.............................................................. 6-53

Figure 6-24 Containment Pressure for Unmitigated STSBO ................................................ 6-56

Figure 6-25 Containment Water Level vs. Volume............................................................... 6-57

Figure 6-26 Surry Containment System ................................................................................ 6-58

Figure 6-27 Pumping Capacity.............................................................................................. 6-59

Figure 7-1 Mean, individual LCF risk per Reactor Year from the Surry unmitigated 

LTSBO scenario for residents within a circular area of specified radius 

from the plant for three values of dose-truncation level...................................... 7-5

Figure 7-2 Mean, individual LNT, LCF risk per Reactor Year from the Surry 

unmitigated LTSBO scenario for residents within a circular area of 

specified radius from the plant for the emergency and long-term phases ........... 7-6

Figure 7-3 Mean, Individual, LCF Risk per Reactor Year from the Surry 

unmitigated STSBO scenario for residents within a circular area of 

specified radius from the plant for three values of dose-truncation level............ 7-8

Figure 7-4 Mean, Individual, LNT, LCF Risk per Reactor Year from the Surry 

unmitigated STSBO scenario for residents within a circular area of 

specified radius from the plant for the emergency and long-term phases ........... 7-9

Figure 7-5 Mean, Individual, LCF Risk per Reactor Year from the Surry, unmitigated 

STSBO with TISGTR scenario for residents within a circular area of 

specified radius from the plant for three choices of dose-truncation level........ 7-11

Figure 7-6 Mean, Individual, LNT, LCF Risk per Reactor Year from the Surry, 

unmitigated, STSBO with TISGTR scenario for residents within a 

circular area of specified radius from the plant for the emergency and 

long-term phases. ............................................................................................... 7-12

Figure 7-7 Mean, Individual, LCF Risk per Reactor Year from the Surry, mitigated, 

STSBO with TISGTR scenario for residents within a circular area of 

specified radius from the plant for three choices of dose-truncation level........ 7-14
xvi

Figure 7-8 Mean, Individual, LNT, LCF Risk per Reactor Year from the Surry 

mitigated STSBO with TISGTR scenario for residents within a circular 

area of specified radius from the plant for the emergency and long-term 

phases................................................................................................................. 7-15

Figure 7-9 Mean, Individual, LCF Risk per Reactor Year from the Surry, 

unmitigated, ISLOCA scenario for residents within a circular area of 

specified radius from the plant for three choices of dose truncation level ........ 7-17

Figure 7-10 Mean, Individual, LNT, LCF Risk per Reactor Year from the Surry, 

unmitigated, ISLOCA scenario for residents within a circular area of 

specified radius from the plant for the emergency and long-term phases ......... 7-18

Figure 7-11 Mean, LNT, latent cancer fatality risks from the SST1 source term for 

residents within a circular area of specified radius from the Surry plant 

for the emergency and long-term phases. .......................................................... 7-23

Figure 7-12 Percentage contribution to total, emergency-phase, and long-term-phase, 

mean, individual risk for the population within 10 miles by chemical 

class for the Surry unmitigated ISLOCA based on the LNT hypothesis........... 7-28

Figure 7-13 Percentage contribution to total, mean, individual risk for the population 

within 10 miles by chemical class for the Surry unmitigated ISLOCA 

based on US BGR dose truncation..................................................................... 7-29

Figure 7-14 Percentage contribution to total, mean, individual risk for the population 

within 10 miles by chemical class for the Surry unmitigated ISLOCA 

based on a truncation level reflecting the HPS Position for quantifying 

health effects...................................................................................................... 7-29

Figure 7-15 Percentage contribution to total, emergency-phase, and long-term-phase, 

mean, individual risk for the population within 20 miles by chemical 

class for the Surry unmitigated ISLOCA based on LNT hypothesis................. 7-30

Figure 7-16 Percentage contribution to total, mean, individual risk for the population 

within 20 miles by chemical class for the Surry unmitigated ISLOCA 

based on US BGR dose truncation..................................................................... 7-31

Figure 7-17 Percentage of contribution to total, mean, individual risk for the 

population within 20 miles by chemical class for the Surry unmitigated 

ISLOCA based on a truncation level reflecting the HPS Position for 

quantifying health effects................................................................................... 7-31

Figure 7-18 Percentage contribution to total, emergency-phase, and long-term-phase, 

mean, individual risk for the population within 50 miles by chemical 

class for the Surry unmitigated ISLOCA based on the LNT hypothesis........... 7-32

Figure 7-19 Percentage contribution to total, mean, individual risk for the population 

within 50 miles by chemical class for the Surry unmitigated ISLOCA 

based on US BGR dose truncation..................................................................... 7-32

Figure 7-20 Percentage contribution to total, mean, individual risk for the population 

within 50 miles by chemical class for the Surry unmitigated ISLOCA 

based on a truncation level reflecting the HPS Position for quantifying 

health effects...................................................................................................... 7-33
xvii

LIST OF TABLES

Table 4-1 Important Design Parameters for Surry ............................................................... 4-3

Table 4-2 Safety Relief Valve Opening Pressure and Flow Capacity ............................... 4-13

Table 4-3 Decay Power in Surry MELCOR Model........................................................... 4-15

Table 4-4 Time versus Temperature Relationship for Fuel Rod Collapse......................... 4-19

Table 5-1 Timing of key events for unmitigated LTSBO.................................................... 5-2

Table 5-2 Timing of key events for mitigated LTSBO...................................................... 5-13

Table 5-3 Comparison of timings of key events for unmitigated LTSBO with and 

without early RCP seal failures.......................................................................... 5-19

Table 5-4 Comparison of the timings of key events for mitigated LTSBO with and 

without early RCP seal failures.......................................................................... 5-27

Table 5-5 Timing of key events for unmitigated STSBO .................................................. 5-32

Table 5-6 Timing of key events for mitigated STSBO ...................................................... 5-43

Table 5-7 Containment conditions at start of hydrogen and carbon monoxide burn......... 5-56

Table 5-8 Timing of key events for unmitigated STSBO TI-SGTR.................................. 5-66

Table 5-9 The timing of key events for mitigated STSBO with TI-SGTR........................ 5-80

Table 5-10 The timing of hot leg failure for SCDAP/RELAP5 simulations with 

TI-SGTR ............................................................................................................ 5-89

Table 5-11 Timing of key events for the Spontaneous SGTR with Expected 

Operator Action ................................................................................................. 5-92

Table 5-12 Timing of key events for the Spontaneous SGTR with Failed Operator 

Action................................................................................................................. 5-99

Table 5-13 The timing of key events for Spontaneous SGTR with Failed Operator 

Action and Faulted Steam Generator SORV ................................................... 5-105

Table 5-14 Sequence of Events for the Unmitigated ISLOCA.......................................... 5-111

Table 5-15 LHSI Piping Aerosol Capture in ISLOCA Separate-effects Calculation 

and Associated DFs.......................................................................................... 5-139

Table 5-16 Comparison of Iodine Spike Source Term to Iodine Source Terms from 

the Other Unmitigated Accidents..................................................................... 5-161

Table 6-1 Scenarios Assessed for Emergency Response ..................................................... 6-3

Table 6-2 Surry Cohort Population Values.......................................................................... 6-5

Table 6-3 Surry Shielding Factors........................................................................................ 6-8

Table 6-4 Unmitigated LTSBO cohort timing ................................................................... 6-15

Table 6-5 Unmitigated STSBO cohort timing ................................................................... 6-19

Table 6-6 Unmitigated STSBO with TI-SGTR cohort timing........................................... 6-22

Table 6-7 Unmitigated STSBO with TI-SGTR cohort timing........................................... 6-24

Table 6-8 Unmitigated ISLOCA cohort timing ................................................................. 6-28

Table 6-9 Sensitivity Case 1 cohort timing........................................................................ 6-33

Table 6-10 Sensitivity Case 2 cohort timing........................................................................ 6-35

Table 6-11 Sensitivity Case 3 cohort timing........................................................................ 6-38

Table 6-12 Description of the Potential Evacuation Failure Locations ............................... 6-41

Table 6-13 Cohort Timing STSBO with TI-SGTR Including Speeds................................. 6-48

Table 6-14 Rated Flowrate of Fire Pumps........................................................................... 6-53
xviii

Table 6-15 Cumulative Water Volume vs. Elevation for Unit 1.......................................... 6-57

Table 7-1 Brief Source-Term Description for Unmitigated Surry Accident Scenarios 

and the SST1 Source Term from the 1982 Siting Study...................................... 7-2

Table 7-2 Mean, Individual LCF Risk per Event (Dimensionless) for Residents 

within the Specified Radii of the Surry Site for the Unmitigated LTSBO 

for a Mean CDF of 2x10-5 pry ............................................................................. 7-5

Table 7-3 Mean, Individual LCF Risk per Reactor Year for Residents within the 

Specified Radii of the Surry Site for Unmitigated LTSBO Scenario for a 

Mean CDF of 2x10-5 pry...................................................................................... 7-5

Table 7-4 Mean, Individual, LCF Risk per Event (Dimensionless) for Residents 

within the Specified Radii of the Surry Site for the Unmitigated STSBO 

Scenario, for a Mean CDF of 1.5x10-6/pry .......................................................... 7-7

Table 7-5 Mean, Individual, LCF Risk per Reactor Year for Residents within the 

Specified Radii of the Surry Site for Unmitigated STSBO Scenario for a 

Mean CDF of 1.5x10-6 pry................................................................................... 7-7

Table 7-6 Mean, Individual, LCF Risk per Event (Dimensionless) for Residents 

within the Specified Radii of the Surry Site for the Unmitigated STSBO 

with TISGTR With a Mean CDF of 4x10-7 pry................................................. 7-10

Table 7-7 Mean, Individual, LCF Risk per Reactor Year for Residents within the 

Specified Radii of the Surry Site for the Unmitigated STSBO with 

TISGTR With a Mean CDF of 4x10-7 pry......................................................... 7-10

Table 7-8 Mean, Individual, LCF Risk per Event (Dimensionless) for Residents 

within the Specified Radii of the Surry Site for Mitigated STSBO with 

TISGTR Scenario with a CDF of 4x10-7 pry..................................................... 7-13

Table 7-9 Mean, Individual, LCF Risk per Reactor-Year for Residents within the 

Specified Radii of the Surry Site for the Mitigated STSBO with TISGTR 

Scenario with a CDF of 4x10-7 pry .................................................................... 7-13

Table 7-10 Mean, Individual LCF Risk per Event (Dimensionless) for Residents 

within the Specified Radii of the Surry Site for the Unmitigated ISLOCA 

Scenario with a CDF of 3x10-8 pry .................................................................... 7-16

Table 7-11 Mean, Individual, LCF Risk per Reactor-Year for Residents within the 

Specified Radii of the Surry Site for the Unmitigated ISLOCA Scenario 

with a CDF of 3x10-8 pry................................................................................... 7-17

Table 7-12 Mean, Individual, Prompt Fatality Risk per Event (Dimensionless) for 

Residents within the Specified Radii of the Surry Site for the ISLOCA 

Scenario with a Mean CDF of 3x10-8 pry.......................................................... 7-18

Table 7-13 Mean, Individual Prompt-Fatality Risk per Reactor-Year for Residents 

within the Specified Radii of the Surry Site for the ISLOCA Scenario 

with a Mean CDF of 3x10-8 pry......................................................................... 7-19

Table 7-14 Effect of Size of Evacuation Zone on Mean, LNT, LCF Risks for 

Residents within the Specified Radii of Surry Site for the Unmitigated 

ISLOCA Scenario ............................................................................................. 7-20
xix

Table 7-15 Mean, Individual, LNT, LCF Risk per Event (Dimensionless) for 

Residents within the Specified Radii of the Surry Site for the unmitigated 

TISGTR Scenario and Comparing the Unmodified Emergency Response 

(ER) and ER Adjusted to Account for the Effect of Seismic Activity on 

Evacuation Routes and Human Response.......................................................... 7-20

Table 7-16 Mean, Individual, LCF Risk per Event (Dimensionless) for Residents 

within the Specified Radii of the Surry Site for the SST1 Source Term 

from the 1982 Siting Study with All Parameters Other than for Source 

Terms Are Taken from the Unmitigated STSBO Scenario ............................... 7-21

Table 7-17 Mean, Individual, LNT, LCF Risk per Event (Dimensionless) for 

Residents within the Specified Radii of the Surry Site for the SST1 Source 

Term from the 1982 Siting Study Using Emergency Response Parameters 

from the STSBO Scenario. Results are Compared with the Unmitigated 

ISLOCA and the Unmitigated STSBO with TISGTR Scenarios ...................... 7-22

Table 7-18 Mean, Individual, Prompt-Fatality Risk per Event (Dimensionless) for 

Residents within the Specified Radii of the Surry Site for the SST1 Source 

Term from the 1982 Siting Study Using Emergency Response Parameters 

from the Unmitigated STSBO Scenario ............................................................ 7-23

Table 7-19 Surface roughness for various land-use categories for the area surrounding 

the Surry site ...................................................................................................... 7-24

Table 7-20 Deposition Velocities Used for the Base Case Calculations and for the 

Surface Roughness Sensitivity Study for Each of the Ten Aerosol Bins in 

the MELCOR Model.......................................................................................... 7-26

xxi

EXECUTIVE SUMMARY

The U.S. Nuclear Regulatory Commission (NRC), the nuclear power industry, and the 

international nuclear energy research community have devoted considerable research over the 

last several decades to examining severe reactor accident phenomena and offsite consequences. 

Following the terrorist attacks of 2001, an NRC initiative reassessed severe accident progression 

and offsite consequences in response to security-related events. These updated analyses 

incorporated the wealth of accumulated research and used more detailed, integrated, and bestestimate modeling than past analyses. An insight gained from these security assessments was 

that the NRC needed updated analyses of severe reactor accidents to reflect realistic estimates of 

the more likely outcomes, considering the current state of plant design and operation and the 

advances in understanding of severe accident behavior.

The NRC initiated the State-of-the-Art Reactor Consequence Analyses (SOARCA) project to 

develop best estimates of the offsite radiological health consequences for potential severe reactor 

accidents for two pilot plants: the Peach Bottom Atomic Power Station in Pennsylvania and the 

Surry Power Station in Virginia. Peach Bottom is generally representative of U.S. operating 

reactors using the General Electric boiling-water reactor (BWR) design with a Mark I 

containment. Surry is generally representative of U.S. operating reactors using the 

Westinghouse pressurized-water reactor (PWR) design with a large, dry (subatmospheric) 

containment. SOARCA results, while specific to Peach Bottom and Surry, may be generally 

applicable to plants with similar designs. Additional work would be needed to confirm this, 

however, since differences exist in plant-specific designs, procedures, and emergency response 

characteristics.

The SOARCA project evaluates plant improvements and changes not reflected in earlier NRC 

publications such as NUREG/CR-2239, “Technical Guidance for Siting Criteria Development,”

NUREG-1150, “Severe Accident Risks: An Assessment for Five U.S. Nuclear Power Plants,”

and WASH-1400, “Reactor Safety Study: An Assessment of Accident Risks in U.S. Commercial 

Nuclear Power Plants.” SOARCA includes system improvements, improvements in training 

and emergency procedures, offsite emergency response, and security-related improvements, as 

well as plant changes such as power uprates and higher core burnup. To provide perspective 

between SOARCA results and more conservative offsite consequence estimates, SOARCA 

results are compared to NUREG/CR-2239, “Technical Guidance for Siting Criteria 

Development,” issued in 1982 and referred to in this report as the Siting Study. Specifically, 

SOARCA results are compared to the Siting Study siting source term 1 (SST1). SST1 assumes 

severe core damage, loss of all safety systems, and loss of containment after 1.5 hours. The 

SOARCA report helps the NRC to communicate its current understanding of severe-accidentrelated aspects of nuclear safety to stakeholders, including Federal, State, and local authorities, 

licensees, and the general public.

The SOARCA project sought to focus its resources on the more important severe accident 

scenarios for Peach Bottom and Surry. The project narrowed its approach by using an accident 

sequence’s possibility of damaging reactor fuel, or core damage frequency (CDF), as a surrogate 

for risk. The SOARCA scenarios were selected from the results of existing probabilistic risk 

assessments (PRAs). Core damage sequences from previous staff and licensee PRAs were 
xxii

identified and binned into core damage groups. A core damage group consists of core damage 

sequences that have similar timing for important severe accident phenomena and similar 

containment or engineered safety feature operability. It is important to note that each core 

damage sequence that belongs to a given core damage group is initiated by a specific cause (for 

example, a seismic event, a fire, or a flood), and that the frequency of each core damage group 

was estimated by aggregating the CDFs of the individual sequences that belong to the 

group. This approach was taken to help ensure that the contributions from all core damage 

sequences were accounted for during the sequence selection process. During the consequence 

analysis, the core damage groups for station blackouts were analyzed as if they were initiated by 

a seismic event. This approach was taken because seismically induced equipment failures occur 

immediately following the seismic event, which produces the most severe challenge to the 

plant. The groups were screened according to their approximate CDFs to identify the most risk 

significant groups. SOARCA analyzed scenarios with a CDF equal to or greater than 10-6 (1 in a 

million) per reactor-year. SOARCA also sought to analyze scenarios leading to an early failure 

or bypass of the containment with a CDF equal to or greater than 10-7 (1 in 10 million) per 

reactor-year, since these scenarios have a potential for higher consequences and risk. This 

approach allowed a more detailed analysis of accident consequences for the more likely, 

although still remote, accident scenarios. 

The staff used updated and benchmarked standardized plant analysis risk (SPAR) models and 

available plant-specific external events information in the scenario-selection process and 

identified two major groups of accident scenarios for analysis. The first group common to both 

Peach Bottom and Surry includes short-term station blackout (STSBO) and long-term station 

blackout (LTSBO). Both types of SBOs involve a loss of all alternating current (ac) power. The 

STSBO also involves the loss of turbine-driven systems through loss of direct current (dc) 

control power or loss of the condensate storage tank and therefore proceeds to core damage more 

rapidly (hence “short term”). The STSBO has a lower CDF, since it requires a more severe 

initiating event and more extensive system failures. SBO scenarios can be initiated by external 

events such as a fire, flood, or earthquake. SOARCA assumes that an SBO is initiated by a 

seismic event since this is the most extreme case in terms of both the timing and amount of 

equipment that fails. Notwithstanding the SOARCA scenario screening process, SBO scenarios 

are commonly identified as important contributors in PRA because of the common cause of 

failure for both reactor safety systems and containment safety systems. 

SOARCA’s second severe accident scenario group, which was identified for Surry only, is the 

containment bypass scenario. For Surry, two containment bypass scenarios were identified and 

analyzed. The first bypass scenario is a variant of the STSBO scenario, involving a thermallyinduced steam generator tube rupture (TISGTR). The second bypass scenario involves an 

interfacing systems loss-of-coolant accident (ISLOCA) caused by an unisolated rupture of lowhead safety injection piping outside containment. The CDF for the ISLOCA, 3×10-8 (3 in 100 

million) per reactor-year, falls below the SOARCA screening criterion for bypass events but it is 

analyzed for completeness because NUREG-1150 identified ISLOCA, in addition to SBO and 

SGTRs, as principal contributors to mean early and latent cancer fatality risks. This scenarioselection process captured the more important internally and externally initiated core damage 

scenarios.
xxiii

SOARCA’s analyses were performed with two computer codes, MELCOR for accident 

progression and the MELCOR Accident Consequence Code System, Version 2 (MACCS2) for 

offsite consequences. The NRC staff’s preparations for the analyses included extensive 

cooperation from the licensees of Peach Bottom and Surry to develop high-fidelity plant systems 

models, define operator actions including the most recently developed mitigation actions, and 

develop models for simulation of site-specific and scenario-specific emergency planning and 

response. Moreover, in addition to input for model development, licensees provided information 

on accident scenarios from their PRAs. Through tabletop exercises of the selected scenarios 

with senior reactor operators, PRA analysts, and other licensee staff, licensees provided input on 

the timing and nature of the operator actions to mitigate the selected scenarios. The licensee 

input for each scenario was used to develop assumed timelines of operator actions and equipment 

configurations for implementing available mitigation measures which include mitigation 

measures beyond those routinely credited in current PRA models. A human reliability analysis, 

commonly included in PRAs to represent the reliability of operator actions, was not performed 

for SOARCA, but instead tabletop exercises, plant walkdowns, simulator runs and other inputs 

from licensee staff were employed to ensure that operator actions and their timings were 

correctly modeled. 

SOARCA modeled mitigation measures, including those in emergency operating procedures 

(EOPs), severe accident management guidelines (SAMGs), and Title 10 to the Code of Federal 

Regulations (10 CFR) 50.54(hh). The 10 CFR 50.54(hh) mitigation measures refer to additional 

equipment and strategies required by the NRC following the terrorist attacks of September 11, 

2001, to further improve each plant’s capability to mitigate events involving a loss of large areas 

of the plant caused by fire and explosions. To assess the benefits of 10 CFR 50.54(hh) mitigation 

measures and to provide a basis for comparison to the past analyses of unmitigated severe 

accident scenarios, the SOARCA project also analyzed each scenario without 10 CFR 50.54 (hh) 

equipment and procedures. The analysis that credits successful implementation of the 10 CFR 

50.54 (hh) equipment and procedures in addition to actions directed by the EOPs and SAMGs is 

referred to as the mitigated case. The analysis without 10 CFR 50.54(hh) equipment and 

procedures is referred to as the unmitigated case (SAMGs were considered but not implemented 

in the unmitigated case). The unmitigated case of the Surry ISLOCA is an exception to this 

general principle because it was necessary to assume that at least one of the EOP actions failed to 

occur for the scenario to lead to core damage. Chapter 3 of NUREG/CR-7110, Volume 1, 

“SOARCA Peach Bottom Integrated Analysis” and Volume 2, “SOARCA Surry Integrated 

Analysis”, details the specific equipment and operator actions credited for each scenario. 

For the LTSBO scenarios for both Peach Bottom and Surry (the most likely severe accident 

scenario for each plant considered in SOARCA) analyzed assuming no mitigation, core damage 

begins in 9 to 16 hours, and reactor vessel failure begins at about 20 hours. Offsite radiological 

release due to containment failure begins at about 20 hours for Peach Bottom (BWR) and at 45 

hours for Surry (PWR). The SOARCA analyses therefore show that time may be available for 

operators to take corrective action and get additional assistance from plant technical support 

centers even if initial efforts are assumed unsuccessful. For the most rapid events (i.e., the 

unmitigated STSBO in which core damage may begin in 1 to 3 hours), reactor vessel failure 

begins at roughly 8 hours, possibly allowing time to restore core cooling and prevent vessel 

failure. In these cases, containment failure and radiological release begins at about 8 hours for 
xxiv

Peach Bottom and at 25 hours for Surry. For the unmitigated Surry ISLOCA, the offsite 

radiological release begins at about 13 hours and in the other bypass event analyzed, the 

TISGTR, the radiological release begins at about 3.5 hours but is shown by analyses to be 

substantially smaller than the 1982 Siting Study SST1 release.

In addition to delayed radiological releases relative to the 1982 Siting Study SST1 case, the 

SOARCA study demonstrates that the amount of radioactive material released is much smaller as 

shown in Figures ES-1 (Iodine-131) and ES-2 (Cesium-137) below. The Surry ISLOCA iodine 

release is calculated to be 16 percent of the core inventory, but the results are more generally in 

the range of 0.5 to 2 percent for iodine and cesium for the other scenarios analyzed. By contrast,

the 1982 Siting Study SST1 case calculated an iodine release of 45 percent and a cesium release 

of 67 percent of the core inventory.

Figure ES-1 Iodine release to the environment for SOARCA unmitigated 

scenarios and the 1982 Siting Study SST1 case
xxv

Figure ES-2 Cesium release to the environment for SOARCA unmitigated 

scenarios and the 1982 Siting Study SST1 case

Past PRAs and consequence studies showed that sequences involving large early releases were 

important risk contributors. For example, the PWR SBO with a TISGTR was historically 

believed to result in a large, relatively early release potentially leading to higher offsite 

consequences. However, MELCOR analysis of Surry performed for SOARCA shows that the 

release is small, because other reactor coolant system piping inside containment (i.e., hot leg 

nozzle) fails soon after the tube rupture and thereby retains the fission products within the 

containment. Additional work would be needed to determine if this result generally applies for 

all types of PWRs.

While this report does not determine the respective likelihoods of the mitigated and unmitigated 

cases of each scenario, the SOARCA results demonstrate the potential benefits of employing 10 

CFR 50.54(hh) mitigation enhancements for the scenarios analyzed. MELCOR analyses were 

used both to confirm the time available to implement mitigation measures and to confirm that 

those measures, once taken, are effective in preventing core damage or significantly reducing 

radiological releases. When successful mitigation is assumed, the MELCOR results indicate no 

core damage for all scenarios except the Surry STSBO and its TISGTR variant. The securityrelated mitigation measures that provide alternative ac power and portable diesel-driven pumps 

are especially helpful in counteracting SBO scenarios. For the Surry STSBO and its TISGTR

variant, the mitigation is sufficient to flood the containment through the containment spray 

system to cover core debris resulting from vessel failure. For the ISLOCA scenario, installed 

equipment unrelated to 10 CFR 50.54(hh) is effective in preventing core damage owing to the 

time available for corrective action.

For scenarios that release radioactive material to the environment, MACCS2 uses site-specific 

weather data to predict the downwind concentration of material in the plume and the resulting 

population exposures and health effects. The analysis of offsite consequences in SOARCA 
xxvi

incorporates the improved modeling capability reflected in the MELCOR and MACCS2 codes as 

well as detailed site-specific public evacuation models. These models were developed for each 

scenario based on site-specific emergency preparedness programs and State emergency response 

plans to reflect timing of onsite and offsite protective action decisions and the evacuation time 

estimates and road networks at Peach Bottom and Surry. Scenarios that are assumed to be 

initiated by a seismic event consider the earthquake’s impact on implementing emergency plans 

from loss of infrastructure (i.e., long-span bridges, traffic signals, sirens). 

The unmitigated versions of the scenarios analyzed in SOARCA have lower risk of early 

fatalities than calculated in the 1982 Siting Study SST1 case. SOARCA’s analyses show 

essentially zero risk of early fatalities. Early fatality risk was calculated to be ~ 10-14 for the 

unmitigated Surry ISLOCA (for the area within 1 mile of Surry’s exclusion area boundary) and 

zero for all other SOARCA scenarios. In comparison, 92 early fatalities for Peach Bottom and 

45 early fatalities for Surry were calculated for the SST1 case in the 1982 Siting Study.

SOARCA results indicate that bypass events (e.g., Surry ISLOCA) do not pose a higher 

scenario-specific latent cancer fatality risk than non-bypass events (e.g., Surry SBO). While 

consequences are greater when the bypass scenario happens, this is offset by the scenario being 

less likely to happen. SOARCA reinforces the importance of external events relative to internal 

events and the need to continue ongoing work related to external events risk assessment.

Offsite radiological consequences were calculated for each scenario expressed as the average 

individual likelihood of an early fatality and latent cancer fatality. Tables ES-1 (Peach Bottom) 

and ES-2 (Surry) show, for both mitigated and unmitigated cases, conditional (on the occurrence 

of the core damage scenario) scenario-specific probabilities of a latent cancer fatality for an 

individual located within 10 miles of the plant. Tables ES-1 and ES-2 show the results using the 

linear no-threshold (LNT) dose-response model, which assumes that the health risk is directly 

proportional to the exposure and even the smallest radiation exposure carries some risk. The 

tables also provide the scenario-specific latent cancer fatality risk for an individual located 

within 10 miles of the plant, taking into account the scenario’s core damage frequency.
xxvii

Table ES-1 Offsite Consequence Results for Peach Bottom Scenarios Assuming Linear 

No- Threshold (LNT) Dose-Response Model

Scenario

Core damage 

frequency 

[CDF]

(per 

reactor-year)*

Mitigated Unmitigated

Conditional 

scenariospecific 

probability of 

latent cancer 

fatality for an 

individual 

located within 

10 miles

Scenario-specific 

risk

[CDF x 

Conditional]

of latent cancer 

fatality for an 

individual located 

within 10 miles 

(per reactor-year)

Conditional 

scenario-specific 

probability of 

latent cancer 

fatality for an 

individual 

located within 

10 miles

Scenario-specific 

risk

[CDF x 

Conditional]

of latent cancer 

fatality for an 

individual located 

within 10 miles 

(per reactor-year)

Long-term SBO 3×10-6 No Core Damage 9×10-5 ~ 3×10-10 ****

Short-term 

SBO with 

RCIC 

Blackstart** 3×10-7

No Core Damage 7×10-5 ~ 2×10-11  

Short-term 

SBO without 

RCIC 

Blackstart

Not Applicable ** * 2×10-4 ~ 6×10-11 ****

* The CDF assumes that 10 CFR 50.54(hh) equipment and procedures were not used.

** Blackstart of the reactor core isolation cooling (RCIC) system refers to starting RCIC without any ac or dc 

control power. Blackrun of RCIC refers to the long-term operation of RCIC without electricity, once it has 

been started. This typically involves using a portable generator to supply power to indications such as reactor 

pressure vessel (RPV) level to allow the operator to manually adjust RCIC flow to prevent RPV overfill and 

flooding of the RCIC turbine. STSBO RCIC blackstart and limited blackrun is credited as an unmitigated case 

for SOARCA purposes because the licensee has included its use in procedures. Past NRC severe accident 

analyses of STSBO scenarios did not credit blackstart of RCIC. A sensitivity calculation without blackstart was 

therefore performed to provide a basis for comparison to past analyses.

*** A scenario with 10 CFR 50.54(hh) mitigation, but without RCIC blackstart was not analyzed.

**** Estimated risks below 1 x 10-7 per reactor year should be viewed with caution because of the potential impact of 

events not studied in the analyses and the inherent uncertainty in very small calculated numbers.
xxviii

Table ES-2 Offsite Consequence Results for Surry Scenarios Assuming LNT DoseResponse Model

Scenario

Core damage 

frequency 

[CDF]

(per 

reactoryear)*

Mitigated Unmitigated

Conditional 

scenariospecific 

probability of 

latent cancer 

fatality for an 

individual 

located within 

10 miles

Scenario-specific 

risk

[CDF x 

Conditional]

of latent cancer 

fatality for an 

individual located 

within 10 miles 

(per reactor-year)

Conditional 

scenariospecific 

probability of 

latent cancer 

fatality for an 

individual 

located within 

10 miles

Scenario-specific 

risk

[CDF x 

Conditional]

of latent cancer 

fatality for an 

individual located 

within 10 miles 

(per reactor-year)

Long-term SBO 2×10-5 No Core Damage 5×10-5 ~ 7×10-10 ****

Short-term SBO 2×10-6 No Containment Failure ** 9×10-5 ~ 1×10-10  

Short-term SBO 

with TISGTR 4×10-7 3×10-4 ** * ~ 1×10-10 **** 3×10-4 ~ 1×10-10  

Interfacing 

systems LOCA 3×10-8 No Core Damage 3×10-4 ~ 9×10-12 ** **

* The CDF assumes that 10 CFR 50.54(hh) equipment and procedures was not used.

** Accident progression calculations showed that source terms in the mitigated case are smaller than in the 

unmitigated case. Offsite consequence calculations were not run, since the containment fails at about 66 

hours. A review of available resources and emergency plans shows that adequate mitigation measures 

could be brought on site within 24 hours and connected and functioning within 48 hours. Therefore 66 

hours would allow ample time for mitigation via measures brought to the site from offsite. 

*** Containment failure is delayed by about 46 hours in the mitigated case relative to the unmitigated case.

Rounding to one significant figure shows conditional LCF probabilities of 3×10-4 for both mitigated and 

unmitigated cases, however the original values were 2.8×10-4 for the mitigated case and 3.2×10-4 for the 

unmitigated case.

**** Estimated risks below 1 x 10-7 per reactor year should be viewed with caution because of the potential 

impact of events not studied in the analyses and the inherent uncertainty in very small calculated numbers.

LCF risks using alternate dose-response models, as well as LCF risks for circular areas out to a 

radius of 50 miles, are also presented. Using a dose-response model that truncates annual doses 

below normal background levels (including medical exposures) results in a further reduction to 

the latent cancer fatality risks (by a factor of 100 for smaller releases and a factor of 3 for larger 

releases). Latent cancer fatality risk calculations are generally dominated by long-term exposure 

to small annual doses (~500 mrem per year corresponding to state return criteria) by evacuees 

returning to their homes after the accident and being exposed to residual radiation over a long 

period of time. SOARCA’s calculated LCF risk results are smaller than extrapolations of 1982 

Siting Study SST1 LCF risk results. However, the difference diminishes when considering 

larger areas, out to a distance of 50 miles from the plant.
xxix

Figure ES-3 compares SOARCA’s scenario-specific latent cancer fatality risks for an individual 

within 10 miles of the plant to the NRC Safety Goal and to an extrapolation of the 1982 Siting 

Study SST11 results. 

Figure ES-3 Comparison of individual LCF risk results for SOARCA mitigated and 

unmitigated scenarios to the NRC Safety Goal and to extrapolations of the 

1982 Siting Study SST1 (plotted on logarithmic scale)

The NRC Safety Goal for latent cancer fatality risk from nuclear power plant operation (i.e., 

2x10-6 or two in one million) is set 1,000 times lower than the sum of cancer fatality risks 

resulting from all other causes (i.e., 2x10-3 or two in one thousand). The calculated cancer 

fatality risks from the selected, important scenarios analyzed in SOARCA are thousands of times 

lower than the NRC Safety Goal and millions of times lower than the general U.S. cancer fatality 

risk.

Comparisons of SOARCA’s calculated LCF risks to the NRC Safety Goal and the average 

annual US cancer fatality risk from all causes are provided to give context that may help the 

 

1 The Siting Study did not calculate LCF risks. Therefore, to compare the Siting Study SST1 case to LCF 

results for SOARCA, the SST1 source term was put into the MACCS2 offsite consequence code files for 

the Peach Bottom and Surry unmitigated STSBO calculations.
xxx

reader to understand the contribution to cancer risks from these nuclear power plant accident 

scenarios. However, such comparisons have limitations for which the reader should be aware. 

Relative to the safety goal comparison, the safety goal is intended to encompass all accident 

scenarios. SOARCA does not examine all scenarios typically considered in a PRA, even though 

it includes the important scenarios. SOARCA represents a mix of limited PRA models with a 

deterministic treatment of various long-term mitigating features. In fact, any analytical 

technique, including PRAs, will have inherent limitations of scope and method. As a result, 

comparison of SOARCA’s scenario-specific calculated LCF risks to the NRC Safety Goal is 

necessarily incomplete. However, it is intended to show that adding multiple scenarios’ low risk 

results in the ~ 10-10 range to approximate a summary risk from all scenarios, would yield a 

summary result that is also below the NRC Safety Goal of 2x10-6 or two in one million.

Relative to the U.S. average individual risk of a cancer fatality comparison, the sources of an 

individual’s cancer risk include a complex combination of age, genetics, lifestyle choices, and 

other environmental factors whereas the consequences from a severe accident at a nuclear plant 

are involuntary and unlikely to be experienced by most individuals.

The SOARCA analyses show that emergency response programs, implemented as planned and 

practiced, reduce the scenario-specific risk of health consequences among the public during a 

severe reactor accident. Sensitivity analyses of seismic impacts on site-specific emergency 

response (e.g., loss of bridges, traffic signals, and delayed notification) at Peach Bottom and 

Surry do not significantly affect LCF risk. 

In summary, the staff believes SOARCA has achieved its objective to develop a body of 

knowledge regarding detailed, integrated, state-of-the-art modeling of the most important severe 

accident scenarios for Peach Bottom and Surry. SOARCA analyses indicate that successful 

implementation of existing mitigation measures can prevent reactor core damage or delay or 

reduce offsite releases of radioactive material. All SOARCA scenarios, even when unmitigated, 

progress more slowly and release much less radioactive material than the 1982 Siting Study 

SST1 case. As a result, the calculated risks of public health consequences from severe accidents 

modeled in SOARCA are very small.

The SOARCA study was nearing completion when the Fukushima Daiichi accident occurred on 

March 11, 2011. The Fukushima accident has many similarities and differences with some of 

the Peach Bottom severe accident scenarios analyzed in SOARCA. While there are significant 

gaps in information and uncertainties regarding what occurred in the Fukushima reactors, an 

appendix to this report compares and contrasts the SOARCA study and the Fukushima accident

based on currently available information for the following topics: (1) operation of the RCIC 

system, (2) hydrogen release and combustion, (3) 48-hour truncation of releases in SOARCA, (4) 

multiunit risk, and (5) spent fuel pool risk.
xxxi

ACKNOWLEDGEMENTS

The contributions from the following individuals in preparing this document are gratefully 

acknowledged. 

Terry Brock Nuclear Regulatory Commission

Richard Chang Nuclear Regulatory Commission

Ata Istar Nuclear Regulatory Commission

Robert Prato Nuclear Regulatory Commission

Mark Orr Nuclear Regulatory Commission

Jason Schaperow Nuclear Regulatory Commission

Abdul Sheikh Nuclear Regulatory Commission

Randolph Sullivan Nuclear Regulatory Commission

Charles G. Tinkler Nuclear Regulatory Commission

Nathan E. Bixler Sandia National Laboratories

Shawn P. Burns Sandia National Laboratories

Randall O. Gauntt Sandia National Laboratories

Joseph A. Jones Sandia National Laboratories

Raymond J. Jun Sandia National Laboratories

Douglas M. Osborn Sandia National Laboratories

Jesse Phillips Sandia National Laboratories

Kyle Ross Sandia National Laboratories

Mark T. Leonard dycoda, LLC

Kenneth C. Wagner dycoda, LLC

xxxiii

ACRONYMS

AC

AFW

ANS

ASME

Alternating Current

Auxiliary Feedwater

American Nuclear Society

American Society of Mechanical Engineers

CCI Core Concrete Interactions

CDF Core Damage Frequency

CST

DC

DOE

Condensate Storage Tank

Direct Current

Department of Energy

EAL

EAS

Emergency Action Levels

Emergency Alert System

ECST

EDG

EOF

EOP

EP

Emergency Condensate Storage Tank

Emergency Diesel Generator

Emergency Operations Facility

Emergency Operating Procedure

Emergency Preparedness

EPA

EPZ

ERO

Environmental Protection Agency

Emergency Planning Zone

Emergency Response Organization

ETE Evacuation Time Estimate

GE

HPS

IPE

IPEE

ISLOCA

LERF

General Emergency

Health Physics Society

Individual Plant Examination

Individual Plant Examination of External Event

Interfacing System Loss of Coolant Accident

Large Early Release Frequency

LLNL

LHSI

LNT

Los Alamos National Lab

Low Head Safety Injection

Linear, No-Threshold

LOCA Loss Of Coolant Accident

LOSW Loss Of Service Water 

LTSBO

MACCS2

MCCI

Long-Term Station Blackout

MELCOR Accident Consequence Code System, Version 2

Molten Corium-Concrete Interactions

MSIV

MSVH

Main Steam Isolation Valve

Main Steam Valve House

NPP

NRC

NRCS

NRF

Nuclear Power Plant

Nuclear Regulatory Commission

Natural Resources Conservation Service

National Response Framework

OREMS Oak Ridge Evacuation Modeling System

ORO Offsite Response Organization
xxxiv

PRT Pressurizer Relief Tank

RCP Reactor Coolant Pump

RCS

RHR

RPV

RVLIS

RWST

Reactor Coolant System

Residual Heat Removal

Reactor Pressure Vessel

Reactor Vessel Level Indication System

Refueling Water Storage Tank

SAE

SAMG

Site Area Emergency

Severe Accident Management Guidelines

SBO

SECPOP

Station Blackout

SECtor POPulation and Economic Estimator

SG Steam Generator

SGTR

SNL

Steam Generator Tube Rupture

Sandia National Laboratories

SOARCA State-of-the-Art Reactor Consequence Analysis Project

SORV Stuck Open Relief Valve

SPAR

SPRA

Standardized Plant Analysis Risk

Seismic Probabilistic Risk Assessment

SRV Safety Relief Valve

STSBO Short-Term Station Blackout

TAF Top of Active Fuel

TDAFW

TSC

Turbine Driven Auxiliary Feedwater

Technical Support Center

UE Unusual Event
1-1

1. INTRODUCTION

This document describes the detailed severe accident analyses (i.e., MELCOR and the MELCOR 

Accident Consequence Code System, Version 2 (MACCS2) code calculations) performed for the 

Surry Power Station as part of the NRC’s State-of-the-Art Reactor Consequence Analyses 

(SOARCA) project. A separate volume of this report describes severe accident analyses for the 

Peach Bottom Atomic Power Station. A summary report describing the formal Peer Review 

Committee activities, comments, and resolutions was published as a separate document entitled 

“Summary Report: Peer Review of the State-of-the Art Reactor Consequence Analyses 

(SOARCA) Project,” (ML121250030).

1.1 Background

The evaluation of accident phenomena and offsite consequences of severe reactor accidents has 

been the subject of considerable research by NRC, the nuclear power industry, and the 

international nuclear energy research community. Most recently, with Commission guidance 

and as part of plant security assessments, updated analyses of severe accident progression and 

offsite consequences were completed using the wealth of accumulated research. These analyses 

are more detailed in terms of the fidelity of the representation and resolution of facilities and 

emergency response, realistic in terms of the use of currently accepted phenomenological models 

and procedures, and integrated in terms of the intimate coupling between accident progression

and offsite consequence models. 

An insight gained from these security assessments was that updated analyses of severe reactor 

accidents were needed to reflect realistic estimates of the more likely outcomes considering the 

current state of plant design and operation and the advances in our understanding of severe 

accident behavior. The SOARCA project evaluates plant improvements and changes (either of 

which can alter safety margins) not reflected in earlier assessments. These include system 

improvements, improvements in training and emergency procedures, offsite emergency response, 

and security-related improvements, as well as plant changes such as power uprates and higher 

core burnup. SOARCA’s more realistic modeling updates the more conservative quantifications 

of offsite consequences found in earlier NRC publications such as NUREG/CR-2239, “Technical 

Guidance for Siting Criteria Development” referred to in this report as the Siting Study.

In addition to the improvements in understanding and calculational capabilities that have resulted 

from these studies, numerous influential changes have occurred in the training of operating 

personnel and the increased use of plant-specific capabilities. These changes include:

• The transition from event-based to symptom-based Emergency Operating Procedures

(EOPs) for the pressurized-water reactor designs.

• The performance and maintenance of plant-specific probabilistic risk assessments (PRAs)

that cover the spectrum of accident scenarios.

• The implementation of plant-specific, full-scope control room simulators to train 

operators.
1-2

• An industry wide technical basis, owners group specific guidance and plant-specific 

implementation of the Severe Accident Management Guidelines (SAMGs).

• Use of additional safety enhancements, described in Title 10, Section 50.54(hh) of the 

Code of Federal Regulations (10CFR50.54(hh)). These enhancements are intended to be 

used to maintain or restore core cooling, containment, and spent fuel pool cooling 

capabilities under the circumstances associated with loss of large areas of the plant due to 

explosions or fire, to include strategies in the following areas:(i) Fire fighting;(ii) 

Operations to mitigate fuel damage; and (iii) Actions to minimize radiological release. 

For the SOARCA scenarios, successful implementation of this equipment and procedures 

would prevent core damage and/or delay or prevent the release.

• Improved phenomenological understanding of influential processes such as:

o in-vessel steam explosions 

o dominant chemical forms for fission products 

o direct containment heating 

o hot leg creep rupture 

o reactor pressure vessel failure, and 

o molten core concrete interactions.

1.2 Objective

The overall objective of the SOARCA project is to develop a body of knowledge regarding the 

realistic outcomes of severe reactor accidents. Corresponding and supporting objectives are as 

follows:

• Incorporate the significant plant improvements and updates not reflected in earlier 

assessments including system improvements, training and emergency procedures, offsite 

emergency response, and recent security-related enhancements described in Title 10, 

Section 50.54(hh) of the Code of Federal Regulations (10 CFR 50.54(hh)) as well as 

plant updates in the form of power uprates and higher core burnup.

• Incorporate state-of-the-art integrated modeling of severe accident behavior that includes 

the insights of some 25 years of research into severe accident phenomenology and 

radiation health effects.

• Evaluate the potential benefits of recent security related mitigation improvements in 

preventing core damage and reducing or delaying an offsite release should one occur.

• Enable NRC to communicate severe accident related aspects of nuclear safety to

stakeholders including Federal, State, and local authorities, licensees, and the public.

• Update quantification of offsite consequences found in earlier NRC publications such as 

NUREG/CR-2239, “Technical Guidance for Siting Criteria Development” [1].
1-3

1.3 Outline of the Report

Section 2 of this report briefly summarizes the method used to select the specific accident 

scenarios subjected to detailed computational analysis. Additional details of this method can be 

found in NUREG-1935. Section 3 describes the results of the accident scenario selection process 

when it was applied to Surry. Section 4 describes the key features of the MELCOR model of the 

Surry Power Station. Section 5 describes the results of MELCOR calculations for severe 

accident progression and radionuclide release to the environment from each accident scenario. 

Section 6 describes the process in which plant-specific emergency response actions were 

represented in the MELCOR Accident Consequence Code System, Version 2 (MACCS2)

calculations of offsite consequences. Section 7 describes the MACCS2 code site-specific 

parameters and the calculations of offsite consequences for each accident scenario. References 

cited in this report are listed in Section 8.

2-1

2. ACCIDENT SCENARIO DEVELOPMENT

The SOARCA project considered accident sequences that have an estimated frequency greater 

than 1x10-6 per reactor-year (pry) operation as candidate sequences for further deterministic 

evaluation. It also considered sequences with frequency as low as 1x10-7 pry if they were judged 

to proceed rapidly enough to have the potential for generating significant early releases of 

radionuclides to the environment or involve a radiological transport pathway from the reactor to 

the environment that bypasses the containment pressure boundary (i.e., bypass sequences). 

Section 2.1 and Section 2.2 summarize the methods used to identify these sequences and the 

screening process for retaining candidate sequences.

Once candidate accident sequences were identified, the analysts evaluated realistic opportunities

for plant personnel to respond to the observed failures of control and safety systems. The 

manner in which mitigation measures were evaluated for each accident sequence is described in 

Section 2.3. 

The end result of this process was a list of accident scenarios (i.e., event sequence plus options 

for mitigation), which were subjected to detailed analysis of plant response including, as 

appropriate, radionuclide release to the environment described in Sections 4 and 5 and offsite 

radiological consequence (see Sections 6 and 7).

2.1 Sequences Initiated by Internal Events

The following scenario selection process was used to determine the scenarios for further 

analyses:

1. Identify candidate accident sequences were identified in analyses using plant-specific

Standardized Plant Analysis Risk (SPAR) models (Version 3.31).

a. Initial Screening – Core damage sequences with low frequencies (less than 

1x10-8 pry) were eliminated from consideration. This step eliminated 4% of the 

overall core damage frequency (CDF) for Surry.

b. Sequence Evaluation – Dominant cutsets for the remaining sequences were 

reviewed to characterize system and equipment availabilities and accident 

sequence timing

c. Sequence Grouping – Sequences cutsets with similar equipment availabilities and 

estimated time for the onset of core damage were aggregated into a single 

‘sequence group’ or ‘scenario.’

2. The availability of containment systems was evaluated using system dependency tables. 

These tables delineate the support systems required for containment systems to function. 

The status of containment systems was then appended to the accident sequence 

description. 

3. Core damage sequences from the licensee’s PRA model were compared with the 

scenarios determined by using the SPAR models. Differences were resolved during 

meetings with licensee staff.
2-2

4. The screening criteria described above were applied to eliminate extremely low 

frequency sequences from further analyses.

This process identified no accident sequences that met the screening criterion of 1x10-6 pry. One

sequence group, however, met the screening criterion of 1x10-7 pry for events that have the 

potential to result in significant early releases to the environment, the spontaneous steam 

generator tube rupture (SGTR) with a frequency of 5x10-7 pry.

An additional sequence was retained for analysis, namely, an interfacing systems loss-of-coolant 

accident (ISLOCA). The ISLOCA scenario analyzed in SOARCA is a catastrophic failure of 

both of the inboard isolation check valve disks within the low-head safety-injection (LHSI) 

piping together with failure to refill the RWST or to cross-connect to the unaffected unit’s 

refueling water storage tank (RWST). For this ISLOCA scenario, the NRC’s SPAR model 

calculated a CDF of 3x10-8 pry, and the NRC’s initial understanding was that the licensee’s PRA 

calculated a CDF of 7x10-7 pry. SOARCA analyses included this scenario because the licensee’s 

PRA for Surry included an ISLOCA frequency of 7x10-7 pry and it has been commonly 

identified as an important contributor in PRA. As described below, based on additional 

discussions with the Surry licensee, it was clarified that the licensee’s latest PRA calculates a 

CDF of 3x10-8 /reactor year for the scenario modeled in SOARCA.

During Surry site visits on January 19, 2011, and October 26, 2011, NRC staff learned that the 

licensee’s current PRA model has the following two ISLOCA scenarios: 

• Scenario 1: Catastrophic failure of one check valve, leak-by of the second check valve, 

and the motor-operated isolation valve (MOV) not being able to close

• Scenario 2 (modeled in SOARCA): Catastrophic failure of two check valves

Scenario 1 would result in a leak between 50 - 300 gallons per minute (gpm) from the RCS. 

Anything less than 50 gpm would be mitigated by a relief valve on the low-pressure side of the 

LHSI injection line; pipe rupture would not occur. The frequency of the catastrophic failure of 

one check valve and the leak-by of the second check valve is 1x10-6 pry. When compounded by 

all the potential failure modes (e.g., operator error and mechanical or electrical failures) of the 

MOV, that lowers the frequency of Scenario 1 to 7x10-7 pry. This frequency does not include 

any consideration of averting core damage by refilling or cross-connecting RWSTs. This is a 

significant conservatism.

 

Scenario 2 would result in a leak above 300 gpm from the reactor coolant system (RCS). The 

licensee’s current PRA model assumes that the probability for the catastrophic failure of both 

isolation check valves is approximately 3x10-8 pry. As with Scenario 1, this frequency does not 

include consideration of averting core damage by refilling or cross-connecting RWSTs.

Scenario 2 does not meet the SOARCA screening criterion of 1x10-7 pry for a bypass event. 

However, we elected to retain it as part of SOARCA because it has been commonly identified as 

an important contributor in PRA.
2-3

This process provides the basic characteristics of each scenario. However, it is necessary when 

calculating a consistent intergrated response to have more detailed information about the 

scenario than is provided in a PRA model. To capture the additional sequence details, the 

SOARCA project conducted further analysis of system descriptions and a review of the normal 

and emergency operation procedures.

2.2 Sequences Initiated by External Events

External events include internal flooding and fire; seismic events; extreme wind, tornado, and 

hurricane related events; and other similar events that may be applicable to a specific site. The 

external event scenarios developed for SOARCA analysis were derived from a review of past 

studies such as NUREG-1150 [2], individual plant examination for external event (IPEEE)

submittals, and other relevant generic information.

Seismic initiators are considered to be limiting for two principal reasons. First, seismic initiators 

are more likely to result in the near immediate failure of systems, whereas fire and flood would 

be expected to result in delayed failures. Secondly, a seismic event may be more likely to fail 

passive components such as water tanks where fire and flood are not. Finally, seismic initiators 

may be more likely to have site-wide impacts. As a result, plant and offsite response to external 

event sequences were assumed to be represented by an earthquake of sufficiently large 

magnitude to result in wide-spread damage to important plant support systems such as electric 

power sources. 

The sequence selection process identified two sequence groups that met the screening criteria of 

1x10-6 pry for containment failure events and one event that met the screening criteria of 

1x10-7 pry for events that have the potential to result in significant early releases to the 

environment:

• long-term station blackout (LTSBO) – 1x10-5 to 2x10-5 pry

• short-term station blackout (STSBO) – 1x10-6 to 2x10-6 pry

• STSBO with thermally induced steam generator tube rupture – 1x10-7 to 8x10-7 pry. This

is a bypass event, which has a screening criterion of 1x10-7 pry. 

2.3 Mitigative Measures 

The site-specific mitigation measures assessments were performed during visits to the Surry 

Power Station site in June 2007 and were supplemented by follow-up telephone conferences and 

correspondence with the licensee later in 2007. The licensee senior reactor operators, PRA 

analysts, and other licensee staff were provided the initial conditions and subsequent failures for 

each of the sequence groups being analyzed. The operator and plant response was subsequently 

evaluated to develop timelines for operator actions and equipment lineup or setup times for the 

implementation of the available mitigation measures. These timelines were developed 

assuming minimum staffing. A result of these assessment and reviews boundary conditions were 
2-4

used to develop the MELCOR boundary conditions that included operator actions and 

applicable mitigation measures.

Mitigation measures considered in the SOARCA analyses include the licensee’s EOPs, SAMGs, 

and mitigation measures and strategies incorporated into plant capabilities in response to the 

terrorist events of September 11, 2001 and codified in 10 CFR 50.54(hh). 

2.3.1 Mitigation of Sequences Initiated by Internal Events

The mitigation measures assessment for internal events also included mitigation measures 

codified in 10 CFR 50.54(hh), but these measures were subsequently shown to be redundant to 

the wide variety of equipment and indications available for mitigating them. The identified 

internal events involve few equipment failures and are controlled by postulated operator errors.

2.3.2 Mitigation of Sequences Initiated by External Events

It was noted earlier that the initiating events for external event sequences were assumed to be 

seismic events, because they are considered to be limiting. For these sequence groups, the 

seismic PRA provided information on the initial availability of installed systems. Next, 

judgments were made concerning the general state of the plant to assess the availability of the

mitigation measures codified in 10 CFR 50.54(hh) and the additional time to 

implement mitigation measures and activate emergency response centers (e.g., Technical 

Support Center (TSC) and Emergency Operations Facility (EOF)).

The seismic events considered in SOARCA result in loss of offsite and onsite AC power (i.e., 

LTSBO) and, for the more severe seismic events, loss of DC power (i.e., STSBO). Under these 

conditions, the use of the turbine driven auxiliary feedwater (TDAFW) system is an important 

mitigation measure. Diverse procedures have been developed for PWRs, including a procedure 

to start and operate the TDAFW system without DC control power, which facilitates a managed 

response to station blackout (SBO) conditions. These procedures were discussed during site 

visits. This is known as TDAFW blackstart. Under 10 CFR 50.54(hh), mitigation measures also 

include the long-term operation of the TDAFW system without electricity (i.e., TDAFW 

blackrun), using a portable generator to supply indications such as steam generator level, to 

allow the operator to manually adjust TDAFW flow to prevent steam generator overfill and 

flooding of the TDAFW turbine. For a LTSBO, TDAFW can be used to cool the core until 

battery exhaustion. After battery exhaustion, TDAFW blackrun can be used to continue to cool 

the core.

The external events PRA does not describe general plant damage and accessibility following a 

seismic event. NUREG/CR-4334 was consulted to assess the potential viability of safety-related 

piping after a 1.0g peak ground acceleration (pga) event [40]. For the short-term station blackout 

(i.e., 0.5-1.0g pga) the damage was assumed to be sufficiently widespread such that accessibility 

would be difficult. The TDAFW system was judged not initially available and was judged not 

recovered under these circumstances prior to fuel damage (i.e., fuel damage in 3 hours) due to 

immediate gross rupture of the emergency condensate storage tank (ESCT). However, 

extrapolating results from NUREG/CR-4334, the low-head safety water injection and 

containment spray safety-related piping were judged to remain intact. Other studies, including a 
2-5

German study that physically simulated ground motion equal to 1.0g pga on an existing plant, 

also supported this evaluation. 

In the less severe long-term station blackout (i.e., 0.3 – 0.5g pga), the TDAFW system was 

available initially and the low-head safety injection and safety-related containment spray piping 

were also judged to remain intact. The integrity of the safety-grade piping provided a connection 

point for a portable, diesel-driven pump to inject water into the RCS or into the containment 

spray systems. 

The 10 CFR 50.54(hh) mitigation measures include the application of portable equipment such 

as portable power supplies for the instrumentation, portable diesel-driven pumps, and portable air 

bottles to open air-operated valves. Applicable procedures have been written to implement these 

mitigative measures under severe accident conditions. The portable injection equipment and the 

site fire truck were observed stored onsite in a structure away from the containment. A 

walk-down of the storage building and pathway to the plant suggested that the operators would 

be able to retrieve the equipment following a seismic event.

The time estimates to implement individual mitigation measures were provided by the licensee 

staff for each sequence group based on the sequence descriptions provided by the NRC. The 

time estimates take into account the plant conditions following the seismic event. The time 

estimates reflect exercises run by licensee staff that provided actual times to move and connect 

the portable, diesel-driven pump. The time estimates for staffing the TSC and the EOF also were 

provided by licensee staff and reflect the possible effects of the seismic event on roads and 

bridges.

The mitigation measures assessment noted the possibility of bringing in equipment from offsite 

(e.g., fire trucks, pumps and power supplies from sister plants or from contractors), but it did not 

quantify the types, amounts, and timing of this equipment arriving and being implemented.

3-1

3. ACCIDENT SCENARIO DEFINITIONS

Sections 3.1 and 3.2 describe the long-term and short-term station blackouts scenarios, 

respectively, which were initiated by a seismic event. The spontaneous SGTR and ISLOCA

accident scenarios are described in Sections 3.3 and 3.4 respectively, which were internal events 

initiated by piping failures. Each Section describes the initiating event, the available systems, 

the pertinent mitigative actions, and the detailed initial and boundary conditions for the severe 

accident code calculations. Section 3.5 provides discussion on the Surry Seismic PRA Study 

and how it compares to the SOARCA project. 

3.1 Long-Term Station Blackout

The LTSBO is initiated by an earthquake (i.e., 0.3–0.5g peak ground acceleration - pga). It has 

an estimated frequency of 1x10-5 to 2x10-5 pry, which meets the SOARCA screening criterion of 

1x10-6 pry. 

Section 3.1.1 describes the initial status of the plant following the seismic event. The key system 

availabilities normally accessible during the course of the accident are summarized in 

Section 3.1.2. The pertinent mitigative measures available to address the accident progression 

are described in Section 3.1.3. Section 3.1.4 describes two scenarios that differ in the assumed 

success (or failure) of the mitigative actions. Mitigated scenarios are defined as those in which 

the mitigative actions are successful. Unmitigated scenarios are defined in which certain key 

mitigative measures are not successfully implemented.

For station blackout scenarios, boiling in the reactor coolant pump (RCP) seal could cause the 

spring-loaded part of the seal to pop open and stay open. As such, MELCOR modeling for Surry 

includes seal failure when conditions in the seal approach saturation. The hole size for this 

failure mode is that which produces a 180 gpm per pump flow rate at normal RCS temperature 

and pressure. Also, it has been hypothesized that seal failure could occur as early as 13 minutes

into a station blackout scenario due to the loss of seal cooling. Seal cooling requires AC power. 

The conditional probability of this early seal failure (i.e., as early as 13 minutes) has been 

estimated by the industry to be 0.2. Applying this 0.2 probability to the Surry LTSBO scenario 

frequency of 1x10-5 to 2x10-5 pry results in an event frequency of 2x10-6 to 4x10-6 pry, which 

meets the SOARCA screening criteria of 1x10-6 pry. While seal failure could occur as early as 

13 minutes into the scenario and could include seal failures in as many as all 3 RCPs, such early 

and multiple seal failures are less likely than single seal failures and would have a lower 

probability. However, to examine the potential range of system response, the project staff 

analyzed with MELCOR, LTSBO mitigated and unmitigated sensitivity cases assuming that the 

seals of all 3 RCP seals failed 13 minutes into the scenario.

3.1.1 Initiating Event

The seismic event results in the loss of offsite power and failure of onsite emergency AC power 

resulting in an SBO event where neither onsite nor offsite AC power are recoverable. All 

systems dependent on AC power are unavailable, including the containment systems (i.e., 

containment spray and fan coolers). It is assumed the Alternate AC diesel generator is 

unavailable. In the long term, the loss of the TDAFW pump may occur due to battery depletion, 
3-2

deletion of the water source and loss of DC power for sensing and control or deletion of the 

water source (i.e., the ECST). Nominal RCP leakage occurs due to the loss of pump seal cooling 

(i.e., initially 21 gpm per pump at normal operating pressure and temperature). The unmitigated 

and mitigated base cases include the potential for a later thermal-mechanical RCP seal failure. In 

addition, unmitigated and mitigated sensitivity cases are performed that include an early RCP 

seal failure (i.e., at 13 minutes). 

3.1.2 System Availabilities

The TDAFW pump is available until the ECST empties. The station batteries give 

instrumentation until they exhaust. Batteries typically last for approximately 2 to 8 hours under 

normal loading conditions depending on the life cycle of the batteries2. At the beginning of its 

life, the battery duration is 8 hours. At the end of its life, the battery duration is 2 hours. It was 

assumed that the battery life for a seismic-initiated LTSBO was 8 hours due to the minimum 

loading conditions caused by the initiating event and the minimum loading expected throughout 

the event due to the limited equipment available. The secondary system power operated relief 

valves (PORVs) are initially available for a manual 100oF/hr system cooldown. The secondary

system PORVs are assumed to close following battery failure. No other systems are available.

3.1.3 Operator Actions and Mitigative Measures

The LTSBO event results in the loss of offsite and onsite AC power. Under these conditions, the 

TDAFW pump is an important mitigation measure. The TDAFW pump is used to cool the core 

until battery exhaustion. After battery exhaustion, blackrun of the TDAFW pump is used to 

remove heat from the primary system. 

The external events PRA does not describe general plant damage and accessibility following a 

seismic event. The damage was assumed to be widespread and accessibility to be difficult, 

consistent with the unavailability of many plant systems. It is assumed the Alternate AC diesel 

generator is unavailable. The ECST initially supplies the TDAFW pump but has finite resources 

(i.e., empty in 5 hours). However, assuming successful mitigation, it was assessed that the 

operators would have adequate time, access, and resources to make up water for injection.

The low-head safety injection and safety-related containment spray piping were judged not likely 

to fail for this scenario. The integrity of this piping provided a connection point for a portable, 

diesel driven pump to inject into the RCS. Licensee staff estimated that transporting the pump 

and connecting it to plant piping takes about two hours. Hence, the availability of the vessel 

injection was assessed to occur at 3.5 hours, or 2 hours after the action was recommended by the 

operators and support staff. Companion unmitigated analyses were also performed to quantify 

the response without successful mitigation by portable pumps.

 2 The Surry DC station batteries are required to last for 2 hours. Initially, the licensee estimated a best-estimate life 

of 8 hours. Following completion of the analysis, 6 hours was thought to be more realistic. However, the ECST 

ran out of water at 5 hours and stopped the TDAFW pump. Consequently, the most significant benefit of the 

station batteries failed at 5 hours, which was less than the best-estimate battery life. 
3-3

In summary, the following actions are credited in the mitigated scenario calculations:

• Provide vessel injection using a portable, high-pressure, diesel driven (Kerr) pump 

through three drain lines on the low head safety injection (LHSI) piping

• Use portable air bottles to operate the steam generator power operated relief valves, 

which allows for depressurization and cooldown of the RCS

• A portable power supply is used to restore SG and RCS level indication

• Manual operation of the TDAFW pump without DC power is credited

• A portable, diesel driven, low-pressure (Godwin) pump is used to supply water to the fire 

header. The firewater can then be supplied via fire hose to the AFW pumps.

While not used in the mitigated scenario calculations, the following additional mitigative 

measures were identified as additional options for consideration.

• Use firewater or pumper truck for cooloing the charging pump oil cooler

• Use an alternative power source for high-head safety injection pump RCS makeup.

3.1.4 Scenario Boundary Conditions

Section 3.1.4.1 lists the sequence of events in the mitigated LTSBO calculation. Section 3.1.4.2

summarizes the sequence of events in the unmitigated LTSBO calculation. Mitigated and 

unmitigated sensitivity cases were also performed that include an early failure of the RCP seals.

3.1.4.1 Mitigated Cases

There is one mitigated base case and one mitigated sensitivity case. The mitigated sensitivity 

case includes an early failure of the RCP seal on all three pumps. The boundary conditions for 

the two cases are listed below. 

Mitigated Case (using portable mitigation equipment)

Event Initiation

• Loss of offsite power followed by a station blackout

• The reactor trips and the MSIVs close

• The DC buses are available, at minimum loading, currently being used for 

instrumentation, PORV operation, and TDAFW operation

• The TDAFW system auto-initiates providing make-up to the steam generators

• The TDAFW system takes suction water from the ECST
3-4

• RCP seal leakage begins at 21 gpm per pump. The RCP seals may subsequently fail due 

to high temperatures causing a leak of 182 gpm per pump (see Section 4.1 for a 

description of the failure model).

• Emergency core cooling systems (ECCSs) are inoperable

• Containment cooling systems are inoperable

• Containment is isolated

• Recovery of offsite power is not expected during the mission time

15 minutes

• Initial Operations assessment of plant status complete, initiate the following actions:

o Attempt manual start of EDGs

o Operation of SG PORVs available for 30 minutes using a dedicated internal 

battery for RCS pressure control

• Station batteries are available

• Steam Generator (SG) level being maintained by AFW, RCS is cooling down, no RCS 

makeup currently available

1 hour

• Manual start of EDGs assumed to be unsuccessful due to initiating event

• The TSC is manned and operational. The primary function of the TSC would be to 

review initiating event, plant status, and operator action to provide guidance on 

alternative mitigation measures.

1.5 hours 

• The offsite EOF is manned. The primary function of the EOF would be to review 

initiating event, plant status, and operator action to provide guidance on alternative 

mitigation measures. The TSC staff members are the primary users of SAMGs and 

mitigation measures codified in 10 CFR 50.54(hh). Shift supervisors and TSC 

supervisors are trained on these procedures.

• Operations initiate a controlled depressurization of the SGs to approximately 120 psi to 

achieve an RCS cooldown of < 100o

F per hour by manually opening the SG PORVs

• TSC and EOF review actions taken by Operations and determine the availability of the 

portable, diesel-driven pumps and the station pumper truck stored outside the protected 

area. Recommend the following actions:

o Connect the portable, high-pressure, diesel-driven (Kerr) pump to three drain lines 

of the LHSI piping for RCS makeup and use portable air bottles for manual 

operation of primary system PORVs, as needed
3-5

o Hook up portable power supply to power instrumentation

o Use the 2 firewater storage tanks (250,000 gallons per tank), the portable, 

low-pressure, diesel-driven (Godwin) pump, and the fire pumper truck to supply 

water via the firewater system for AFW suction, if necessary 

o Setup to provide the firewater system or a portable, low-pressure, diesel-driven 

(Godwin) pump to the containment spray header in preparation for containment 

cooling 

1.75 hours 

• Operations assesses and concurs with TSC and EOF recommendations. Operations 

prioritizes recommendation based on plant conditions and begins implementation.

3.5 hours 

• The Kerr pump provides emergency 65 gpm makeup flow to the RCS3,4

• A portable power supply provides power to the instrumentation

• TDAFW pump maintaining SG level

• Pre-staging and lineups are ongoing for other mitigation measures:

o Setup to provide the firewater system or a portable, low-pressure, diesel-driven 

(Godwin) pump to the firewater system in preparation for containment cooling

o Use the 2 firewater storage tanks (250,000 gallons per tank), the portable, 

low-pressure, diesel-driven (Godwin) pump, and the fire pumper truck to supply 

water via the firewater system for AFW suction, if necessary 

Mitigated Case (using portable mitigation equipment) + early RCP seal failure

Identical sequence of events as the unmitigated base case but includes an early RCP seal failure 

at 13 minutes.

13 minutes

• All three RCP seals fail and leak at a nominal rate of 182 gpm per pump

 3

 An effective flowrate of 65 gpm at 500 rpm was determined using pump data provided by the Kerr Pump 

Corporation that includes 95% and 92% mechanical and volumetric efficiencies. 4 Per SOARCA request in August 2010, Surry sent a timeline to connect the Kerr pump. The timeline showed 

30 minutes to recognize the event and perform the initial damage assessment per procedures. At 30 minutes, 

Operations would begin deploying the Kerr pump and be ready for injection at 150 minutes from the initiating 

event. It was recognized that post-seismic conditions could complicate the process. The start of injection at 

3.5 hours versus 150 min includes a 1-hour conservatism.
3-6

3.1.4.2 Unmitigated Cases

There is one unmitigated base case and one unmitigated sensitivity case. The unmitigated 

sensitivity case includes early failures of the RCP seals. The boundary conditions for the two 

cases are listed below. 

Unmitigated Case (without portable mitigation equipment)

Event Initiation

• Loss of offsite power followed by a station blackout

• The reactor trips and the main steam isolation valves (MSIVs) close

• The DC buses are available, at minimum loading, currently being used for 

instrumentation, PORV operation, and TDAFW system operation

• The TDAFW system auto-initiates providing make-up to the SGs

• The TDAFW system takes suction water from the ECST

• RCP seal leakage begins at 21 gpm per pump. The RCP seals may subsequently fail due 

to high temperatures causing a leak of 182 gpm per pump (see Section 4.1 for a 

description of the failure model).

• ECCSs are inoperable

• Containment cooling systems are inoperable

• Containment is isolated

• Recovery of offsite power is not expected during the mission time

15 minutes

• Initial Operations assessment of plant status complete, initiate the following actions:

o Attempt manual start of EDGs

o Operation of SG PORVs available for 30 minutes using a dedicated internal 

battery for RCS pressure control

• Station batteries are available

• The SG level is being maintained by AFW, RCS is cooling down, no RCS makeup 

currently available

1 hour

• Manual start of EDGs assumed to be unsuccessful due to initiating event
3-7

• The TSC is manned and operational. The primary function of the TSC would be to 

review the initiating event, plant status, and operator action to provide guidance on 

alternative mitigation measures.

1.5 hours 

• The EOF is manned. The primary function of the EOF would be to review initiating 

event, plant status, and operator action to provide guidance on alternative mitigation 

measures. The TSC staff members are the primary users of SAMGs and mitigation 

measures codified in 10 CFR 50.54(hh). Shift supervisors and TSC supervisors are 

trained on these procedures.

• Operations initiate a controlled depressurization of the SGs to approximately 120 psi to 

achieve an RCS cooldown of <100oF per hour by manually opening the SG PORVs

• The TSC and EOF review actions taken by Operations and determine the availability of 

the portable, diesel-driven pumps and the station pumper truck stored outside the 

protected area. Recommend the following actions:

o Connect the portable, high-pressure, diesel-driven (Kerr) pump to three drain lines 

of the LHSI piping for RCS makeup and use portable air bottles for manual 

operation of primary system PORVs, as needed.

o Hook up portable power supply to power instrumentation

o Use the 2 firewater storage tanks (250,000 gallons per tank), the portable, 

low-pressure, diesel-driven (Godwin) pump, and the fire pumper truck to supply 

water via the firewater system for AFW suction, if necessary 

o Setup to provide the firewater system or a portable, low-pressure, diesel-driven 

(Godwin) pump to the containment spray header in preparation for containment 

cooling 

1.75 hours 

• Operations assesses and concurs with TSC and EOF recommendations. Operations 

prioritizes recommendations based on plant conditions.

>1.75 hours 

• All mitigative actions are unsuccessful including connecting a portable, diesel-driven 

pump for vessel injection, refilling the water supply for the TDAFW (i.e., the ECST), and 

maintaining instrumentation using a portable power supply

8 hours 

• DC station batteries are exhausted5

• SG PORVs reclose

• Loss of control and instrumentation for the TDAFW

 5

 The Surry DC station batteries are required to last for 2 hours. Initially, the licensee estimated a best-estimate life 

of 8 hours. Following completion of the analysis, 6 hours was thought to be more realistic. However, the ECST 

ran out of water at 5 hours and stopped the TDAFW pump. Consequently, the most significant benefit of the 

station batteries failed at 5 hours, which was less than the best-estimate battery life.
3-8

Unmitigated Case (without portable mitigation equipment) + early RCP seal failure

Identical sequence of events as the unmitigated base case but includes an early RCP seal failure 

at 13 minutes.

13 minutes

• All three RCP seals fail and leak at a nominal rate of 182 gpm per pump

3.2 Short-Term Station Blackout

The STSBO is initiated by an earthquake (0.5–1.0g pga). It is more severe than the LTSBO and 

has an estimated frequency of 1x10-6 to 2x10-6 pry, which meets the SOARCA screening 

criterion of 1x10-6 pry. 

Section 3.2.1 describes the initial status of the plant following the seismic event. The key system 

availabilities during the course of the accident are summarized in Section 3.2.2. The pertinent 

mitigative measures available to address the accident progression are described in Section 3.2.3. 

Section 3.2.4 describes two scenarios that differ in the assumed success (or failure) of the 

mitigative actions. Mitigated scenarios are defined as those in which mitigative actions are 

successful. Unmitigated scenarios are defined as those in which certain key mitigation measures 

are not successfully implemented. In addition, mitigated and unmitigated scenarios are defined 

that include a thermally-induced steam generator tube rupture (TI-SGTR).

3.2.1 Initiating Event

The seismic event results in a loss-of onsite power and failure of onsite emergency AC power 

resulting in a SBO event where neither onsite nor offsite AC power are recoverable. All systems 

dependent on AC power are unavailable, including all active ECCSs and the containment 

engineered safety systems (e.g., the containment sprays and fan coolers). The seismic event also 

causes a loss of DC power, which makes it impossible to remotely control the TDAFW pump. 

The RCS and containment are undamaged and the containment is isolated. No instrumentation is 

available. Significant structural damage is judged to have occurred, including structural failure 

of the turbine building and loss of access to the condenser blow down valves. Additionally, the 

seismic event has caused a failure of the ECST, the source of water for the TDAFW system. 

Auxiliary building accessibility is difficult, due to fallen piping and cabling, steam and water 

leaks, and damaged stairways. Following the loss of the seal cooling flow, the RCP seals will 

nominally leak at 21 gpm (i.e., at normal operating pressure and temperature). The RCP seals 

may fail later in the accident if the RCP seal region heats to saturated conditions. 

Both unmitigated and mitigated sensitivity cases are performed that include a TI-SGTR(s). 

Thermally-induced steam generator tube ruptures are a known risk contributor and have been 

investigated by industry and the NRC. The SBO has an estimated frequency of 1x10-6 to 2x10-6

pry, and the conditional probability of tube rupture have been estimated by the NRC to be in the 

range of 0.1 to 0.4 [30]. Therefore, the overall frequency of this sequence group is 1x10-7 to 

8x10-7 pry, which meets the SOARCA screening criterion of 1x10-7 pry for bypass events. In the 

context of the short-term station blackout sequence evaluations, sensitivity studies are performed 

to examine the response with a TI-SGTR.
3-9

3.2.2 System Availabilities

Systems available include the ECCS accumulators, portable power supplies, portable air bottles, 

and portable high-pressure (Kerr) and low-pressure (Godwin) diesel driven pumps. Containment 

spray and firewater piping is assumed to remain intact.

3.2.3 Mitigative Actions

The TDAFW system was assumed not to be available initially and was judged not recovered 

under these circumstances prior to fuel damage (i.e., fuel damage in 3 hours) due to immediate 

gross rupture of the ECST and lack of alternative suction sources that could be aligned within 

3 hours. However, there was significant time, access, and resources to establish containment 

sprays with the portable emergency pump by 8 hours. This action both mitigates the release and 

delays containment failure.

NUREG/CR-4334 was consulted to assess the potential viability of safety-related piping after a 

1.0g pga event [40]. Extrapolating results from NUREG/CR-4334, the low-head safety water 

injection and safety-related containment spray piping were judged to remain intact. Other 

studies, including a German study that physically simulated ground motion equal to 1.0g pga on 

an existing plant, also supported this evaluation. The integrity of the safety-grade piping 

provided a connection point for a portable, diesel-driven pump to inject water into the RCS or 

into the containment spray systems. The licensee staff estimated that transporting the pump and 

connecting it to plant piping takes about two hours. Because of difficult accessibility, the set-up 

of the containment spray system following a large seismic event was assumed to require 8 hours. 

Hence, water provided to the diesel-driven portable emergency pump was assumed to be 

available only after vessel failure (i.e., the MELCOR results indicate 3 hours to core damage and 

7 hours to lower head failure). Thus, it was provided only to the containment sprays. Additional

unmitigated analyses were performed to quantify the response without successful mitigation by a 

portable pump.

3.2.4 Scenario Boundary Conditions

Section 3.2.4.1 lists the sequence of events in the unmitigated STSBO calculation. 

Section 3.2.4.2 summarizes the sequence of events in the mitigated STSBO calculation that 

credits one additional manual action. Sensitivity cases for the mitigated and unmitigated TISGTRs are also described.

3.2.4.1 Unmitigated Cases

There is one unmitigated base case and two unmitigated sensitivity cases. The unmitigated 

sensitivity cases include thermally induced steam generator tube ruptures prior to creep rupture 

in any other RCS location. Since the sensitivity cases include a stuck open secondary system 

relief valve, there is an open containment bypass pathway to release radionuclides to the 

environment. In the base case, the secondary system relief valve closes when the pressure falls 

below the closing setpoint. Additionally, operators are assumed to not connect the portable, 

diesel driven (Godwin) pump in all three cases. 
3-10

Unmitigated base case

Event Initiation

• Loss of offsite power followed by the failure of all diesel generators and a station 

blackout

• Successful reactor trip and Main Steam isolation Valves (MSIVs) close

• RCS and containment are undamaged and the containment isolates

• Failure of TDAFW system due to failure of the ECST

• An early RCP seal failure during subcooled conditions is not included in this scenario, 

but late RCP seal failures may occur (see Section 4.1 for a discussion on RCP seal 

failure)

• Active ECCS equipment is inoperable due to electrical and system damage 

• Containment cooling systems are inoperable due to electrical and system damage

• Recovery of offsite and onsite power is not expected during the mission time

30 minutes

• Initial Operations assessment of plant status complete; Operations initiates the following 

action:

o Attempt manual start of the EDGs and SBO diesel generator

• RCS pressure being maintained by code safety valves, PORVs not currently available 

because of loss of instrument air and backup air

1 hour

• Manual start of EDGs and SBO diesel generator assumed to be unsuccessful due to 

initiating event

• Offsite EOF is manned and operational. The primary function of the EOF is review of

initiating event, plant status, and operator action to provide guidance on alternative 

mitigation measures. The TSC staff members are the primary users of SAMGs and 

mitigation measures codified in 10 CFR 50.54(hh). Shift supervisors and TSC 

supervisors are trained on these procedures.
3-11

1.5 hours 

• Offsite EOF reviews actions taken by operations. Recommend the following actions:

o Connect portable power supply for instrumentation 

o Connect the portable, high-pressure, diesel-driven (Kerr) pump to three drain lines of 

the LHSI piping for RCS makeup and use portable bottles for manual operation of SG 

PORVs, as needed

o Connect the portable, diesel-driven (Godwin) pump for containment spray or 

containment flooding 

1.75 hours

• Operations assesses and concurs with offsite EOF recommendations. Operations 

prioritizes recommendations based on plant conditions and begins implementation.

2 hours

• The TSC is manned and operational. Because of the magnitude of the seismic event, a 

one-hour delay in reporting of TSC members was assumed. The primary function of the 

TSC would be to review the initiating event, plant status, and operator action to provide 

guidance on alternative mitigation measures.

3.5 hours

• Determine the availability of the portable, diesel-driven (Godwin) pump, portable air 

bottles, and portable power supply 

• Portable air bottles connected to the steam generator PORVs are available for 

depressurizing RCS

• The portable diesel-driven pumps are being positioned and the connections are being 

assessed

>6.5 hours

• Unable to connect portable injection systems

• No other mitigation attempts are successful 

Unmitigated sensitivity cases with TI-SGTRs

The unmitigated sensitivity cases have an identical sequence of events as the unmitigated base 

case but include a stuck open relief valve on the secondary side that leads to a TI-SGTR.

3 hours

• The lowest-pressure safety relief valve sticks open on the steam generator with the tube 

rupture
3-12

At a time calculated by MELCOR to be 3 hr 33 min

• A TI-SGTR is assumed to occur when the hot leg C cumulative creep damage index 

exceeds 5%. A hot leg creep rupture index of 100% is used to predict hot leg failure. 

Consequently, the TI-SGTR is specified to occur prior to hot leg creep rupture failure.

o Sensitivity Case 1 – rupture area is the equivalent of 100% of the tube flow area

o Sensitivity Case 2 – rupture area is the equivalent of 200% of the tube flow area

3.2.4.2 Mitigated Cases

There is a mitigated base case and a mitigated sensitivity case, both involving the use of a diesel 

driven (Godwin) pump for containment sprays. The mitigated sensitivity case includes a 

thermally induced steam generator tube rupture prior to any other RCS creep rupture failure. 

Since the sensitivity case includes a stuck open secondary system relief valve, there is an open 

containment bypass pathway to release radionuclides to the environment. In the base case, the 

secondary system relief valve closes when the pressure falls below the closing setpoint.

Mitigated base case

Event Initiation

• Loss of offsite power followed by a station blackout

• Successful reactor trip and MSIVs close

• RCS and containment undamaged and the containment isolated

• Failure of TDAFW pump due to failure of the ECST

• An early RCP seal failure during subcooled conditions is not included in this scenario, 

but late RCP seal failures may occur (see Section 4.1 for a discussion on RCP seal 

failure)

• ECCSs are inoperable due to electrical and system damage 

• Containment cooling systems are inoperable due to electrical and system damage

• Recovery of offsite power is not expected during the mission time

30 minutes

• Initial Operations assessment of plant status complete; Operations initiates the following 

action:

o Attempt manual start of EDGs and SBO diesel generator

• The RCS pressure is maintained by code safety valves. The pressurizer PORVs are not 

currently available because of loss of instrument air and backup air.
3-13

1 hour

• Use portable power supply to restore minimum instrumentation (e.g., RCS level, RCS 

pressure, SG level)

• Manual start of EDGs and SBO diesel generator assumed to be unsuccessful due to 

initiating event

• The EOF is manned. The primary function of the EOF would be to review initiating 

event, plant status, and operator action to provide guidance on alternative mitigation 

measures. The TSC staff members are the primary users of SAMGs and mitigation 

measures codified in 10 CFR 50.54(hh). Shift supervisors and TSC supervisors are 

trained on these procedures.

1.5 hours 

• Offsite EOF reviews actions taken by operations. Recommend the following actions:

o Maintain portable power supply for instrumentation 

o Connect the portable, high-pressure, diesel-driven (Kerr) pump to three drain lines of 

the LHSI piping and use portable bottles for manual operation of SG PORVs, as 

needed

o Connect the portable, diesel-driven (Godwin) pump for containment spray or 

containment flooding 

1.75 hours

• Operations assesses and concurs with offsite EOF recommendations. Operations 

prioritizes recommendations based on plant conditions and begins implementation.

2 hours

• The TSC is manned and operational. Because of the magnitude of the seismic event, a 

one-hour delay in reporting of TSC members was assumed. The primary function of the 

TSC would be to review the initiating event, plant status, and operator action to provide 

guidance on alternative mitigation measures.

• EOF is manned and operational

3.5 hours

• Determined the availability of the remotely located portable, diesel-driven (Godwin) 

pump, portable air bottles, and portable power supply

• Portable air bottles to the steam generator PORVs are available for depressurizing RCS

• The portable diesel-driven pumps are being positioned and the connections are being 

assessed.
3-14

8 hours

• The portable, diesel-driven (Godwin) pump is staged at the discharge canal and pumps 

water through the established piping and into the fire protection system. Fire hose would 

take the water from the hydrants to the special fitting on the containment spray pumps.

6

Mitigated case with TI-SGTR

The mitigated sensitivity case has an identical sequence of events as the mitigated base case but 

includes a stuck open relief valve on the secondary side that leads to a TI-SGTR.

3 hours

• The lowest-pressure safety relief valve sticks open on the steam generator with the tube 

rupture.

At a time to be calculated by MELCOR (which was 3 hr 33 min)

• A TI-SGTR is assumed to occur when the hot leg C cumulative creep damage index 

exceeds 5%. A hot leg creep rupture index of 100% is used to predict hot leg failure. 

Consequently, the TI-SGTR is specified to occur prior to hot leg creep rupture failure.

• The steam generator tube rupture area is the equivalent of 100% of the tube flow area

3.3 Spontaneous SGTR

Section 3.3.1 describes the initial status of the plant following the tube rupture. The key system 

availabilities during the course of the accident are summarized in Section 3.3.2. The pertinent 

mitigative measures available to address the accident progression are described in Section 3.3.3. 

Section 3.3.4 delineates various scenarios based on the success of the mitigative actions. In 

particular, a mitigated scenario is defined where the mitigative actions are successful. Two 

unmitigated scenarios are also defined where certain key operator actions are not successfully 

performed.

3.3.1 Initiating Event

This sequence group consists of a spontaneous rupture of a steam generator tube equivalent to 

100% of the tube flow area, 0.47 in2

. The leak is assumed to occur near the steam generator 

inlet-side tube sheet. 

3.3.2 System Availabilities

The full complement of systems is considered functional in this scenario including all systems 

associated with engineered safeguards and instrumentation and control as well as all auxiliary 

 6 Per SOARCA request in August 2010, Surry sent a timeline to connect the Godwin pump. The timeline showed 

30 minutes to recognize the event and perform the initial damage assessment per procedures. At 30 minutes, 

operations would begin deploying the Godwin pump and be ready for injection at 120 minutes from the initiating 

event. It was recognized that post-seismic conditions could complicate the process. The start of injection at 

8 hours versus 2 hours includes a 6-hour conservatism.
3-15

and emergency systems. The operators fail to 1) isolate the faulted SG, 2) depressurize and 

cooldown the RCS, and 3) refill the RWST or cross-connect to the unaffected unit’s RWST. 

3.3.3 Mitigative Actions

The SPAR model and the licensee’s PRA concluded that the spontaneous SGTR event proceeds 

to core damage because of the above errors. However, the PRA models do not appear to have 

credited the significant time available for the operators to correct their mistakes. They also do 

not appear to credit technical assistance from the TSC and the EOF. The subsequent accident 

simulation showed that 27 to 46 hours are available for mitigative actions before the core 

damage begins (see Section 5.4). Therefore, the licensee provided realistic estimates of the times 

by which the operators would respond to the event. These time estimates included consideration 

of indications that the operators would have of the bypass accident, operator training on plant 

procedures for dealing with bypass accidents and related drills, and assistance from the TSC and 

EOF, which were estimated to be manned and operational by 1 to 1.5 hours into the event.

Operator actions in this scenario are essentially those expected per training and procedure. 

Specifically, the operators are trained to perform the following actions to mitigate the sequence:

• Secure AFW delivery to the steam generator with the broken tube (the faulted steam 

generator) 

• Secure 1 of the 3 total high head safety injection (HHSI) pumps

• Isolate the faulted steam generator, i.e., close the MSIVs serving the faulted steam 

generator

• Secure the remaining HHSI pumps once the faulted generator is isolated, which will end 

the RCS leakage

• Perform a 100°F/hr cool-down of the RCS 

• Establish long-term cooling with residual heat removal

The following other mitigation measures were identified but not used.

• Use the pressurizer PORVs to depressurize the RCS to get an accumulator injection at 

low pressure

• Cross-connect to the unaffected unit’s RWST 

• Use firewater makeup to RWST from the firewater header at ~300 gpm from the two 

300,000 gallons firewater storage tanks, then the James River

• The portable, low-pressure, diesel-driven (Godwin) pump is available to makeup to the 

RWST and the CST at ~2000 gpm at 120 psi
3-16

• ~190,000 gallons available from the spent fuel pool for rapid RWST makeup

• Procedures exist to align firewater to the suction of the AFW pump via installed piping 

and valves from firewater storage tanks and the James River

• Two portable, high-pressure, diesel-driven (Kerr) pumps are available to inject into RCS 

using water from the RWST at 2.5 hours (i.e., assumes guidance from TSC and EOF at 

1.5 hours and an hour to implement)

3.3.4 Boundary Conditions

Section 3.3.4.2 lists the sequence of events to be prescribed in the mitigated spontaneous steam 

generator tube rupture where the operator successfully performs the actions described in 

Section 3.3.3. Section 3.4.4.1 summarizes the sequence of events in two unmitigated scenarios 

where the operator does not successfully perform the actions described in Section 3.3.3. The 

second unmitigated scenario uses the same failed operator actions but also includes the failure of 

the steam generator secondary system relief valve to create a sustained containment bypass 

pathway for fission products.

3.3.4.1 Unmitigated Cases

There are two unmitigated cases. No other successful operator actions are credited after 

2.5 hours. 

Unmitigated Case 1

2.5 hours 

• Fail to isolate the faulted steam generator

• Fail to depressurize and cool down the RCS

• Fail to extend the available duration of ECCS injection by refilling the RWST or cross

connecting to the other unit’s RWST

Unmitigated Case 2

Exactly the same boundary conditions as Unmitigated Case One but include an additional 

equipment failure.

At a time to be calculated by the severe accident analysis code which was 44 minutes

• Fail the secondary system relief valve open when water first reaches the valve. The 

stuck-open valve creates an open bypass containment pathway to the environment (see 

Note below).

2.5 hours 

• Fail to isolate the faulted steam generator
3-17

• Fail to depressurize and cool down the RCS

• Fail to extend the available duration of ECCS injection by refilling the RWST or cross 

connecting to the other unit’s RWST

Note: There was some uncertainty whether water could reach the secondary system relief valve. 

The utility stated that the secondary system would not fill up completely due to the large 

volume of piping and 12 steam traps (i.e., eight 1.5” lines and four 1” lines) open to the 

main condenser. The MELCOR model did not represent the steam traps or steam dump 

valves to the condenser. The calculation conservatively neglected any leakage pathways 

for water from the steam line except for the cycling relief valve. 

3.3.4.2 Mitigated Case

There is one mitigated base case. Although the operator initially fails to implement the correct 

procedures, the errors are eventually identified by the technical support groups and the correct 

procedures are followed. The boundary conditions are listed below.

Event Initiation

• Spontaneous tube rupture equivalent to 100% of the tube flow area (i.e., 0.47 in2)

• The reactor trips

• The turbine stop valves automatically close

• The 8 steam dump valves automatically go to the full open position and then throttle open 

and close to maintain RCS Tave at 547 °F

• Containment Phase 1 isolation auto-initiates

• The HHSI auto-initiates and all three pumps start and operate as designed. The operator 

secures one charging pump early in the event as required by procedure. The water source

is the RWST (380,000 gallons).

• The one turbine-driven and two motor-driven auxiliary feedwater pumps automatically 

start on a low-level actuation signal. The initial water source is the ECST 

(110,000 gallons) but can be refilled from the CST, which has 300,000 gallons.

• Reactor coolant pumps continue to run

• Operators fail to: 1) isolate the faulted SG, 2) depressurize and cooldown the RCS, and 

3) refill the RWST or cross-tie to the unaffected unit’s RWST.

10 minutes

• Initial Operations assessment of plant status complete
3-18

15 minutes

• RCS level being maintained by HHSI, operator secures one of the three HHSI pumps per 

procedure

• Operator takes control of AFW to maintain level in the SGs

• When level in the faulted SG reaches the top of fill range, AFW flow will be stopped to 

that SG

30 minutes

• Damaged SG continues to fill, overflowing into the TDAFW pump turbine causing it to 

shut down

• The two MDAFW pumps provide makeup to non-faulted SGs

1 hour

• The TSC is manned and operational. The primary function of the TSC would be to 

review initiating event, plant status, and operator action to provide guidance on 

alternative mitigation measures.

1.5 hours

• RCS and SG levels being maintained by HHSI and AFW, respectively

• Offsite EOF is manned. The primary function of the EOF would be to review initiating 

event, plant status, and operator action to provide guidance on alternative mitigation 

measures. The TSC staff members are the primary users of SAMGs and mitigation 

measures codified in 10 CFR 50.54(hh). Shift supervisors and TSC supervisors are 

trained on these procedures.

• The TSC and EOF recognize that the damaged SG is not isolated, the operators are not 

implementing procedure E-3, “Steam Generator Tube Rupture” and the RCS is not being 

cooled down and depressurized. Recommends to the operators that they implement the 

following actions:

o Implement procedure E-3

o Isolate the damaged S/G

o Cooldown and depressurize the RCS

1.75 hours 

• Operations assesses TSC and EOF diagnoses, concurs with their determination, and 

implements procedure E-3, 

2.5 hours 

• Within 45 minutes the damaged SG is isolated, HHSI is secured, and the RCS is 

undergoing a normal cooldown
3-19

Event Termination 

• Establish long-term cooling using the residual heat removal (RHR) system (i.e., 

closed-circuit cooling system)

o RCS at 400-450 psi and ~350  ̊F for RHR entry conditions

o Operators verify RCS is 30  ̊F sub-cooled, pressure stabilized, pressurizer level in 

normal band and stabilized, and non-affected SG levels in normal band and stabilized

3.4 Interfacing Systems LOCA

This sequence group is initiated by a common mode failure of both low-head safety injection

(LHSI) inboard isolation check valve disks. The open pathway pressurizes and ruptures the 

low-pressure piping outside the containment, which opens a containment bypass LOCA. This 

sequence group consists of the bypass LOCA followed by operator failure to refill the RWST, or 

cross-connect to the unaffected unit’s RWST. 

Section 3.4.1 describes the initial status of the plant following the pipe rupture. The key system 

availabilities during the course of the accident are summarized in Section 3.4.2. The pertinent 

mitigative measures available to address the accident progression are described in Section 3.4.3. 

Section 3.4.4 describes two scenarios that differ in the assumed success (or failure) of the 

mitigative actions. Mitigated scenarios are defined as those in which mitigative actions are 

successful. Unmitigated scenarios are defined as those in which certain key mitigation measures 

are not successfully implemented.

3.4.1 Initiating Event

The ISLOCA initiates with failure of both of the inboard isolation check valve disks resulting in 

over-pressurization and failure of the LHSI discharge side piping outside of containment in the 

Safeguards Building. The resulting double-ended guillotine pipe break permits back-flow of the 

high-pressure RCS water into the Safeguards Building.7

 Water will also spill into the Safeguards 

Building via forward flow through the LHSI pumps to the pipe break. The broken LHSI line has 

a number of flow restrictions, including a 2.57” venturi between the RCS and the break that will 

limit the break flow. 

3.4.2 System Availabilities

The full complement of systems is considered functional in this scenario including all systems 

associated with engineered safeguards and instrumentation and control as well as all auxiliary 

and emergency systems. 

3.4.3 Mitigative Actions

The SPAR model and the licensee’s PRA concluded that the ISLOCA proceeds to core damage. 

However, the PRA models do not appear to have credited the significant time available for the 

 7 A double-ended break was specified in the scenario description. A detailed separate effects model of the LHSI 

piping was developed that included flow restrictions, bends, elevation changes and other losses. The break in the 

low-pressure piping was expected to occur between the end of the high-pressure piping and a third check valve. 

If the break was under water, it could offer some fission product scrubbing benefit. The results of the break 

location analyses are reported in Section 4.12.
3-20

operators to respond adequately. The PRA model also does not appear to credit technical

assistance from the TSC and EOF. The more realistic analysis of thermal hydraulics in 

Section 5.5.3 subsequently estimated 6 hours until the RWST is empty and 13 hours until the 

fission product releases begin, providing considerable time for the operators to respond. The 

ISLOCA time estimates are based on a double-ended pipe rupture, which drains the RWST at the 

maximum rate.

Operator actions in this scenario are essentially those expected per training and procedure. 

Specifically, the operators are trained to perform the following actions to mitigate the sequence:

1. Per the Surry EOPs, only two HHSI pumps are required. All HHSI pumps will start but 

one is secured. Per results of the table-top scenario development with Surry operations 

using the emergency operating procedures, the redundant HHSI pump would be isolated 

in 15 minutes.

2. Per the Surry EOPs, a LOCA outside the containment would be identified and the LHSI 

pumps would be isolated.8

3. Per the Surry EOPs, the operators would isolate the leakage from the RWST side of the 

break into the Safeguards Building (i.e., by closing the LHSI pump suction valves). 9

4. Per results of the table-top scenario development with Surry operations using the EOPs, 

the operators will take control of the AFW pumps to maintain normal level in the steam 

generators after 15 minutes.

5. Per results of the table-top scenario development with Surry operations, the operators will 

shift HHSI injection from the cold leg to the hot leg to minimize backflow leakage to the 

Safeguards Building by 1 hour and 45 minutes. An additional HHSI pump can be 

secured if an adequate water level can be maintained to minimize the spill rate into the 

Safeguard Building.

6. Per results of the table-top scenario development with Surry operations, the operators will 

start a cooldown at 1 hour. In order to minimize break flow, the operators would 

completely depressurize the steam generators.

The following other mitigation measures were identified but not used in the unmitigated 

ISLOCA calculations. 

• Per the Surry EOPs, Operations would establish residual heat removal (RHR) cooling of 

the RCS once entry conditions are established.

 8

 Per best-estimate timing for the ISLOCA in the Surry full-scope training simulator following EOPs, LHSI Pump 

A would be isolated in 6 minutes and 17 seconds and the LHSI Pump B would be isolated at 15 minutes and 44 

seconds. Surry Operations obtained these timings from a training exercise that simulated the initial portion of the 

subject ISLOCA. 9

 Per best-estimate timing for the ISLOCA in the Surry full-scope training simulator following EOPs, the LHSI 

pump suction would be isolated in 16 minutes and 18 seconds. Surry Operations obtained these timings from a 

training exercise that simulated the initial portion of the subject ISLOCA.
3-21

• The RWST can be refilled using firewater makeup from the firewater header at ~300 gpm 

from two 250,000 gallon firewater storage tanks, then from the James River.

• The portable, low-pressure, diesel-driven (Godwin) pump is available to makeup to the 

RWST and the CST at ~1200 gpm 

• 190,000 gallons available from the spent fuel pool are available for rapid RWST makeup

• Operations could align the unaffected unit’s HHSI pumps and RWST to the affected unit 

through a series of operator actions.10

3.4.4 Boundary Conditions

Section 3.4.4.1 lists the sequence of events to be prescribed in the unmitigated ISLOCA, which 

credits all operator actions identified in Section 3.3.3, but does not credit any of the additional 

mitigative actions identified at the end of the section (e.g., RHR). Section 3.4.4.2 summarizes 

the sequence of events in the mitigated ISLOCA, which credits additional operator actions. 

3.4.4.1 Unmitigated Interfacing Systems LOCA 

Event Initiation

• The LHSI inboard isolation check valves fail causing a pipe break in the low pressure 

piping in the Safeguards Building

• The reactor trips on low pressure

• Containment Phase-1 isolation auto-initiates

• All three HHSI pumps auto-initiate on the ECCS injection signal.

• LHSI initiates on the ECCS injection signal, which pumps water into the Safeguards 

Building through the pipe break until the LHSI pump motors become submerged or are 

isolated

• The MSIVs close

• The RCPs trip or are shutdown once two-phase conditions develop at the pump

• The one turbine-driven and two motor-driven auxiliary feedwater pumps automatically 

start on a low-level actuation signal. The initial water source is the ECST 

(110,000 gallons) but can be refilled from the CST, which has 300,000 gallons. 

 10 The shift to the unaffected unit’s HHSI and RWST is not in the normal emergency procedures but is well known 

to the operators and is a redundant design feature of the Surry ECCS. Since it would temporarily affect the other 

unit’s resources, it is a 10CFR50.54 decision (i.e., a directive that the operators can go outside of their procedures 

if necessary to ensure the safety of the plant).
3-22

3 minutes

• The LHSI outboard isolation valve (i.e., Valve 1890C) is flooded and becomes 

inaccessible. This flooding is shown in the MELCOR results in Section 5.5. Therefore, 

the ISLOCA cannot be isolated from the RCS.

5 minutes

• Initial Operations assessment of plant status is complete. A LOCA outside the 

containment identified.

6 minutes 18 sec

• LHSI Pump A is isolated

15 minutes

• 3 HHSI pumps confirmed as running, one pump is isolated 

15 minutes and 44 seconds

• LHSI Pump B is isolated

16 minutes and 18 seconds

• Operations isolate the LHSI pump suctions. The action ends the RWST spillage to 

Safeguards Building. 

45 minutes

• Operations transfers HHSI injection to the RCS hot legs

50 minutes

• The TSC is manned. Primary function would be to review initiating event, the plant 

status, and the operator action to provide guidance on alternative mitigation measures. 

The TSC staff is the primary users of SAMGs and extreme damage mitigation guidelines 

(EDMGs). Shift supervisors and TSC supervisors are trained on SAMGs and EDMGs.

• The EOF is manned. Primary function would be to review the initiating event, the plant 

status, and the operator action to provide guidance on alternative mitigation measures. 

1 hour

• Operators begin RCS cooldown

1.25 hours

• The TSC is operational

1.5 hours 

• The EOF is operational

• The TSC and EOF review and concur with actions taken by operations. They

recommend:
3-23

o Using RHR to terminate the accident

o Refilling the RWST

o Reducing the RWST draindown by terminating another HHSI if level can be 

maintained above the top of active fuel 

o Shifting to the unaffected unit’s RWST, if necessary, to maintain injection.

1.75 hours 

• Secure second HHSI pump

>1.75 hours 

• Operations do not successfully implement any further mitigative actions. All injection 

will terminate when the RWST empties.

3.4.4.2 Mitigated Interfacing Systems LOCA 

The mitigated case has an identical sequence of events until 1.75 hours, the time assessed to 

implement the TSC and EOF recommendations. The operator successfully initiates the 

following actions, starting at 1.75 hours.

1.75 hours 

• Operations assesses TSC and EOF recommendation to lineup the unaffected unit’s 

RWST to provide makeup to the RWST while continuing to provide RCS makeup with 

the same RWST.

• Operations reviews and prepares to swap HHSI flow to unaffected unit’s RWST, if 

necessary. At 1.75 hours, approximately 160,000 gallons are still available. 

Note: The HHSI pumps could trip off line if the Auxiliary Building is allowed to flood to 

approximately 5 feet above the Auxiliary Building basement floor without 

mitigation measures. The volume of the Auxiliary Building basement that will 

result in flooding of the HHSI pumps is 530,000 gallons. Another mitigation 

option is to use portable submersible pumps to pump out the Auxiliary Building 

basement to preclude flooding of the HHSI pumps. This option is recognized by 

the licensee but is not included in plant procedures.

• Operations recommends establishing long-term cooling using the RHR system per 

emergency operating procedures (i.e., closed-circuit cooling system)

• Secure 3rd HHSI when level is stable and RHR cooling is established

3.5 Surry Seismic PRA Study

Late in the SOARCA project, well after the process of scenario selection and accident analysis 

was completed, the team became aware of a new relevant seismic PRA (SPRA) study that was 

conducted for the Surry plant. This study, entitled Surry SPRA Pilot Plant Review, was 

sponsored by the Electric Power Research Institute (EPRI), Simpson Gumpertz and Heger 

(SGH) and Dominion Resources. 
3-24

The objectives of the study were:

• To evaluate the process and requirements involved in updating a seismic PRA originally 

developed for the IPEEE program in order to meet the intent of the new ASME/ANS 

Level 1 Seismic PRA Standard for Capability Category II [56], and 

• To review the requirements of the seismic PRA standard for clarity and reasonableness 

given the current state of the art in performing seismic PRAs.

The Surry SPRA study started with the Surry seismic assessment developed for the IPEEE and 

updated various elements of that seismic assessment in order to meet the author’s interpretation 

of the requirements in the new ASME/ANS Seismic PRA Standard. While the Surry SPRA

study represents an updated examination of the Surry plant, including a new seismic hazard and 

calculation of the site response to the new hazard, the primary purpose of the Surry SPRA study 

was to examine issues attendant to implementation of the new ASME/ANS Standard. Surry 

served as a representative plant for the SPRA study. In some instances not essential to the 

objective of the Surry SPRA study, the evaluation relied upon generic or expected fragility 

parameters that may be different from the Surry-specific values. Thus, the CDF and the large 

early release frequency (LERF) results presented in the Surry SPRA study may not be reflective 

of the Surry plant. Nevertheless, certain results and assumptions in this SPRA study may have a 

bearing on the SOARCA project.

The Surry SPRA study produced a total CDF, which is comparable to the CDF of the SOARCA 

external events on the order of 2x10-5

/yr, and the risk profile peaks at a seismic interval 

corresponding to approximately 0.4g pga. However, the dominant scenario, comprising 50% of 

the total CDF, was identified as a loss of service water (LOSW). Such an event was not 

analyzed in SOARCA and thus some discussion of its significance relative to SOARCA events is 

needed. 

The LOSW event in the Surry SPRA study occurs as a result of failure of the turbine building’s

steel superstructure that is assumed to damage the cables that power the circulating water 

isolation valves leaving the valves in the open position. With the valves in the open position, the

intake canal, the source of service water, gravity drains because the seismic event causing a loss 

of offsite power which causes failure of the circulating water pumps which supply the intake 

canal. The intake canal is estimated to drain in a time interval ranging from 45 minutes to 

6 hours, depending on the number of circulating water isolation valves assumed to fail open. 

The intake canal is the ultimate heat sink for the plant. Thus, its loss results in a loss of the 

component cooling water system, which cools high-pressure injection pumps and reactor coolant 

pump seals. The timing of core damage is not specified in the Surry SPRA study other than to 

note that recovery of offsite power is not expected within the 24-hour mission time. In addition, 

the Surry SPRA study assumes the ECST and the fire protection water tanks are not available 

due to their low capability to withstand seismic loading. If indeed these tanks are assumed to fail 

catastrophically this would result in immediate loss of auxiliary feedwater (AFW) unless other 

sources of water for the AFW system can be aligned. 
3-25

In terms of frequency, the LOSW described in the Surry SPRA study most closely matches the 

SOARCA scenario described as a long-term SBO (i.e., 2x10-5/yr). In certain respects, the 

LTSBO scenario may qualitatively serve as a surrogate for the LOSW event. In both instances,

the plant is undergoing a loss of heat removal transient with potential reactor coolant pump seal 

leakage. If the intake canal does not drain for several hours then the availability of service water 

together with injection capability (e.g., feed and bleed) may delay core damage in much the same 

way that AFW delays core damage in the LTSBO scenario. In the Surry SPRA study, the EDGs 

would be expected to be functional for some time for the less severe seismic events. Thus, core 

cooling could be achieved by primary system depressurization via relief valves coupled with 

low-pressure injection, which does not require component cooling water since it has its own 

radiator cooling. The RWST is available for some of the dominant (i.e., 0.4g pga) seismic 

events. Further, if the ECST and fire protection water tanks do not fail catastrophically, AFW 

may be available for some period as is the case in the LTSBO scenario. In addition, an 

alternative water source exists in the emergency condensate makeup tank together with AC 

powered emergency condensate booster pumps. Further, the emergency condensate makeup 

tank NPSH without the booster pumps may provide some capability. Finally, the B.5.b Godwin 

pump may be used to provide feedwater to the steam generators via the firewater/feedwater 

connection using the discharge canal as the water source. Thus, for some number of the LOSW 

permutations the timing of core damage would be comparable to or longer than that of the 

LTSBO scenario. It is clear that the worst-case LOSW scenario as identified in the Surry SPRA 

study combined with the loss of the ECST would still be bounded by the unmitigated short-term

SBO scenario consequences, since that scenario credits no primary side injection or auxiliary 

feedwater.

4-1

4. MELCOR MODEL OF THE SURRY PLANT

The Surry MELCOR model applied in this report was originally generated at Idaho National 

Engineering Laboratories (INEL) in 1988 [8]. The model was updated by Sandia National 

Laboratories (1990 to present) for the purposes of testing new models, advancing the 

state-of-the-art in modeling of PWR accident progression, and providing support to 

decision-makers at the NRC for analyses of various issues that may affect operational safety. 

Significant changes were made during the last twenty years in the approach to modeling core 

behavior and core melt progression, as well as the nodalization and treatment of coolant flow 

within the RCS and reactor vessel. Detailed reports have been prepared to discuss this model 

evolution as part of the MELCOR code development program [10], and these discussions will 

not be repeated here. It is simply noted that the model described herein is a culmination of these 

efforts and represents the state-of-the-art in modeling of potential PWR severe accidents. 

In preparation for the SOARCA analyses described in this report, the model was further refined 

and expanded in two areas. The first area is an upgrade to MELCOR Version 1.8.6 core 

modeling. These enhancements include:

• A hemispherical lower head model that replaces the flat bottom-cylindrical lower head 

model,

• New models for the core former and shroud structures that are fully integrated into the 

material degradation modeling, including separate modeling of debris in the bypass 

region between the core barrel and the core shroud,

• Models for simulating the formation of molten pools both in the lower plenum and the

core region, crust formation, convection in molten pools, stratification of molten pools 

into metallic and oxide layers, and partitioning of radionuclides between stratified molten 

pools,

• A reflood quench model that separately tracks the component quench front, quench 

temperature, and unquenched temperatures,

• A control rod silver aerosol release model, and 

• An application of the CORSOR-Booth release model for modern high-burn-up fuel.

The second area focused on the addition of user-specified models to represent a wide spectrum 

of plant design features and safety systems to broaden the capabilities of MELCOR to a wider 

range of severe accident sequences. These enhancements included:

• Update of the containment leakage/failure model (see Section 4.7),

• Update of core degradation modeling practices,
4-2

• Modeling of individual primary and secondary system relief valves with failure logic for 

rated and degraded conditions,

• Update of the containment flooding characteristics,

• Heat loss from the reactor to the containment,

• Separate motor and turbine-driven auxiliary feedwater models with control logic for plant 

automatic and operator cooldown responses,

• New turbine-driven auxiliary feedwater models for steam flow, flooding failure, and 

performance degradation at low pressure,

• Nitrogen discharge model for accumulators, 

• Update of the fission product inventory, the axial and radial peaking factors, and an 

extensive fission product tracking control system, and

• Improvements to the natural circulation in the hot leg and steam generator and the 

potential for creep rupture (see Section 4.4).

Table 4-1 provides a brief summary of plant design parameters that are helpful in comparing the 

configuration of Surry to other reactors of interest.

The model description is subdivided into description of the vessel and reactor coolant system 

(Section 4.1), primary and secondary system relief valve modeling (Section 4.2), the decay heat 

power modeling (Section 4.3), the natural circulation modeling (Section 4.4), the core 

degradation modeling (Section 4.5), the containment model (Section 4.6), the containment 

leakage model (Section 4.7), and the auxiliary building model (Section 4.8). 

Section 4.9 summarizes the best modeling practices applied to accident progression analyses 

conducted under the SOARCA project. The best practices include discussions of the base case 

approach to modeling key phenomena that have significant importance to the progression of the 

accident and uncertainty in their response. The Safeguards Area, Contaiment Spray Pump Area, 

and Main Steam Valve House are described in Section 4.10. The Safeguards ventilation system 

is described in Section 4.11, and the low head safety injection piping is described in 

Section 4.12. Section 4.13 describes the radionuclide deposition model for the low head safety 

injection piping, and Section 4.14 describes the methodology used for the two MELCOR models 

involving the low head safey injection piping.
4-3

Table 4-1 Important Design Parameters for Surry

Parameter Value

(SI units)

Value

(British units)

Rated Core Power [MWth] 2,546

Number of Fuel Assemblies in Core 157

Rod Array 15 x 15

Fuel Rods per Assembly 204

Fuel (UO2) Mass [kg / lb] 79,650 175,600

Zircaloy Mass in Fuel Cladding [kg / lb] 16,465 36,300

RPV Inner Diameter [m / ft] 1.994 6.542

RPV Height and Closure [m / ft] 12.319 40.417

Pressurizer Relief Valves [kg/s / lbm/hr] 2 x 26.46 2 x 210,000

Pressurizer Safety Valves [kg/s / lbm/hr] 3 x 36.96 3 x 293,330

Pressurizer Relief Tank Volume [m3 / ft3] 36.8 1300

Pressurizer Relief Tank Liquid Volume [m3 / ft3] 25.5 900

Pressurizer Relief Tank Design Pressure [bar / psig] 6.89 100

Reactor Inlet / Outlet Temperature [oC/oF] 282 / 319 540 / 606

RCS Coolant Flow [kg/s / lbm/hr] 12,738 101.1x106

Nominal RCS Pressure [MPa /psia] 15.5 2,250

Number of Steam Generators 3

Steam Generator Recirculation Rate 3.4

Steam Generator Heat Transfer Area [m2 / ft2] 4,785 51,500

Secondary Pressure [MPa /psia] 15.5 2,250

Secondary Side Water Mass [kg / lbm] 41,640 91,800

Secondary Side Volume [m3 / ft3] 166 5,868

Emergency Condensate Storage Tank Water Volume 

(ISLOCA / Other Scenarios) [L / gal]

416,395 /

363,400

110,000* /

96,000**

Refueling Water Storage Tank Water Volume [L / gal] 1,511,893 399,400

Turbine-driven Aux. Feedwater Pump [m3/s / gpm] 1 x 0.442 @ 

832 m

1 x 700 @

2,730 ft

Motor-driven Aux. Feedwater Pump [m3/s / gpm] 2 x 0.221 @ 

832 m

2 x 350 @

2,730 ft

Containment Design Pressure [MPa / psig] 0.31 45

Containment Volume [m3 / ft3] 50,970 1,800,000

Containment Operating Pressure [bar, psia] 0.62 to 0.71 9 to 10.3

Containment Operating Temperature [oC/oF] 24 to 52 75 to 125

Accumulator Water Volume [m3/ft3] 3 x 27.6 3 x 975

Accumulator Pressure [bar / psig] 4.14 to 4.59 600 to 665

High Head Safety Injection [m3/s / gpm] 3 x 0.0095 @ 

1,768 m

3 x 150 @ 

5,800 ft

Low Head Safety Injection [m3/s / gpm] 2 x 0.189 @

69 m

2 x 3,000 @ 

225 ft

* best estimate for ISLOCA scenario

** minimum amount required by technical specifications
4-4

4.1 Vessel and Reactor Coolant System

Figure 4-1 illustrates the configuration of the hydrodynamic model for the Surry RCS. The 

model includes explicit representation of the entire reactor coolant system including each of the 

three reactor coolant loops ̧ steam generators, and reactor coolant pumps, the steam lines out to 

the isolation valves and associated safety and power-operated relief valves. On Loop C, the 

pressurizer and associated safety and power-operated relief valves, and the pressurizer relief tank 

are modeled. Boundary conditions are used to represent the turbine pressure and feedwater flow 

to allow direct calculation of the nominal, full-power steady state operating conditions.

Logic models with mass and energy sources and sinks model the accumulators, the ECCSs, the 

main feedwater, and the motor and turbine-driven auxiliary feedwater. Separate logic models are 

used to represent the plant control systems such as the reactor scram logic, the emergency core 

cooling signal, the main feedwater control and trip logic, the turbine control valve isolation logic, 

reactor trip logic, reactor pump trip and failure logic, the containment spray actuation, the 

containment recirculation spray and the residual heat removal, the containment fan cooler 

actuation, and the plant station batteries.

Following a loss of seal cooling, water will leak through the pump seals. Under high 

temperature and two-phase degraded accident conditions, the pumps seals could fail and create a 

large leak. For each pump, three flow paths model the pump seal leakage. These leak paths 

describe chronic leaks from the RCS pump seals that are estimated to leak at 21 gpm at full 

reactor pressure [19]. The leakage model is also set up to mimic the seal failures in the pump 

using guidance from the utility’s probabilistic pump seal leakage model [20]. For example, the 

failure of the second stage seals was modeled to occur coincidently with loss of liquid 

subcooling in the RCP pump (i.e., voiding of the RCP). The model is set up to include the 

following leak rates for each of the three loops:

• 21 gpm nominal leakage at 15.5 MPa with failure of the seal cooling system 

(i.e., no AC power)

• 182 gpm at 15.5 MPa (failure of the #1 and #2 seal following change to saturated 

conditions in pump)11

• 480 gpm at 15.5 MPa (i.e., blowout of the seal internals with flow being 

controlled by the Labyrinth seal upstream of the seal package)

MELCOR’s choked flow model will predict the change in seal leakage flowrate as a function of 

pressure, quality, and liquid and gas temperature.

 11 Upon failure of the #1 seal, the #2 seal is also expected to immediately fail [Dominion, ET-CME-05-0020].
4-5

Figure 4-2 shows a detailed illustration of the reactor vessel hydrodynamic nodalization and the 

corresponding spatial divisions of the core. The core is represented by five concentric rings of 

hydrodynamic control volumes and core structures (i.e., fuel assemblies, control rods, and the 

supporting steel internal structures). Each ring is divided into five vertically stacked 

hydrodynamic control volumes. The axial length of the fuel in each ring is represented by ten 

axial cells in each ring. The outer ring (i.e., Ring 5) in the active fuel region is further 

subdivided into two regions. The inside region of Ring 5 models the peripheral assemblies of the 

core. The outer region of Ring 5 models the bypass region between the core shroud around the 

fuel and the core barrel. The detailed nodalization was required to simulate evolving, 

two-dimensional natural circulation flow as the core level dropped and a more accurate and 

continuous representation of the fuel power profile and subsequent degradation. 

As shown on the left-hand side of Figure 4-2 a 6-ring by 7-axial level nodalization is used in the 

lower plenum, offering a detailed radial spatial representation of the bottom of the vessel and 

associated structures. Ring 6, which is not included in the active fuel region, represents the outer 

radial region beneath the vessel downcomer. Separate axial levels represent the core plate, the 

flow mixer, and the lower core plate. Between the core supporting structures are the support 

columns, which transmit the load within the core to the lower support plate. The vessel lower 

head is subdivided into 10-radial by 6-azimuthal segments for a two-dimensional conduction 

solution. The lower head failure is evaluated using a one-dimensional mechanical response 

model that determines the stresses and strains in the lower head to predict creep-rupture failure. 

The lower head structural creep (i.e., plastic strain) failure is calculated using the default 

Larson-Miller lifetime damage model.

A matrix of axial and radial flow paths simulates two-dimensional flow patterns in the core 

region. Each flow path in the core and lower plenum nodalization simulates the effects of flow 

blockages and changes in resistance during core degradation. Ring 5 also uses special flow paths 

to represent the hydraulic openings following the failure of the core shroud if such failure is 

predicted.

The five ring radial hydrodynamic nodalization from the core extends upward into the upper 

plenum of the vessel. The upper plenum is divided into two axial levels with radial flow 

between each ring. Each ring also includes a representation of the guide tubes. Gas or water can 

flow through the control rod guide tubes between the upper plenum and the upper head. In the 

outer radial ring, there are three axial levels to separate the natural circulation flow outward to 

the hot legs (CV-154) versus the returning flow (CV-153). The leakage pathways between the 

downcomer and the upper plenum and from the downcomer to the upper head are also 

represented. The upper plenum to downcomer leakage path is an important fission product gas 

pathway in the ISLOCA when the reactor coolant system loop seals are filled and residual vessel 

water blocks reverse gas flow from the core to the downcomer.

As indicated by the different colored regions in Figure 4-3, the core was subdivided into five

regions. Each core region or “ring” (i.e., the terminology used in MELCOR) models the 

response of the included fuel assemblies. Figure 4-3 also shows the relative power of the fuel 

assemblies in each ring. The radial power profile in the center of the core is relatively flat. 
4-6

However, the peripheral region has a sharp decrease in the assembly powers. The inner four 

rings were defined to provide some resolution in the power profiles and are similarly sized to the 

outer ring (i.e., important for the thermal response). The 5-ring nodalization balances the 

objectives of representing of the radial power variations in the core versus excessive complexity 

for computational efficiency. Once core degradation begins, the 5-ring nodalization provides a 

good representation of the regional fuel collapses and flow blockages. 

The steam generator nodalizations are shown in Figure 4-4 and Figure 4-5. The red flowpaths

are only active in natural circulation conditions. Both the hot legs and the steam generator tubes 

are split into two halves to permit counter-current natural circulation flows (see Section 4.4). 

The steam generator includes explicit modeling of the primary-side tubes, the steam generator 

inlet and outlet plenums, the secondary side of the steam generator, the steam lines, and the 

safety and power-operated relief valves. The hot leg and steam generator nodalization is 

somewhat complicated because it must simulate conditions ranging from (a) normal operating 

conditions, (b) single-phase liquid and two-phase accident conditions, and (c) single-phase gas 

natural circulation conditions. As will be discussed in Section 4.4, special flow paths are 

activated to simulate some of the natural circulation phenomena.

The model includes the heat loss from the reactor system to the containment. Each external 

structure of the vessel, the recirculation looping, the steam generators, and the steam lines 

transfer heat to the containment. These heat structures are coupled to the appropriate control 

volumes representing different regions of containment. The total heat loss to the containment at 

rated conditions is 0.08% (1.97 MW), (see Table 5.3-2 in reference [47]).
4-7

CV530 - PRT

Reactor Vessel

(see detailed 

drawings)

Steam Generator

 1-RC-E-1B

(see hot leg/steam 

generator Loop B detail)

Steam Generator

 1-RC-E-1A

(see hot leg/steam 

generator Loop A detail)

Steam Generator

 1-RC-E-1C

(see hot leg/steam 

generator Loop C detail)

CV220 CV235

CV101

CV153

CV201 CV / 222

CV202 CV / 221

CV401/CV422

CV402/CV421

CV437 CV440

237 CV 240 CV

435 CV 420 CV

500 CV

CV520

Pressurizer

FL241

FL242

FL243FL244

FL442

FL443 FL444

FL441

FL201/FL240

FL202/FL292

FL231

FL431

CF206

Accumulator

(Mass 

Source)

Accumulator

(Mass 

Source)

CF206

HHSI LHSI

CF222

CF218RWST Suction (Mass Source)

Containment Sump Suction (Mass Source)CF942

FL401/FL440

FL402/FL492

LHSI

HHSI

CF222

CF218

RWST Suction (Mass Source)

Containment Sump 

Suction (Mass Source)

CF942

Rupture 

Disk

FL540

PRT 

Cubicle

CV042

SRV-3

FL538

SRV-2FL537

PORV-1, FL530

PORV-2, FL531

SRV-1

FL535

FL505

FL510

CV301/CV322 CV302/CV321

FL301/

FL340

FL202/FL292

320 CV 335 CV

CV337

CV340

FL342FL331FL341

FL343

FL344

Accumulator

(Mass 

Source) CF206

HHSI

LHSI

CF218RWST Suction (Mass Source)

Containment Sump Suction (Mass Source)CF942

CF222

RC

Pump

RC

Pump

RC

Pump

RCP Seal Leakage to SG A Cubicle – CV020

FL901 (21 gpm)

FL911 (182 gpm)

FL921 (500 gpm)

FL903 (21 gpm)

FL913 (182 gpm)FL923 (500 gpm)

FL902 (21 gpm)

FL912 (182 gpm)

FL922 (500 gpm)

Aux. Bldg.

CV817

FL997 - Letdown Line Bypass

RCP Seal Leakage to SG C Cubicle – CV040

RCP Seal Leakage to SG B Cubicle – CV030

Pressurizer 

Surge Line

Loop B

Loop C

Loop A

FL283

270 CV 280 CV

CV281 

Turbine

MSIV

FL286

Steam TrapFL287

SRV-2/3/4/5, FL264

SRV-1, FL263

PORV, FL262

CV998

Environment

Flow RestrictorFL285

FL288 CV295

TDAFW Pump

FL483

480 CV 470 CV

CV281 

Turbine

MSIV

FL486

Steam TrapFL487

Flow RestrictorFL485

SRV-1, FL463

PORV, FL462

CV998

Environment

FL488 CV295

TDAFW Pump

SRV-2/3/4/5, FL464

380 CV 370 CV

CV281 

Turbine

MSIV

FL386

Steam TrapFL387

Flow RestrictorFL385

SRV-1, FL363

PORV, FL362

CV998

Environment

FL388 CV295

TDAFW Pump

SRV-2/3/4/5, FL364

FL383

FL893

Surge Line Creep Rupture Basement

CV005

Figure 4-1 Surry Reactor Coolant System Hydrodynamic Nodalization
4-8

Figure 4-2 Surry Reactor Vessel Core, Lower Plenum, and Upper Plenum and Steam 

Dome Hydrodynamic and COR Structure Nodalization

CV110

Lower 

Plenum

710 CV

720 CV

730 CV

740 CV

750 CV

701 CV

712 CV

722 CV

732 CV

742 CV

752 CV

714 CV

724 CV

734 CV

744 CV

754 CV

716 CV

726 CV

736 CV

746 CV

756 CV

718 CV

728 CV

738 CV

748 CV

758 CV

FL710 FL720 FL730 FL740 FL750

FL760 FL770 FL780 FL790

FL712 FL722 FL732 FL742 FL752

FL762 FL772 FL782 FL792

FL764 FL774 FL784 FL794

FL766 FL776 FL786 FL796

FL768 FL778 FL788 FL798

FL714 FL724 FL734 FL744 FL754

FL716 FL726 FL736 FL746 FL756

FL718 FL728 FL738 FL748 FL758

FL112 FL122 FL132 FL142 FL152 FL706

CV112 CV122 CV132 CV142 CV152

FL701

Bypass

FL707

Westinghouse 3-Loop PWR

CVH/FL Nodalization

FL705FL704

FL703

FL702

0.0 m

0.93924 m

1.76396 m

2.0503 m

2.3800 m

2.4181 m

3.01985 m

3.0643 m

3.4301 m

3.7958 m

4.1615 m4.5273 m

4.8930 m

5.2588 m

5.6245 m

5.9902 m

6.3560 m

6.7217 m

108 208 308 408 508

109 209 309 409 509

110 210 310 410 510

111 211 311 411 511

112 212 312 412 512

113 213 313 413 513

114 214 314 414 514

115 215 315 415 515

116 216 316 416 516

117 217 317 417 517

106 206 306 406 506

104 204 304 404 504

103 203 303 403 503

102 202 302 402 502 602

101 201 301 401 501

Instrument

Thimble Tubes(MELCOR NS),

HLST =

1.76396 m

Lower Core Plate(MELCOR SS -

PLATE)

Lower Support Columns(MELCOR SS -

COLUMN)

Flow Mixer Plate(MELCOR SS -

PLATEG)

Upper Support 

Columns(MELCOR SS -

COLUMN)

Bottom Nozzles & 

Lower Core Plate(MELCOR NS & 

MELCOR SS -PLATEG)

Active Fuel and Control Rods (Top Nozzles in Axial Level 17, 

Core Shroud and Former in Ring 5)

Westinghouse 3-Loop PWR

Core Nodalization

DZLH = 0.13 m

R1 = 

0.27134 m

R2 = 

0.60715 m

R3 = 

0.94846 m

R4 = 

1.33587 m

R5 = 

1.7005 m

R6 = 

1.994 m

Core CellsLower Head 

Nodalization

Flow Path Opens at Shroud Failure

Flow Path 

Controlled by 

COR Blockage Model

Legend

CV748

CV160

Upper 

Plenum

114 CV

FL112 FL122 FL152 FL706

CV718 CV728 CV738 CV758

Westinghouse 3-Loop PWR

Upper RPV CVH/FL Nodalization

Downcomer 101 CV

FL100

CV701

FL117

112 CV 114 FL FL116

FL

124 CV

127

122 CV 124 FL FL126

FL164

FL

134 CV

132

132 CV 134 FL FL136

FL137

FL

144 CV

142

142 CV 144 FL FL146

FL147

154 CV 154 FL 153 CV FL156

FL157

152 CV

Control Rods 116 CV

Control Rods 126 CV

Control Rods 136 CV

Control Rods 146 CV

Control Rods 156 CV

FL192

FL162 FL172 FL182

FL193

FL199

FL174 FL184 FL194

FL153
4-9

Figure 4-3 Surry Reactor Core Radial Power Profile and Nodalization

R P N M L K J H G F E D C B A

1

2

3

4

5

6

7

8

9

1

1

1

1

1

1

MELCOR COR Radial Ring 1, Power Factor = 1.226

MELCOR COR Radial Ring 2, Power Factor = 1.293

MELCOR COR Radial Ring 3, Power Factor = 1.301

MELCOR COR Radial Ring 4, Power Factor = 1.110

MELCOR COR Radial Ring 5, Power Factor = 0.321
4-10

Figure 4-4 Surry Steam Generator A Hydrodynamic Nodalization

RPV

CV209

CV211

CV212

CV215

CV213

CV217

CV216

CV214

CV225

CV226

FL233

FL294

FL218 FL228

FL216 FL215

FL227 FL217 FL295 FL265

CV201 CV202

CV222 CV221

CV153

CV152

FL291

FL296

FL231

FL292

FL202 FL240

FL201 FL210

FL211

FL250 CV260

CV265CV260

CV250

CV250

FL830 & FL831

FL252

FL253

FL254

FL283

SG Tube Creep Rupture

FL205

FL873 Creep Rupture

SG A

Cubicle

CV020

FL206 FL293

FL266

FL245

Main Feedwater

(Mass Source) CF251

Turbine Driven Aux. Feedwater

(Mass Source) CF253

Motor Driven Aux. Feedwater

(Mass Source) CF1325

to CV270

and Turbine

to CV220

Hot Leg

Cold Leg

Loop A

Natural Circulation Flow Path Definitions

(identified in red on diagram)

*Natural circulation entry conditions defined as (a) 

hot leg CVs <5% water, (b) >10 K super heat in hot 

leg, and (c) recirculation pumps tripped.

*Maintain natural circulation flow paths when (a) 

hot leg CVs <10% full of water, (b) pumps are off, 

(c) no major creep rupture failures, and (d) loop 

seal flow is <20% of HL flow. (CF5672)

*FL240, FL205 and FL206 are open and FL245 is 

closed during non-natural circulation conditions. 

FL240, FL205 and FL206 are closed and FL245 is 

open in natural circulation conditions.

*FL293 and FL295 are open and FL211 and FL266 

are closed in non-natural circulation conditions. 

FL293 and FL295 are closed and FL211 and 

FL266 are open in natural circulation conditions.

*FL211/FL210 and FL266/FL265 are adjusted to 

give mixing ratio of 15%/85%.

*FL202 pressure drop adjusted to give CD = 0.12 

(from FLUENT).

*FL218 and FL227 pressure drop adjusted to give 

Tube_flow/HL_flow (Mratio) = 2.
4-11

RPV 309 CV

311 CV

312 CV CV315

CV313

CV317

CV316

CV314

CV325

CV326

333 FL

394 FL

FL318 FL328

FL315 FL316

FL317 FL327

395 FL

365 FL

CV302 CV301

CV322 CV321

CV153CV152

FL391

FL396

FL331

FL392FL302 FL340

FL301 FL310

FL311

FL350 CV360

CV365

360 CV

350 CV

350 CV

FL828 & FL829

FL352

FL353

FL354

FL383

SG Tube Creep Rupture

FL305

FL853

Creep Rupture

SG B

Cubicle

CV030

FL306 FL393

366 FL

FL345

Main Feedwater

(Mass Source) CF351

Turbine Driven Aux. Feedwater

(Mass Source) CF353

Motor Driven Aux. Feedwater

(Mass Source) CF1335

to CV370

and Turbine

to CV320

Hot Leg

Cold Leg

Loop B

Natural Circulation Flow Path Definitions

(identified in red on diagram)

*Natural circulation entry conditions defined as (a) hot leg CVs <5% water, (b) >10 K super heat in hot leg, and (c) recirculation pumps tripped.*Maintain natural circulation flow paths when (a) hot leg CVs <10% full of water, (b) pumps are off, (c) no major creep rupture failures, and (d) loop seal flow is <20% of HL flow. (CF5673)*FL340, FL305 and FL306 are open and FL345 is closed during non-natural circulation conditions. FL340, FL305 and FL306 are closed and FL345 is open in natural circulation conditions.*FL393 and FL395 are open and FL311 and FL366 are closed in non-natural circulation conditions. FL393 and FL395 are closed and FL311 and FL366 are open in natural circulation conditions.*FL311/FL310 and FL366/FL365 are adjusted to give mixing ratio of 15%/85%.*FL302 pressure drop adjusted to give CD = 0.12 (from FLUENT).*FL318 and FL327 pressure drop adjusted to give Tube_flow/HL_flow (Mratio) = 2.

RPV 409 CV

411 CV

412 CV CV415

CV413

CV417

CV416

CV414

CV425

CV426

433 FL

494 FL

FL418 FL428

FL415 FL416

FL417 FL427

495 FL

465 FL

CV402 CV401

CV422 CV421

CV153CV152

FL491

FL496

FL431

FL492FL402 FL440

FL401 FL410

FL411

FL450 CV460

CV465

460 CV

450 CV

450 CV

FL826 & FL827

FL452

FL453

FL454

FL483

SG Tube Creep Rupture

FL405

FL833

Creep Rupture

SG C

Cubicle

CV040

FL406 FL493

466 FL

FL445

Main Feedwater

(Mass Source) CF451

Turbine Driven Aux. Feedwater

(Mass Source) CF453

Motor Driven Aux. Feedwater

(Mass Source) CF1345

to CV470

and Turbine

to CV420

Hot Leg

Cold Leg

500 CV

505 FL

507 FL

Loop C

Natural Circulation Flow Path Definitions

(identified in red on diagram)

*Natural circulation entry conditions defined as (a) hot leg CVs <5% water, (b) >10 K super heat in hot 

leg, and (c) recirculation pumps tripped.*Maintain natural circulation flow paths when (a) hot leg CVs <10% full of water, (b) pumps are off, (c) no major creep rupture failures, and (d) loop seal flow is <20% of HL flow. (CF5664)*FL440, FL405 and FL406 are open and FL445 is closed during non-natural circulation conditions. FL440, FL405 and FL406 are closed and FL445 is open in natural circulation conditions.

*FL493 and FL495 are open and FL411 and FL466 

are closed in non-natural circulation conditions. 

FL493 and FL495 are closed and FL411 and FL466 are open in natural circulation conditions.*FL411/FL410 and FL466/FL465 are adjusted to give mixing ratio of 15%/85%.

*FL402 pressure drop adjusted to give CD = 0.12 (from FLUENT).*FL418 and FL427 pressure drop adjusted to give Tube_flow/HL_flow (Mratio) = 2.*FL505 is open and FL506 and FL507 are closed during non-natural circulation conditions. FL505 is closed and FL506 and FL507 are open in natural circulation conditions. 

Surge Line

506 FL

Figure 4-5 Surry Steam Generator B and C Hydrodynamic Nodalization
4-12

4.2 Primary and Secondary System Relief Valve Modeling

Special models were included to simulate the failure of the primary and secondary system relief 

valves. Each valve was individually modeled to accurately characterize its operational 

characteristics. The potential for failure under normal operating conditions and failure at high 

temperature, severe accident conditions was considered. Section 4.2.1 summarizes the primary 

system valve modeling on the pressurizer and Section 4.2.2 discusses the secondary system 

modeling. A peer review committee member noted there was a difference in the PWR versus 

BWR valve failure modeling under rated conditions. The explanation for the differences is 

discussed in Section 4.2.3.

4.2.1 Primary System Relief Valves

Each of the three safety relief valves (SRVs) on the pressurizer is represented separately in the 

MELCOR model. The valves are individually sized to flow 293,330 lb/hr (steam) at 2,485 psig

[12]. Their opening pressures are staggered by 14.50 psi with the lowest opening pressure set to 

2,485 psig. The valves close when pressure drops below 96% of their opening pressure [13]. 

The SRV with the lowest opening pressure is configured to fail open using the following criteria.

• A per demand failure probability of 0.0027 [13] and failure at a cumulative distribution 

function value of 0.5 (i.e., failure after 256 cycles by the relation: P(n) = 1-(1-Pd)n where 

P(n) is the cumulative distribution function value, Pd is the per-demand failure probability, 

and n is the number of cycles), or

• 10 cycles above 1,000 K

Each valve is represented separately.

Each of the two PORVs on the pressurizer is also represented separately in the MELCOR model.

The valves are individually sized to flow 210,000 lb/hr at 2,335 psig [12]. Their opening 

pressures are staggered by 14.50 psi with the lowest opening pressure set to 2,335 psig. The 

valves are defined to close when pressure drops below 96% of their opening pressure. The 

PORV with the lowest opening pressure is configured to fail open using the following criteria.

• A per demand failure probability of 0.0028 [13] and failure at a cumulative distribution 

function value of 0.5 (i.e., failure after 247 cycles), or

• 10 cycles above 1,000 K

The PORVs and SRVs empty into the Pressurizer Relief Tank (PRT), which is modeled as a 

separate control volume within the PRT cubicle in the containment. The PRT is 1300 ft3

(36.8 m3

) and has a water volume of 900 ft3 (24.4813 m3). Included in the PRT are “rupture 

disks” that fail at a pressure of 100 psig and have a capacity of 900,000 lb/hr (113.4 kg/s) [12].

4.2.2 Secondary System Relief Valves

The PORVs on the main steam lines (i.e., one PORV on each of the three lines) are represented 

separately in the MELCOR model. The valves are sized to flow 373,000 lb/hr at 1,035 psig [12].
4-13

The valves are set to open at 1,035 psig, set to close at 96% of their opening pressure, and are 

configured to fail open using the following criteria.

• A per demand failure probability of 0.0058 [13] and failure at a cumulative distribution 

function value of 0.5 ( i.e., failure after 119 cycles), or

• 10 cycles above 1,000 K

The three smaller SRVs (i.e., one per main steam line) of the 15 total SRVs in the main steam 

system (i.e., five SRVs per main steam line) are modeled individually in the MELCOR model. 

The smaller valves are sized to flow 361,750 lb/hr at 1,085 psig [13]. The valves are set to open 

at 1,085 psig, set to close at 96% of their opening pressure, and are configured to fail open using 

the following criteria.

• A per demand failure probability of 0.0027 [13] and failure at a cumulative distribution 

function value of 0.5 (i.e., failure after 256 cycles), or

• 10 cycles above 1,000 K

The SRVs other than the smaller ones are modeled jointly in the case of each steam line, i.e., in 

the case of each steam line, the four SRVs other than the one smaller SRV are represented 

jointly. The joint representation is based on the information shown in Table 4-2 regarding 

opening pressure and flow capacity [13].

Table 4-2 Safety Relief Valve Opening Pressure and Flow Capacity

Opening 

pressure (psig)

Capacity at 

opening pressure 

(lb/hr)

1,095 826,060

1,110 837,235

1,120 844,680

1,135 855,850

Accordingly, the SRVs in the MELCOR model jointly representing the four larger SRVs on a 

main steam line are sized to flow 3,423,400 lb/hr (i.e., 4 x 855,850 lb/hr) at 1,135 psig and are 

set to open at 1,115 psig. The valves are defined to close at 96% of their opening pressure. 

There are three of these valves in the MELCOR model, one for each main steam line.

4.2.3 PWR versus BWR Valve Failure Modeling

The PWR analysis selected a median or 50% failure probability for rated conditions at the 

beginning of the SOARCA. A 90% failure probability was used in the initial BWR calculations 

to represent a ‘high confidence’ level for an event that was perceived to be a ‘benevolent failure.’ 

These different modeling approaches were developed independently of each other, and the 

inconsistency was recognized later as a consequence of questions raised by the Peer Review 

panel. When the SOARCA analysis was revised to address these and other Peer Review 
4-14

comments, the differences in failure criteria narrowed, but were also found to be unimportant to 

results, as explained below.

The approach used to model stochastic failure of an SRV to reclose in the BWR analysis was 

replaced by a more 'best estimate' approach based on early Peer Review comments. The revised 

criterion for stochastic SRV failure was defined based on the ‘expected value’ for the number of 

cycles a valve would experience at the time of failure. The ‘expected value’ is calculated as 

1/Failure-Rate. If one translates this approach to a cumulative probability at the time of failure, 

the value corresponds to a 63% confidence level for BWRs, which is closer to, but still different 

from, the assumed 50% probability used in the PWR analysis. The calculated number of cycles 

experienced by primary and secondary coolant system relief and safety valves in the PWR is 

much less than the number corresponding to the median (i.e., 50%) failure probability.12 

Therefore, stochastic failure never occurs in the PWR calculations. Confidence in this 

observation would only increase if the failure condition were shifted from the median failure 

probability to the probability corresponding to the ‘expected value’ (i.e., 63%). 

Due to the low sensitivity of the valve failure characteristics to the SOARCA sequences, new 

PWR calculations were not performed to explore valve failure characteristics on the progression 

of events (i.e., except as noted in footnote 12 for the SGTR). Future PWR calculations will 

adopt the ‘expected value’ as the recommended base case value. In contrast, the BWR results 

were sensitive to the valve failure characteristics and the impact of failure variations are 

examined in NUREG/CR 7110, Volume 1.

4.3 Decay Heat Power Modeling

Full-power steady state reactor power is 2546 MWth. The decay heat data for Surry was from a 

recent NRC project analyzing accident source terms for on high burnup (HBU) cores [51]. The 

core power profile was based on plant-specific, cycle-specific nuclear design reports obtained 

from the licensees. Three different recent cycles were examined to ensure that significant 

cycle-to-cycle variations were not observed. The Surry decay power was based on information 

for Unit 2 cycles 16 through 18 [52][53][54]. The core decay heat and fission product 

inventories were calculated using results from a SCALE/ORIGEN decay heat of the Surry 

core [55]. In the SCALE/ORIGEN analysis, the plant specific data from the recent cycles was 

extrapolated slightly to a burnup of 59 GWd/t for the lead assembly. The resultant HBU decay 

heat and fission product decay heat were slightly conservative relative to best-estimate values 

and significantly larger than low burnup values [51]. The decay heats, masses, and specific 

activities as a function of time were processed and applied as input data to MELCOR to define 

decay heat and the radionuclide inventory. Values used in the MELCOR calculations 

corresponded to those generated for equilibrium conditions, in the middle of an operating cycle. 

A summary of the total core-wide decay power generated by this process is listed in Table 4-3.

 12 The reactor coolant pump seal leakage in the station blackout sequences created a depressurization mechanism 

that reduced the requirement for relief valve flow. The ISLOCA depressurized due to the break and did not 

require primary system or secondary system relief, except in a controlled manner for the cooldown. The SGTR 

sequence did include considerable primary and secondary system cycling. An additional failure mechanism due 

to damage following solid liquid flow was examined as a sensitivity calculation. None of the PWR sequences 

resulted in any challenges to the high temperature, severe accident failure criteria (i.e., 10 cycles at >1000 K).
4-15

Table 4-3 Decay Power in Surry MELCOR Model

Time Decay Power (MW)

0.0 sec 179.3

1.0 sec 165.8

3.0 sec 151.3

7.0 sec 136.5

13.0 sec 125.0

27.0 sec 111.7

54.0 sec 99.5

1.8 min 87.4

3.7 min 76.8

7.4 min 67.4

14.8 min 57.9

29.8 min 47.8

60.0 min 38.3

2.0 hr 30.9

12.0 hr 19.5

24.0 hr 16.1

48.0 hr 13.0

4.4 Natural Circulation Modeling

Three natural circulation flow patterns can be expected during a severe accident; (1) in-vessel 

circulation, (2) countercurrent hot leg flow, and (3) loop natural circulation (see Figure 4-6 [5]). 

Natural circulation is important in severe accident sequences because circulating steam from the 

core to upper reactor internals, the hot leg, and the SGs; (1) transfers heat away from the core, 

(2) changes the core melt progression, and (3) changes in-vessel fission product distribution. 

More importantly, the resultant heating of the external piping could progress to a thermal stress 

(i.e., creep rupture) failure of the primary pressure boundary and a resulting depressurization 

prior to lower head failure. For example, a high-pressure station blackout accident is not 

expected to result in full loop natural circulation flow (i.e., natural circulation pattern 3 shown on 

the left-hand side of Figure 4-6) at the start of the core degradation phase of the accident because 

the loop seal is not cleared. Consequently, the prediction of the first two natural circulation flow 

patterns is most critical [5]. The first two natural circulation flow patterns have been studied 

experimentally in the 1/7th scale natural circulation test program by Westinghouse Corporation 

for the Electric Power Research Institute (EPRI) [14] [15], computationally using the FLUENT 

computational fluid dynamics computer program [6] [7], and with plant application analyses 

using SCDAP/RELAP5 [9]. Subsequently, MELCOR was used to model the 1/7th-scale natural 

circulation tests [18]. The reader is referred to References [5] through [7] for detailed 

discussions of natural circulation behavior.

More recently, NRC has continued improving natural circulation modeling as part of the steam 

generator tube integrity program [16] [17]. The natural circulation modeling techniques used in 

MELCOR plant models were based on work performed as part of the code assessment of the 

1/7th scale tests [18], which closely followed the previous work performed by Bayless [5]. The 
4-16

natural circulation MELCOR modeling approach in the Surry model was updated for the 

SOARCA project to incorporate some of the recent modeling advances used by Fletcher with the 

SCDAP/RELAP5 severe accident analysis code [16].

The key features of the updated MELCOR natural circulation models are the following:

• 5 radial rings in the vessel and upper plenum for natural circulation

o Separate axial and radial flow paths throughout the core and upper plenum

o Radial and axial blockage models in the core during degradation

• Modeling important modes of heat transfer in the internal vessel,

o Convective heat transfer

o Gas-structure radiation in the upper plenum

o Structure to structure thermal radiation within the core

o Variable Zircaloy emissivity as a function oxide layer thickness

o Variable steel emissivities in the core as a function temperature

• Hot leg counter-current natural circulation tuned to a Froude Number correlation using 

results from a NRC FLUENT CFD analysis [6] [7],

Q = CD [ g (Δρ / ρ ) D

5 

] 

1/2

where: g is the acceleration due to gravity.

Q is the volumetric flow rate in a horizontal duct

ρ is the average fluid density (ρ) 

Δρ is the density difference between the two fluids

CD is the hot leg discharge coefficient

D is the pipe hydraulic diameter 

o Hot leg split into upper and lower halves

o CD from FLUENT = 0.12

• Steam generator mixing fractions based on FLUENT CFD analysis [6] [7]

o Inlet plenum subdivided into 3 regions for hot, mixed, and cold regions from plume 

analyses

o Flow ratio from the inlet SG plenum into the hot SG tubes is 15% from the hot, 

unmixed plume and 85% from the mixed region 

o Flow ratio into the lower (and cooler) portion of the hot leg piping from the inlet SG 

plenum is 15% from the cold SG tubes and 85% from the mixed region

o The SG is nodalized to have 50% of the SG tubes in upflow and 50% in downflow13

• Modeling important modes of heat transfer in the hot leg and steam generator

 13 Boyd, et al., and Fletcher and Beaton [7][16] used a 41%/59% flow split of hot tubes to cold tubes in the steam 

generator for natural circulation conditions. For simplicity, in non-natural circulation conditions, a 50/50 split 

was used in the MELCOR model.
4-17

o Convective heat transfer

• Augmented in hot leg based on FLUENT turbulence evaluations

o Gas to structure radiative exchange in the hot leg and steam generator tubes

o Heat loss through the piping and insulation

• Steam generator tube to hot leg flow ratio tuned to results from the FLUENT CFD 

analysis [6] [7]

o The ratio of the upward SG flow rate in the tubes to the horizontal flow rate from the 

reactor vessel into the hot leg piping was set to a value of 2 as in the FLUENT CFD 

analysis

• The pressurizer and steam generator PORV and safety valves were modeled individually 

to accurately represent the flow disruption to the natural circulation flow when individual 

valves opened

• Creep rupture modeling

o Hot leg nozzle carbon safe zone

o Hot leg piping

o Surge line

o Steam generator inlet tubes

The complexities of time-varying buoyant flows are impossible to resolve using MELCOR. 

Consequently, special flow paths are introduced to simulate natural circulation conditions 

measured in experiments and calculated using computational fluid dynamics codes. The red 

flow paths in Figure 4-4 and Figure 4-5 show the special flow paths in the hot legs and steam 

generators. As indicated in the legend, special flow paths are activated during natural circulation 

conditions to achieve the desired flow patterns. In particular, valves and additional head/drag 

terms are applied to match the desired phenomena. During natural circulation conditions 

(i.e., single-phase gas flow into the hot leg and steam generator), the red flow paths are activated. 

The result is a counter-current circulation flow pattern in the hot leg that matches the Froude 

Number correlation, a counter-current tube flow rate that is twice the hot leg flow, and 85% to 

15% flow mixing between the mixture and hot and cold streams entering and leaving the steam 

generator inlet plenum. However, if conditions change that would preclude the natural 

circulation flow pattern (e.g., flooding by the accumulators or an injection system, a creep 

rupture piping failure, operation of multiple relief valves, etc.), the control logic reactivates 

MELCOR’s normal two-phase thermal-hydraulic model with the base nodalization (i.e., the 

‘black’ flow paths in Figure 4-4 and Figure 4-5).
4-18

Figure 4-6 Natural Circulation Flow Patterns in a PWR

4.5 Core Degradation Modeling

The core support plate of the Surry reactor is column supported with the assemblies loading the 

plate between the columns. The lower core support structures are modeled using the plate and 

column structures. The core support plate is modeled as a grid supported support structure. The 

core support plate failure can fail due to stress, failure of the supporting columns, or can fail over 

time by creep at stresses below the yield stress, which is represented using a Larson-Miller 

creep-rupture model. Similarly, the support columns can fail by yielding and the failure of 

columns by buckling. The lower support plate is modeled as an edge supported plate, which can 

fail due to stress or can fail over time by creep at stresses below the yield stress. Non-supporting 

structures are used to represent the control rod tubes within the core in axial levels six through 

fifteen. These structures fail locally based on the thickness of the steel in the component. 

The core melt progression modeling options have been set to be consistent with current 

best-practices guidelines, which are generally default models (i.e., see reference [10]). The fuel 

rod cladding ruptures at relatively low temperature (i.e., modeled at 800°C in MELCOR) and 
4-19

releases fission gases from the fuel-cladding gap. As the fuel temperature increases, an oxide 

shell forms on the outer surface of the fuel cladding. Since the oxide shell has a higher melting 

temperature than the unoxidized Zircaloy inside of the fuel rod, the Zircaloy on the interior of the 

cladding will become molten once the temperature rises above the melting temperature (see 

Figure 4-7). Based on observations from Phebus tests, MELCOR includes a molten Zircaloy 

breakout model as the oxidized Zircaloy loses structural integrity. The molten Zircaloy flows 

through cracks in the cladding and relocates downward, which leaves a thin Zircaloy oxide shell 

holding the fuel pellets. Following the relocation of the molten Zircaloy, the local power due to 

Zircaloy oxidation ceases. The subsequent local thermal response is governed by decay heat and 

any relocation of molten material from above. The calculated failure rod collapse mechanisms 

include (a) failure collapse due to melting the oxidized shell or (b) failure collapse of the 

supporting structure, and (c), a time-at-temperature model that calculates the failure collapse of 

the oxidized Zircaloy shell holding the fuel rods. The time-at-temperature model acknowledges 

a thermal-mechanical weakening of the oxide shell as a function of temperature. As the 

temperature rises above Zircaloy melting temperature (i.e., represented as 2098 K in MELCOR) 

towards 2500 K, a thermal lifetime function linearly accrues increasing damage from 10 hours to 

1 hour until a predicted local thermal-mechanical failure, respectively (see Table 4-4). 

Table 4-4 Time versus Temperature Relationship for Fuel Rod Collapse

Temperature Time to Failure

2000 K

2090 K

2100 K

2500 K

2600 K

2700 K

Infinite

10 days

10 hr

1 hr

5 min

30 sec
4-20

Figure 4-7 Depiction of the Fuel Rod Degradation

4.6 Containment

The containment is divided into a total of nine control volumes and seventeen flow paths. 

Figure 4-8 and Figure 4-9 show the hydrodynamic nodalization of the containment. The control 

volumes represent the basement, the cavity under the reactor, the three separate steam generator 

cubicles, the pressurizer cubicle, the pressurizer relief tank (PRT) cubicle, the lower dome, and 

the upper dome. The basement region includes the bottom part of the containment as well as the 

surrounding cavity that lies between the outer wall and internal crane wall. 

The walls, floors, ceilings, and equipment in the containment are modeled as heat conducting 

structures. The structures will absorb and release heat during the course of an accident 

simulation. Fission products can deposit on any structure, however, gravitational settling only 

occurs on horizontal structures. The major walls include the outer walls of the containment that 

are shared with the environment (1.37 m thick), the wall separating the reactor cavity and 

basement (1.37 m thick), the wall separating the pressurizer cubicle and the outer cavity (0.61 m 

thick), and the PRT cubicle floor (0.3 m thick). Additional major walls include the outer wall 

separating the upper dome from the environment (1.37 m thick) and the wall separating the lower 

dome from the upper dome (0.76 m thick). Figure 4-8 shows these two wall sections have 

different thickness. The containment dome has a hemispherical geometry and is approximately 

0.762 m thick. 

ZrO2 oxide

Shell

Oxidizing Zr

Metal held under

Oxide shell

Release of 

Molten Zr (2400K)

ZrO2 oxide

Shell

Oxidizing Zr

Metal held under

Oxide shell

Release of 

Molten Zr (2400K)
4-21

The reactor cavity is represented using special physics models for core concrete interactions 

(CCI). The concrete floor is a combination of limestone aggregate and common sand concrete 

and has a 0.135 mass fraction of iron rebar. This concrete has an ablation temperature of 1650 K 

and an initial temperature of 311 K. The reactor cavity is represented with a flat-bottom 

cylindrical cavity that has an inner radius of 4.28 m and an outer radius of 5.58 m. The thickness 

of the concrete below the bottom of the cavity is 3.04 m. 

The reactor cavity connects to the basement through a 12” diameter hole bored through the 

shield wall at elevation -25’-0” (centerline).14 The centerline of this hole is located 2’-7” above 

the containment floor. Water in the basement or the cavity will flow through this hole when it is 

greater than 2’-7” deep. This has significance in the long-term boil-off of water when debris is 

located in the reactor cavity. 

 14 The containment model is based on Unit 1. Note: Unit 2 does not have this hole. 
4-22

Figure 4-8 Containment Hydrodynamic Nodalization
4-23

Figure 4-9 Containment Hydrodynamic Nodalization, Plan View

4.7 Containment Leakage Model

Extensive research and scale model testing of reinforced and pre-stressed concrete containments 

to determine behavior at beyond design basis accident pressure has been performed in the last 

25 years at SNL [36] and the Central Electricity Generating Board (CEGB) [37]. Testing has 

shown that concrete containments start to leak at leak rates much higher than design leakage and 

well before a large rupture or gross failure would occur. This leakage could preclude the large 

rupture or failure. The relationship between containment leakage and internal pressures for 

reinforced concrete and pre-stressed concrete containment model tests is described in References 

4 and 5. The details of the containment performance model developed by the NRC staff for use 

in this analysis are described in detail in Appendix A. The concrete containments start to leak 

appreciably once the liner plate yields and tears. The rate of leakage when the liner plate yields 

and tears is about 10 times more than normal leakage of 0.10 percent of containment air mass 

per day at the containment design pressure. The leakage rate increases appreciably with further 
4-24

increases in test pressure. Once the rebar yields, the leakage rate is about 10-15 percent per day. 

Thereafter, the leakage rate continues to increase and reaches to about 60-65 percent per day

when the strain in the rebar is about 1-2 percent. The containment pressure does not increase 

significantly after the leakage rate exceeds 60-65 percent per day. The liner welds and concrete 

crack after the rebar and liner plate yields to create a path for leakage. The leakage occurs in 

areas such as equipment hatch, personnel airlocks, and penetrations where local strains are 

substantially higher than the global strains. All leakage was assumed to occur through the 

equipment hatch (i.e., the largest penetration into the containment), which conservatively 

releases all fission products to the environment. Leakage through most of the other penetrations 

would allow some fission product deposition in the connected buildings (e.g., the auxiliary and 

safeguards buildings). 

 

The results in [38] and [39] are for scale model tests of two concrete containments. Rebar and 

concrete crack spacing, and aggregate size can affect the leakage rates in full size containments. 

However, based on the results of testing and analyses presented above, it is reasonable to 

conclude that all concrete containments start to leak once the rebar and liner plate yield. In 

addition, leakage becomes excessive once the strains in the reinforced and pre-stressed concrete 

containments reach about 2 and 1 percent, respectively. Based on information from the 

containment model test and analyses, it is reasonable to assume that containment leakage is 

about one percent of the containment mass per day when the liner plate yields. This increases to 

13 percent of containment mass per day when rebar yield. Similarly, a leakage rate of 62 percent 

can be used in severe accident analysis when the containment global strains are 1-2 percent. The 

uncertainty in the leakage rate can be accounted for by conservatively reducing the yield and 

failure pressure calculated by simplified analysis to 85 percent of the calculated value.

The location of the leakage can have a significant effect on the results of the severe accident 

analysis and dose rates. For instance, if the containment leakage occurs through penetrations that 

are located inside adjoining plant buildings, the fission product release into atmosphere would be 

significantly less as compared to direct leakage to the environment. Previously, some of the 

severe accident analyses were based on the assumption that the leakage takes place at the top of 

the containment dome. A more realistic approach is to consider leakage to occur at the 

equipment hatch, which was done in SOARCA. Leakage through the equipment hatch 

discharges into the environment from the side of the containment dome.

The implementation into the Surry MELCOR model uses two containment failure mechanisms.

1. Nominal leakage per design specifications - 0.1% volume/day at PDesign, see Figure 4-10.

2. Containment overpressure leakage as described above - see Figure 4-11.

The nominal leakage is always active but very small. The containment overpressure failure 

occurs at 2.17 times the design pressure, or 0.775 MPa (112.4 psia). This estimate of 2.17 times 

the design pressure is derived from a curve fit of the three data points shown in Figure 4-11. The 

leakage starts very small but grows as the pressure increases. If the containment pressure 

subsequently decreases, the leakage area will not decrease from the maximum value.
4-25

Figure 4-10 Nominal Containment Leakage Model

Figure 4-11 Containment Failure Leakage Model

Nominal Containment Leakage

Design Specification 0.1% Vol/day, PD = 45 psi

0.00%

0.05%

0.10%

0.15%

0.20%

0.25%

0.30%

0.0 0.5 1.0 1.5 2.0 2.5 3.0

P/PD

%Vol / Day

Containment Failure Leakage

NRC Surry Failure Model, PD = 45 psi

0%

50%

100%

150%

200%

250%

300%

350%

400%

0.0 0.5 1.0 1.5 2.0 2.5 3.0

P/PD

%Mass/Day

MELCOR

Scaled from NUREG/CR 5121
4-26

4.8 Auxiliary Building

A total of 9 control volumes and 17 flow paths represent the Auxiliary Building (see 

Figure 4-12). The auxiliary building is modeled on a floor-by-floor basis beginning with the 

basement floor and rising up through the main floors up to the fourth floor. The first floor, at a 

2’0” elevation, is broken up into four control volumes. The first floor is subdivided to represent 

major rooms at that elevation. The HHSI pumps and motors are located at this elevation. 

The second floor, at 13’ elevation, is divided into 3 control volumes. A large room in the middle 

of the second floor contains boric acid transfer pumps as well as part of the boric acid tanks. The 

other two rooms contain the cable vault, electrical tunnel, and electrical vault. The middle room 

is connected to the side rooms by doorways. This floor also connects with the first floor by 

three separate stairwells. The third floor, at a 27’-6” elevation, is represented by a single control 

volume. The third floor contains the volume control tanks and part of the boric acid tanks. The 

third floor connects with the second floor by the stairwell located next to the elevator. The 

fourth floor, at a 45’-10” elevation is also represented by a single control volume. The fourth 

floor is where the personnel hatches are located along with the heating and ventilation 

equipment. There are many potential leakage locations to the environment on the fourth floor 

through ventilation ducting and the blowout panels. The leakage is represented as a 0.65 m2

(7 ft2) flow path to the environment.

Representations have been included in the Surry MELCOR model of the Safeguards Area, 

Contaiment Spray Pump Area, and Main Steam Valve House, as described in Section 4.10. 

Details of the Safeguards ventilation system and low head safety injection piping are described in 

Sections 4.11 and 4.12, respectively. Other buildings that are directly connected or could be 

involved as a fission product pathway (i.e., the turbine building) were not represented because 

they were not relevant fission product pathways for the sequences analyzed.
4-27

Figure 4-12 Auxiliary Building Hydrodynamic Nodalization

4.9 Best Modeling Practices

The SOARCA project is intended to provide a body of knowledge regarding the realistic 

outcomes of severe reactor accidents. To accomplish this objective, the SOARCA project used

integrated modeling of accident progression and offsite consequences using both state-of-the-art 

computational analysis tools and best modeling practices drawn from the collective wisdom of

the severe accident analysis community. 

The MELCOR 1.8.6 computer code [10] embodies much of this knowledge and was used for the 

accident and source-term analysis. MELCOR includes capabilities to model the two-phase 

thermal-hydraulics, core degradation, fission product release, transport, deposition, and 

containment response. The SOARCA analyses include operator actions and equipment 

performance issues as prescribed by the sequence definition and mitigative actions. The 

MELCOR models are constructed using plant data and the operator actions were developed 

based on discussions with operators during site visits. The code models and user-specified 

modeling practices represent the current best practices.

Uncertainties remain in our understanding of the phenomena that govern severe accident 

progression and radionuclide transport. Consistent with the best-estimate approach in SOARCA, 

all phenomena were modeled using best-estimate characterization of uncertain phenomena and 

events. Important severe accident phenomena and the proposed approach to modeling them in 

the SOARCA calculations were presented to an external expert panel during a public meeting 
4-28

sponsored by the NRC on August 21 and 22, 2006 in Albuquerque, New Mexico. A summary of 

this approach is described in Section 4.9.1. These phenomena are singled out because they are 

important contributors to calculated results and have uncertainty. Section 4.9.2 briefly describes 

the two other topics, steam explosions, and direct containment heating that had been previously 

included in lists of uncertain phenomena. Finally, a systematic evaluation of phenomenological 

uncertainties for a particular sequence is a separate task and not discussed in this report. That 

task will evaluate the importance and impact of alternative settings or approaches for key 

uncertainties.

4.9.1 Approach to Modeling Important Phenomena 

A review of severe accident progression modeling for the SOARCA project was conducted at a 

public meeting in Albuquerque, New Mexico on August 21-22, 2006 [11]. This review focused 

primarily on best modeling practices for the application of the severe nuclear reactor accident 

analysis code MELCOR for realistic evaluation of accident progression, source term, and offsite 

consequences. The scope of the meeting also included consideration of potential enhancements 

to the MELCOR code as well as consideration of the SOARCA project in general.

The review was conducted by five panelists with demonstrated expertise in the analysis of severe 

accidents at commercial nuclear power plants. The panelists were drawn from private industry,

the Department of Energy national laboratory complex, and a company working on behalf of 

German Ministries. The review was coordinated by Sandia National Laboratories and Nuclear 

Regulatory Commission staff.

The following important uncertain modeling practices were presented to the peer review panel. 

The review panel provided written comments and suggestions, which were incorporated into the 

subsequent analyses. Base case approaches were identified for these uncertain and typically 

important parameters. 

• Safety relief valve cycling and failure

Mean opening and reclosing failure probabilities for the pressurizer and steam generator 

power operated relief (PORV) and safety valves (SV) were applied in the calculations. A 

high temperature thermal failure model was also applied. This is discussed in further 

detail in Section 4.2.

• Pump seal leakage and blowout

The base case pump seal leakage model described in Section 4.1 was identified as the 

base case modeling approach. In addition, early seal failure sensitivity calculations were 

performed for the LTSBO.
4-29

• Loop seal clearing and effects on the accident progression

The most important impact from this event is an increased vulnerability of the steam 

generator tubes for failure due to a full loop circulation of hot gases from the core during 

core degradation. MELCOR has basic thermal-hydraulic modeling for calculating loop 

seal clearing. However, it is recognized that loop seal clearing is related to other 

complex and uncertain events, such sensitive system hydrodynamic pressure balances 

during core degradation events and pump seal leakage. NRC has a separate research 

program examining thermally-induced steam generator tube failure. Due to the potential 

importance of steam generator tube failure (i.e., the most important consequence of loop 

seal clearing), calculations were performed that included steam generator tube failure.

• Fuel degradation and relocation treatment

An additional model has been added to characterize the structural integrity of the fuel 

rods under highly degraded conditions. The new failure model acknowledges a 

thermal-mechanical weakening of the oxide shell as a function of time and temperature. 

As the local cladding oxide temperature increases from the Zircaloy melting temperature 

(i.e., represented as 2098 K in MELCOR) towards 2500 K, a thermal lifetime function 

accrues increasing damage from 10 hours to 1 hour until a local failure of the oxide shell.

• Lower plenum debris/coolant heat transfer

Following the fuel-debris slump into the lower plenum, there may be fuel-coolant

Interactions and multi-dimensional cooling of the resultant debris bed. The lower plenum 

heat transfer settings were updated to disable the 0-dimensional Lipinski counter-current 

flow limiting (CCFL) [10]. The 0-dimensional CCFL model prevented multidimensional heat transfer in the lower plenum. The resultant behavior resulted in multidimensional film and nucleate boiling debris heat transfer if there was a pool in the lower 

plenum. The debris interaction with water somewhat cooled the debris and delayed 

failure of the lower head.

• Core plate failure

The timing of core plate failure affects the relocation of the degraded core materials from 

the core region into the lower plenum. The local thermal-mechanical failure of the lower 

core plate, the flow mixer plate, and the lower support forging are calculated within 

MELCOR using the Roark engineering stress formulae. The yield stress is calculated 

based on the loading and local temperature.

• Fission product release, speciation, and volatility

First, the CORSOR-Booth diffusion model was used to calculate the release of 

radionuclides from fuel.

Second, the predominant speciation of cesium was changed based on detailed analysis of 

the deposition and transport of the volatile fission products in the Phebus facility tests. 

The analysis revealed molybdenum combined with cesium and formed cesium 

molybdate. Previously, the default predominant chemical form cesium was cesium 

hydroxide. As consistent with past studies, all the released iodine combines with the 

cesium. Applications of this information to the MELCOR models used in the SOARCA 
4-30

calculations are described in SAND2010-1633, “Synthesis of VERCORS and Phebus 

Data in Severe Accident Codes and Applications.”

• RCS natural circulation treatment

The base case RCS natural circulation models described in Section 4.4 were identified as 

the base case modeling approach.

• Vessel lower head failure and debris ejection

The base case approach of modeling the vessel lower head failure and debris ejection 

included some modifications in MELCOR. First, all the solid debris in the lower plenum 

is in contact with water, if present. Previously, a restrictive one-dimensional 

counter-current flooding limitation criterion prevented penetration of water into the 

debris bed. Second, the vessel lower head fails using a creep rupture model. A 

Larson-Miller failure criterion is calculated based on the one-dimensional conduction and 

stress profile through the lower head. The failure of a lower head penetration prior to 

gross head failure was judged unlikely based on observations from experimental studies 

at Sandia National Laboratories lower head failure (LHF) tests.

• Ex-vessel phenomena - CCI 

The default model’s ex-vessel debris surface heat flux to an overlying pool of water was 

enhanced to replicate the magnitude observed in the MACE tests. The default model did 

not include multi-dimensional effects of fissures, other surface non-uniformities, and side 

heat fluxes.

• Ex-vessel phenomena - Hydrogen combustion

The default MELCOR ex-vessel combustion model was used with the modeling options 

to include horizontal and vertical propagation of burns and the time delay for the flame 

front to span the width of the control volume.

4.9.2 Early Containment Failure Phenomena

Two phenomenological issues not included in the best-estimate approach used in SOARCA 

include: (1) alpha-mode containment failure and (2) direct containment heating leading to 

containment failure. These severe phenomena leading to an early failure of the containment 

were included in some of the first studies to quantify the risks from nuclear reactors. However, 

they are not included in the SOARCA analyses because they are either extremely low likelihood 

or are physically unfeasible as described below.

The alpha-mode event is characterized by the supposition that an in-vessel steam explosion 

might be initiated during core meltdown by molten core material falling into the water-filled 

lower plenum of the reactor vessel. The concern was that the resulting steam explosion could 

impart sufficient energy to separate the upper vessel head from the vessel itself and form a 

missile with sufficient energy to penetrate the reactor containment. This would produce an early 

failure of the containment building at a time when the largest mass of fission products is released 

from the reactor fuel. In the following years, significant research was focused on characterizing 

and quantifying this hypothesized response in order to attempt to reduce the significant 

uncertainty. A group of experts ultimately concluded in a position paper published by the 
4-31

Nuclear Energy Agency’s Committee on the Safety of Nuclear Installations that the alpha-mode 

failure issue for Western-style reactor containment buildings can be considered resolved from a 

risk perspective, posing little or no significance to the overall risk from a nuclear power plant.

Similarly, direct containment heating (DCH) was another important event identified to cause 

early containment failure. NUREG-1150 [2] was an important risk study that included DCH as 

an early containment failure phenomenon. Extensive research was performed with the goal of 

characterizing DCH. Additionally, research was performed concerning other phenomena that 

can preclude an early, energetic failure of the containment (e.g., natural circulation leading to 

creep rupture of the RCS boundary, see Section 4.4). First, the extensive natural circulation 

research shows that RCS failure prior to vessel failure due to RCS creep rupture is most likely. 

In the unlikely event there is a high-pressure vessel failure (i.e., not within SOARCA’s 

objectives for best-estimate evaluations), the resolution of the DCH issue found early 

containment failure to be very unlikely [22]. The issue resolution utilized a probabilistic 

framework that decomposes the DCH problem into three probability density functions that 

reflect the most uncertain initial conditions (i.e., UO2 mass, zirconium oxidation fraction, and 

steel mass). Uncertainties in the initial conditions are significant, but the quantification approach 

established reasonable bounds that are not unnecessarily conservative. The phenomenological 

models in the probabilistic model were compared with an extensive database including recent 

integral simulations at two different physical scales (1:10-scale in the Surtsey facility at Sandia 

National Laboratories and 1:40-scale in the COREXIT facility at Argonne National Laboratory). 

The loads predicted by these models were significantly lower than those from previous 

parametric calculations. The containment load distributions do not intersect the containment 

strength (fragility) curve in any significant way, resulting in containment failure probabilities 

less than 10-3 for all scenarios considered. Sensitivity analyses did not show any areas of large 

sensitivity. Consequently, DCH is not a likely accident progression event and therefore not 

within SOARCA’s best-estimate approach guidelines. 

4.10 Safeguards Area, Containment Spray Pump Area, and Main Steam Valve House

Specifically for importance in the ISLOCA accident scenario, representations have been included 

in the Surry MELCOR model of the Safeguards Area, Containment Spray Pump Area, and Main 

Steam Valve House. These interconnected buildings, referred to collectively in this document as 

Safeguards or Safeguards buildings, are located outside and adjacent to each unit’s reactor 

containment building. In the ISLOCA scenario postulated for Surry, the RCS blows down into 

the Safeguards Area and fission products are released to the environment through the Safeguards 

buildings.

Figure 4-13 illustrates the MELCOR nodalization constructed to represent the Safeguards 

buildings. The Safeguards Area and Containment Spray Pump Area are served by a safetyrelated filtered ventilation system. This system is represented in the MELCOR modeling of 

Safeguards. The nodalization employed for it is shown in Figure 4-14.

Access to the Safeguards Area is from the yard at approximately grade elevation. The building 

houses two of its reactor unit’s four containment recirculation spray pumps and both of its unit’s 

low-head safety injection pumps. The building has three levels. Ladders lead down from grade 
4-32

elevation to the spray pumps and safety injection pumps, each pump residing in a separate 

cubicle.

Adjacent to the Safeguards Area is the Containment Spray Pump Area. A unit’s two 

containment spray pumps are located in its Containment Spray Pump Area. A unit’s instrument 

air compressors are located in the basement of its Containment Spray Pump Area.

Adjacent to the Containment Spray Pump Area is the Main Steam Valve House. A unit’s two 

electric-motor-driven auxiliary feedwater pumps and one steam-turbine-driven auxiliary 

feedwater pump are located in its Main Steam Valve House. Grating floors above the auxiliary 

feedwater pumps provides access to the main steam safety valves, steam generator power 

operated relief valves, and main steam trip valves.

A basement area common between the Containment Spray Pump Area and the Main Steam 

Valve House contains the auxiliary feedwater booster pumps and service water piping.

A noteworthy aspect of the construction of the Safeguards buildings is that these buildings do not 

have an integral back wall. The buildings back up to Containment and the Containment cylinder 

serves as a back wall to the buildings. Where the walls and roofs of the Safeguards buildings 

approach the Containment cylinder, there are 3” gaps (shaker spaces) covered with flashing or 

angle iron. 
4-33

Figure 4-13 Safeguards Buildings MELCOR Nodalization
4-34

Figure 4-14 Safeguards Buildings Ventilation MELCOR Nodalization
4-35

4.10.1 Building Interconnectivity

A detailed accounting of the interconnectivity between the Safeguards buildings is presented in 

Appendix D. Key penetrations, doorways, etc., and their influences are described below.

Pipe penetration between Safeguards Area and Containment Spray Pump Area:

There is a rectangular opening in the wall between the Safeguards Area and the Containment 

Spray Pump Area that would limit the depth of the pool that could develop on the Safeguards 

Area floor given an ISLOCA. The depth of the pool would be important because the ISLOCA 

pipe break could potentially be submerged such that fission products emanating from the break 

could be captured (i.e., scrubbed) by the pool. A pool in the Safeguards Area would overflow to 

the Containment Spray Pump Area through the subject opening. The opening is 21” tall and 5’ 

long. The base of the opening is 4’-9” off the Safeguards Area floor. The opening looks to be 

~50% full of piping. It is represented by Flow Path 625.

The flooding in the Safeguards Area would submerge the non-submersible motor of isolation 

valve MOV 1890C making the ISLOCA unisolable. The flooding would not be sufficiently high 

to wet the LHSI pump motors. As the LHSI pumps are largely sheltered from the most likely 

pipe break and as the electrical wiring to their motors drops down from above, the LHSI pumps 

are judged unsusceptible to flooding from an ISLOCA.

Water level in the building would remain at the 4’-9” mark until water (liquid) stopped spilling 

from the break. From then on, the level would slowly fall as the pool in the Safeguards Area 

leaked to the Containment Spray Pump Area past the flashing covering the shaker space between 

the two buildings. The ISLOCA pipe break, assumed centered 1’-9” off the floor, would 

potentially be submerged during core degradation. Submergence would support scrubbing of 

fission product aerosols and vapors by the pool. MELCOR addresses this phenomenon. 

Shaker space between Safeguards Area and Containment Spray Pump Area:

As described earlier, the Safeguards buildings do not have integral back walls. The buildings 

back up to Containment and the Containment cylinder serves as a back wall to the buildings. 

Where the wall between the Safeguards Area and the Containment Spray Pump Area approaches 

the Containment cylinder, there is a 3” shaker space. The shaker space is covered by flexible 

metallic flashing on both sides of the wall. The flashing is close fitting but may not be 

watertight. Consequently, a pool in the Safeguards Area would leak past the flashing into the 

Containment Spray Pump Area and the depth of the pool would be influenced by the leakage as 

time progressed. As described previously, the depth of the pool would be important because the 

ISLOCA pipe break could potentially be submerged such that fission products emanating from 

the break could be scrubbed by the pool. 

Appendix D includes a cross sectional drawing of the flashing. The flashing is represented by 

Flow Path 636. The flashing was assumed to have an effective 1/32”-wide leakage gap along its 

length. The flow-path distance through the gap (from the Safeguards area to the Containment 

Spray Pump Area) was assumed to be 6” based on the cross-section dimension of the angle iron. 

Flow Path 636, therefore, was defined with the flow area of a 1/32” gap, a length of 6”, and a 

hydraulic diameter of 1/32”. Entry and exit form losses of 0.5 and 1.0 were assumed on both 

sides of the wall for a combined form loss of 3.0. Uncertainty exists in the assumed 1/32” 
4-36

effective leakage gap of the flashing given that that the walls of Safeguards Area are not 

especially flat and that the angle iron mating the flashing to the walls is rigid.

Open doorway and taller opening joining the basements of the Containment Spray Pump Area 

and the Main Steam Valve House:

The basements of the Containment Spray Pump Area and the Main Steam Valve House are 

largely open to each other via a doorway and a taller and wider opening. Water and air could 

flow freely between the basements. Flow Path 628 accomplishes this interconnectivity. 

Water overflowing the Safeguards Area to the Containment Spray Pump Area would move 

readily across the largely open floor to the adjoining floor of the Main Steam Valve House and 

quickly fill a 9’ pit there. A pipe tunnel leads from the pit to the Auxiliary Building. The tunnel 

is filled with fire-resistant penetration sealant. Water would pool on the Main Steam Valve 

House and Containment Spray Pump Area floors high enough for hydrostatic pressure to 

dislodge this sealant opening the tunnel. Water issuing from the ISLOCA pipe break and 

through the tunnel would ultimately flood the basement of the Auxiliary Building.

Pipe tunnel between Main Steam Valve House and Auxiliary Building:

As noted above, there is a pipe tunnel between the Main Steam Valve House and Auxiliary 

Building. The tunnel opens on the Main Steam Valve House side to a pit in the floor. The 

tunnel cross section is 18” (high) x 7’-7”. The tunnel is not long, traversing only the thickness of 

the wall between the Main Steam Valve House and Auxiliary Building. The tunnel is 

crisscrossed with occasional rebar and filled with a sprayed penetration sealant. The sealant is

thought to have little capacity to tolerate a pressure differential and so is assumed to dislodge in 

the ISLOCA scenario when the pit in the Main Steam Valve House floods. Water level would be 

1’-3” above the top of the tunnel at this point. The question arises as to whether the sealant 

might survive 1’-3” of water. The question becomes mute however realizing that if the tunnel 

were blocked to flow, water level (in an ISLOCA) would rise to many feet in the Main Steam 

Valve House. It is not credible that the sealant could withstand many feet of water. The subject 

tunnel is represented by Flow Path 634.

4.10.2 Existing Potential Fission Product Release Pathways

Several potential pathways exist for the release of fission products from the Safeguards buildings 

given an ISLOCA resulting in core damage. These pathways are identified below.

Safeguards exhaust ventilation system:

The filtered exhaust ventilation system serving Safeguards would be a pathway for fission 

product release should the filters fail. The system and the modeling of it are described in 

Section 4.11. The filters could fail from exposure to excessive temperatures or excessive 

differential pressure. Excessive temperatures (> 250 oF) could arise from hot gasses passing 

through the filters or from the heat produced by the decay of fission products captured by the 

filter media. Excessive differential pressure (> 25 in H2O) could develop across the filters as 

they load with aerosol and impede flow.
4-37

Safeguards supply ventilation system:

The supply ventilation system serving Safeguards would be a release pathway should Safeguards 

pressure become greater than ambient pressure. The system is fitted with continuous roll filters 

and steam heating coils for cold-weather space heating. The roll filters are not high efficiency 

aerosol filters. The system has a 16,000 cubic feet per minute (cfm) fan, but to ensure inward 

leakage to Safeguards, the fan is not operated. Section 4.11 describes the system and how it is 

modeled in MELCOR.

Safeguards Area personnel door:

The lone personnel door for accessing the Safeguards Area is centrally located in the top floor of 

the building at ground level. The door opens outward and is equipped with a closer. The door 

opens to a modest push by hand. It has no latch. Flow Path 611 shown in Figure 4-13 represents

this door in the MELCOR model. Given a meaningful elevation of Safeguards Area pressure 

(i.e., 1 in H2O gauge), this flow path is opened to the environment. The flow path is reclosed 

upon loss of the elevated pressure.

Abandoned penetrations in the Containment Spray Pump Area:

Two abandoned penetrations exist in the exterior wall of the upper level of the Containment 

Spray Pump Area. The penetrations lead directly to the environment. They are approximately 1’ 

in diameter and 12’ off the floor. Flow Path 629 shown in Figure 4-13 represents these 

penetrations.

Abandoned exhaust fan in the Containment Spray Pump Area:

An abandoned exhaust fan resides in the roof of the Containment Spray Pump Area. The 

opening in the roof associated with the fan was estimated to be 2’ x 2’. Flow Path 626 shown in 

Figure 4-13 represents this opening.

Main Steam Valve House exhaust fan:

The upper level of the Main Steam Valve House has a wall-mounted exhaust fan that runs 

continuously. The fan flows 13,500 cfm. There is no filtering of the air exhausted by this fan, as 

the Main Steam Valve House is not considered a potentially contaminated area. Flow Path 631

shown in Figure 4-13 represents this fan.

Main Steam Valve House ventilation intake:

An air intake exists in the exterior wall of the upper level of the Main Steam Valve House. The 

intake is near the floor. It is 8’ wide x 4’-6” wide and louvered. Flow Path 630 shown in 

Figure 4-13 represents this intake.

4.10.3 Potential Building Boundary Over-Pressure Failures

An ISLOCA at Surry and consequential RCS blowdown into the Safeguards Area would 

pressurize the Safeguards buildings. The buildings are reinforced concrete structures largely 

below grade with corrugated steel roofs. The construct of the buildings has been scrutinized and 

likely building boundary failures have been identified. The likely failures and the modeling of 

them are described below. Subsequent hydrogen burns in the Safeguards Area could be expected 

to further damage the building’s structural boundary. 
4-38

Tearing of Safeguards Area and/or Containment Spray Pump Area roof flashing:

The Safeguards buildings do not have integral back walls. The buildings back up to containment 

and the containment cylinder serves as a back wall to the buildings. The steel roofing on the 

buildings extends to within a few inches of the containment wall leaving a gap that is closed with 

flashing. This flashing has been judged as a weak point in the pressure boundaries of the 

Safeguards buildings. The flashing has been specified to tear given a pressure differential across 

it of 37.5 psf. The basis for this criterion is presented in Appendix D. The flashing is 

represented with Flow Paths 612 and 643 shown in Figure 4-13 for the Safeguards Area and the 

Containment Spray Pump Area, respectively. Flow Path 612 reflects a gap 3” wide by 60’-4” 

long while Flow Path 643 reflects a gap 3” wide by 22’-9” long.

Gross failure of Safeguards Area and/or Containment Spray Pump Area roofs:

The tearing of roof flashing described above would significantly vent the Safeguards buildings 

given an ISLOCA and consequential RCS blowdown into them. The venting, however, might 

not be sufficient to curtail further damage to the boundaries of the buildings from the blowdown 

or from subsequent hydrogen burns occurring within the buildings.

The Safeguards buildings are reinforced concrete structures largely below grade with corrugated 

steel roofs. From considerations presented in Appendix D, the roofs are identified as the weakest 

boundaries of the buildings and are taken to fail given a pressure differential of 75 psf across 

them. The failure is considered gross in that a large opening in the roof results. Flow Paths 613 

and 647 shown in Figure 4-13 represent gross roof failures in the Safeguards Area and the 

Containment Spray Pump Area, respectively. These flow paths have an area of 10 m2 each.

4.11 Safeguards Ventilation System

The Safeguards Area and the Containment Spray Pump Area are normally exhausted unfiltered 

by dual 6,000 cfm fans. As these buildings have recognized contamination potential, they are 

instead exhausted by a filtered safety-related ventilation system given a LOCA. On a safety 

injection signal, the normal exhaust fans are automatically isolated and the safety-related exhaust 

system is automatically started. The safety-related exhaust has particulate and iodine filtration 

and dual parallel redundant fans. The fans and filters serving the Safeguards Area and 

Containment Spray Pump Area reside in the Auxiliary Building. The fans exhaust through a 

stack above the roof of the Auxiliary Building. Figure 4-14 illustrates the configuration of the 

Surry Safeguards Ventilation System and the MELCOR representation of it. Fresh air to the 

Safeguards Area and Containment Spray Pump Area is supplied by ducting routed from a 

common intake point in the exterior wall of the Containment Spray Pump Area. As originally 

constructed, an intake fan forced fresh air through the ducting. To ensure negative pressures 

within the Safeguards buildings, this fan is no longer operated. Air is instead drawn though the 

intake ducting in response to the negative pressures developed in the Safeguards buildings by the 

exhaust ventilation system. Intake air is continually cleaned by roll filters and heated by steam 

coils in the winter. The roll filters are not high efficiency.

The safety-related exhaust fans and filters serving the Safeguards areas are not dedicated to these 

areas. Instead, the fans and filters serve many areas at Surry having recognized contamination 

potential. Of the combined 72,000 cfm drawn by the fans through the filters, 11,000 cfm would 
4-39

be drawn from Unit 1 or Unit 2 Safeguards following the generation of a safety injection signal 

in whichever unit. A flow-restricting orifice in the exhaust ducting leading from each unit’s 

Safeguards area limits flow. 

Pressure switches installed in the inlets of the safety-related exhaust fans trip the fans should 

inlet pressure reduce to less than 21 inches H2O gauge. Excessive aerosol loading on the filters 

could cause this condition. This functionality is captured in the MELCOR modeling. Heavy 

loading on the filters is not expected to result in tearing of the filters as the filter vendor has 

stated that the filters could easily withstand the maximum pressure head the fans can develop.

The particulate filters in the safety-related exhaust system are high-efficiency particulate air 

(HEPA) filters. Two banks of filters are installed, one associated with each fan. Each filter bank 

is constructed of 30 individual filter cells.

The filter units in the safety-related exhaust system, one filter unit for each fan, are constructed 

from individual filter cells. Each filter cell comprises a pre-filter, backed by a HEPA filter, 

backed by two parallel charcoal filters. The frontal area of a filter cell is 2’ x 2’. The HEPA 

filters are 1’ thick. Each filter unit is 10 filter cells wide by 3 filter cells high. The MELCOR 

modeling of the filter units represents:

• A 99.5% aerosol capture efficiency by the HEPA filters

• The clean flow resistance of the HEPA filters

• The added flow resistance of the HEPA filters due to aerosol loading within them

• Tearing of the HEPA filters from excessive pressure difference across them

• A 99.0% iodine and cesium iodide vapor capture efficiency by the charcoal filters

• The clean flow resistance of the charcoal filters

The pre-filters are not represented in the MELCOR modeling. The temperatures to which the 

HEPA filters are exposed are monitored in the ISLOCA calculations (i.e., relative to the 250 o

F 

maximum continuous temperature rating for the filters) but no over temperature failure criterion 

is implemented in the modeling. While tearing of the HEPA filters on excessive differential 

pressure is implemented, the filters are not expected to tear given that the filter vendor concluded 

they could easily withstand the maximum pressure the fans could develop. 

Substantial amounts of fission product aerosols could be expected to lodge in the HEPA filters of 

the safety-related exhaust system serving Safeguards given an ISLOCA at Surry. The potential 

exists for loading on the filters to be excessive in that the fans of the system shut down. A 

shutdown of the fans would leave the fission products lodged in the filters without a cooling flow 

of air. Decay heat generated by the fission products could be an issue with respect to 

overheating of the filters, the filter plenum, etc. The MELCOR modeling of the phenomena that 

could lead to filter loading shutting down the fans is complete and so the MELCOR calculations 

should indicate well a loss-of-cooling threat to the filters.

Appendix D describes the Surry Safeguards Ventilation System and the MELCOR modeling of it 

in detail.
4-40

4.12 Low Head Safety Injection Piping

The LHSI piping that would be subjected to RCS pressure, should the two serial check valves in 

any one of the three cold leg injection lines fail, extends backwards from the check valves into 

the Safeguards Area and then through the Containment Spray Pump Area and Main Steam Valve 

House well into the Auxiliary Building. Two MELCOR representations of LHSI piping were 

used in the ISLOCA analysis – a simple representation where the fluid volume and metal mass of 

the piping was unaccounted for and a detailed representation where these parameters were taken 

into account. In the simple representation, frictional losses, form losses, and critical flow areas 

were represented in a single flow path. No control volumes or heat structures were included in 

the simple representation. In the detailed representation, flow losses and critical flow areas were 

portioned among several flow paths and several control volumes and heat structures were 

included to account for the fluid volume and metal mass of the piping. The simple 

representation was used in the MELCOR full-plant ISLOCA calculation. The detailed 

representation was used in a MELCOR separate-effects calculation to estimate decontamination 

factors (DFs) associated with aerosol deposition for use in the full-plant calculation. Figure 4-15

illustrates the detailed MELCOR model of the LHSI piping and roughly identifies the attributes 

of the physical piping.

The location of the LHSI piping rupture could be in the Safeguards Area, the Containment Spray 

Pump Area, the Main Steam Valve House, or even in the Auxiliary Building. The break could 

be above the flood level or below it. If the break were in the Safeguards Area, some of the LHSI 

piping would be submerged throughout core degradation. If the break were elsewhere, none of 

the LHSI would be submerged. A submerged break location would support scrubbing of fission 

product aerosols and vapors by the pool. Submerged piping would stay relatively cold 

promoting fission product deposition in the piping.

From the check valves to isolation motor operated valve (MOV) 1890C, the 6” and 10” piping is 

Schedule 160 with 0.718” and 1.125” wall thicknesses, respectively, and a pressure rating higher 

than RCS operating pressure. From MOV 1890C back, none of the piping is rated strong enough 

to withstand RCS operating pressure. Between MOV 1890C and Flow Element (FE) 1945 for 

LHSI Train A and FE 1946 for LHSI Train B, the piping (10”) is Schedule 40 having a 0.365” 

wall thickness. The 8” piping in the Safeguards buildings and all of the way back to isolation 

points within the Auxiliary Building is Schedule 40 having a wall thickness of 0.322”. Between 

FEs 1945 and 1946 and the LHSI pump discharge check valves, the piping (10”) is Schedule 10 

with a 0.165” wall thickness. It is these relatively short sections of thin-walled 10” piping that 

are judged as most susceptible to rupture given the dual check valve failure of the postulated 

ISLOCA. This piping ranges in centerline elevation from 13’-9” to 15’-8” (i.e., 1’-9” to 3’-8” 

off the floor). Significantly more of this piping exists outside the pump cubicles than inside 

them and slightly more of the piping is at the 13’-9” elevation than at the 15’-8” elevation. 

Figure 4-16 shows the location of the different LHSI pipe sections in the Safeguards Area.

The ISLOCA break is assumed to occur centered 1’-9” above the floor with an area equivalent to 

the orifice area of FE 1945 or FE 1946 each of which are of diameter 7.1469”. Only one of the 

thin-walled 10” piping sections is assumed to rupture not both. The break is assumed to happen 

outside of a LHSI pump cubicle.
4-41

Optional MELCOR pool scrubbing logic (SPARC) was enabled in the flow path representing the 

LHSI piping (Flow Path 812) in the simple representation. This logic removes radionuclide 

aerosols and vapors from a gas as it flows through a pool of water. The flow area of Flow 

Elements 1945/1946 (7.1469” diameter) was specified. A single vent hole with horizontal 

orientation was called out.

Long standing experiments presented in Section 4.13 suggest that much of the fission-product 

aerosol carried into the LHSI piping, as a consequence of an ISLOCA, would deposit in the 

piping. The deposition would be by means of turbulent deposition and impaction. As identified 

in Section 4.13, MELCOR has been modified as part of the ISLOCA analysis effort to account 

for these mechanisms. A consequence of large deposits of radioactive material in the LHSI 

piping is the associated decay heat. MELCOR appropriately associates decay heat by 

radionuclide class and local mass wherever material is carried throughout the construct of a 

model. As such, the heating of heat structures representing LHSI piping segments, and hence the 

heating of radionuclide deposits on them, is accounted for as the radionuclides decay. The result 

of this heating, as evidenced by MELCOR, is substantial revaporization of more volatile 

radionuclides (e.g., CsI) from deposits in the LHSI piping. 

The energy emitted by fission product deposits in the LHSI piping would be in the form of beta 

particles and gamma radiation. The beta particles would be readily absorbed by the vapor in the 

pipe and by the steel pipe wall. The gamma radiation would be partially absorbed by the pipe 

wall. Trivial gamma radiation would be absorbed by the vapor. In modeling the LHSI piping, 

consistent with MELCOR’s default assumption, it was assumed that half of the energy emitted 

by fission product deposits is associated with beta particles and half with gamma radiation. It 

was further assumed that half of the beta particles are absorbed by the vapor and half by the pipe 

wall. In evaluating the gamma absorption in the pipe wall, gamma ray attenuation was assumed 

given by [58]:

I = I0e−μx

where,

I = attenuated radiation exposure rate

I0 = original radiation exposure rate

μ = linear attenuation coefficient (cm-1)

x = absorber thickness (cm-1)

Given the half-value layer for steel associated with gamma radiation from cesium-137 equal to 

1.6 cm [58], substituting into the equation above yields:

0.5 = e−μ ×1.6

Solving this equation for μ:

ln(0.5) = −μ × 1.6

μ = 0.43322
4-42

For the steel pipe then:

I

I0

= e−0.43322x

where x is the thickness of the pipe wall.

For the 6” Schedule 40 (0.719” wall) LHSI piping and the 10” Schedule 40 (0.365” wall) LHSI 

piping in Safeguards, I/Io is equal to 0.453 and 0.669, respectively. These values relate to 

gamma absorption fractions of 0.547 and 0.331 for the 6” and 10” pipe, respectively. 

Considering these gamma absorption fractions, the equal partitioning of decay energy between 

beta particles and gamma radiation, and the equal partitioning of beta deposition between the 

vapor in the pipe and the pipe wall, the distribution of decay energy produced by fission product 

deposits in the LHSI piping was specified in the MELCOR modeling as:

• 25% to the vapor in the pipe, 52.3% to the pipe wall, and 22.7% to the environment in the 

case of 6” pipe

• 25% to the vapor in the pipe, 41.6% to the pipe wall, and 33.4% to the environment in the 

case of 10” pipe

Appendix D describes the Surry LHSI piping and the MELCOR modeling of it in further detail.
4-43

Figure 4-15 Low Head Safety Injection Piping MELCOR Nodalization
4-44

Figure 4-16 Low Head Safety Injection Pipe Sections in Safeguards Area
4-45

4.13 Radionuclide Deposition in LHSI Piping by Turbulent Deposition and Impaction 

Deposition in the LHSI Piping

During an ISLOCA, fission product aerosols and non-radioactive aerosols released from the core 

and core debris will transport and deposit along the following path:

• the reactor vessel and reactor coolant system,

• the low head safety injection (LHSI) piping, and

• the Safeguards Area.

From the Safeguards Area, the fission products can reach the environment through openings in 

the Safeguards Area or through the ESF ventilation system.

MELCOR calculates aerosol deposition using the following models: gravitational settling, 

thermophoresis, diffuseophoresis, and Brownian motion. This is sufficient for flow regimes 

typically encountered in severe accident analysis where the flow of steam and hydrogen is slow 

and wafting (e.g., 1 meter/second). However, a peer review committee member commented that, 

in an ISLOCA, flow rates in the LHSI piping which has an internal diameter of 5.2 inches can be 

high turbulent and turbulent aerosol deposition and inertial deposition from flow irregularities 

can be important.

Turbulent deposition results when particle trajectories depart from flow path-lines. This 

departure is caused by the mass of an aerosol particle, which causes it to accelerate or decelerate 

more slowly than a gas particle (molecule). Because of the chaotic motions of a highly turbulent 

flow, turbulent deposition can dominate over other deposition mechanisms.

Deposition induced by flow irregularities is generally by impaction of aerosol particles against 

surfaces in the vicinity of the geometrical irregularity. Deposition is caused by inertial effects; 

aerosol particles are unable to follow flow streamlines because the aerosol is denser than the 

surrounding fluid. Some of the aerosol particles thus collide with a surface and deposit. 

Generally, larger aerosol particles deposit preferentially over smaller ones, which are able to 

follow fluid streamlines more closely.

As a result of this peer review comment, new aerosol deposition models were added to

MELCOR. Models for turbulent deposition in straight pipes and inertial deposition in elbows 

were taken from modeling developed by INEL [59] and from the VICTORIA code [60]. Adding 

models from these two sources enabled benchmarking of the two models against each other in 

addition to validating the results against experiments. Models for inertial deposition in a sudden 

contraction and a vena contracta (e.g., a venturi) were also added to MELCOR from the 

VICTORIA code.

The new MELCOR models for turbulent deposition in straight pipes and inertial deposition in 

elbows were validated against the results of the Light Water Reactor (LWR) Aerosol 

Containment Experiments (LACE) Project. The LACE project, organized by Electric Power 

Research Institute (EPRI), performed large-scale experiments to investigate aerosol behavior 

under simulated LWR accident conditions to provide a database for testing containment-aerosol 
4-46

and related thermal-hydraulic computer codes. The tests studied aerosol behavior under 

postulated severe accidents conditions not adequately addressed by previous test programs. The 

studied conditions included total containment bypass (i.e., ISLOCA). Individual LACE tests that 

studied ISLOCA conditions were CB-1, CB-2, CB-3, LA1, LA3A, LA3B, and LA3C. LACE 

reports for these tests are the following:

• “Aerosol Behavior Under LWR Containment Bypass Conditions—Results of Tests CB-1, 

CB-2 and CB-3,” LACE TR-001, November 1986

• “Aerosol Behavior in LWR Containment Bypass Piping—Results of LACE Test LA3,” 

LACE TR-011, July 1987

• “Summary of Posttest Aerosol Code Comparisons for LWR Aerosol Containment 

Experiment (LACE) LA1,” LACE TR-022, ORNL/M-365, October 1987

• “Summary of Posttest Aerosol Code-Comparisons Results for LWR Aerosol 

Containment Experiment (LACE) LA3,” LACE TR-024, ORNL/M-492, June 1988

A summary of the LACE project is given in “The LWR Aerosol Containment Experiments 

(LACE) Project, Summary Report” [61]. Validation of the new MELCOR models against the 

LACE tests that studied ISLOCA conditions is documented in Appendix D.

The new MELCOR models were applied in a separate-effects calculation that used boundary 

conditions from the integral full-plant calculation to estimate aerosol retention in the LHSI 

piping. The estimated aerosol retention for each fission product class was in turn used to specify 

a DF for each MELCOR radionuclide class (see Appendix B, Table B-1) as input to the integral 

full-plant calculation which represented the LHSI piping as a single junction. The nodalization 

and the results of the separate-effects calculation are described in Section 4.14 and Section 5.5.2, 

respectively. 

4.14 Analysis Methodology Involving Two MELCOR Models for LHSI Piping

Two MELCOR representations of the LHSI piping were utilized in the ISLOCA analysis:

1. A simple representation where the fluid volume and metal mass of the piping were

unaccounted, and 

2. A detailed representation where these parameters were taken into account. 

The simple representation, which amounted to a single flow path, was defined as part of the 

overall MELCOR model representing the RCS, Containment, Safeguards Area, etc. The detailed 

representation consisting of several small control volumes, flow paths, and heat structures was 

defined as a much smaller standalone MELCOR problem (i.e., separate-effects problem) absent 

of an RCS, Containment, Safeguards Area, etc. The detailed representation was developed to 

best exercise the aerosol deposition (i.e., turbulent deposition and impaction) modeling newly 

incorporated into MELCOR as part of the ISLOCA analysis effort. It allowed sensitivity 

investigations into phenomena such as heat loss from the piping and gamma ray transmission 
4-47

through it to be accomplished efficiently. Boundary conditions were imposed on the separateeffects problem reflective of the conditions in an initial overall ISLOCA calculation. 

Decontamination factors (DFs) determined from the separate-effects problem were then imposed 

on a final overall MELCOR calculation to account for the phenomena of turbulent deposition 

and impaction in the LHSI piping. The sequential steps taken to accomplish the ISLOCA 

calculation using the two MELCOR models were:

1. An initial overall ISLOCA calculation was carried out without particulate deposition 

accounted for in the LHSI piping. The calculation wrote files containing thermal 

hydraulic and aerosol transport information to be imposed on a separate-effects

(deposition) calculation. The specific information written to the files consisted of:

• Pressure in RCS Cold Leg A (CV 240)

• Vapor temperature in Cold Leg A

• Steam partial pressure in Cold Leg A

• Mole fractions of oxygen, nitrogen, hydrogen, carbon monoxide, and methane in 

the noncondensable gas content of Cold Leg A

• Pool level in the Safeguards Area

• Pool temperature in the Safeguards Area

• Integral masses of radionuclides by class and size bin transported into the LHSI 

piping

• Mass transport rates of nonradioactive aerosols released from control rods (e.g., 

Ag, In, and Cd) and from zirconium cladding (i.e., Sn) into the LHSI piping. 

These rates were not written by size bin since the allocation of aerosols is not 

available (i.e., not reported by MELCOR) by size bin for nonradioactive aerosols.

2. A separate-effects calculation was carried out with the thermal hydraulic information 

identified above imposed as boundary conditions. The integral radioactive aerosol 

masses transported into the LHSI piping in the initial overall ISLOCA calculation were 

differentiated to form rates, and radioactive aerosols were sourced into the separateeffects calculation at these rates by class and size bin. Nonradioactive aerosols were 

sourced into the separate-effects calculation at the rates from the overall ISLOCA 

calculation assuming a lognormal size distribution with a 2-micron mass mean diameter 

and a geometric standard deviation of two. Figure 4-17 through Figure 4-19 illustrate 

the imposition of thermal hydraulic conditions. Figure 4-20 and Figure 4-21 show the 

sourcing of aerosols managed in the separate-effects calculation.

3. DFs were determined from the results of the separate-effects calculation reflecting the 

efficiency of aerosol deposition in the LHSI piping by radionuclide class. The DFs were 
4-48

determined as the ratio of aerosol mass entering the piping to the difference between 

aerosol mass entering the piping and aerosol mass retained in the piping.

4. The DFs determined from the results of the separate-effects calculation by radionuclide 

class were imposed in a final overall ISLOCA calculation on the flow path representing 

the LHSI piping.

Figure 4-17 Cold Leg Pressure and Temperature Imposed on ISLOCA Deposition 

Calculation

300

400

500

600

700

800

900

1000

1100

1200

1300

0

50

100

150

200

250

300

350

400

6 12 18 24 30 36 42 48 54

Pressure (psia)

Temperature (K)

Time (hr)

Pressure

Temperature
4-49

Figure 4-18 Cold Leg Partial Pressure Imposed on ISLOCA Deposition Calculati

0.1

1

10

100

1000

6 12 18 24 30 36 42 48 54

Pressure (psia)

Time (hr)

Steam

O2

N2

H2

CO

CO2
4-50

380370360350340330320310300 12

12.5

13

13.5

14

14.5

15

15.5

16

16.5

6 12 18 24 30 36 42 48 54

Temperature (K)

Elevation (ft)

Time (hr)

Safeguards water lvlPipe bottom/topBreak bottom/topSafeguards floorSafeguards temperatureSaturation temperature

Figure 4-19 Safeguards Area Pool Level and Temperature Imposed on ISLOCA 

Deposition Calculation
4-51

Figure 4-20 Radioactive Aerosol Sourcing into ISLOCA Deposition Calculation

Figure 4-21 Masses of Radionuclides Sourced into the ISLOCA Deposition Calculation 

by Size

0.001

0.01

0.1

1

10

100

1000

6 12 18 24 30 36 42 48 54

Mass (kg)

Time (hr)

Cs

Ba

Te

Ru

Mo

Ce

La

U

Cd

Sn

CsI

Cs2MoO4

0

20

40

60

80

100

120

140

160

6 12 18 24 30 36 42 48 54

Mass (kg)

Time (hr)

1.00 - 1.48 microns

1.48 - 2.19

2.19 - 3.23

3.23 - 4.78

4.78 - 7.07

7.07 - 10.46

10.46 - 15.47

14.47 - 22.87

22.87 - 33.82

33.82 - 50.00

5-1

5. INTEGRATED THERMAL HYDRAULICS, ACCIDENT

PROGRESSION, AND RADIOLOGICAL RELEASE ANALYSIS

This Section describes the integrated self-consistent analysis of each scenario using the 

MELCOR code. The analysis includes calculations to confirm the table top exercise results to 

ensure that the timing and capacity of mitigation measures are capable of preventing core 

damage or delaying or reducing fission product releases. This analysis also includes sensitivity 

calculations without B.5.b mitigation measures. Version 1.8.6YR of the MELCOR severe

accident analysis code was used for the Surry analysis [10]. 

5.1 Long-Term Station Blackout

The long-term station blackout is assumed to be initiated by a seismic event. Section 5.1.1

presents the results of an unmitigated scenario with initially successful operator actions to 

depressurize the RCS and maintain TD-AFW flow.15 However, once the DC station batteries 

fail at 8 hours, no more operator actions are successful. For the mitigated scenario in 

Section 5.1.2, a portable emergency pump is connected to the RCS at 3.5 hours and a continuous 

supply of water is maintained.

The Surry SPAR model assessed an early RCP seal failure with the credited operator actions as 

less likely than a late failure (i.e., 20% versus 80% likely). Sections 5.1.3 and 5.1.4 present the 

results of unmitigated and mitigated sensitivity calculations that assess the impact of an early 

RCP seal failure.

5.1.1 Unmitigated Long-Term Station Blackout 

Table 5-1 summarizes the timing of the key events in the unmitigated LTSBO. As described in 

Section 3.1, the accident scenario initiates with a complete loss of all onsite and offsite AC 

power but the DC station batteries are available. The reactor successfully scrams and the 

containment isolates but all powered safety systems are unavailable except the TD-AFW. The 

timings of the key events are discussed further in Sections 5.1.1.1 and 5.1.1.2. However, it is 

worth noting that the fission product releases from the fuel do not begin until 16 hr and 

significant fission product releases to the environment do not begin until after 45 hr. 

 15 Some successful operator actions were credited but actions to successfully connect the emergency diesel pumps 

for vessel injection were not. Hence, unmitigated refers to failure of specific critical actions that increased the 

severity of the sequence.
5-2

Table 5-1 Timing of key events for unmitigated LTSBO

Event Description

Time

(hh:mm)

Initiating event

Station blackout – loss of all onsite and offsite AC power 00:00

MSIVs close

Reactor trip

RCP seals initially leak at 21 gpm/pump

00:00

TD-AFW auto initiates at full flow 00:01

First SG SRV opening 00:03

Operators control TD-AFW to maintain level 00:15

Operators initiate controlled cooldown of secondary at 

~100oF/hr 01:30

Upper plenum water level starts to decrease 01:57

Accumulators begin injecting 02:25

Vessel water level begins to increase 02:30

SG cooldown stopped at 120 psig to maintain TD-AFW flow 03:35

Emergency CST empty 05:08

DC Batteries Exhausted 08:00

S/G PORVs reclose 08:00

Pressurizer SRV opens 13:06

PRT failure 13:40

Start of fuel heatup 14:16

RCP seal failures (calculated) 14:46

First fission product gap releases 16:04

Creep rupture failure of the C loop hot leg nozzle 17:06

Accumulator empty 17:06

Vessel lower head failure by creep rupture 21:08

Debris discharge to reactor cavity 21:08

Cavity dryout 21:16

Containment at design pressure (45 psig) 28:00

Start of increased leakage of containment (P/Pdesign = 2.18) 45:32
5-3

5.1.1.1 Thermal-hydraulic Response

The responses of the primary and secondary pressures are shown in Figure 5-1. At the start of 

the accident sequence, the reactor successfully scrams in response to the loss of power. The 

main steam line and containment isolation valves close in response to the loss of power. The 

reactor coolant and main feedwater pumps also trip due to the loss of power. In response to the 

loss of the main feedwater, the turbine-driven auxiliary feedwater automatically starts. The 

TD-AFW initiates at full flow but is subsequently controlled by the operator after 15 min to 

maintain level. The TD-AFW restores the steam generator liquid levels by about 30 min and is 

throttled thereafter. After the closure of the main steam isolation valves, the secondary system 

quickly pressurizes to safety relief valve opening pressure, which causes the safety relief valves 

to open and then subsequently close when the closing pressure criterion is achieved. The relief 

flow through the SG SRVs is the principle primary system energy removal mechanism in the 

first 90 min. 

The heat removal through the steam generator depressurizes the primary system to 10.3 MPa by 

90 min. At 90 min, the operator starts a controlled (100o

F/hr) cooldown of the primary system 

by opening a steam generator power-operated relief valve (PORV). As the secondary pressure 

decreases, the saturation temperature of the water in the boiler section of the steam generator also 

decreases, which cools the primary system fluid. At about 3.5 hr, the steam generators reached 

0.93 MPa (120 psig), where the secondary system pressure was stabilized. The TD-AFW can 

achieve full flow (700 gpm) at 600 psig, but degrades thereafter. It is described to work below 

600 psig with an estimated lower limit of operability at 120 psig. Even with degraded 

performance at 120 psig, the TD-AFW adequately maintained the steam generator level until 

5 hr 8 min when the ECST empties. In the unmitigated sequence, no operator actions were 

credited to replenish the ECST inventory. After 5 hr 8 min, the steam generator level starts to 

decrease and is empty by 12 hr 18 min.

By depressurizing the primary system to 120 psig via the secondary system cooldown, several 

beneficial results were achieved. First, the leakage through the RCP seals decreased from 

21 gpm per pump at full operating pressure conditions to less than 7 gpm per pump. 

Furthermore, if a RCP seal should fail under these conditions, then the resulting leakage flow 

would be much lower than if the primary system pressure was not actively controlled to low 

pressure. Second, the accumulators begin injecting at 600 psig (4.1 MPa). The accumulators are 

a source of cold water to replace the losses due to RCP seal leakage and the volume shrinkage 

during the cooldown. By 8 hr, about 4500 gal had been discharged from each accumulator, or 

about two-thirds of the water inventory. Consequently, as shown in Figure 5-2, the inventory 

loss was minor during the first 8 hr.

At 8 hr, the station batteries were estimated to fail. At the same time, the steam generator relief 

valves closed and were no longer actively controlled. In response to the steam generator valve 

closure, both the primary and secondary systems rapidly pressurized to the secondary safety 

relief valve opening pressure. The primary system remained just above this pressure until about 

12 hr 18 min, when the steam generators boiled dry. Subsequently, the primary system 

pressurized to the pressurizer safety relief valve opening set point and began to relieve steam and 

water. The fluid in the vessel heated to saturation conditions and then swelled in response to the 

heatup. Once the pressurizer safety relief valves began cycling, a significant amount of fluid is 
5-4

vented out of the RCS and the vessel level dropped quickly (see Figure 5-2). The top of the fuel 

was uncovered by 14.3 hr and the core heatups began (see Figure 5-3). 

Shortly after the start of the core uncovery, the RCP seals failed when saturated water started 

flowing through the loop seals. The effective leak rate increased from 21 gpm at full operating 

pressure and temperature to 182 gpm at full operating pressure and temperature.16 Once the 

two-phase water level drops below the core plate, the decrease in the vessel two-phase level 

slows because the water level is below the bottom of the fuel (see Figure 5-2).

Similar to the STSBO (see Section 5.2.1), an in-vessel natural circulation flow develops between 

the hot fuel in the core and the cooler structures in the upper plenum. Hot gases rise out of the 

center of the core and rise into the upper plenum and return down the cooler peripheral sections 

of the core. Simultaneously, a natural circulation circuit develops between the vessel and the 

steam generator. Due to its close proximity to the hot gases exiting in the vessel, the hot leg 

nozzle at the carbon steel interface region to the stainless steel piping was first predicted to fail 

by creep rupture at 17 hr 6 min.

17

Following the accumulator injection, the decay heat from the fuel boiled away the injected water. 

By 18.2 hr, a large debris bed had formed in the center of the core. The debris bed continued to 

expand until 18.9 hr when all the fuel had collapsed and was resting on the core plate. The hot 

debris failed the lower core plate and fell onto the lower support plate, which failed at 19.9 hr. 

Following the lower support plate failure, the debris relocated onto the lower head. The small 

amount of remaining water in the lower head boiled away. As shown in Figure 5-4, the hot 

debris heated the inner surface of the lower head toward its melting temperature. As the heat 

transferred through the lower head, the lower head weakened and failed at 21 hr 8 min due (i.e., 

primarily due the thermal stress component in the thermal-mechanical failure criteria, due to the 

low differential pressure).

By 21.3 hr, nearly all the hot debris relocated from the vessel into the reactor cavity under the 

reactor vessel. The hot debris immediately boiled away the water in the reactor cavity and 

started to ablate the concrete. The ex-vessel core-concrete interactions (CCI) continued for 

the remainder of the calculation, which generated non-condensable gases. In addition, the hot 

gases exiting the reactor cavity and the radioactive heating from airborne and settled fission 

products steadily evaporated the water on the containment floor outside the reactor cavity from 

21.1 hr to 67 hr. The resultant non-condensable and steam production pressurized the 

containment (see Figure 5-5). At 45.5 hr, the containment failed due to liner tearing near the 

containment equipment hatch at mid-height in the cylindrical region of the containment. The 

containment continues to pressurize until the leakage flow balanced the steam and 

 16 The leak model is tuned to these values at normal operating conditions. In a transient calculation, the leakage 

flow rate changes as a function of subcooling, quality, and pressure. 17 Alternate failure locations could include the pressurizer surge line and the steam generator tubes. In the 

MELCOR calculation, the RCP seals had failed so hot gases were no longer flowing out the pressurizer when the 

core exit temperatures were hottest. Due to the relatively high pressure in the steam generators’ secondary side, 

the resultant thermal-mechanical stresses across the steam generator tubes were less severe than the hot leg 

nozzle. Consequently, the most vulnerable location was calculated to be the hot leg nozzle. The initial failure 

location for this scenario is also part of an on-going investigation of another research program in the NRC.
5-5

non-condensable gas generation. By 67 hr (2.8 days), all the water on the floor has evaporated. 

The containment depressurized thereafter due to only a smaller gas loading from the 

non-condensable gas generation. The conservatively assumed failure location was around the 

equipment hatch, which is located on the side of the containment without a surrounding building 

(e.g., the auxiliary or safeguards buildings) other locations such as personnel airlocks and 

penetrations would result in lower releases due to a transport and deposition inside adjacent 

buildings. Consequently, any released fission products are released directly to the environment.

A summary of the containment gas concentrations is shown in Figure 5-6. The steam 

concentration in the containment rapidly increases to ~52% following the pressurizer relief tank 

failure at 13:40. The steam concentration subsequently rises during the core degradation with a 

gradual build-up of hydrogen. At hot leg failure, the steam increases above 60% and remains 

above this level until the end of the calculation. Following vessel failure, water in the 

containment basement is evaporated, which further increases the steam concentration. At a 

steam concentration of 55% or higher, it is essentially impossible to ignite hydrogen or sustain a 

burn (see combustion modeling in Section 4.9). Similarly, the same experimental research 

showed that a minimum concentration of hydrogen and oxygen is also needed before hydrogen 

will ignite and burn [31]. The minimum hydrogen concentration for combustion with an active 

source (7%) and the maximum steam concentration (55%) are also shown in Figure 5-6 [13]. An 

active source for combustion might include debris, hot metal surfaces, hot aerosols, etc. 

Although the in-vessel hydrogen production is very significant, combustible conditions did not 

exist in the containment through 72 hours. The steam concentration is above the minimum 

threshold (55%) for combustion whenever any significant amount of hydrogen is present and the 

hydrogen concentration is always below the minimum ignition threshold (7%). 
5-6

Figure 5-1 Unmitigated LTSBO primary and secondary pressure history

Figure 5-2 Unmitigated LTSBO vessel two-phase coolant level history

Primary and Secondary Pressures

LTSBO - No Mitigation With Portable Equipment

0

2

4

6

8

10

12

14

16

18

0 6 12 18 24

Time (hr)

Pressure (MPa)

Pressurizer

SG A

SG C

SG B

Hot leg creep rupture

S/G PORVs 

reclose

Start RCS 

cooldown

Batteries 

exhausted Emerg CST 

empty

SG dryout

Vessel Water Levels

LTSBO - No Mitigation with Portable Equipment

-4

-2

0

2

4

6

8

10

0 3 6 9 12 15 18 21 24

Time (hr)

Two-Phase Level (m)

Accumulators

Hot leg creep 

rupture failure

Batteries exhausted S/G dryout

Start RCS 

cooldown ECST Empty

RCP Seal Failures

PORVs open

Accumulators

BAF

TAF

Lower head

Vessel top
5-7

Figure 5-3 Unmitigated LTSBO core temperature history

Figure 5-4 Unmitigated LTSBO lower head inner and outer temperature history

Peak Temperatures

LTSBO - No Mitigation with portable equipment

300

800

1300

1800

2300

2800

3300

0 6 12 18 24

Time (hr)

Temperature (K)

All fuel rods 

collapsed

Vessel failure 

at 21 hours

Peak debris 

temperature

Peak intact fuel 

cladding temperature

Hot leg failure 

at 17 hours

Lower Head Temperatures

LTSBO - No Mitigation with portable equipment

300

500

700

900

1100

1300

1500

1700

1900

2100

0 6 12 18 24

Time (hours)

Temperature (K)

Inside

Outside

Vessel failure 

at 21 hours

Hot leg failure 

at 17 hours
5-8

Figure 5-5 Unmitigated long-term station blackout containment pressure history

Figure 5-6 Unmitigated long-term station blackout containment gas concentrations

Containment Pressure

LTSBO - No Mitigation, Calculated RCP Seal Failure

0

100000

200000

300000

400000

500000

600000

700000

800000

900000

1000000

0 1 2 3 4 5 6 7

Time (days)

Pressure (Pa)

0

15

29

44

58

73

87

102

116

131

145

Pressure (psia)

Start leakage 

(45 h)

RCS hot leg 

creep 

rupture 

failure

Vessel 

failure

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 24 48 72

Mole Fraction (-)

Time (hr)

Surry Unmitigated Long-Term Station Blackout

Containment Gas Concentrations

Steam

Oxygen

Nitrogen

Hydrogen

Carbon Monoxide

Max Steam

Min Hydrogen
5-9

5.1.1.2 Radionuclide Release

The fission product releases from the fuel started following the first failures of the fuel cladding 

in the hottest rods at 16 hr 4 min, or about 1 hr 40 min after the uncovery of the top of the fuel 

rods. The in-vessel fission product release phase continued through vessel failure at 21.1 hr. 

Initially, the fission product releases from the fuel circulated through the primary system as well 

as being released to the containment through the pressurizer safety relief valves. The PRT

rupture disk opened about 30 min before the start of the fission product releases. Subsequently, 

the fission product releases exiting through the pressurizer relief valves were not well retained in 

the PRT because the pool was nearly saturated and the PRT rupture disk was open. Following 

vessel failure, the fission product releases continued from the ex-vessel fuel in the reactor cavity.

Figure 5-7 and Figure 5-8 show the fission product distributions of the iodine and cesium 

radionuclides that were released from the fuel, respectively. Approximately 99% of the iodine 

and cesium were released from the fuel prior to vessel failure while the remaining amount was 

released ex-vessel. At the time of the hot leg failure, only a small portion of the volatile 

radionuclides (4.3% of the noble gases and ~1% of the cesium and iodine) had been released 

from the fuel. The resultant blowdown of the vessel immediately discharged the majority of the 

release to the containment. Following the RCS blowdown after the hot leg nozzle failure, more 

radionuclides were released from the fuel as the core further degraded. At low pressure 

conditions, the fission products continued to circulate within the vessel and to the steam 

generators with a portion being deposited on the structural surfaces (i.e., 10% and 15% of the 

iodine and cesium are retained in the RCS). However, as shown in the figures, the majority of 

the released radionuclides went to the containment. Within 36 hr, most of the airborne fission 

products in the containment settled on surfaces. This was significant because the containment 

failure occurred at 45 hr 32 min. Consequently, there was little airborne mass that would be 

released to the environment.

The chemical form of the released iodine was cesium iodide, which was more volatile than the 

predominant form of the released cesium, which was cesium-molybdate (Cs2MoO4). As shown 

in the iodine distribution figure (see Figure 5-7), the in-vessel iodine mass was slightly

decreasing following vessel failure through 4 days. The slight decrease of mass represents a 

revaporization process of previously deposited radionuclides. The late in-vessel revaporization 

continued after containment failure and had a contribution to the environmental release. The 

primary thermal mechanisms for the revaporization of the iodine came from a natural circulation 

flow of hot gas. Very hot gases (i.e., from 990 K to >1616 K) flowed from the reactor cavity 

through the failed vessel lower head, through the reactor vessel, and out the failed hot leg nozzle. 

The combination of the decay heat and hot gases heated the deposited cesium iodide, which led 

to vaporization of some of the deposited cesium iodide. In addition, the cesium iodide 

dissociates when the cesium reacts with stainless steel surfaces. The residual elemental iodine 

vaporizes and contributes to the release. The natural circulation flow pattern subsequently 

vented the vaporized iodine radionuclides from the RCS to the containment and led to a small 

increase in the environmental source term. 

In contrast to the iodine response in Figure 5-7, the deposited cesium-molybdate was less volatile 

and remained deposited in the RCS. Except for inside the reactor cavity, the containment was 
5-10

cooler than vessel and well below conditions that would vaporize settled radionuclides. 

Consequently, none of the deposited radionuclides in the containment vaporized. 

Finally, Figure 5-9 summarizes the releases of the radionuclides to the environment. At 4 days, 

80% of the noble gases, 2.3% of the tellurium, 0.6% of the iodine, 0.75% of the radioactive 

cadmium, 0.08% of the cesium, 0.08% of the barium, and 0.04% of the radioactive tin had been 

released to the environment. All other releases were less than 0.02% of the initial inventory. As 

shown in the figure, there were some environmental releases prior to the containment failure at 

45.5 hr due to nominal leakages (i.e., design specification of 0.1% vol/day at the design 

pressure). After the failure of the containment, the releases to the environment increased 

sharply. Over the first day after containment failure, 50% of the airborne noble gases in the 

containment were released. Over the next day, only 30% more was released. 

Figure 5-7 Unmitigated LTSBO iodine fission product distribution history

Iodine Distribution

LTSBO - No Mitigation, Calculated RCP Seal Failure

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

Containment Deposited

Containment Airborne

In-Vessel Total

Env. Release

92% deposited 

in containment

Revaporization of 

deposited in-vessel 

0.6% environmental 

release

Revaporization of 

deposited cavity
5-11

Figure 5-8 Unmitigated LTSBO cesium fission product distribution history

Figure 5-9 Unmitigated LTSBO environmental release history of all fission products

Cesium Distribution

LTSBO - No Mitigation, Calculated RCP Seal Failure

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

Containment Deposited

Containment Airborne

In-Vessel Total

Env. Release

85% deposited 

in containment

0.08% environmental 

release

RCS Total = 15%

Fission Product Release to the Environment

LTSBO - Not mitigated with portable equipment

0.000001

0.00001

0.0001

0.001

0.01

0.1

1

0 1 2 3 4

Time (days)

Fraction release (-)

NG

I

Cs

Ba

Te

Ru

Mo

Ce

La

U

Cd

Sn

Containment 

failure
5-12

5.1.2 Mitigated Long-Term Station Blackout

Table 5-2 summarizes the timing of the key events in the mitigated LTSBO. As described in 

Section 3.1, the accident scenario initiates with a complete loss of all onsite and offsite power. 

The reactor successfully trips and the containment isolates but all powered safety systems are 

unavailable. The timings of the key events are discussed further in Section 5.1.2.1. Unlike the 

unmitigated LTSBO described in Section 5.1.1, the mitigated LTSBO credits the successful 

connection of the portable, diesel-driven (Kerr) pump to three drain lines of the LHSI piping to 

the RCS.18 The Kerr pump is a positive displacement pump with a flow rate of 65 gpm at 

500 rpm, which was determined with pump data obtained from the Kerr Pump Corporation.

19 

The Kerr pump takes suction from the refueling water storage tank (RWST), which has a 

387,000 gal capacity. The refueling water storage tank could be refilled as necessary. The 

sequence of events is identical to the unmitigated LTSBO until 3 hr 30 min when the Kerr pump 

starts operating. The Kerr pump operation starts prior to any core degradation. The 

emergency Kerr pump is effective at maintaining the vessel water level above the top of the fuel 

for the duration of the sequence. In fact, the pump was throttled to a small fraction of its rated 

flow. 

 18 The utility has a 3-way connection from the Kerr pump to the three drain lines of the LHSI piping to the RCS. 19 An effective flowrate of 65 gpm at 500 rpm was determined using pump data provided by the Kerr Pump 

Corporation that includes 95% and 92% mechanical and volumetric efficiencies.
5-13

Table 5-2 Timing of key events for mitigated LTSBO

Event Description

Time

(hh:mm)

Initiating event

Station blackout – loss of all onsite and offsite AC power 00:00

MSIVs close

Reactor trip

RCP seals initially leak at 21 gpm/pump

00:00

TD-AFW auto initiates at full flow 00:01

Operators control TD-AFW to maintain level 00:15

Vessel water level drains into upper plenum 00:30

Operators initiate controlled cooldown of secondary at 

~100oF/hr 01:30

Upper plenum water level starts to decrease 01:57

Accumulators begin injecting 02:25

Vessel water level beings to increase 02:30

Start emergency diesel pump injection into RCS 03:30

SG cooldown stopped at 120 psig to maintain TD-AFW flow 03:35

DC station batteries fail but operator actions continue to control 

the secondary pressure at 120 psi and maintain TD-AFW flow 08:00

Level maintained at the cold leg elevation with emergency pump 

throttled to 5% >03:30

5.1.2.1 Thermal-Hydraulic Response

The progression of events in the mitigated LTSBO is identical to the unmitigated LTSBO as 

described in Section 5.1.1 through the first 3 hr 30 min. In particular, the operators take actions 

to throttle the TD-AFW to maintain a normal level in the steam generators and perform a cool 

down of the RCS using the steam generator relief valves. Similar to the unmitigated case, the 

accumulators begin injecting at 2 hr 25 min in response to the decrease in the primary system 

pressure. It is estimated that the operators could begin RCS injection using the portable, 

diesel-driven Kerr pump by 3 hr 30 min. 

At the time the emergency pump is ready for injection, the primary system pressure is 2.0 MPa

(278 psig) or well within the pressure head capacity of the Kerr pump (see Figure 5-10). Similar 

to the unmitigated LTSBO, the secondary system is depressurized to 120 psi, or the lower limit 

of operability for the TD-AFW. Due to the RCP seal leakage and liquid volume shrinkage from 

the cool down, the vessel level initially decreased but started to recover after 2 hr 25 min

following the start of the accumulator injection (see Figure 5-11). The peak fuel cladding 

temperature and vessel lower head followed the primary system liquid temperature, which was 

steadily cooled down by the steam generators (see Figure 5-12 and Figure 5-13, respectively). 
5-14

At 3 hr 30 min, the emergency injection begins and supplements the RCS inventory make-up 

with the accumulators. For the purposes of the calculation, a simple control system was created 

to ramp the Kerr pump flow (i.e., maximum flow rate of 65 gpm at 500 rpm) based on the reactor 

vessel level indication system. The flow increases from 5% (3.2 gpm) to 65 gpm (100%) if the 

vessel water level drops below the hot leg elevation. Since the level remained above the bottom 

of the hot leg, the flow was throttled to 3.2 gpm for the entire transient (see Figure 5-15). 

Although there was a small mismatch between the leakage flow and the emergency injection 

flow, the swollen level in the vessel remained relatively constant (see Figure 5-11). Because of 

the relatively low level in the vessel, there was minimal excess spillage of the injected water into 

the cold leg, which would eventually be lost due to seal leakage.

 

Finally, since heat removal from the RCS was maintained for this sequence, there was not a 

significant challenge to the containment. As seen in Figure 5-14, the containment pressure only 

rose slightly due to heat losses from the RCS.

There were several lessons learned while investigating the mitigation of the LTSBO. First, the 

operator action to reduce the primary system pressure to the threshold of the TD-AFW operation 

allowed the maximum injection from the accumulators (i.e., two thirds of their liquid inventory). 

The accumulator flow was significant in the short-term restoration of the vessel liquid inventory. 

Hence, the depressurization to 120 psi maintained TD-AFW flow but allowed for significant 

accumulator flow. Second, the reduction of the primary system pressure reduced the RCP seal 

leakage flow from 21 gpm per pump to less than 3 gpm per pump. Furthermore, if a RCP seal 

should fail under these conditions, then the resulting leakage flow would be much lower than if 

the primary system pressure was not actively controlled to low pressure. Third, the emergency 

pumps have excess capacity to maintain long-term make-up. However, only a small fraction of 

the rated capacity is needed to maintain the level in the vessel at the hot leg elevation. Fourth, 

operator actions were required to replenish the water supply to the ECST for the TD-AFW (i.e., 

exhausted after 5.2 hours and required 545,000 gal for 4 days) but not the RWST for the 

emergency RCS injection (i.e., at 87% inventory after 4 days). Finally, if successful operation 

actions were taken to replenish the ECST water supply and maintain the steam generator 

pressure at 120 psi, then considerably more time is available to establish the RCS emergency 

injection. 
5-15

Figure 5-10 Mitigated long-term station blackout primary and secondary pressure 

history

Figure 5-11 Mitigated long-term station blackout vessel two-phase coolant level

0

290

580

870

1160

1450

1740

2030

2321

2611

0

2

4

6

8

10

12

14

16

18

0 6 12 18 24

Pressure (MPa)

Pressure (psi)

Time (hr)

Primary and Secondary Pressures

LTSBO - Mitigated with Portable Equipment

Pressurizer

SG A

SG B

SG C

Start RCS 

cooldown

at 1.5 hr

Start RCS 

injection

at 3.5 hr

-4

-2

0

2

4

6

8

10

0 6 12 18 24

Two-Phase Level (m)

Time (hr)

Vessel Water Levels

LTSBO – Mitigation with Portable Equipment

Accumulators

BAF

TAF

Lower head

Vessel top

Start RCS 

injection 

with portable 

pump

Start RCS 

cooldown
5-16

Figure 5-12 Mitigated long-term station blackout core temperature history

Figure 5-13 Mitigated long-term station blackout lower head inner and outer 

temperature history

300

400

500

600

700

800

900

1000

0 6 12 18 24

Temperature (K)

Time (hr)

Peak Cladding Temperatures

LTSBO - Mitigated with Portable Equipment

300

350

400

450

500

550

600

0 6 12 18 24

Temperature (K)

Time (hr)

Vessel Lower Head Temperatures

LTSBO - Mitigated with Portable Equipment

Inside

Outside
5-17

Figure 5-14 Mitigated long-term station blackout containment pressure history

Figure 5-15 Mitigated LTSBO vessel emergency make up and pump seal leakage flows

5.1.2.2 Radionuclide Release

There was no fission product release for the mitigated LTSBO scenario.

0

4

9

13

17

22

0

0.03

0.06

0.09

0.12

0.15

0 6 12 18 24

Pressure (MPa)

Pressure (psi)

Time (days)

Containment Pressure 

LTSBO - Mitigated with Portable Equipment

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0 6 12 18 24

Flow Rate (kg/s)

Time (hr)

Seal Leakage versus Emergency Diesel Injection Flow

LTSBO - Mitigated with Portable Equipment

Total Seal Leakage

Emergency Diesel Injection (Kerr)

Start RCS injection at 3.5 hr

Kerr pump throttled to 5% 

of capacity
5-18

5.1.3 Unmitigated Long-Term Station Blackout with Early RCP Seal Failure

Table 5-3 summarizes the timing of the key events in the unmitigated LTSBO with early and late 

RCP seal failure. As described in Section 3.1, the accident scenario initiates with a complete 

loss of all onsite and offsite AC power but the DC station batteries are available. The reactor 

successfully scrams and the containment isolates but all powered safety systems are unavailable 

except the TD-AFW. 
5-19

Table 5-3 Comparison of the timings of key events for unmitigated LTSBO with and 

without early RCP seal failures

Event Description

Early Seal 

Failure

Time

(hh:mm)

Late Seal 

Failure

Time

(hh:mm)

Initiating event

Station blackout – loss of all onsite and offsite AC power 00:00 00:00

MSIVs close

Reactor trip

RCP seals initially leak at 21 gpm/pump

00:00 00:00

TD-AFW auto initiates at full flow 00:01 00:01

First SG SRV opening 00:03 00:03

Assumed early RCP seal failures, 182 gpm/pump at full-pressure 00:13 n/a

Operators control TD-AFW to maintain level 00:15 00:15

Operators initiate controlled cooldown of secondary at 

~100oF/hr 01:30 01:30

Upper plenum water level starts to decrease 00:33 01:57

Accumulators begin injecting 02:13 02:25

Vessel water level begins to increase n/a 02:30

SG cooldown stopped at 120 psig to maintain TD-AFW flow 03:35 03:35

Emergency CST empty 05:29 05:08

DC Batteries Exhausted 08:00 08:00

S/G PORVs reclose 08:00 08:00

Pressurizer SRV opens n/a 13:06

PRT failure n/a 13:40

Start of fuel heatup 09:38 14:16

RCP seal failures (calculated) n/a 14:46

First fission product gap releases 11:03 16:04

Creep rupture failure of the C loop hot leg nozzle 12:29 17:06

Accumulator empty 12:29 17:06

Vessel lower head failure by creep rupture 14:26 21:08

Debris discharge to reactor cavity 14:26 21:08

Cavity dryout 14:36 21:16

Containment at design pressure (45 psig) 42:08 28:00

Start of increased leakage of containment (P/Pdesign = 2.18) 55:40 45:32
5-20

5.1.3.1 Thermal-hydraulic Response

The responses of the primary and secondary pressure systems are shown in Figure 5-16. The 

system pressure responses through the first 10 hours are very similar to the unmitigated case 

without an early RCP seal failure (see Figure 5-1). However, as will be discussed below, the 

early RCP seal failure had an important impact on primary system pressure response after 

10 hours. As shown in Figure 5-17, the early RCP seal failures occur at 13 minutes versus 

14 hours 46 minutes in the late failure case. Consequently, the early leakage flowrate is much 

higher than the late failure case. The high RCP seal leakage causes a faster decrease in the 

vessel water inventory (see Figure 5-18), earlier core degradation and start of fission product 

release, and earlier hot leg creep rupture failure and vessel failure (see Table 5-3). 

The water level decreased more quickly in the early seal failure case but stabilized near 5.2 m 

after the accumulators discharged. At 5.2 m, the water level in the core is below the hot and cold 

leg elevations, so the seal leakage had a minor effect of the vessel water inventory. The 

discharge of the accumulator water partially subcooled the core. The fluid in the core began 

boiling again but reflux cooling in the SG kept the pressure and water level constant. Once the 

SG PORVs closed at 8 hours with the DC battery failure, the RCS pressurized; the inventory loss 

out the break increased; and the vessel water level began decreasing.

A comparison of the containment pressure responses for the cases with and without early RCP 

seal failures is shown in Figure 5-19. Although vessel failure occurs earlier in the case with 

early RCP leakage, the subsequent containment pressurization is slower. The long-term 

containment pressurization is due to two components: (1) non-condensable gas production from 

core-concrete interactions and (2) vaporization of water in the containment. The 

non-condensable gas production was similar in the two cases. However, the case with late 

failure of the RCP seals had more water available on the containment floor for vaporization. 

Since all the water on the containment floor eventually vaporized due to connectivity with the 

ex-vessel debris in the cavity, the total steam production was higher in the late RCP seal failure 

case. 

Although the total amount of water discharged into the containment was approximately equal for 

both cases, the distribution of the water was different. In the early seal failure case, 217,000 kg

of water was discharged through the RCP seals versus only 65,000 kg in the late seal failure 

case. Due to the interconnectivity of the containment compartments and the specific energy of 

the fluid being released, a significant portion of the water in the early seal failure case was held 

up in the pressurizer relief tank compartment, which did not directly communicate with the 

debris in the cavity (i.e., the location of the long-term, ex-vessel core debris heat source). In 

contrast, once the PRT rupture disk failed at 13.6 hr in the late RCP seal case, continued high 

energy steam from pressurizer safety valve cycling dumped 122,000 kg into the containment, 

which condensed on structures in the containment and mostly drained to the containment floor 

(i.e., rather than being retained in the pressurizer relief tank compartment). Therefore, more 

water was available in the late seal failure case for long-term boiling by the debris in the cavity.

It is also interesting to look at similarities and differences in the events that affected the 

short-term containment pressurization. In the late RCP seal failure case, the containment 

pressurization had three events that caused step increases in the containment pressure. First, the 
5-21

failure of the pressurizer relief tank caused the first significant pressurization of the containment 

at 13 hr 40 min (0.57 days). Some hydrogen burns also cause short-term pressurizations 

following this event. However, the containment pressure returned to the pre-burn pressure after 

the burn. The early RCP seal leakage case did not have this event (i.e., failure of the pressurizer 

relief tank). Second, the hot leg failed by creep rupture and suddenly increased the pressure in 

the containment in both cases (i.e., the net pressure increase following the coincidental hydrogen 

burn at hot leg failure). Since the hot leg failure in the late RCP seal failure case occurred from 

16.2 MPa versus 6.4 MPa in the early RCP seal failure case, significantly more energy was

released to the containment as evidenced by the rapid pressure response rise following hot leg 

creep rupture (see Figure 5-19). Finally, both cases had similar pressurizations following vessel 

failure when the core debris rapidly evaporated the water in the reactor cavity. In summary, the 

first rapid containment pressurization event only occurred in the late RCP failure case and the 

second event was more severe in the late RCP failure case (i.e., the net pressure increase after the 

hydrogen deflagration spike). The third event was longer duration in the late RCP failure case.

In an integral sense, a portion of the overall system energy generated in the late RCP seal failure 

case was released to the containment 4.5 hours after the early RCP failure case (i.e., the 

difference in the timings of the hot leg creep rupture failures, see Table 5-3)

20. The additional 

energy storage in the primary system in the late seal failure case results in a higher pressure and 

temperature of the gas in the primary system at the time of the leg creep rupture. In contrast, a 

portion of the early RCP seal failure decay energy generated over the time period between hot 

leg creep ruptures in the two cases (i.e., 12 hr 29 min to 17 hr 6 min) is absorbed into the 

concrete by core-concrete interactions in the early RCP seal failure case (i.e., from 14 hr 26 min

to 17 hr 6 min). The result was a higher heat load to the containment atmosphere in the early 

RCP seal failure case, which resulted in a faster evaporation rate of the water on the containment 

floor. 

In addition to the faster rate of evaporation, more water was available to evaporate in the late 

RCP seal failure case (see Figure 5-20). Consequently, the resulting pressurization was faster 

and longer. The timings of the containment failure were 55.7 and 45.5 hours for the early and 

late RCP seal failure cases, respectively. Due to the additional water for vaporization in the late 

seal failure case, the resulting containment pressurization was not only faster but also resulted in 

a higher pressure (see Figure 5-19). The importance of the water vaporization is demonstrated 

by comparing Figure 5-19 and Figure 5-20. The containment depressurization does not begin 

until all the water is evaporated. Hence, the water vaporization was a significant component of 

the containment pressurization relative to the non-condensable CCI gas generation. 

Finally, due to higher peak containment pressure in the late RCP seal failure case, the resultant 

containment leakage area was also slightly higher, 0.0053 m2 (8.3 in2

) versus 0.0028 m

2 (4.4 in2

). 

Consequently, the long-term depressurization rate following the containment failure was faster in 

the late RCP seal failure case.

 20 The heat removal by the secondary system ended at approximately the same time in both cases. The energy 

released through the seal failure was lower in the late RCP seal failure until 14 hr 46 min, when the seal failed.
5-22

Figure 5-16 Unmitigated long-term station blackout primary and secondary 

pressure history

Figure 5-17 Comparison of the unmitigated long-term station blackout RCP seal leakages 

with and without early RCP seal failures

Primary and Secondary Pressures

LTSBO - No Mitigation, RCP Seal Failure at 13 min

0

2

4

6

8

10

12

14

16

18

0 12 24 36 48

Time (hr)

Pressure (MPa)

Pressurizer

SG A

SG C

SG B

Hot leg C creep rupture

S/G PORVs 

reclose

Total Seal Leakage

LTSBO - No Mitigation, Early and Late RCP Seal Failure Cases

0

50

100

150

200

250

300

350

400

450

500

0 6 12 18 24

Time (hr)

Flow Rate (gpm)

Early Seal Failure

Seal failure at 13 min Late Seal failure

Seal failure at 14.8 hr
5-23

Figure 5-18 Comparison of the unmitigated long-term station blackout vessel level 

responses with and without early RCP seal failures

Figure 5-19 Comparison of the unmitigated long-term station blackout containment 

pressure responses with and without early RCP seal failures

Vessel Water Level

LTSBO with and without Early RCP Seal Failure

-4

-2

0

2

4

6

8

10

0 3 6 9 12 15 18 21 24

Time (hr)

Two-Phase Level (m)

Accumulators

Hot leg creep 

rupture failure

ECST Empty

RCP Seal Failures at 13 min

BAF

TAF

Lower head

Vessel top

Vessel 

failure

Late RCP Seal Failure Early RCP 

Seal Failure

Hot leg creep 

rupture failure

Batteries exhausted 

Start RCS 

cooldown

Vessel 

failure

Containment Pressure

LTSBO with and without Early RCP Seal Failure

0

100000

200000

300000

400000

500000

600000

700000

800000

900000

1000000

0 1 2 3 4 5 6 7

Time (days)

Pressure (Pa)

0

15

29

44

58

73

87

102

116

131

145

Pressure (psia)

Start leakage

RCS hot leg 

creep 

rupture 

failure

Vessel 

failure

Late RCP Seal Failure

Early RCP Seal Failure

Pressurizer relief tank failure
5-24

Figure 5-20 Comparison of the unmitigated long-term station blackout containment 

water pool masses with and without early RCP seal failures

5.1.3.2 Radionuclide Release

The environmental releases for the noble gases, iodine, and cesium for the two cases are shown 

in Figure 5-21 through Figure 5-23, respectively. The radionuclide behavior was similar to the 

late RCP seal failure responses described in Section 5.1.1. However, the magnitude and timing 

of the releases from the late RCP seal failure case bounded the response from the early RCP seal 

failure. The early RCP failure releases were smaller and delayed because: (a) the containment 

failure was later (2.3 days versus 1.9 days), (b) the containment failure area and associated 

depression rate were smaller (0.0028 m

2 versus 0.0053 m2

), and (c) the time for settling between 

hot leg failure and containment failure was longer (43 hours versus 28 hours). However, both

cases had small environmental releases of iodine and cesium due to the late containment failure 

timing.

Containment Floor Water Mass

LTSBO with and without Early RCP Seal Failure

0

50000

100000

150000

200000

0 1 2 3

Time (days)

Mass (kg)

Late RCP Seal Failure

 Early RCP Seal Failure

PRT failure

Hot leg creep 

rupture failure

Hot leg creep 

rupture failure

Vessel failure

Vessel failure
5-25

Figure 5-21 Comparison of the unmitigated long-term station blackout noble gas releases 

to the environment with and without early RCP seal failures

Figure 5-22 Comparison of the unmitigated long-term station blackout iodine releases to 

the environment with and without early RCP seal failures

Noble Gas Release to the Environment

LTSBO with and without Early RCP Seal Failure

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4 5 6 7

Time (days)

Fraction of Inventory (-)

Late RCP Seal Failure

Early RCP Seal Failure

Late RCP Seal Failure Early RCP Seal Failure

Iodine Release to the Environment

LTSBO with and without Early RCP Seal Failure

0

0.01

0.02

0.03

0.04

0.05

0 1 2 3 4 5 6 7

Time (days)

Fraction of Inventory (-)

Late RCP Seal Failure

Early RCP Seal Failure

Late RCP Seal Failure

Early RCP Seal Failure
5-26

Figure 5-23 Comparison of the unmitigated long-term station blackout cesium releases to 

the environment with and without early RCP seal failures

5.1.4 Mitigated Long-Term Station Blackout with Early RCP Seal Failure

Table 5-4 summarizes the timing of the key events in the mitigated LTSBO. As described in 

Section 3.1, the accident scenario initiates with a complete loss of all onsite and offsite power. 

The reactor successfully trips and the containment isolates but all powered safety systems are 

unavailable. The loss of seal-cooling to the RCP causes relatively warm RCS fluid to flow into 

the RCP seal cavity. At 13 minutes, the RCP seal cavity fills with hot RCS water and the SPAR 

model assess a 20% probability of RCP seal failure, which is used in this case. Consequently, all 

three RCPs are specified to have seal failures that increases their leakage to a nominal value of 

181 gpm/RCP-pump at full-power conditions. The mitigated LTSBO credits the successful 

connection of the portable, diesel-driven Kerr pump to three drain lines of the LHSI piping to the 

RCS. The Kerr pump is a positive displacement pump with a flow rate of 65 gpm at 500 rpm,

which was determined with pump data obtained from the Kerr Pump Corporation.

21 The Kerr 

pump takes suction from the refueling water storage tank, which has a 387,000 gal capacity and 

could be refilled as necessary. In addition, a portable power supply was available to maintain the 

secondary cooldown after the DC batteries fail.

The sequence of events is similar to the unmitigated LTSBO with early RCP seal failure until 

3 hr 30 min when the Kerr pump starts operating. The Kerr pump operation starts prior to any 

core degradation (see Table 5-4) and is effective at maintaining the vessel water level above the 

 21 An effective flowrate of 65 gpm at 500 rpm was determined using pump data provided by the Kerr Pump 

Corporation that includes 95% and 92% mechanical and volumetric efficiencies.

Cesium Release to the Environment

LTSBO with and without Early RCP Seal Failure

0

0.01

0.02

0.03

0.04

0.05

0 1 2 3 4 5 6 7

Time (days)

Fraction of Inventory (-)

Late RCP Seal Failure

Early RCP Seal Failure

Late RCP Seal Failure

Early RCP Seal Failure
5-27

top of the fuel for the duration of the sequence. Initially, the Kerr pump flow rate is 3.2 gpm. 

However, at 11.6 hr into the event sequence, the water level in the vessel falls below the bottom 

of the hot leg as observed on the reactor vessel level instrumentation system (RVLIS). 

Subsequently, the Kerr pump flow rate was increased up to 65 gpm and provided sufficient 

make-up to balance the loss of coolant through RCP seal leakage. The timings of the key events 

are discussed further in Section 5.1.2. 

Table 5-4 Comparison of the timings of key events for mitigated LTSBO with and 

without early RCP seal failures

Event Description

Early Seal 

Failure

Time

(hh:mm)

Late Seal 

Failure

Time

(hh:mm)

Initiating event

Station blackout – loss of all onsite and offsite AC power

00:00 00:00

MSIVs close

Reactor trip

RCP seals initially leak at 21 gpm/pump

00:00 00:00

TD-AFW auto initiates at full flow 00:01 00:01

RCP pump seals fail, 182 gpm/pump at full-pressure 00:13 n/a

Operators control TD-AFW to maintain level 00:15 00:15

Vessel water level drains into upper plenum 00:30 00:30

Operators initiate controlled cooldown of secondary at 

~100oF/hr 01:30 01:30

Upper plenum water level starts to decrease 01:13 01:57

Accumulators begin injecting 02:15 02:20

Vessel water level begins to increase n/a* 02:25

Emergency diesel pump injection to RCS 03:30 03:30

SG cooldown stopped at 120 psig to maintain TD-AFW flow 03:35 03:35

DC station batteries fail but operator actions continue to control 

the secondary pressure at 120 psi and maintain TD-AFW flow 08:00 08:00

Water level falls below the bottom of the hot leg

Emergency diesel pump injection to RCS increases to 65 gpm

11:36 n/a

Water level stabilizes at 8” above the bottom of the hot leg 14:00 n/a

* In the case early RCP seal leakage, the vessel water level stops decreasing following accumulator injection and 

remains relatively steady at ~1 m above the top of the fuel to the end of the calculation. The emergency pumps 

are adequate to maintain level at the bottom of the cold and hot legs. However, the RCS leakage through the three

failed RCP seals is too high to raise the level above the hot and cold leg elevation.
5-28

5.1.4.1 Thermal-Hydraulic Response

The responses of the primary and secondary pressures are shown in Figure 5-24. The system 

pressure responses are very similar to the mitigated case without an early RCP seal failure (see 

Figure 5-10). The impact of the early RCP seal failure did not have a significant impact on 

pressure response. More important was the successful operator actions to depressurize the 

primary system using the secondary system PORVs. As shown in Figure 5-25, the early RCP 

seal failures occur at 13 minutes. However, the total RCP seal leakage was less than 70 gpm

after 2 hours because of the successful RCS depressurization. The RCP seals were not predicted 

to fail in the mitigated case without early failure due to adequate subcooling.

As discussed in Section 5.1.2, the operators take actions to throttle the TD-AFW to maintain a 

normal level in the steam generators and perform a cooldown of the RCS using the steam

generator relief valves. The accumulators begin injecting at 2 hr 15 min following the decrease 

in the primary system pressure below the pressure of the accumulators. The operators stage the 

portable, diesel-driven Kerr pump and begin injecting at 3 hr 30 min.

At the time the emergency pump is ready for injection, the primary system pressure is 2.0 MPa

(278 psig) or well within the pressure head capacity of the Kerr pump (see Figure 5-24). 

The secondary system is depressurized to 120 psi, or the lower limit of operability for the 

TD-AFW. As shown in Figure 5-26, the vessel level decrease in the early RCP failure case was 

more severe than the late RCP seal failure case. However, accumulator injection maintained the 

vessel level in the upper plenum until the emergency portable in injection started. Consequently, 

there was no early fuel uncovery or fuel heatups.

At 3 hr 30 min, the emergency injection begins and supplements the RCS inventory make-up 

with the accumulators. For the purposes of the calculation, a simple control system was created 

to ramp the Kerr pump flow (maximum flow rate of 65 gpm at 500 rpm) based on the reactor 

vessel level indication system. The flow increases from 5% (3.2 gpm) to 65 gpm (100%) if the 

vessel water level drops below the hot leg elevation. When the level dropped below the bottom 

of the hot leg at 11.6 hours, the flow was increased to 65 gpm for the remainder of the transient 

(see Figure 5-15). The emergency diesel pump restored the vessel water level to ~8 inches above 

the bottom of the hot leg by 14 hours. Because of the relatively low level in the vessel, there 

was minimal excess spillage of the injected water into the cold leg, which would eventually be 

lost due to seal leakage.

 

The lessons learned that were described in Section 5.1.2 are also applicable in the mitigated early 

RCP failure case. There are the following quantitative and timing differences for the early RCP 

seal failure case versus the late RCP seal failure case insights that were presented in 

Section 5.1.2. First, the reduction of the primary system pressure reduced the total RCP seal 

leakage flow from a nominal value of 182 gpm per pump to approximately the pumping capacity 

of the high-head emergency diesel pump (i.e., 65 gpm) after 4 hours. The reduction in RCP 

leakage delays core damage and increases the likelihood that the operators will be able to prevent 

core damage. The depressurization also allows use of a lower head pump to provide RCS 

make-up. Second, operator actions were required to replenish the water supply to the ECST for 

the TD-AFW (i.e., exhausted after 5.8 hours and required 420,000 gal for 72 hours) and the 
5-29

RWST for the emergency RCS injection (i.e., exhausted after 57 hours). These actions were 

necessary to maintain primary and secondary coolant injection. 

Figure 5-24 Primary and secondary pressure responses for the mitigated long-term 

station blackout with early RCP seal failure

0

290

580

870

1160

1450

1740

2030

2321

2611

0

2

4

6

8

10

12

14

16

18

0 6 12 18 24

Pressure (MPa)

Pressure (psi)

Time (hr)

Primary and Secondary Pressures

LTSBO - Mitigated with Portable Equipment

Pressurizer

SG A

SG B

SG C

Start RCS 

cooldown

at 1.5 hr

Start RCS 

injection

at 3.5 hr
5-30

Figure 5-25 Comparison of the mitigated long-term station blackout RCP seal leakages 

with and without early RCP seal failures

Figure 5-26 Comparison of the mitigated long-term station blackout vessel level with and 

without early RCP seal failures

0

100

200

300

400

500

600

0 6 12 18 24

Flow Rate (gpm)

Time (hr)

Total Seal Leakage

LTSBO - Mitigation with Portable Equipment, Early and Late RCP Seal Failures

Early Seal Failure

Late Seal Failure

Start 150 gpm 

RCS injection 

at 3.5 hr

-4

-2

0

2

4

6

8

10

0 8 16 24 32 40

Two-Phase Level (m)

Time (hr)

Vessel Water Level

LTSBO with and without Early RCP Seal Failure

Accumulators

ECST refilled

RCP Seal Failures at 13 min

BAF

TAF

Lower head

Vessel top

8 hr battery life No Early RCP Seal Failure

Early RCP Seal Failure
5-31

Figure 5-27 Comparison of the mitigated long-term station blackout portable pump 

injection rates with and without early RCP seal failures

5.1.4.2 Radionuclide Release

There was no fission product release for the mitigated LTSBO scenario with early RCP failures.

5.2 Short-Term Station Blackout 

The STSBO is assumed to be initiated by a seismic event. Section 5.2.1 presents the results of 

an unmitigated scenario with no successful operator actions. For the mitigated scenario in 

Section 5.2.2, a portable emergency pump is connected to the containment spray system at 

8 hours and available to inject 1,000,000 gallons.

5.2.1 Unmitigated Short-Term Station Blackout 

Table 5-5 summarizes the timing of the key events in the unmitigated STSBO. As described in 

Section 3.2.1, the accident scenario initiates with a complete loss of all onsite and offsite power

and failure of the ECST. The reactor successfully trips and the containment isolates but all 

powered safety systems are unavailable. The timings of the key events are discussed further in 

Sections 5.2.1.1 and 5.2.2.1. However, it is worth noting that fission product releases from the 

fuel do not begin until 2 hr 57 min and significant fission product releases to the environment do 

not begin until 25 hr 32 min. Section 5.2.1.1 summarizes the thermal-hydraulic response of the 

reactor and containment while Section 5.2.2.2 summarizes the associated radionuclide release 

from the fuel to the environment.

0

5

10

15

20

25

0 6 12 18 24

Flow Rate (kg/s)

Time (hr)

Total Seal Leakage versus Emergency Pump Injection

LTSBO - Mitigation with Portable Equipment, Early and Late RCP Seal Failures

Early Seal Failure

Late Seal Failure

Early Failure Injection

Late Failure Injection

Start 150 gpm 

RCS injection 

at 3.5 hr
5-32

Table 5-5 Timing of key events for unmitigated STSBO

Event Description

Time

(hh:mm)

Initiating event

Station blackout – loss of all onsite and offsite AC and DC 

power

00:00

MSIVs close

Reactor trip

RCP seals initially leak at 21 gpm/pump

TD-AFW starts but fails to inject due to ECST rupture

00:00

First SG SRV opening 00:03

SG dryout 01:16

Pressurizer SRV opens 01:27

Pressurizer relief tank rupture disk opens 01:46

Start of fuel heatup 02:19

RCP seal failures 02:45

First fission product gap releases 02:57

Creep rupture failure of the C loop hot leg nozzle 03:45

Accumulators start discharging 03:45

Accumulators are empty 03:45

Vessel lower head failure by creep rupture 07:16

Debris discharge to reactor cavity 07:16

Cavity dryout 07:27

Containment at design pressure (45 psig) 11:00

Start of increased leakage of containment (P/Pdesign = 2.18) 25:32

Containment pressure increase slows 32:00

Containment pressure stops decreasing 44:14

End of calculation 48:00

5.2.1.1 Thermal-Hydraulic Response

The responses of the primary and secondary pressure systems are shown in Figure 5-28. At the 

start of the accident sequence, the reactor successfully scrams in response to the loss of power. 

The main steam line isolation and containment isolation valves close in response due to the loss 

of power. The reactor coolant and main feedwater pumps also trip to the loss of power. Once 

the main steam lines close, the normal mechanism of heat removal from the primary system is 

unavailable. Consequently, both the primary and secondary system pressures rise. 

The secondary system quickly pressurizes to the safety relief valve opening pressure, which 

results in the safety relief valves to open and then subsequently close when the closing pressure 
5-33

criterion is achieved. The relief flow through the SG SRVs is the principle primary system 

energy removal mechanism in the first hour. There is also energy removal through the RCP seal 

leakage, but the energy flow is small relative to the SG SRV flow. 

Due to the complete loss of all feedwater at the start of the calculation, the water inventory in the 

steam generators decreased very rapidly and was completely boiled away by 1 hr 16 min. 

Although the steam generators relief valves continue to cycle and release steam, the associated 

heat removal is inadequate and the primary system sharply increases to the pressurizer safety 

relief valve opening pressure. The safety valves on the pressurizer begin opening and closing 

to remove excess energy. However, the pressurizer relief valve flow causes a steady decrease in 

the primary system coolant inventory (see Figure 5-29). The fuel starts to uncover at 2 hr 19 min

(see Figure 5-30). The fuel cladding fails at 2 hr 59 min, which starts the release of fission 

products from the fuel. The fuel rods starts to degrade above 2400 K as the molten zirconium 

breaks through the oxidized shell of the cladding on the fuel rods and eventually collapse due a 

thermal-mechanical weakening of the remaining oxide shell at high temperature. As shown in

Figure 5-30, the peak fuel-debris temperature reaches the fuel-zirconium oxide eutectic melting 

temperature of 2800 K.

Following the uncovery of the fuel, an in-vessel natural circulation flow develops between the 

hot fuel in the core and the cooler structures in the upper plenum. Hot gases rise out of the 

center of the core rise into the upper plenum and return down the cooler peripheral sections of 

the core. Simultaneously, a natural circulation circuit develops between the hot gases in the 

vessel and the steam generator [5]. Hot gases from inside the vessel flow along the top of the hot 

leg and into the steam generator. The hot gases flow through the steam generator in 

approximately half the tubes and return through the remaining tubes. The large masses of the hot 

leg nozzle, hot leg piping, and the steam generator tubes absorb the heat from the gases exiting 

the vessel. The cooler gases leaving the steam generator return to the vessel along the bottom of 

the hot leg. Due to its close proximity to the hot gases exiting in the vessel, the hot leg nozzle at 

the carbon steel interface region to the stainless steel piping was predicted to fail by creep 

rupture at 3 hr 45 min.

22

Upon creep failure of the hot leg nozzle, a large hole opened that rapidly depressurized the RCS 

(i.e., like a large break loss-of-coolant accident). The RCS depressurization permitted a 

complete accumulator injection at low-pressure (i.e., water level rise at 3 hr 45 min on

Figure 5-29). Although the water filled above the core region, the hottest fuel in the 

core remained in film boiling and continued to collapse and degrade (see Figure 5-30). The 

lower temperature regions on the periphery of the core quenched but subsequently reheated once

the water level decreased into the core.

 22 Alternate failure locations could include the pressurizer surge line and the steam generator tubes. There was 

some residual water in the pressurizer that cooled the surge line. Due to the relatively high pressure in the steam 

generator secondary side, the resultant thermal-mechanical stresses across the steam generator tubes were less 

severe than the hot leg nozzle. Consequently, the most vulnerable location was calculated to be the hot leg 

nozzle. The variation of the scenario was evaluated assuming a SG tube was the initial failure location.
5-34

Following the accumulator injection at 3 hr 45 minute, the decay heat from the fuel boiled away 

the injected water. By 4.3 hr, a large debris bed had formed in the center of the core. The debris 

continued to expand until 5.8 hr when all the fuel had collapsed and was resting on the core 

plate. The hot debris failed the core support plate and fell onto the lower core support plate, 

which failed at 6.6 hr. Following the lower core support plate failure, the debris bed relocated 

onto the lower head. The small amount of remaining water in the lower head was quickly boiled 

away. As shown in Figure 5-32, the hot debris heated the inner surface of the lower head toward 

its melting temperature. As the heat transferred through the lower head, the lower head 

weakened and failed at 7 hr 16 min due (i.e., primarily due the thermal stress component in the 

thermal-mechanical failure criteria, due to the low differential pressure).

By 7.5 hr, nearly all the hot debris relocated from the vessel into the reactor cavity in the 

containment under the reactor vessel. The hot debris boiled away the water in the reactor cavity 

started to ablate the concrete. The ex-vessel core-concrete interactions (CCI) continued for 

the remainder of the calculation, which generated non-condensable gases. In addition, the hot 

gases exiting the reactor cavity and the radioactive heating from airborne and settled fission 

products steadily evaporated the water on the containment floor outside the reactor cavity from 

7.3 hr to 44 hr. The resultant non-condensable gas and steam generation pressurized the 

containment (see Figure 5-33). At 25.5 hr, the containment failed due to liner tearing near the 

containment equipment hatch at mid-height in the cylindrical region of the containment (i.e.,

containment leakage area in Figure 5-31). The containment continues to pressurize until the 

leakage flow balanced the steam and non-condensable gas generation (see Figure 5-34). By 

44 hr, all the water on the floor has evaporated. The containment depressurized thereafter due to 

only a smaller gas loading from the non-condensable gas generation.

The containment failure location was around the equipment hatch, which is located on the side of 

the containment without a surrounding building (e.g., not adjacent to the auxiliary or safeguards 

buildings). Consequently, all released fission products are released directly to the environment.

A summary of the containment gas concentrations is shown in Figure 5-35. The steam 

concentration in the containment rapidly increases to ~51% following the pressurizer relief tank 

failure at 1:46. The steam concentration remains relatively constant until hot leg failure. At hot 

leg failure, the steam increases to ~68% and remains above this level until the end of the 

calculation. At a steam concentration of 55% or higher, it is essentially impossible to ignite 

hydrogen or sustain a burn (see combustion modeling in Section 4.9). Similarly, the same 

experimental research showed that a minimum concentration of hydrogen and oxygen is also 

needed before hydrogen will ignite and burn [31]. The hydrogen concentration for combustion 

with an active source (7%) and the maximum steam concentration (55%) are also shown in 

Figure 5-35 [13]. Although the in-vessel hydrogen production is very significant, combustible 

conditions did not exist in the containment through 48 hours. The steam concentration is above 

the minimum threshold (55%) for combustion whenever any significant amount of hydrogen is 

present and the hydrogen concentration is always below the minimum ignition threshold (7%).
5-35

Figure 5-28 Unmitigated STSBO primary and secondary pressures history

Figure 5-29 Unmitigated short-term station blackout vessel two-phase coolant level

Primary and Secondary Pressures

STSBO - No Mitigation

0

2

4

6

8

10

12

14

16

18

0 1 2 3 4 5 6 7 8

Time (hr)

Pressure (MPa)

Pressurizer

SG A

SG C

SG B

Hot leg C creep rupture

SG dryout

Vessel Water Level

STSBO - No Mitigation with Portable Equipment

-4

-2

0

2

4

6

8

10

0 1 2 3 4 5 6 7 8

Time (hr)

Two-Phase Level (m)

Accumulators

BAF

TAF

Lower head

Vessel top

Vessel failure
5-36

Figure 5-30 Unmitigated STSBO core temperature history

Figure 5-31 Unmitigated STSBO containment leakage area

300

800

1300

1800

2300

2800

3300

0 1 2 3 4 5 6 7 8

Temperature (K)

Time (hr)

Peak Cladding/Debris Temperature

STSBO - No Mitigation

Vessel failure at 

7.3 hours

All fuel rods 

collapsed

Hot leg creep 

rupture failure

Fuel and clad oxide melting = 2800 K

0.00

0.01

0.02

0.03

0.04

0.05

0.06

0.07

0.08

0.09

0.10

0 1 2 3 4

Leak Area (m2)

Time (days)

Containment Failure Leakage Area

STSBO - No Mitigation
5-37

Figure 5-32 Unmitigated STSBO lower head inner and outer temperature history

Figure 5-33 Unmitigated short-term station blackout containment pressure history

Lower Head Temperatures

STSBO - No Mitigation

300

700

1100

1500

1900

0 2 4 6 8

Time (hr)

Temperature (K)

Inside Surface

Outside Surface

Vessel failure

Containment Pressure

STSBO - No Mitigation

0

100000

200000

300000

400000

500000

600000

700000

800000

900000

1000000

0 1 2 3 4 5 6 7

Time (days)

Pressure (Pa)

0

15

29

44

58

73

87

102

116

131

145

Pressure (psia)

Start of high 

leakage 

(25.5 h)

RCS hot leg 

creep 

rupture 

failure

Vessel 

failure
5-38

Figure 5-34 Unmitigated short-term station blackout containment leakage area

Figure 5-35 Unmitigated short-term station blackout containment gas concentrations

0.00

0.01

0.02

0.03

0.04

0.05

0.06

0.07

0.08

0.09

0.10

0 1 2 3 4

Leak Area (m2)

Time (days)

Containment Failure Leakage Area

STSBO - No Mitigation

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 12 24 36 48

Mole Fraction (-)

Time (hr)

Surry Unmitigated Short-Term Station Blackout

Containment Gas Concentrations

Steam

Oxygen

Nitrogen

Hydrogen

Carbon Monoxide

Max Steam

Min Hydrogen
5-39

5.2.1.2 Radionuclide Release

The fission product releases from the fuel started following the first thermal-mechanical failures 

of the fuel cladding in the hottest rods at 2 hr 57 min, or about 38 min after the uncovery of the 

top of the fuel rods. The in-vessel fission product release phase continued through vessel failure 

at 7.3 hr. Initially, the fission product releases from the fuel circulated through the primary 

system as well as being released to the containment through the pressurizer safety relief valves. 

The PRT rupture disk opened about 1 hour before the start of the start of the fission product 

releases. Subsequently, the fission product releases exiting through the pressurizer relief valves 

were not well retained in the PRT because the pool was nearly saturated and in the PRT rupture 

disk was open. Following vessel failure, the fission product releases continued from the 

ex-vessel fuel in the reactor cavity.

Figure 5-36 and Figure 5-37 show the fission product distributions of the iodine and cesium 

radionuclides that were released from the fuel, respectively (see Appendix B for a detailed 

radionuclide core inventory). Approximately 97% and 98% of the iodine and cesium, 

respectively, were released from the fuel prior to vessel failure while the remaining amount was 

released ex-vessel. At the time of the hot leg failure, approximately 40% of these volatile 

radionuclides had been released. The resultant blowdown of the vessel immediately discharged

the airborne fission products to the containment. However, about 6% of the iodine and 5% of the 

cesium remained in the vessel. Most of the radionuclides retained in the RCS were deposited in 

the steam generators during the natural circulation phase of the accident. Following the RCS 

blowdown after the hot leg nozzle failure, more radionuclides were released from the fuel as the 

core further degraded. At low pressure conditions, the fission products continued to circulate 

within the vessel and to the steam generators with a substantial portion being depositing on the 

structural surfaces. However, as shown in the figures, the majority of the released radionuclides 

were transported to the containment. Within the first day, most of the airborne fission products 

in the containment settled on surfaces. This was significant because the containment failure 

occurred at 25 hr 32 min. Consequently, there was little airborne mass that could be released to 

the environment.

The chemical form of the released iodine was cesium-iodine, which was more volatile than the 

predominant form of the released cesium, which was cesium-molybdate (Cs2MoO4). As shown 

in the iodine history figure (see Figure 5-36), the in-vessel iodine mass was decreasing following 

vessel failure until approximately 2.9 days. The decrease of mass represents a vaporization 

process of previously deposited radionuclides. The late in-vessel vaporization release was 

significant because it continued after containment failure and had a significant contribution to the 

overall environmental release. The thermal mechanisms for the vaporization of the iodine were 

from two sources. First, the fission product decay of the settled radionuclides heated the 

structures. Second, a natural circulation flow of very hot gases (i.e., from 1050 K to >1600 K) 

flowed from the reactor cavity, through the failed vessel lower head, through the reactor vessel, 

and out the failed hot leg nozzle. As the deposited cesium-iodine heated, gaseous iodine was 

released and the cesium remained chemisorbed to the stainless steel surfaces. The natural 

circulation flow pattern also effectively vented the gaseous iodine from the RCS to the 

containment. 
5-40

In contrast to the iodine response in Figure 5-36, the deposited cesium molybdate was less 

volatile and remained deposited in the RCS (see Figure 5-37). Except for inside the reactor 

cavity, the containment was cooler than vessel and well below conditions that would vaporize 

settled radionuclides. Consequently, none of the deposited radionuclides (i.e., including iodine) 

in the containment vaporized. 

Finally, Figure 5-38 summarizes the releases of the radionuclides to the environment. At 4 days, 

92% of the noble gases, 1.7% of the tellurium, 1.0% if the iodine, 0.75% of the radioactive 

cadmium, 0.4% of the cesium, and 0.1% of the barium and radioactive tin had been released to 

the environment. All other releases were less than 0.1% of the initial inventory. There were 

some environmental releases prior to the containment failure at 25.5 hr due to nominal leakages 

(i.e., design specification of 0.1% vol/day at the design pressure). The releases to the 

environment increased sharply after the failure of the containment. Between 25.5 hr, 48 hr, 53% 

of the airborne noble gases in the containment were released. Over the next 2 days, ~40% more 

was released.

Figure 5-36 Unmitigated STSBO iodine fission product distribution history23

 23 In-vessel refers to the entire primary-side of the RCS.

Iodine Distribution

STSBO - No Mitigation

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

Containment Deposited

Containment Airborne

In-Vessel Total

Env. Release

In-vessel revaporization

Environmental 

release = 1%

Hot leg creep 

rupture failure

Revaporization in cavity 
5-41

Figure 5-37 Unmitigated STSBO cesium fission product distribution history 24

Figure 5-38 Unmitigated STSBO environmental release history of all fission products

 24 In-vessel refers to the entire primary-side of the RCS.

Cesium Distribution

STSBO - No Mitigation

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

Containment Deposited

Containment Airborne

In-Vessel Total

Env. Release

In-vessel revaporization

Environmental 

release = 0.4%

Hot leg creep 

rupture failure

Fission Product Release to the Environment

STSBO - No Mitigation

0.000001

0.00001

0.0001

0.001

0.01

0.1

1

0 1 2 3 4

Time (days)

Fraction release (-)

NG

I

Cs

Ba

Te

Ru

Mo

Ce

La

U

Cd

Sn
5-42

5.2.2 Mitigated Short-Term Station Blackout 

Table 5-6 summarizes the timing of the key events in the mitigated STSBO. As described in 

Section 3.2, the accident scenario initiates with a complete loss of all onsite and offsite power. 

The reactor successfully trips and the containment isolates, but all powered safety systems are 

unavailable. The timing of key events are discussed further in Sections 5.2.2.1 and 5.2.2.2. 

Unlike the unmitigated STSBO described in Section 5.2.1, the mitigated STSBO credits the 

successful connection of the portable, low-pressure, diesel-driven (Godwin) pump to the 

containment spray system at 8 hours. The Godwin pump is a high-flow, low-head pump with a 

design capacity of 2,000 gpm at 120 psi.25 A reliable source of water is maintained while 

1,000,000 gallons is injected into the containment through the containment sprays. In particular, 

the core has degraded and failed the vessel lower head prior to the spray actuation (see 

Table 5-6). The emergency containment sprays are effective at reducing the containment 

pressure and scrubbing of aerosols from the containment atmosphere by spray droplets while 

they are operating. However, the containment subsequently pressurizes after the sprays are 

terminated to the failure pressure. While not investigated, intermittent operation of the sprays 

and deeper flooding could have further delayed failure of the containment.

 25 The rated containment pump spray flowrate was 3,200 gpm. It was judged that the portable Godwin pump would 

pressurize the system and develop the spray droplet flow patter.
5-43

Table 5-6 Timing of key events for mitigated STSBO

Event Description

Time

(hh:mm)

Initiating event

Station blackout – loss of all onsite and offsite AC and DC 

power

00:00

MSIVs close

Reactor trip

RCP seals initially leak at 21 gpm/pump

TD-AFW starts but fails to inject due to ECST rupture

00:00

First SG SRV opening 00:03

SG dryout 01:16

Pressurizer SRV opens 01:27

Pressurizer relief tank rupture disk opens 01:46

Start of fuel heatup 02:19

RCP seal failures 02:45

First fission product gap releases 02:57

Creep rupture failure of the C loop hot leg nozzle 03:45

Accumulators start discharging 03:45

Accumulators are empty 03:45

Vessel lower head failure by creep rupture 07:16

Debris discharge to reactor cavity 07:16

Cavity dryout 07:27

Start of containment sprays 8:00

End of containment sprays (1,000,000 gal) 15:02

Containment at design pressure (45 psig) 40:00

Start of increased leakage of containment (P/Pdesign = 2.18) 66:30

5.2.2.1 Thermal-Hydraulic Response

The progression of events in the mitigated STSBO is identical to the unmitigated STSBO as 

described in Section 5.2.1 through the first 8 hr, which includes core degradation and vessel 

failure. Consequently, the system pressure and peak fuel temperature responses though the first 

8 hr is identical between the two sequences (i.e., compare Figure 5-39 and Figure 5-28, 

Figure 5-41 and Figure 5-30, and Figure 5-42 and Figure 5-32). However, after 8 hr, there are 

some key differences. For example, the long-term reactor vessel level shows different behavior 

after 8 hr. Although the water in the vessel boils away by 6.6 hr when the core relocates onto the 

lower head, the vessel water level starts to recover at 13.7 hr as shown in Figure 5-40. After 

5.7 hr of containment spray operation, the water level in the containment was calculated to fill 
5-44

above the bottom of the vessel.26 At 15 hr when the containment spray was terminated, the 

containment water had flooded ~1.3 m into the vessel. 

The reactor cavity of Unit 1 of the Surry containment connects to the surrounding lower regions 

of the containment through: (a) a 12” hole in the reactor cavity wall at 2’-7” above the bottom of 

the floor, (b) a penetration at 24’-3” above the floor, and (c) the holes in the cavity wall for the 

RCS piping (i.e., nearly 40’ above the bottom of the floor). In the case of this scenario, the lower 

hole into the reactor cavity was flooded whereas the upper openings were well above the water 

level. Hence, there was no natural circulation of water from the containment basement into the 

reactor cavity and out the gaps for the RCS piping penetrations. While the containment sprays 

were operating, a significant portion of the spray water drained into the reactor cavity from 

refueling pool. The resultant water flow through the reactor cavity removed the heat from the 

fuel debris. Once the spray flow stopped, the water in the cavity heated to saturation conditions 

and started to boil. The resulting steam load from the boiling pressurized the containment to the 

failure pressure (see Figure 5-43). Although there was 1,000,000 gallons of water in the 

containment, the core debris was only in thermal contact with ~40,000 gallons in the reactor 

cavity. Consequently, the containment pressurized to failure conditions much faster than if all 

1,000,000 gallons were being heated. As stated in Section 4.2, intermittent spray operation 

and/or flooding above the RCS piping penetrations would have substantially delayed 

containment failure. Although the containment sprays did not prevent containment failure, they 

delayed containment failure by over 40 hr relative to the unmitigated case.

Although the exact conditions following a severe seismic event were not known, it was estimated 

that portable sprays could be started by 8 hr. A summary of the containment gas concentrations 

is shown in Figure 5-44. As seen in the mitigated short-term station blackout response, the steam 

concentration in the containment rapidly increases to ~51% following the pressurizer relief tank 

failure at 1:46. The steam concentration remains relatively constant until hot leg failure. At hot 

leg failure, the steam concentration increases to ~68% and remains above this level until the 

initiation of the emergency containment sprays at 8 hours. As shown previously in Sections 

5.1.1.1 and 5.2.1.1, the unmitigated long- and short-term station blackouts remained steaminerted. However, the mitigation using emergency sprays results in a much different response. 

Although the spray operation reduces the containment pressure by condensing steam, it increases 

the relative concentrations of the hydrogen and oxygen. The containment reaches combustible 

conditions and burns (i.e., which releases energy, rapidly increases the pressure, and makes 

steam). The steam condenses and more hydrogen (and carbon monoxide from the MCCI) burns. 

This process repeats until the containment sprays are terminated causing sharp hydrogen 

deflagration spikes in the containment pressure (Figure 5-45). Following the termination of the 

emergency sprays, the ex-vessel debris heats the spray water in the containment basement and 

eventually steam-inerts the containment at ~26 hr. The peak hydrogen deflagration pressure 

spikes during the emergency spray operation are well below the containment failure pressure 

(~110 psia). The uncertainties in the combustion behavior during emergency spray operation are 

further examined in Section 5.2.3.Based on the pressure response of the unmitigated STSBO, the 

 26 At the time of the calculation, the exact flooding water level characteristics of the Surry containment were not 

known. Subsequently, information was obtained from the plant that shows ~1,160,000 gal are needed to fill to 

the bottom of the vessel. Consequently, the calculated water level response of this scenario is actually consistent 

with a slightly higher integrated spray flow.
5-45

containment sprays must be started before 15.6 hr while the containment pressure was below the 

shutoff head of the portable pump. Once the containment sprays are initiated, they are effective 

in quickly reducing the containment pressure. Consequently, there was almost eight

additional hours from the assumed starting time to establish containment sprays, or 15.6 hr after 

the start of the scenario. 

The selection of the containment sprays as a mitigation technique for this scenario was 

particularly beneficial for several reasons. First, the containment sprays were extremely 

effective in scrubbing of the airborne aerosols from the containment atmosphere by spray 

droplets into the large containment pool. Second, the sprays delayed containment failure for an 

additional 41 hr. In contrast, the alternate strategies of containment flooding or vessel injection 

would not be expected to be as beneficial for this scenario (i.e., assuming an initiation time after 

vessel failure). The high-head portable pump used for vessel injection can only provide 65 gpm

versus 2000 gpm for the portable containment spray pump. Consequently, the time to deep flood 

the containment would be significantly longer. More importantly, the water would merely fall 

out of the failed vessel and not reduce the containment pressure (i.e., versus the highly effective 

heat and mass transfer from the containment spray system). In fact, the small amount of water 

flooding onto the ex-vessel core debris would enhance the pressurization of the containment 

versus a dry reactor cavity. Furthermore, the injection flow would not directly knockdown the 

airborne aerosol radionuclides. Similarly, direct containment injection using the high flow pump 

would not depressurize the containment nor reduce any airborne fission products.

Figure 5-39 Mitigated STSBO primary and secondary pressure history

Primary and Secondary Pressures

STSBO

0

2

4

6

8

10

12

14

16

18

0 1 2 3 4 5 6 7 8

Time (hr)

Pressure (MPa)

Pressurizer

SG A

SG B

SG C

Hot leg C creep rupture

SG dryout
5-46

Figure 5-40 Mitigated short-term station blackout vessel two-phase coolant level

Figure 5-41 Mitigated short-term station blackout core temperature history

Vessel Water Level

STSBO - Mitigation with Portable Equipment

-4

-2

0

2

4

6

8

10

0 3 6 9 12 15 18 21 24

Time (hr)

Two-Phase Level (m)

Accumulators

Start emerg. diesel 

containment sprays BAF

TAF

Lower head

Vessel top

Cavity floods into the 

failed vessel

Peak Cladding Temperatures

STSBO - Mitigated with portable equipment

300

800

1300

1800

2300

2800

3300

0 1 2 3 4 5 6 7 8

Time (hr)

Temperature (K)

Vessel failure at 

7.3 hours

All fuel rods 

collapsed

Hot leg creep 

rupture failure

Fuel and clad oxide melting = 2800 K
5-47

Figure 5-42 Mitigated STSBO lower head inner and outer temperature history

Figure 5-43 Mitigated short-term station blackout containment pressure history

Lower Head Temperatures

STSBO - Mitigation with portable equipment

300

700

1100

1500

1900

0 2 4 6 8

Time (hr)

Temperature (K)

Inside Surface

Outside Surface

Vessel failure

Containment Pressure

STSBO - Mitigation with Portable Equipment

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4 5

Time (days)

Pressure (Pa)

0

15

29

44

58

73

87

102

116

131

145

Pressure (psi)

Start of 

high 

leakage

Sprays 

off

Sprays 

on

H2 burns
5-48

Figure 5-44 Mitigated short-term station blackout containment gas concentrations

Figure 5-45 Mitigated and unmitigated short-term station blackout containment gas 

concentrations

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 12 24 36 48

Mole Fraction (-)

Time (hr)

Surry Mitigated Short-Term Station Blackout

Containment Gas Concentrations

Steam

Oxygen

Nitrogen

Hydrogen

Carbon Monoxide

Max Steam

Min Hydrogen

0

25

50

75

100

125

150

0 12 24 36 48

Pressure (psia)

Time (hr)

Surry Short-Term Station Blackout

Containment Pressure Response

Unmitigated

Mitigated

Emergency sprays operation

(spikes are hydrogen burns 

from 7% concentration)
5-49

5.2.2.2 Radionuclide Release

The radionuclide response of the mitigated STSBO is identical to the unmitigated response 

described in Section 4.1.2 for the first 8 hr, or through vessel failure until the start of the 

containment sprays. Following the start of the emergency containment sprays at 8 hr

(0.25 days), the airborne aerosols of iodine and cesium rapidly decease (see Figure 5-46 and 

Figure 5-47, respectively). By the time the sprays are terminated at 15 hr (0.63 days), almost all 

of the airborne aerosols have been captured in the pool on the containment floor. Since the 

containment failure was delayed until 66 hr 30 min (2.8 days), the amount of airborne mass 

available for release was insignificant. The environmental release of iodine and cesium was very 

small (i.e., 0.007% and 0.003%, respectively).

Due to the deep flooding in the reactor cavity by the spray operation, the bottom of the failed 

vessel lower head is submerged in water. Therefore, the natural hot circulation flow that 

promoted vaporization of deposited radionuclides in the unmitigated STSBO is not present. 

Instead, the water pool in the reactor cavity cools the bottom of the vessel. Due to the boiling in 

the cavity, relatively cool steam flows through the vessel and out the failed hot leg nozzle 

location, which removes heat and inhibits vaporization of deposited radionuclides in the upper 

vessel and hot leg. Consequently, the vaporization of the in-vessel deposited fission products 

(i.e., especially cesium-iodine) that was seen in the unmitigated STSBO (i.e., characteristic of 

vaporization), was negligible in the mitigated case through 4 days.

Finally, Figure 5-48 summarizes the releases of the radionuclides to the environment. At 4 days, 

60% of the noble gases, 0.0046% of the tellurium, 0.0065% of the iodine and cadmium, 0.0027% 

of the cesium had been released to the environment. Except for the noble gases, all the releases 

were less than 0.01% of the initial inventory. As shown in the figure, the initial releases to the 

environment were due to the nominal leakages (i.e., design specification of 0.1% vol/day at the 

design pressure) prior to the containment failure at 66 hr 30 min (2.8 days). Following 

containment failure at 2.8 days until 4 days, the noble gas release went from 0.24% to 60%, 

which represents a significant flushing of the containment gas space to the environment. There 

is some evidence of increased leakage of the other radionuclides after containment failure. 

However, the response is exaggerated in the figure due to the semi-logarithmic scale. The 

absolute magnitude of the releases was small relative to the unmitigated response.
5-50

Figure 5-46 Mitigated short-term station blackout iodine fission product distribution 

history 27

Figure 5-47 Mitigated STSBO cesium fission product distribution history

 27 In-vessel refers to the entire primary-side of the RCS.

Iodine Distribution

STSBO - Mitigated with Portable Equipment

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

Containment Deposited

Containment Airborne

In-Vessel Total

Env. Release

Environmental 

release = 0.007%

88% deposited in containment

RCS total = 11%

Cesium Distribution

STSBO - Mitigated with Portable Equipment

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

Containment Deposited

Containment Airborne

In-Vessel Total

Env. Release

Environmental 

release = 0.003%

 84% deposited in containment

RCS total = 15%
5-51

Figure 5-48 Mitigated STSBO environmental release history of all fission products

5.2.3 Uncertainties in the Hydrogen Combustion in the Mitigated Short-Term Station 

Blackout

During the peer review of the mitigated STSBO, there was concern about the undesirable effects 

of combustion following an emergency spray actuation in the containment. Although emergency 

sprays mitigated the accident, there was severe damage to the fuel with considerable hydrogen 

production. On the other hand, another peer review comment concerned the time it took to 

transport a diesel-driven pump and connect it to plant piping following a large seismic event. 

The peer reviewer noted that while the licensee estimated 2 hours would be required, SOARCA 

assumed that 8 hours would be required. This comment also supports the idea that the mitigated 

case could have involved injection into the RCS resulting in no core damage and therefore no 

hydrogen combustion. It should be noted that further discussion with the plant staff revealed that 

the decision to install and operate the emergency sprays is made by the TSC and is also 

uncertain. The TSC will use calculation aids to assess the benefits of containment sprays versus 

the adverse effects of enhancing conditions for hydrogen burns. Consequently, the mitigated 

STSBO calculation described in Section 3.2.3 and the following sensitivity calculations examine 

the potential system responses assuming the use of emergency sprays as a mitigative action, 

which is uncertain. Alternate responses include: (a) no successful mitigative actions (i.e., the 

unmitigated response described in Section 5.1.1), (b) emergency RCS injection prior to vessel 

failure, (c) emergency AFW injection that greatly delays core damage, and (d) containment 

flooding, which does not enhance hydrogen combustion.

In the mitigated case described in Section 3.2.3, there was severe damage to the fuel with 

considerable hydrogen production due to the delay in connecting the diesel-driven pump to plant 

Fission Product Release to the Environment

STSBO - Mitigated with portable equipment

0.000001

0.00001

0.0001

0.001

0.01

0.1

1

0 1 2 3 4

Time (days)

Fraction release (-)

NG

I

Cs

Ba

Te

Ru

Mo

Ce

La

U

Cd

Sn
5-52

piping. However, the severity of the accident was reduced by covering the core debris with 

water, scrubbing airborne fission products, and reducing containment pressure. As shown in

Figure 5-49, 200 kg of hydrogen was produced by the time of the hot leg failure (i.e., 3:45 hr). 

There was a hydrogen burn coincident with hot leg rupture. Subsequently, the hydrogen 

production continued and an additional 150 kg of hydrogen was produced by vessel lower head 

rupture failure (i.e., 7:16 hr). 

Figure 5-49 Comparison of the mitigated short-term station blackout containment 

pressure history versus the dome hydrogen mass and total hydrogen 

production

Following vessel failure, steam production by debris falling into water on the containment floor

kept the containment steam-inerted.28 Consequently, there were no further hydrogen burns in the 

unmitigated case. In the mitigated case, emergency containment sprays subsequently condensed 

the steam in the containment atmosphere and reduced the containment pressure. The net effects 

of the emergency spray operation between 8 to 15 hours were high hydrogen and oxygen 

concentrations, good mixing, and a low steam concentration (i.e., conditions suitable for 

combustion). Hydrogen combustion was modeled whenever the gaseous concentrations reached 

the specified levels in MELCOR’s default combustion model (i.e., Xsteam <55%, Xhydrogen>10%, 

 28 As described in Section 4.9, MELCOR’s default hydrogen combustion model was used, which identifies steam 

concentrations greater than 55% to be steam-inerted and unable to support a hydrogen burn. Following vessel 

failure, the ex-vessel debris boiled water on the containment floor to maintain a high steam concentration.

0

50000

100000

150000

200000

250000

300000

350000

400000

450000

500000

0

100

200

300

400

500

600

700

800

900

1000

0 6 12 18 24

Pressure (Pa)

Mass (kg)

Time (hr)

Hydrogen Combustion in Mitigated STSBO

Total In- and Ex-Ves H2

Total H2 Burned

Dome H2 Mass

Pressure Containment 

sprays
5-53

and Xoxygen >5%). This resulted in several smaller burns as shown in Figure 5-49. Due the 

uncertainty of the ignition source, SOARCA peer reviewers inquired about the consequences of a 

later but larger burn. The MELCOR burn package defaults of 7% and 10% hydrogen 

concentrations are used for ignition with an identified ignition source and spontaneous ignition, 

respectively. It was conservatively assumed that there were no ignition sources to maximize the 

combustion loads. Thus, a 10% hydrogen ignition concentration criterion was used for all 

SOARCA scenarios.

Several facets were investigated relative to the SOARCA peer reviewer’s comments. First, the 

potential ignition sources were reviewed. The most likely sources are the hot gases exiting the 

failed hot leg (i.e., a hot jet or hot pipe) and debris when the vessel fails. Two locations were

examined, the containment cavity (see Figure 5-50) and the containment dome (i.e., see

Figure 5-51, representative of the bulk conditions due to high mixing during spray operation). 

Hot, hydrogen-rich gases discharge into the cavity following hot leg failure. The temperature of 

the gases is well above the auto-ignition temperature for a hydrogen jet (i.e., 950-1100 K), as 

shown in Figure 5-50. Hence, ignition is likely, which occurred in the base calculation. The 

specific conditions at the hot leg failure are summarized in the figure. However, due to the 

subsequent discharge of the accumulators and full depressurization of the primary system, the 

high temperature jet flow stopped shortly thereafter. Furthermore, steam from the vessel quickly 

inerted the cavity atmosphere to >90%.29 Following vessel failure, very hot debris fell from the 

vessel into the cavity, which also created a likely ignition source. However, once the hot debris 

reached the water pool on the cavity floor, there was a sustained, rapid steam production. The 

additional steaming re-inerted the containment and increased the containment pressure, which 

lowered the relative hydrogen concentration.

 29 The high steam source through the hot leg failure location into the containment occurred as accumulator water

 boiled in the hot vessel debris.
5-54

Figure 5-50 Comparison of the mitigated short-term station blackout containment cavity 

gas concentration history and potential ignition source temperatures

Figure 5-51 Comparison of the mitigated short-term station blackout containment dome 

gas concentration history and potential ignition source temperature

Hot Leg Creep Rupture Gas Jet Temperature versus Cavity Gas Concentration

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

0 4 8 12 16

Time (hr)

Gas Concentration (-)

500

600

700

800

900

1000

1100

1200

1300

1400

1500

Gas Temperature (K)

Cavity H2

Cavity O2

Cavity Steam

HL Jet

Jet autoignition 

temperature 

range

HL Failure

XH2, cav = 0.07

XH2, jet = 0.20

X O2 = 0.13

XH2O = 0.35

TJet = 1300 K

Hot jet over

After VF

XH2 = 0.03 - 0.07

X O2 = 0.0 - 0.02

XH2O = ~0.9

Tdeb = >1700 K

Vessel 

failure

Hot 

leg 

Hot Leg Creep Rupture Gas Jet Temperature versus Dome Gas Concentrations

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

0 4 8 12 16

Time (hr)

Gas Concentration (-)

500

600

700

800

900

1000

1100

1200

1300

1400

1500

Gas Temperature (K)

Bulk H2

Bulk O2

Bulk Steam

HL Jet

Jet autoignition 

temperature 

range

HL Failure

XH2, cav = 0.03

XH2, jet = 0.20

X O2 = 0.10

XH2O = 0.52

TJet = 1300 K

Hot jet over

After VF

XH2 = 0.03 - 0.07

X O2 = 0.0 - 0.02

XH2O = ~0.9

Tdeb = >1700 K

Vessel 

failure

Hot 

leg 
5-55

For the regions outside the cavity (i.e., characterized by the dome region in Figure 5-51), the 

response was similar to the cavity at hot leg failure. A high temperature, hydrogen-rich gas jet 

exited into the dome for a short period until the accumulators discharged. Hence, ignition is 

likely in the dome within the jet. However, as shown in both the cavity and dome figures, the 

steam concentration rapidly increased to inerting conditions (i.e., Xsteam>55%) once the 

accumulators discharged. Subsequently, the debris exiting the vessel after vessel failure could 

entrain hot debris particles and small aerosols outside the cavity or a burn in the cavity could 

propagate into the surrounding regions. As indicated in Figure 5-51, the hydrogen concentration 

is above the default ignition criterion when an ignition source is present (i.e., >7%). However, 

the oxygen concentration is low (i.e., below the default ignition criterion) and the steam 

concentration is high (i.e., above the default ignition criterion). 

In summary, ignition would be expected at hot leg failure, which occurred. However, the 

amount of hydrogen available for combustion is limited at this phase of the accident and the 

pressurization during the burn was well below the pressure capacity of the containment. An 

ignition later than hot leg creep rupture failure (e.g., following vessel creep rupture failure, which 

is the next clear ignition source) is unlikely due to high steam inerting throughout the 

containment.

In the second facet of hydrogen uncertainty examined, the variability in the timing of the

combustion was examined. The spray operation led to conditions where hydrogen combustion 

was possible (i.e., see multiple burns during the spray operation in Figure 5-49). However, the 

presence of a clear ignition source after hot leg failure is uncertain. Consequently, a sensitivity 

case was run to investigate a delayed burn with a larger amount of hydrogen accumulation. In 

this case, three conservatisms were applied relative to the base case. First, all combustion in the 

containment was prevented until the emergency spray operation completed (i.e., ~15 hr). This 

includes preventing combustion at the time of hot leg failure when an ignition source is present. 

Second, ignition sources were assumed to be present in the containment at 15 hr, even though 

station blackout conditions existed. Third, ignition sources were activated simultaneously in all 

regions of the containment at 15 hr, which initiated burns without any propagation delay. The 

pressure response in the sensitivity case is shown in Figure 5-52. Due to robust circulation flows 

during spray operation, the hydrogen concentration was relatively uniform between 18-20% 

throughout the various regions in contamination. Following the specified deflagration at 

15 hours, the peak pressure was 60 kPa (8.7 psi) above the containment pressure that would 

result in increased containment leakage (i.e., 779 kPa, 113 psi).
5-56

Figure 5-52 Mitigated short-term station blackout containment pressure history for the 

sensitivity calculation with delayed ignition

In the sensitivity calculation, MELCOR’s default deflagration model considers the relative gas 

concentrations and the overall geometry for a best-estimate calculation. It not only includes 

combustion of hydrogen, but also combustion of carbon monoxide from the ex-vessel 

core-concrete interactions. The conditions at the start of the burn at 15 hours are shown in

Table 5-7.

Table 5-7 Containment conditions at start of hydrogen and carbon monoxide burn

Condition in the Containment Dome Value

Pressure 118.2 kPa

Temperature 314 K

Hydrogen mole fraction 19.0%

Carbon monoxide mole fraction 14.6%

Steam mole fraction 4.5%

Oxygen mole fraction 12.7%

Nitrogen mole fraction 49.2%

As shown in Figure 5-53, the combined concentration of the hydrogen and carbon monoxide 

exceeded the stoichiometric amount of oxygen after 10.3 hours. Hence, the deflagration was 

oxygen limited. Since the heat of combustion of hydrogen with oxygen and carbon dioxide with 

0

100

200

300

400

500

600

700

800

900

1000

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

0 4 8 12 16

Hydrogen Production (kg)

Pressure (MPa)

Time (hr)

Containment Pressure Response

Mitigated Surry STSBO - Delayed Hydrogen Burn

Sprays

Delayed 

hydrogen 

burn

Increased Containment Leakage

Hot leg 

failure

Vessel 

failure
5-57

oxygen is approximately the same, any additional hydrogen gas generation would not 

substantially change the maximum peak pressure from deflagration. 

Figure 5-53 Mitigated short-term station blackout containment dome gas concentrations 

versus stoichiometric

A set of six other delayed burn sensitivity calculations were performed. In addition to delaying 

combustion until 15 hours, additional sensitivity cases were performed delaying combustion to 

9.5 and 11.5 hours. As described earlier, a hydrogen burn due to jet auto-ignition following the 

hot leg failure and then further burns were suppressed until 9.5, 11.5, and 15 hours. These 

calculations were performed to investigate the benefit of pre-burning some of the hydrogen at a 

low concentration. Finally, a sensitivity case was calculated with the spray initiation and 

termination occurring 2 hours earlier (i.e., the emergency sprays ran from 6 hr to 13 hr). the 

hydrogen burn was delayed until 13 hr. in all the delayed combustion sensitivity calculations, 

ignition source was simultaneously applied at the control volumes at the specified time.

The containment pressure histories for all the delayed combustion sensitivity calculations are 

shown in Figure 5-54. Only the cases where combustion was delayed to the end of the spray 

operation (i.e., 15 hours, or 13 hours in the 6 hr start case) was the pressure criterion for 

increased leakage exceeded.

30 Hence, three of the sensitivity cases resulted in increased 

containment leakage and four cases remained below the failure pressure criterion through 

 30 Although Figure 5-53 shows the containment dome exceeds the stochiometric concentration of oxygen by

10.3 hours, the combustible gas concentration in the lower reigons of the containment was still increasing until

15 hours. Hence, a larger and more uniform burn occurred.

0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.35

0.40

0 4 8 12 16

Mole Fraction (-)

Time (hr)

Hydrogen and Carbon Monoxide Mole Fractions vs Oxygen Concentration

Surry STSBO - Delayed Hydrogen Burn

H2 + CO

H2

O2

Stoichiometric

Oxygen limited
5-58

24 hours. The resultant environmental source term of cesium and iodine are summarized in 

Figure 5-55 and Figure 5-56, respectively. Through 24 hr, the cesium and iodine environmental 

source terms were very small (i.e., <0.01%). It is difficult to distinguish the nominal leakage 

from cases with increased containment leakage from the cases with some increased leakage. In 

summary, the containment sprays were very effective at reducing the source term and the driving 

pressure leakage. The negative consequences of possible burns due to a spray operation were 

offset by the scrubbing of containment sprays. 

Figure 5-54 Comparison of the containment pressure responses for the sensitivity 

calculations with delayed ignition

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

0 4 8 12 16

Pressure (MPa)

Time (hr)

Containment Pressure Response

Mitigated Surry STSBO - Delayed Hydrogen Burn

Increased Leakage Pressure

Burn at 9.5 hr

Burn at 11.5 hr

Burn at 15.0 hr

Burn early & 9.5 hr

Burn early & 11.5 hr

Burn early & 15.0 hr

Sprays at 6 hr + burn at 13 hr

Sprays

Delayed 

hydrogen 

burns

Increased Containment Leakage

Hot leg 

failure
5-59

Figure 5-55 Comparison of the cesium release to the environment for the sensitivity 

calculations with delayed ignition

Figure 5-56 Comparison of the iodine release to the environment for the sensitivity 

calculations with delayed ignition

0.00001

0.0001

0.001

0.01

0.1

1

0 6 12 18 24

Release Fraction (-)

Time (hr)

Cesium Environmental Release Fraction

Mitigated Surry STSBO - Delayed Hydrogen Burn

Burn at 9.5 hr

Burn at 11.5 hr

Burn at 15.0 hr

Burn early & 9.5 hr

Burn early & 11.5 hr

Burn early & 15.0 hr

Sprays at 6 hr + burn at 13 hr

0.00001

0.0001

0.001

0.01

0.1

1

0 6 12 18 24

Release Fraction (-)

Time (hr)

Iodine Environmental Release Fraction

Mitigated Surry STSBO - Delayed Hydrogen Burn

Burn at 9.5 hr

Burn at 11.5 hr

Burn at 15.0 hr

Burn early & 9.5 hr

Burn early & 11.5 hr

Burn early & 15.0 hr

Sprays at 6 hr + burn at 13 hr
5-60

To confirm the MELCOR estimate of peak pressure, the adiabatic, isochoric, complete 

combustion (AICC) pressure was calculated [27]. The complete combustion assumption refers 

to the participation of the reactants only (i.e., hydrogen, carbon monoxide, oxygen). Some small 

fraction of the fuel will always exist in equilibrium with the combustion products. The AICC 

assumptions result in the highest possible equilibrium pressures. Inclusion of best-estimate heat 

transfer, volume expansion, and incomplete combustion will result in lower pressures (e.g., the 

default MELCOR combustion model). If the AICC process assumptions are met, then at 

equilibrium, simple deflagrations, accelerated flames, and detonations reach the same final AICC 

pressure.31 A number of hydrogen deflagration and detonation studies were conducted as part of 

this analysis. Delayed hydrogen deflagration and detonation were shown to be a threat to 

containment integrity. 

Figure 5-57 shows the bulk hydrogen and oxygen concentrations in the containment. At 15 hr, 

the peak hydrogen concentration was ~20% and the peak oxygen concentration was 12%. The 

AICC pressure for a 20% hydrogen concentration shown in Figure 5-58. The peak pressure was 

just slightly below the containment pressure that would result in increased containment leakage. 

However, since both hydrogen and carbon monoxide were present in the containment, the 

maximum equivalent hydrogen concentration for stoichiometric burn for the available oxygen 

would be a 24% hydrogen concentration. The peak AICC pressure from a 24% equivalent 

hydrogen concentration was slightly above the containment pressure that would result in 

increased containment leakage and approximately equal to the peak pressure calculated in the 

MELCOR delayed burn sensitivity calculation. 

 31The difference between deflagrations and detonations in a confined volume is the transient pressure-time histories 

between ignition and final equilibrium.
5-61

Figure 5-57 Mitigated short-term station blackout containment gas concentration history 

for the sensitivity calculation with delayed ignition

Figure 5-58 Mitigated short-term station blackout containment pressure history for the 

sensitivity calculation with delayed ignition

0.0

0.1

0.2

0.3

0.4

0 4 8 12 16

Concentration (-)

Time (hr)

Range of Hydrogen Concentrations (except Cavity)

H2 Bulk Max

H2 Bulk Min

O2 Bulk Max

O2 Bulk Min

0.00

0.58

1.15

1.73

0

5

10

15

0 4 8 12 16

Presssure (MPa)

Pressure (P/Ppre-burn)

Time (hr)

Containment Pressure Response

Surry STSBO - Delayed Hydrogen Burn

CJ (Detonation)

AICC at 24%

AICC at 20%

Increased Containment Leakage

MELCOR -Dome

Default MELCOR 

deflagration response

model after sprays complete
5-62

The potential for a detonation event was also raised by the peer review committee. While the 

dynamics pressure of a freely propagating detonation would exceed the equilibrium AICC 

pressure, this dynamic pressure load may not be large enough to damage a reinforced concrete 

containment. The detonation pressure can be estimated using the Chapman-Jouguet (CJ) model

[28]. The CJ model is derived from conservation of mass, momentum, and energy across a 

one-dimensional flow discontinuity. The shock wave is assumed to be sonic. It is known that 

gaseous detonation waves are three-dimensional and are not discontinuously thin. However, the 

CJ model predicts the measured detonation pressure within about 15%. The CJ pressure 

represents the one-dimensional average of the actual pressure in a detonation wave. At 15 hours, 

the CJ pressure ratio (i.e., Pcj/Ppre-burn, or the peak detonation pressure divided by the pre-burn 

pressure) was calculated to be 12.5, or slightly less than twice the AICC pressure. Consequently, 

a detonation would momentarily exceed the pressure where increased containment leakage

would be predicted for a static load. As stated above, the final equilibrium pressure following 

the detonation wave is the AICC pressure.

To assess the impact of the increased containment leakage from an earlier combustion event, the 

timing of the burn and the benefits of the emergency spray airborne radionuclide scrubbing 

should be simultaneously considered. Figure 5-59 shows the airborne concentration of cesium

and iodine aerosols as a function of the containment hydrogen concentration. Following the 

actuation of the sprays, the airborne aerosol mass decreases rapidly. Before the hydrogen 

concentration reaches a quantity that, if combusted, could result in increased containment 

leakage, there is a negligible mass of airborne aerosols for release. 

Figure 5-59 Mitigated short-term station blackout containment airborne radionuclide 

history versus hydrogen concentrations for the sensitivity calculation with 

delayed ignition
5-63

In summary, spray operation for an extended period with no early deflagration is needed to 

achieve a combustible or detonable mixture potentially capable of causing increased containment 

leakage. However, if ignition is delayed until the spray termination (i.e., 15 hours, then both 

MELCOR and the AICC model predict a peak pressure above the containment pressure criterion 

for increased leakage. Similarly, the CJ model was used to estimate the detonation pressure. 

The peak pressure predicted by the CJ model is approximately twice as large as the AICC value. 

The precise criteria for detonation are difficult to determine. However, detonations have been 

experimentally achieved at hydrogen concentrations as low as 14%, then the release of fission 

products could occur as early as 9.3 hours. However, the sprays are effective at settling airborne 

aerosols before detonable quantities could be formed that could fail the containment. The 

resulting fission product release would consist of only noble gases and would not be expected to 

substantially increase the offsite health consequences relative to the base case calculation. The 

conditions that potentially lead to severe combustion or detonable events (i.e., emergency spray 

operation) also include enhanced scrubbing of the airborne aerosols, which minimizes the impact 

of increased containment leakage due to a combustion event. Consequently, the best-estimate 

response reported in Section 5.2.2 is a reasonable representation of the source term for this 

scenario.

5.3 Short-Term Station Blackout with Thermally-Induced SGTR

The STSBO with thermally-induced SGTR scenario is assumed to be initiated by a large seismic 

event. Section 5.3.1 presents the results of an unmitigated scenario with no successful operator 

actions. For the mitigated scenario in Section 5.3.2, a portable emergency pump is connected to 

the containment spray system at 8 hours and available to inject 1,000,000 gallons.

The SOARCA sequences were selected based on screening criteria applied to SPAR evaluations 

for Surry. While hot leg rupture is generally predicted deterministically prior to 

thermally-induced SGTR (TI-SGTR), an induced tube rupture sequence was selected as a variant 

of the Surry STSBO analysis. The most significant competing challenges occur between hot leg 

(HL) creep rupture failure versus a TI-SGTR; the parameters that govern the timing of HL creep 

rupture relative to the TI-SGTR were examined. Section 5.3.3 examines the sensitivity of the 

timing of HL creep rupture failure to the TI-SGTR. NUREG/CR-6995 shows that a TI-SGTR is 

only likely if the reactor coolant pump seals do not fail and there is failure of the steam generator 

safety valves or the main steam isolation valve leaks to depressurize the secondary. 

Based on the calculated MELCOR results, the requirement for other failures, and the extensive 

research in NUREG/CR-6995, high-pressure reactor pressure vessel failure was not included as 

representative accident signature for SOARCA. Based on many analyses of STSBO sequences, 

the lower head failure prior to hot leg failure is rarely if ever observed based on current best 

modeling practices. Based on a considerable number of analyses performed by the NRC in the 

context of the SGTR integrity studies (i.e., see NUREG/CR-6995), it has been a consistent 

finding that HL creep rupture precedes lower head failure by a significant margin. It is 

recognized that the lower head failure at high pressure may retain some residual low likelihood 

in current PRA practice; however, this is considered to be of such low likelihood that it should 

not be considered in the context of the SOARCA process for best estimate analysis.
5-64

A considerable amount of work has been done by the NRC analyzing the potential for thermally 

induced steam generator tube rupture (e.g., NUREG/CR-6995). The SOARCA project 

incorporated the findings from these studies to include the potential for TI-SGTR. Two cases 

were considered; a single tube rupture and a rupture of two tubes. The failures were assumed to 

occur near the steam generator inlet plenum tube sheet where the high temperature fluid plume

enters the tube bundle. The one and two tube TI-SGTR cases showed interesting and divergent 

effects of enhancing oxidation and providing additional core cooling, respectively. It was 

assumed that the most vulnerable tube defect was exposed to the highest temperature gas

entering the steam generator. It was believed that the SOARCA approach was conservative in 

the timing and the location of the TI-SGTR. Additional realism could be introduced by 

reviewing data from plant inspections to examine the location and magnitudes of defects. This 

data could be cross-correlated against the CFD work done by the NRC (i.e., NUREG-1788) 

against the likelihood of those tubes receiving the highest temperature gas from the inlet plenum.

In order to assess the potential for a TI-SGTR, it is necessary to estimate the time difference 

between when a tube is predicted to undergo creep rupture and when depressurization of the 

RCS might occur by another means. Previous analyses such as those documented in NUREG1570, “Risk Assessment of Severe Accident-Induced Steam Generator Tube Rupture” [30] have 

shown that the hot leg pipes or nozzles, or possibly the surge line, would most likely fail first 

when the RCS is at high pressure and the steam generators have dried out following failure of all 

feedwater systems (the high/dry condition). Moreover, for Westinghouse plants, the secondary 

side would also have to be depressurized, either from operator action or from failure of a MSSV 

to reseat (the high/dry/low condition). Even then, considerable tube degradation would be 

necessary before the tubes would be calculated to fail first.

The probability of a TI-SGTR has previously been assessed to be 0.25 [30]. Consequently, 

calculations are performed in Section 5.3 with a TI-SGTR to supplement the calculations 

described in Section 5.2 without tube failures. More recent research has investigated the timing 

of the TI-SGTR relative to creep rupture failure of the hot leg with mechanistic simulations of 

natural circulation flow patterns [16][17][31][32]. The results of the research show comparable

timings for hot leg creep rupture failure and thermally-induced steam generator tube failure, for a 

flawed tube at maximum thermal stress conditions, with the former slightly preceding the later 

for most conditions. MELCOR also predicts failure of a hot leg prior to any steam generator 

tubes (i.e., potential failures are monitored at both locations) but only models unflawed SG tubes. 

Similar to the approach in NUREG/CR-6995, the calculations presented in Section 5.3 increased 

the mechanical stress across the tubes by prescribing a stuck-open safety relief valve and an 

increase in the thermal stress by inducing tube failure at a lower criterion than the default model. 

The timing of the TI-SGTR was selected to occur at a hot leg creep index of 5%, which was 14 

min before hot leg failure (i.e., corresponds to a SG stress multiplier >3, see Table 5-10). The 

5% hot leg creep index condition occurred prior to the start of the rapid radionuclide release from 

the fuel, which allowed an evaluation of the impact of the TI-SGTR prior to the development of 

a competing radionuclide release pathway through the hot leg. The hot leg subsequently failed at 

a 100% creep index. The primary system depressurized following the failure of the hot leg, 

which dramatically decreased flow through the TI-SGTR and discharged any airborne 

radionuclides into the containment. The majority of the radionuclides subsequently released 
5-65

from the fuel flowed through the failed hot leg to the containment. The new hot leg release 

pathway mitigated the magnitude of the potential release of radionuclides that bypass the 

containment. 

Alternatively, a TI-SGTR could have been assumed to occur earlier or later relative to the hot leg 

failure. An earlier TI-SGTR failure was not credible due to the low thermal heating to the SG 

tubes and was not expected to significantly increase the environmental source term (i.e., although 

the first cladding gap failure occurred at 02:57, only 1.5% of the iodine inventory was released 

by the TI-SGTR at 03:33). A pressure-induced rupture subsequent to the secondary side 

depressurization would be possible, however, if the flawed tube is sufficiently degraded.

SOARCA could have assumed the TI-SGTR occurs later, when Zr oxidation takes off. 

Observation of Figure 5-64 shows that for the single SGTR case, the intense oxidation phase 

would start around 3.7 hrs and peak around 3.75 hrs. If the SGTR occurred at 3.75 hrs, only 

about 2 min would elapse prior to the hot leg creep rupture. 

Finally, Section 5.3.3 examines and quantifies the impact of no hot leg failure following a TISGTR. Without hot leg failure, the iodine release to environment only increased from 0.2% to 

0.6% before the hot leg creep index reached 1800% (i.e., a very strong indicator of a creep 

rupture failure). Consequently, it was concluded that the SOARCA approach reasonably 

characterized the source term of a TI-SGTR prior to hot leg failure with the 5% hot leg creep 

index failure criterion. An additional sensitivity study is presented in 5.3.3 that assessed the 

impact of a delayed hot leg failure.

 

5.3.1 Unmitigated Short-Term Station Blackout with Thermally-Induced Steam 

Generator Tube Rupture

Table 5-8 summarizes the timings of the key events for the unmitigated STSBO with a TI-SGTR

scenario. Unlike the unmitigated STSBO described in Section 5.2.1, either one (i.e., equivalent 

of 100% flow area) or two (i.e., equivalent of 200% flow area) steam generator tubes fail prior to 

any other RCS creep rupture failures along with a stuck open secondary safety relief valve. 

Consequently, there is a containment bypass pathway for fission products once the steam 

generator tubes fail. As described in Section 3.2, the accident scenario initiates with a complete 

loss of all onsite and offsite power. The reactor successfully trips and the containment isolates 

but all powered safety systems are unavailable. The timings of the key events are discussed 

further in Sections 5.3.1.1 and 5.3.1.2. Similar to the unmitigated STSBO, the fission product 

releases from the fuel do not begin until 2 hr 57 min. However, since a steam generator SRV

sticks open at 3 hr and the tubes fail at 3 hr 33 min, fission product releases to the environment 

can begin earlier than in the unmitigated STSBO described in Section 5.2.1 (i.e., 3 hr 33 min

versus 25 hr 32 min). Two cases were performed to examine the sensitivity of the tube failure 

size to the magnitude of the fission product release to the environment and the potential for 

preventing hot leg creep rupture failure. Section 5.3.1.1 summarizes the thermal-hydraulic 

response of the reactor and containment while Section 5.3.1.2 summarizes the associated 

radionuclide release from the fuel to the environment.
5-66

Table 5-8 Timing of key events for unmitigated STSBO TI-SGTR

Event Description

100% 

TI-SGTR

Time

(hh:mm)

200% 

TI-SGTR

Time

(hh:mm)

Station blackout – loss of all onsite and offsite AC and DC 

power

MSIVs close

Reactor trip

RCP seal leak at 21 gpm/pump

TD-AFW fails

00:00 00:00

First SG SRV opening 00:03 00:03

SG dryout 01:14 01:14

Pressurizer SRV opens 01:27 01:27

PRT failure 01:47 01:47

Start of fuel heatup 02:19 02:19

RCP seal failures 02:46 02:46

First fission product gap releases 02:57 02:57

Stuck open SG SRV 03:00 03:00

SGTR 03:33 03:33

Creep rupture failure of the Loop C hot leg nozzle 03:47 03:49

Accumulator discharges 03:47 03:49

Accumulator empty 03:47 03:49

Vessel lower head failure by creep rupture 07:30 06:51

Debris discharge to reactor cavity 07:30 06:51

Cavity dryout 07:54 07:21

Containment at design pressure (45 psig) 12:34 13:36

Start of increased leakage of containment 

(P/Pdesign = 2.18) 27:54 30:14

Containment pressure stops decreasing 40:18 40:20

5.3.1.1 Thermal-hydraulic Response

The responses of the primary and secondary pressure systems are shown in Figure 5-60 for the 

100% and 200% TI-SGTR cases. The initial response through 3 hr is identical to the 

unmitigated STSBO (see Section 5.2.1). At 3 hr, a safety valve on steam generator C (SG-C) 

fails open.32 SG-C subsequently depressurizes to near atmospheric conditions and creates a large 

 32 The valve failure was a specified boundary condition to develop a high differential pressure drop across the steam 

generator tubes and a direct bypass flow path to the environment. The valve failure occurred after the majority of 

the safety valve cycles but before the predicted time of the hot leg failure (i.e., to promote a steam generator tube 

failure).
5-67

differential pressure across the steam generator tubes. During the core damage phase, hot gases 

circulate through the steam generator and increase the thermal stress across the tubes (see 

Figure 5-61). The equivalent of a 100% or 200% tube area failure occurs at 3 hr and 33 min (see

Figure 5-62), or about 12 min before the previously predicted creep rupture failure of the hot leg 

in the unmitigated STSBO (see Section 5.2.1). MELCOR predicts hot leg nozzle creep rupture 

failure prior to tube rupture in the unmititaged STSBO. However, this sequence assumes the 

presence of a defect in one or two steam generator tubes that increases the thermal-stress. 

Consequently, although the hot steam generator tube creep stress index is ~0 (see Figure 5-61), 

one or two tubes were specified to fail. The steam generator tube must fail prior to the hot leg 

nozzle because the resulting primary system depressurization would preclude any possibility of a 

thermally-induced SGTR. The criterion for a thermally-induced SGTR was a hot leg creep 

rupture index of 5%, which was about 12 min before the calculated hot leg nozzle failure. The 

combination of the TI-SGTR and the leakage through the failed RCP seals (2 hr and 45 min) 

causes a slow depressurization of the primary system. At 3 hr 45 min and 3 hr 47 min, 

respectively for the 100% and 200% TI-SGTR cases, the hot leg nozzle also fails due to a 

thermally-induced creep rupture. The failure of the hot leg nozzle leads to a rapid 

depressurization of the primary system and injection of the accumulator water. Following the 

depressurization of the RCS, the TI-SGTR flowrate drops to <0.2 kg/s through vessel failure at 

7 hr 30 min and 6 hr 51 min, for the 100% and 200% TI-SGTR cases respectively.

There were some differences in the timing of events for the 100% versus the 200% TI-SGTR 

cases following the opening of the TI-SGTR. As shown in Figure 5-62, the flow rate through the 

200% tube rupture case was approximately twice as large as the flow through the 100% tube 

rupture case. The net effect was: (a) increased heat removal from the core (see Figure 5-63), 

(b) a higher flow of gas past the hot leg nozzle, (c) a reduction in the zirconium oxidation rate in 

the 200% case (see Figure 5-64), and (d) a slightly faster depressurization rate in the 200% case 

versus the 100% case. The first two effects increased the heat flow past hot leg nozzle while 

the second two effects reduced the core exit temperature and the mechanical stress across the hot 

leg nozzle. The net effect was a slightly later hot leg creep rupture failure in the 200% case 

relative to the 100% case. Hence, the 100% case represented a condition that enhanced the core 

rate oxidation and accelerated core damage whereas the 200% case increased core cooling and 

decreased oxidation.

The vessel water level is shown in Figure 5-65. In response to flow out of the pressurizer safety 

relief valve, the pump seal leakage, and leakage flow through the TI-SGTR, the vessel water 

level drops into the core and uncovers the fuel. The fuel heatup leads to a natural circulation 

phase that fails a steam generator tube(s) and eventually a hot leg nozzle creep rupture failure. 

The RCS pressure drops rapidly once the hot leg nozzle fails and the accumulators dump to refill 

the core with water. As discussed above, the timing to the RCS hot leg failure is not 

significantly different between the 100% and 200% cases. 

Prior to the quench of the fuel by the accumulator water, the 100% case had more oxidation than 

the 200% case. Hence, the oxide layer thickness and potential for further oxidation following the 

core accumulator reflood was lower in the 100% case than in the 200% case. As shown in 

Figure 5-66, the zirconium cladding in the 200% case oxidizes at a higher rate than in the 100% 

case following the hot leg failure. Due to higher oxidation power in the 200% case in the 
5-68

post-reflood phase, the fuel degradation, the debris relocation to the lower head, and the failure 

of the vessel occurred faster in the 200% case. The fuel relocated to the lower plenum at 6.5 hr

and 6 hr in the 100% and 200% cases, respectively. The vessel failure occurred at 7 hr 30 min

and 6 hr 51 min in the 100% and 200% cases, respectively (see Figure 5-67).

Following vessel failure, the debris dropped into the reactor cavity under the reactor vessel. The 

hot debris immediately boiled away the water in the reactor cavity and started to ablate the 

concrete. The ex-vessel core-concrete interactions (CCI) continued for the remainder of the 

calculation, which generated non-condensable gases. In addition, the hot gases exiting the 

reactor cavity and the radioactive heating from settled fission products steadily evaporated the 

water on the containment floor outside the reactor cavity from the time of vessel failure to 

1.7 days (i.e., 41 hr in both cases). The resultant non-condensable and steam production 

pressurized the containment (see Figure 5-68). However, due to the TI-SGTR, there was a 

leakage pathway from the containment through the vessel. The TI-SGTR slowed the 

pressurization of the containment relative to the unmitigated STSBO. Due to the larger 

TI-SGTR leakage area in the 200% case, the containment failure area in the 200% case was 

smaller than in the 100% case (see relative leakage areas in Figure 5-69). In both cases, the 

containment continued to pressurize until the leakage flow balanced the steam and 

non-condensable gas generation. Hence, the containment failure leakage area increased in each 

case until the sum of the TI-SGTR and containment failure leakage areas balanced the gas 

generation. By 44 hr, all the water on the floor was evaporated and the steam generation 

stopped. The containment depressurized thereafter without any steam generation.

The conservatively assumed containment failure location was around the equipment hatch, which 

is located on the side of the containment without a surrounding building (e.g., the auxiliary or 

safeguards buildings); other locations, such as personnel airlocks and penetrations, would result 

in lower releases due to transport and deposition inside adjacent buildings. Consequently, the 

fission products that leaked from the containment are released directly to the environment.
5-69

Figure 5-60 Unmitigated 100% and 200% TI-SGTR STSBO primary and secondary 

pressures histories

Figure 5-61 Comparison of the hot leg nozzle and steam generator tube structure 

temperatures and creep rupture indices for the STSBO with TI-SGTR

Primary and Secondary Pressures

STSBO + 100% & 200% TI-SGTR - No Mitigation

0

2

4

6

8

10

12

14

16

18

0 1 2 3 4 5 6 7 8

Time (hr)

Pressure (MPa)

100% SGTR

200% SGTR

SG A

SG B

SG C - 100% SGTR

SG C - 200% SGTR

Hot leg C creep rupture

SG dryout

Thermally-induced SGTR

SG C SORV

Hot leg C creep rupture

Vessel Failure

100% 

SGTR

200% 

SGTR

0.00

0.25

0.50

0.75

1.00

300

400

500

600

700

800

900

1000

1100

0 1 2 3 4

Creep Rupture Index (-)

Temperature (K)

Time (hr)

Unmitigated Short-term Station Blackout

with Thermal-Induced SGTR

Hot Leg Nozzle

SG Tube - Hot stream

Hot Leg Creep Index

SG Tube Creep Index

SG TI-SGTR was specified at 3.5 hr to examine 

consequences. MELCOR would normally not 

predict a TI-SGTR until a creep index = 1.

HL nozzle failed due to creep 

rupture when the creep index = 1 
5-70

Figure 5-62 Unmitigated 100% and 200% TI-SGTR STSBO primary and TI-SGTR 

flowrate histories

Figure 5-63 Unmitigated 100% and 200% TI-SGTR STSBO vessel convective 

heat removal rate from the fuel

SGTR Flowrate

STSBO + 100% & 200% TI-SGTR - No Mitigation

0

2

4

6

8

10

12

0 6 12 18 24

Time (hr)

Flowrate (kg/s)

100% SGTR

200% SGTR

Core Convection Rate to Fluid

STSBO + 100% & 200% TI-SGTR - No Mitigation

0

25

50

75

3.50 3.55 3.60 3.65 3.70 3.75 3.80 3.85

Time (hr)

Power (MW)

200%

100%

Decay heat

Thermally-induced SGTR

Hot leg failure
5-71

Figure 5-64 Unmitigated 100% and 200% TI-SGTR STSBO fuel oxidation power before 

the RCS hot leg failure

Figure 5-65 Unmitigated 100% and 200% TI-SGTR short-term station blackout vessel 

two-phase coolant level

Comparison of Oxidation Powers

STSBO + 100% & 200% TI-SGTR - No Mitigation

0

25

50

75

3.50 3.55 3.60 3.65 3.70 3.75 3.80 3.85

Time (hr)

Power (MW)

100% SGTR

200% SGTR

Intense oxidation 

phase in 100% 

TI-SGTR

Hot leg failure

Decrease in 

oxidation rate in 

200% TI-SGTR

TI-SGTR

Vessel Water Level

STSBO + 100% & 200% TI-SGTR - No Mitigation

-4

-2

0

2

4

6

8

10

0 1 2 3 4 5 6 7 8

Time (hr)

Two-Phase Level (m)

Accumulators

BAF

TAF

Lower head

Vessel top

Vessel Failure

100% 

SGTR

200% 

SGTR

100% TI-SGTR

200% TI-SGTR
5-72

Figure 5-66 Unmitigated 100% and 200% TI-SGTR short-term station blackout fuel 

oxidation power after the RCS hot leg failure

Figure 5-67 Unmitigated 100% and 200% TI-SGTR short-term station blackout lower 

head inner and outer temperature histories

Comparison of Oxidation Powers

STSBO + 100% & 200% TI-SGTR - No Mitigation

0

25

50

75

3.75 4.00 4.25 4.50 4.75 5.00 5.25

Time (hr)

Power (MW)

100% SGTR

200% SGTR

Intense oxidation 

phase in 100% SGTR

Intense oxidation 

phase in 200% SGTR

Post acumulator dump intense 

oxidation phase in 200% SGTR
5-73

Figure 5-68 Unmitigated 100% and 200% TI-SGTR short-term station blackout 

containment pressure histories

Figure 5-69 Unmitigated 100% and 200% TI-SGTR short-term station blackout 

containment and TI-SGTR leakage areas

0

73

145

0

500000

1000000

1500000

0 1 2 3 4

Pressure (psi)

Pressure (Pa)

Time (days)

Containment Pressure response

STSBO + 100% & 200% TI-SGTR - No Mitigation

100% SGTR

200% SGTR

Emerg Spray

Cont failure

Emerg Godwin Spray 

Shutoff head

512 kPa (74 psi)

Start of 

containment high 

leakage

776 kPa (113 psi)

TI-SGTR and Containment Failure Leakage Areas

STSBO + 100% & 200% TI-SGTR - No Mitigation

0.0000

0.0005

0.0010

0.0015

0.0020

0.0025

0.0030

0.0035

0.0040

0 1 2 3 4

Time (days)

Leak Area (m2

)

100% SGTR

200% SGTR

100% TI-SGTR

200% TI-SGTR

TI-SGTR leakage areas

Containment failure leakage areas
5-74

5.3.1.2 Radionuclide Release

The fission product releases from the fuel started following the first thermal-mechanical failure

of the fuel cladding in the hottest rods at 2 hr 57 min, or 38 min after the uncovery of the top of 

the fuel rods. At 3 hr, the secondary safety relief valve on SG-C sticks open and allows the 

steam generator to depressurize to near atmospheric conditions. Prior to the TI-SGTR, any 

fission products leaving the RCS would flow out the pressurizer safety relief valve to the 

pressurizer relief tank in the containment. However, the PRT over-pressurized and failed prior to 

the start of the fission product releases. Hence, any fission products vented to the containment 

were not scrubbed in the PRT. 

Due to the complete loss of all feedwater at the start of the calculation, the water inventory in the 

steam generators decreased very rapidly and was completely boiled away by 1 hr 16 min. 

Consequently, there is no water on the secondary side of the steam generator after the TI-SGTR 

at 3 hr 33 min. Furthermore, since the steam generator relief valve stuck open at 3 hr, the 

released fission products can flow directly out the failed generator tube and through the stuck 

open relief valve to the environment. The flow of fission products through the tube rupture into 

the steam generator is very complicated and beyond the current modeling capabilities in 

MELCOR. Several decontamination mechanisms such as: (a) impaction, vena contracta effects 

at the tube rupture, (b) deposition in bends, and (c) capture by the secondary side tube grid 

spacers are not addressed by the MELCOR aerosol deposition models. It was estimated from 

ARTIST tests that the steam generator aerosol decontamination in a full-scale steam generator 

would be between 4.7 and 9 [29]. The normal aerosol capture and settling models were disabled 

on the secondary side in MELCOR and the secondary side decontamination factor was 

prescribed to be seven (i.e., approximately the average of 4.7 and 9).

Figure 5-70 and Figure 5-71 show the fission product distributions of the iodine radionuclides for 

the 100% and 200% TI-SGTR cases, respectively. The basic trends of the two cases were 

similar. The resultant distribution of iodine was partitioned between the RCS (i.e., including the 

vessel and the primary side of the steam generator tubes), the secondary side of the steam 

generators, the containment, and the environment. During the high release phase of the accident, 

the iodine is simultaneously released to the containment via the pressurizer safety relief valve, 

the secondary side of the steam generator and the environment via the TI-SGTR, or retained in 

the RCS. At the time of the TI-SGTR at 3 hr 33 min (0.15 days), only 8.4% of the iodine had 

been released from the fuel. About 1% was discharged to the containment via the pressurizer 

safety relief valve with the 7.4% retained in the RCS. 

Between the timing of the 100% TI-SGTR and vessel failure, 98% of the iodine was released 

with 80% in the containment, 15% retained in the RCS, 3.1% in the SG secondary, and 0.5% in 

the environment. The overall steam generator and steam line decontamination factor was ~7 

(i.e., the specified value). Eighty percent of the iodine transported to the containment versus 

only 3.6% in the steam generator secondary or the environment. The numbers were similar for 

the 200% TI-SGTR case with 97% released, 80% in the containment, 10% in the RCS, 5.3% in 

the steam generator secondary, and 0.8% in the environment. Due to the larger leak rate through 

the TI-SGTR, the 200% case had about twice the environmental release by vessel failure. The 

trends are similar for cesium, which are shown in Figure 5-72 and Figure 5-73.
5-75

The flow rate through the TI-SGTR decreased rapidly following hot leg failure at ~3.8 hr (see 

Figure 5-62), which slowed the release of the fission products to the faulted steam generator. 

Subsequently, the fission products moved from the reactor coolant system via the TI-SGTR 

rupture and the failed hot leg piping via natural circulation processes until the vessel lower head 

failure. As shown in Figure 5-70 through Figure 5-73, the releases to the containment or 

retention in the RCS increased most rapidly following hot leg failure. Prior to vessel failure, the 

fission product releases to the environment through the failed SGTR tube was roughly 

proportional to the size of the TI-SGTR leakage hole for the two cases (see Figure 5-74 and 

Figure 5-75). 

After the lower head vessel failure, the releases to the environment for the 100% TI-SGTR were 

faster than the 200% case. By 4 days, the iodine releases to the environment were almost 

identical between the two cases (see Figure 5-74) and the cesium releases were much closer than 

at vessel failure (Figure 5-75). As shown in Figure 5-69, the TI-SGTR leakage areas were 

smaller than the containment leakage areas. However, the 100% TI-SGTR case needed a larger 

containment failure area to remove energy than the 200% case. Consequently, there was more 

leakage from the containment in the 100% TI-SGTR case than in the 200% TI-SGTR case. Most 

of the releases through the TI-SGTR rupture were retained in the secondary side of the steam 

generator (i.e., a DF~7). In contrast, the fission products released through the containment 

failure went directly to the environment without any local retention or deposition in the leakage 

pathway through the tear in the containment wall. Since the 100% TI-SGTR case had more flow 

out the containment failure, the releases to the environment after the containment failure in the 

100% case were higher than in the 200% case. This non-intuitive trend eventually led to similar 

environmental releases for the two cases, which is evident in Figure 5-74 and Figure 5-75.

Finally, Figure 5-76 and Figure 5-77 summarize the releases of the radionuclides to the 

environment for the 100% and 200% cases, respectively. At 4 days, 95% of the noble gases, 

4.2% of the molybdenum, 1.5% of the iodine, 0.7-0.8% of the cesium, 2.7% (100%) and 1.5% 

(200%) of the tellurium, and 0.2% of the barium had been released to the environment.
5-76

Figure 5-70 Unmitigated 100% TI-SGTR STSBO iodine fission product distribution 

history

Figure 5-71 The unmitigated 200% TI-SGTR short-term station blackout iodine fission 

product distribution history

Iodine Distribution

STSBO + 100% TI-SGTR - No Mitigation

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

Total Release

Containment Total

RCS Total

SG Secondary

Env. Release

Environmental 

release = 1.5%

 1.5%

 4.5%

 8.5%

100%

85%

Iodine Distribution

STSBO + 200% TI-SGTR - No Mitigation

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

Total Release

Containment Total

SG Secondary

RCS Total

Env. Release

Environmental 

release = 1.5%

1.5%

6.3%

7.3%

99%

84%
5-77

Figure 5-72 The unmitigated 100% TI-SGTR short-term station blackout cesium fission 

product distribution history

Figure 5-73 The unmitigated 200% TI-SGTR short-term station blackout cesium fission 

product distribution history

Cesium Distribution

STSBO + 100% TI-SGTR - No Mitigation

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

Total Release

Containment Total

RCS Total

SG Secondary

Env. Release

Environmental 

release = 0.7%

100%

78%

18%

4%

0.7%

Cesium Distribution

STSBO + 200% TI-SGTR - No Mitigation

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

Total Release

RCS Total

SG Secondary

Containment Total

Env. Release

Environmental 

release = 0.85%

100%

77%

18%

4%

0.8%
5-78

Figure 5-74 The unmitigated 100% and 200% TI-SGTR short-term station blackout 

iodine fission product distribution history

Figure 5-75 The unmitigated 100% and 200% TI-SGTR short-term station blackout 

cesium fission product distribution history

Iodine Release to the Environment

STSBO with TI-SGTR - No Mitigation

0

0.02

0.04

0.06

0.08

0.1

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

100% TISGTR

200% TISGTR

1.5%

Cesium Release to the Environment

STSBO with TI-SGTR - No Mitigation

0.00

0.02

0.04

0.06

0.08

0.10

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

100% TISGTR

200% TISGTR

0.85%

0.67%
5-79

Figure 5-76 The unmitigated 100% TI-SGTR short-term station blackout environmental 

release history of all fission products

Figure 5-77 The unmitigated 200% TI-SGTR short-term station blackout environmental 

release history of all fission products

Fission Product Release to the Environment

STSBO + 100% TI-SGTR - No Mitigation

0.00001

0.0001

0.001

0.01

0.1

1

0 1 2 3 4

Time (days)

Fraction release (-)

NG

I

Cs

Ba

Te

Ru

Mo

Ce

La

U

Cd

Sn

95%

4.2%

2.7%

1.6%

0.7%

0.9%

0.2%

0.7%

Fission Product Release to the Environment

STSBO + 200% TI-SGTR - No Mitigation

0.00001

0.0001

0.001

0.01

0.1

1

0 1 2 3 4

Time (days)

Fraction release (-)

NG

I

Cs

Ba

Te

Ru

Mo

Ce

La

U

Cd

Sn

95%

4.2%

1.8%

1.5%

1.3%

0.2%

0.8%
5-80

5.3.2 Mitigated Short-Term Station Blackout with Thermally Induced Steam Generator 

Tube Rupture

Table 5-9 summarizes the timings of the key events in the mitigated STSBO with a TI-SGTR. 

One (i.e., equivalent of 100% flow area) steam generator tube failed prior to any other RCS 

creep rupture failures along with a stuck open secondary safety relief valve. Consequently, there 

is a containment bypass pathway for fission products once the steam generator tube fails. As 

described in Section 3.2, the accident scenario initiates with a complete loss of all onsite and 

offsite power. The reactor successfully trips and the containment isolates but all powered safety 

systems are unavailable. The mitigated STSBO credits the successful connection of the portable, 

low-pressure, diesel-driven (Godwin) pump to the containment spray system at 8 hr. The 

Godwin pump is a high-flow, low-head pump with a design capacity of 2000 gpm at 120 psi. A 

reliable source of water is maintained while 1,000,000 gallons is injected into the containment 

through the containment sprays. At the time of the analysis, there was no guidance in the 

emergency procedures for the duration of the spray operation or termination, so the 

1,000,000 gallons amount was somewhat arbitrarily selected. The sequence of events is identical 

to the unmitigated STSBO with a TI-SGTR until 8 hr. In particular, the core has degraded and 

failed the vessel lower head prior to the spray actuation (see Table 5-9). The emergency 

containment sprays are effective at reducing the containment pressure and scrubbing of airborne 

fission products from the containment atmosphere by spray droplets while they are operating. 

However, the containment subsequently pressurizes after the sprays are terminated to the failure 

pressure. While not investigated, intermittent operation of the sprays and deeper flooding could 

have further delayed failure of the containment. Section 5.3.2.1 summarizes the 

thermal-hydraulic response of the reactor and containment while Section 5.3.2.2 summarizes the 

associated radionuclide release from the fuel to the environment.

Table 5-9 The timing of key events for mitigated STSBO with TI-SGTR

Event Description

Time

(hh:mm)

Station blackout – loss of all onsite and offsite AC and DC 

power

MSIVs close

Reactor trip

RCP seal leak at 21 gpm/pump

TD-AFW starts but fails to inject due to ECST rupture

00:00

First SG SRV opening 00:03

SG dryout 01:14

Pressurizer SRV opens 01:27

PRT failure 01:47

Start of fuel heatup 02:19

RCP seal failures 02:46

First fission product gap releases 02:57

Stuck open SG PORV 03:00
5-81

Event Description

Time

(hh:mm)

SGTR 03:33

Creep rupture failure of the Loop C hot leg nozzle 03:47

Accumulator discharges 03:47

Accumulator empty 03:47

Vessel lower head failure by creep rupture 07:30

Debris discharge to reactor cavity 07:30

Cavity dryout (temporary) 07:54

Start of containment sprays 08:00

End of containment sprays (1,000,000 gal) 15:02

Containment at design pressure (45 psig) 44:10

Start of increased leakage of containment (P/Pdesign = 2.18) 74:48

5.3.2.1 Thermal-Hydraulic Response

The progression of events in the mitigated STSBO with TI-SGTR is identical to the unmitigated 

STSBO with TI-SGTR as described in Section 5.3.1 through the first 8 hr, which includes core 

degradation and vessel failure (e.g., compare the system pressure from Figure 5-78 and the 100% 

case in Figure 5-60 or the vessel level from Figure 5-79 and the 100% case in Figure 5-65). The 

portable emergency pump was connected to the containment spray system at 8 hours and begins 

injection. By 15 hours, 1,000,000 gallons were sprayed into the containment and the emergency 

injection was terminated. There was no guidance in the emergency procedures for the duration 

of the spray operation or termination; therefore, 1,000,000 gallons was an assumed value.

After the containment sprays terminated at 15 hours, the containment water was flooded to 

~0.1 m below the bottom of the vessel (see Figure 5-79). The water levels in the reactor cavity 

and the containment basement were approximately equal due to the hydraulic connection through 

the 12” hole in the reactor cavity wall at 2’-7” above the bottom of the floor. The reactor cavity 

also connects to the containment basement via a penetration at 24’-3” above the floor and the 

holes in the cavity wall for the RCS piping (i.e., nearly 40’ above the bottom of the floor). 

Similar to the response seen in mitigated STSBO (Section 5.2.23.2.4.2), the water level was too 

low to allow natural circulation from the containment basement into the reactor cavity and out 

the gaps and RCS piping penetrations present. Since the reactor cavity contains the fuel debris 

from the failed reactor vessel, the water heated to boiling once the sprays terminated. 

Intermittent spray operation and/or flooding above the RCS piping penetrations would have 

substantially delayed containment failure. Although the containment sprays did not prevent 

containment failure, they delayed containment failure by over ~46 hr relative to the unmitigated 

case.

The containment sprays are effective in quickly reducing the containment pressure. As shown in

Figure 5-80, the containment pressure would reach the shutoff head of the emergency portable 
5-82

pump by 17.5 hours. Based on the containment pressurization rate, it is estimated that there 

would be considerable additional time to connect the spray system. However, without additional

spray flow above the initial 1,000,000 gal, the containment will pressurize above the emergency 

pump shutoff head by 2.2 days (52 hours) and to failure conditions by 3.1 days (74.8 hours). See 

the long-term containment pressure response in Figure 5-81. 

The selection of the containment sprays as a mitigation technique for this scenario was 

particularly beneficial for several reasons, as previously discussed in Section 3.2.4.2. These 

benefits included aerosol knockdown in the containment, delaying containment failure by almost 

2 days, and deep flooding and cooling the ex-vessel debris. The spray operation reduced the 

flow out the failed steam generator tube to the environment (i.e., a containment bypass leakage 

path prior to containment failure). The impact of these benefits on the source term is discussed 

in Section 5.3.2.2.

Figure 5-78 The mitigated STSBO primary and secondary pressure history

Primary and Secondary Pressures

STSBO + 100% TI-SGTR - Mitigation with Portable Equipment

0

2

4

6

8

10

12

14

16

18

0 1 2 3 4 5 6 7 8

Time (hr)

Pressure (MPa)

Pressurizer

SG A

SG B

SG C

Hot leg C creep rupture

SG dryout

Thermally-induced SGTR

SG C SORV

Vessel failure
5-83

Figure 5-79 The mitigated short-term station blackout vessel two-phase coolant level

Figure 5-80 The mitigated short-term station blackout containment pressure history

Vessel Water Level

STSBO + 100% TI-SGTR - Mitigation with Portable Equipment

-8

-6

-4

-2

0

2

4

6

8

10

0 4 8 12 16

Time (hr)

Two-Phase Level (m)

Accumulators

BAF

TAF

Lower head

Vessel top

Vessel 

failure

Start emerg. diesel 

containment sprays

Terminate emerg. diesel 

containment sprays

Cavity water level

Vessel water level

Containment Pressure

STSBO + 100% TI-SGTR - Mitigation with Portable Equipment

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

0 12 24 36 48

Time (hours)

Pressure (MPa)

0

15

29

44

58

73

87

102

116

131

145

Pressure (psia)

RCS hot leg 

creep 

rupture 

failure

Vessel 

failure

Emerg Godwin 

Spray Shutoff head

512 kPa (74 psi)

Time to 512 kPa 

in unmitigated 

STSBO 

+ 100% TI-SGTR

(17.5 hr)
5-84

Figure 5-81 The mitigated short-term station blackout containment pressure history

5.3.2.2 Radionuclide Release

The radionuclide response of the mitigated STSBO with a TI-SGTR is identical to the 

unmitigated response described in Section 5.3.1.2 for the first 8 hr, or through vessel failure until 

the start of the containment sprays. Following the start of the emergency containment sprays at 

8 hr (0.33 days), the airborne aerosols of iodine and cesium rapidly decease (see Figure 5-82 and

Figure 5-83, respectively). By the time the sprays are terminated at 15 hr (0.63 days), almost all 

of the airborne aerosols have been captured in the pool on the containment floor. Since the 

containment failure was delayed until 74 hr 48 min (3.1 days), natural settling of the airborne 

mass in the containment was also significant, which is reflected in the small environmental 

release of iodine and cesium (i.e., 0.5% and 0.4%, respectively). However, natural settling was 

also effective in the unmitigated case, which does not occur until 27 hr 54 min. As will be 

discussed next, the spray water was important in preventing revaporization. This was the most 

significant difference between the mitigated and unmitigated cases.

Due to the deep flooding in the reactor cavity by the spray operation, the bottom of the failed 

vessel lower head is at the top of the water level in the reactor cavity.

33 Therefore, the natural 

hot circulation flow that promoted revaporization in the unmitigated STSBO is not present. 

Instead, the water pool in the reactor cavity cools the bottom of the vessel. Due to some boiling 

in the cavity, a relatively cool flow of steam passes through the vessel and out the failed hot leg 

nozzle location, which also removes heat and inhibits revaporization of deposited radionuclides 

 33 Although the level is 0.1 m below the inside of reactor vessel, the level covers the bottom of the outside surface 

of the lower head, which is 0.13 m thick. Hence, the water blocks the flow of air into the reactor vessel through 

the lower head failure hole.

Containment Pressure

STSBO + 100% TI-SGTR - Mitigation with Portable Equipment

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4

Time (days)

Pressure (MPa)

0

15

29

44

58

73

87

102

116

131

145

Pressure (psia)

Start of high 

leakage 

(74.8 hr)

RCS hot leg creep rupture failure

Vessel 

failure

Containment 

spray 

operation
5-85

in the upper vessel and hot leg. Consequently, the revaporization of the in-vessel deposited 

fission products (i.e., especially cesium-iodine) that was seen in the unmitigated STSBO with a 

TI-SGTR is characteristic of revaporization) was negligible in the mitigated case through 4 days

(see Figure 5-84 and Figure 5-85).

Figure 5-82 The iodine distribution in the containment for short-term station blackout 

with a 100% thermally-induced SGTR with spray mitigation

Iodine Distribution

STSBO + 100% TI-SGTR - MitigatIion with Portable Equipment

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

Containment Deposited

Containment Airborne

Env. Release

Environmental 

release = 0.5%

82% deposited in containment

Start of containment sprays

Containment failure

End of containment sprays
5-86

Figure 5-83 The cesium distribution in the containment for short-term station blackout 

with a 100% thermally-induced SGTR with spray mitigation

Figure 5-84 The short-term station blackout with a 100% thermally-induced SGTR with 

and without spray mitigation iodine environmental release

Cesium Containment Distribution

STSBO + 100% TI-SGTR - MitigatIion with Portable Equipment

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

Containment Deposited

Containment Airborne

Env. Release

Environmental 

release = 0.4%

 79% deposited in containment

Start of containment sprays

Containment failure

End of containment sprays

Iodine Release to the Environment

STSBO + 100% TI-SGTR - With and Without Mitigation with Portable Equipment

0.00

0.02

0.04

0.06

0.08

0.10

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

Without sprays

With sprays

1.5%

0.5%
5-87

Figure 5-85 The short-term station blackout with a 100% thermally-induced SGTR with 

and without spray mitigation cesium environmental release

5.3.3 Uncertainties in the Failure of the Thermally-Induced Steam Generator Tube 

versus the Hot Leg 

During the peer review of the unmitigated STSBO with a thermally-induced tube rupture, there 

were questions about the competing events of a thermally-induced steam generator tube rupture 

versus the hot leg creep rupture failure. The probability of a TI-SGTR has previously been 

assessed to be 0.25 [30]. Consequently, calculations were performed in Section 5.3 with a TISGTR to supplement the calculations described in Section 5.2 without tube failures. More recent 

research has investigated the relative timing of the TI-SGTR relative to creep rupture failure of 

the hot leg with mechanistic simulations of natural circulation flow patterns[16][17][31][32]. 

The results of the research show comparable timings for hot leg creep rupture failure and 

thermally-induced steam generator tube failure, for a flawed tube at maximum thermal stress 

conditions, with the former slightly preceeding the later for most conditions. 

MELCOR also predicts failure of a hot leg prior to failure of any steam generator tubes 

(i.e., potential failures are monitored at both locations). Consequently, the calculations presented 

previously in Section 5.3 increased the mechanical stress across the tubes by prescribing a 

stuck-open safety relief valve and an increase in the thermal stress by inducing tube failure at a 

lower criterion than the default model. Subsequent to the failure of the steam generator tube, the 

hot leg failed and mitigated the magnitude of the potential release of radionuclides that bypass 

the containment. 

To investigate the relative vulnerability of the hot leg to a TI-SGTR, a sensitivity calculation was 

performed with MELCOR where the failure of the hot leg was prevented. Figure 5-86 shows the 

Cesium Release to the Environment

STSBO + 100% TI-SGTR - With and Without Mitigation with Portable Equipment

0.00

0.02

0.04

0.06

0.08

0.10

0 1 2 3 4

Time (days)

Fraction of Inventory (-)

Without sprays

With sprays

0.7%

0.4%
5-88

creep rupture damage index of the hot leg. The steam generator tube failed at 3.55 hr. Hot leg 

failure was predicted 14 min later at 3.8 hr when the failure index reached a lifetime value of 1. 

Vessel failure was calculated to occur at 5.3 hr in the sensitivity calculation. Between 3.8 and 

5.3 hr, the damage index increased from 1 to greater than four orders of magnitude larger. The 

creep index is highly sensitive to the thermal response of the hot leg nozzle as very hot gases 

continue to flow from the core (see hot leg temperature responses in Figure 5-87 ). 

Figure 5-88 includes the iodine release to the environment for the failure and no failure case. As 

discussed in Section 5.3.1.2, there is a direct pathway for radionuclide releases to the 

environment through the failed steam generator tube prior to hot leg failure. However, the iodine 

release to the environment essentially stopped once the hot leg failed. Between 3.8 hours and 

4 hours, the hot leg creep failure index in the no hot leg failure sensitivity case increased more 

than an order of magnitude (i.e., a factor of 18) above the best-estimate failure value. The iodine 

release to the environment increased by a factor of three during this time to a 0.6% release. 

Consequently, there is a high sensitivity to the creep failure damage index at high temperatures 

that quickly increases the index above the failure threshold.

In summary, it is not credible that the hot leg would not fail by creep rupture in the examined 

scenarios. The conditions that lead to the TI-SGTR are the same conditions that promote hot leg 

failure. As discussed in Section 5.3.1, the TI-SGTR increased heat removal from the core and 

the heat flow past the hot leg nozzle. The best-estimate creep rupture damage index is rapidly 

increasing near the time of the TI-SGTR. Within 10 minutes after the best-estimate failure time 

of the hot leg nozzle, the creep rupture damage index has increased by an order of magnitude due 

to the strong dependence of the nozzle strength to temperature. There is a factor of three

increase in the iodine release to the environment while the creep rupture index increases to an 

order of magnitude larger. However, the release of iodine to the environment was only 0.6% at 

the order-of-magnitude higher damage value.

Three sensitivity calculations were also performed using the SCDAP/RELAP5 code and 

associated natural circulation severe accident model [16][17][31]. The best-estimate parameters 

in the SCDAP/RELAP5 calculation were based on the latest FLUENT CFD research [32]. 

Unlike the MELCOR calculation, which used a specified criterion to create the TI-SGTR 

(i.e., specified to occur ~10 min prior hot leg failure timing from the STSBO in Section 5.2.1 ), 

the SCDAP/RELAP5 simulation tied the TI-SGTR to stress enhancing vulnerabilities due to 

flaws developed during in-service operation. The three SCDAP/RELAP5 cases examined: (1) a 

TI-SGTR in the hottest portion of the natural circulation plume and a stress multiplier of two, (2) 

a TI-SGTR in the hottest portion of the natural circulation plume and a stress multiplier of three, 

and (3) multiple tube failures with a stress multiplier of two. The results of the 

SCDAP/RELAP5 study (i.e., shown in Table 5-10) confirmed that: (a) TI-SGTR will not 

preclude hot leg creep rupture failure, and (b) hot leg creep rupture failure occurs within minutes 

of the TI-SGTR for a range of tube stress conditions. 
5-89

Table 5-10 The timing of hot leg failure for SCDAP/RELAP5 simulations with TI-SGTR

Case

Delay of Hot Leg Failure 

after TI-SGTR

(min)

1. Steam generator tube stress multiplier of 2 1.2

2. Steam generator tube stress multiplier of 3 8.8

3. Multiple steam generator tubes w/stress multiplier of 2 1.3

Figure 5-86 The hot leg creep rupture failure index in the short-term station blackout 

sensitivity case with a 100% thermally-induced SGTR and no hot leg failure

0.1

1

10

100

1000

10000

0 1 2 3 4 5 6

Failure Index (-)

Time (hr)

STSBO + TI-SGTR

Hot Leg Creep Rupture Failure Index

Larson Miller Creep 

Rupture Failure Criterion

Thermally-induced SGTR

Vessel failure in no HL 

failure sensitivity
5-90

Figure 5-87 The hot leg temperature response in the thermally-induced steam generator 

tube rupture cases with and without hot leg failure

Figure 5-88 The hot leg creep rupture failure index and iodine release to environment for 

the thermally-induced SGTR cases with and without hot leg failure

5.4 Spontaneous SGTR

The spontaneous SGTR sequence is a double-ended-guillotine rupture of a single steam 

generator tube occurring while the reactor system is operating at normal conditions. 

0.0001

0.001

0.01

0.1

1

0.01

0.1

1

10

100

0 1 2 3 4

Fraction of Inventory (-)

Failure Index (-)

Time (hr)

Hot Leg Creep Rupture Failure Index

Larson Miller Creep 

Rupture Failure Criterion

Thermally-induced SGTR

Vessel failure in base case

Iodine release to environment with HL failure

Iodine release to environment 

with no HL failure sensitivity

0.006

0.002

18
5-91

Section 5.4.1 presents the results of a mitigated scenario where the expected operator actions are 

successful. For the unmitigated scenario in Section 5.4.2, the operator fails to isolate the faulted 

steam generator or cooldown the RCS using the two intact generators. Finally, the unmitigated 

scenario in Section 5.4.3 has the same failed operator actions as the previous unmitigated 

scenario. In addition, the relief valve on the faulted generator is assumed to fail open when 

water from the SGTR fills the steam generator and flows through the valve.

5.4.1 Mitigated Spontaneous Steam Generator Tube Rupture

Table 5-11 summarizes the timing of the key events in the mitigated spontaneous steam 

generator tube rupture with expected operator actions. As described in Section 3.3.1, the 

accident scenario initiates with a spontaneous double-ended guillotine rupture of one steam 

generator tube. After about three minutes, the reactor successfully trips and the containment 

isolates. The full complement of systems at Surry Power Station is considered functional in this 

scenario including all systems associated with engineered safeguards and instrumentation and 

control as well as all auxiliary and emergency systems. Operator actions are successful at 

isolating the faulted steam generator and cooling the reactor system to permit operation of the 

residual heat removal (RHR) system. Section 5.4.1.1 summarizes the thermal-hydraulic 

response of the reactor and containment while Section 5.4.1.2 summarizes the associated 

radionuclide release from the fuel to the environment.
5-92

Table 5-11 Timing of key events for the Spontaneous SGTR with Expected Operator 

Action

Event Description

Time

(hh:mm)

Spontaneous SGTR 00:00

Reactor scram 00:03

Turbine stop valves close 00:03

Steam dump valves open and modulate Not 

accomplished*

Steam dump valves close (RCS temperature < 547 °F) Not 

accomplished*

HHSI initiated (3 pumps) 00:03

First AFW delivery 00:03

Operators take control of AFW (as the SGs are overfilling) 00:15

AFW delivery to faulted steam generator secured** 00:15

1 of 3 HHSI pumps secured 00:15

Faulted steam generator flooded 00:20

TDAFW fails (turbine floods) 00:20

Faulted steam generator PORV 1st lifts 00:23

Faulted steam generator isolated 02:30

HHSI secured 02:30

Leakage through faulted steam generator PORV stopped*** 02:30

100 °F/hr cool-down initiated 02:30

RHR entry pressure (400 – 450 psig) achieved 03:18

(450 psig)

RHR entry temperature (350°F) achieved 03:43

* The automatic operation of steam dump valves was not represented in the Surry MELCOR model. 

The thermal-hydraulic signature in the subject calculation suggests that the valves might be active 

for the first 6 min following scram but at no other time. Valve action would reduce RCS 

temperature by ~25 °F for the first few min and by a few °F in the next few min. The differences are 

thought to be inconsequential.

** AFW to the faulted steam generator is stopped on account of runaway high level in the generator.

*** Isolating the faulted steam generator, i.e., closing the MSIVs serving it, in combination with 

securing HHSI, stops the leakage through the faulted steam generator PORV.
5-93

5.4.1.1 Thermal-hydraulic Response

Figure 5-89 through Figure 5-95 present the thermal hydraulic response predicted by MELCOR 

for spontaneous SGTR at Surry where reactor systems operate as designed and reactor operators 

respond as expected per training and procedure. System pressure histories are shown in 

Figure 5-89. Figure 5-90 shows RCS conditions relative to RHR entry conditions. Figure 5-91

and Figure 5-92 show reactor pressure vessel level and fuel cladding and reactor vessel lower 

head temperatures, respectively. RWST and ECST inventories are tracked in Figure 5-93 while 

containment pressure is tracked in Figure 5-94. Steam generator level is tracked in Figure 5-95. 

Operator actions in this scenario are those expected per training and procedure. Specifically, the 

operators:

• Secure AFW delivery to the steam generator with the broken tube (the faulted steam 

generator) at 00:15:00, i.e., 15 min after the SGTR event, in response to runaway high 

level in the generator.

• Secure 1 of the 3 total HHSI pumps at 00:15:00

• Isolate the faulted steam generator, i.e., close the MSIVs serving the faulted steam 

generator at 2:30:00

• Secure HHSI at 2:30:00 (which in combination with the above actions ends the RCS 

leakage)

• Initiate a 100° F/hr cool-down of the RCS at 2:30:00

• Accomplish RHR entry at 03:43:00 (predicted)

The spontaneous SGTR quickly leads to a reactor scram, turbine stop valve closure, HHSI 

actuation, and AFW delivery. The flow of primary system coolant through the tube rupture into 

the secondary side of the faulted steam generator results in a fairly sustained leak to the 

environment beginning at 23 min when the PORV serving the generator first lifts.

Once level control of AFW is initiated in the MELCOR calculation (i.e., at 15 min simulating the 

operators taking control), HHSI stably removes core decay heat up to the time when operator 

action to end the leakage of coolant to the environment is represented (i.e., at 2.5 hr) and 

operator actions to depressurize and cool down the RCS are modeled to begin (i.e., also at 

2.5 hr). There is no heat removal by the intact steam generators during this time as evidenced in 

Figure 5-93 where no drawdown on ECST inventory shows to occur. Note that pressure in the 

intact steam generators during this period is low relative to the setpoint of the main steam line 

PORVs such that no venting of the generators occurs and hence no heat removal is accomplished 

by them. Coinciding flow out the tube rupture is governed by the head-versus-flow 

characteristics of the HHSI pumps, the flow resistance at the tube rupture, and the set point of the 

PORV on the faulted steam generator. Operator action to secure HHSI, modeled to occur at 

2.5 hrs, allows the PORV to reseat shortly thereafter ending leakage through the tube rupture. 

The 100°F/hr cool-down initiated at 2.5 hrs gradually brings the temperature of the RCS down to 

RHR entry temperature. RHR entry pressure is reached somewhat earlier, however, the timing 
5-94

may be accelerated given that no active pressure control was represented in the MELCOR 

calculation, i.e., no pressurizer heater operation was modeled. The 100°F/hr cool-down in the 

MELCOR calculation was accomplished realistically in that the intact steam generators were 

vented in a controlled fashion while AFW was delivered as needed to maintain level. 

The results of the MELCOR calculation simulating an SGTR with expected operator action show 

that RHR entry conditions would be achieved without challenging the RWST inventory and 

without immoderately draining the ECST. No uncovering or overheating of the reactor core 

would occur and no damage to the core would result. 

Figure 5-89 SGTR with Operator Action – System Pressures

 

0

2

4

6

8

10

12

14

16

18

0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5

Time (hr)

Pressure (MPa)

SG A (faulted SG)

SG B

SG C

Pressurizer

HHSI secured

Broken SG isolated

Leak through faulted SG PORV stopped

100 °F/hr cool-down initiated

Operators take control of AFW

 - AFW to SG A (faulted SG) secured

 - AFW to SGs B & C interrupted due to high level
5-95

Figure 5-90 SGTR with Operator Action – RCS Conditions Relative to RHR Entry 

Conditions

Figure 5-91 SGTR with Operator Action – RPV Level

 

0

2

4

6

8

10

12

14

16

18

0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5

Time (hr)

Pressure (MPa)

300

350

400

450

500

550

600

Tavg (K)

Pressure

Tavg

HHSI secured

Broken SG isolated

Faulted SG PORV leakage stopped

100 °F/hr cool-down initiated

RHR entry temp (350 degF)

RHR entry press (450 psig)

 

0

2

4

6

8

10

12

14

0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5

Time (hr)

Level (m)

Vessel top

Vessel bottom

BAF

TAF
5-96

Figure 5-92 SGTR with Operator Action – Maximum Cladding and Lower 

Head Temperatures

Figure 5-93 SGTR with Operator Action – RWST and ECST Inventories

 

300

350

400

450

500

550

600

650

0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5

Time (hr)

Temperature (K)

Cladding

Lower head outer

Lower head inner

Saturation

 

0

50000

100000

150000

200000

250000

300000

350000

400000

0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5

Time (hr)

Inventory (gal)

RWST

ECST
5-97

Figure 5-94 SGTR with Operator Action- Containment Pressure

Figure 5-95 SGTR with Operator Action- Steam Generator Level

5.4.1.2 Radionuclide Release

No fission product releases from the reactor core occurred in the spontaneous SGTR with 

expected operator actions. 
5-98

5.4.2 Unmitigated - Spontaneous SGTR with Failed Operator Action

Table 5-12 summarizes the timing of the key events in the spontaneous steam generator tube 

rupture with failed operator actions. As described in Section 3.3.1, the accident scenario initiates

with a spontaneous double-ended-guillotine rupture of one steam generator tube. After about 

three minutes, the reactor successfully trips and the containment isolates. The full complement

of systems at Surry Power Station is considered functional in this scenario including all systems 

associated with engineered safeguards and instrumentation and control as well as all auxiliary 

and emergency systems. However, the operator actions are not successful at isolating the faulted 

steam generator or cooling the reactor system to permit operation of the RHR system. 

Eventually, the refueling water storage tank and the ECST are exhausted (i.e., after ~11 hr and 

~33.5 hr, respectively) leading to core uncover (i.e., beginning at ~43.75 hr) and core damage. 

Due to the long amount of time to core damage, it was judged unlikely that the operators would 

not correct missed actions (i.e., failure to isolate the faulted SG, failure to cool down and 

depressurize, and failure to refill the RWST or connect to the unaffected unit’s RWST). 

Section 5.4.2.1 summarizes the thermal-hydraulic response of the reactor and containment while 

Section 5.4.2.2 summarizes the associated radionuclide release from the fuel to the environment.
5-99

Table 5-12 Timing of key events for the Spontaneous SGTR with Failed Operator 

Action

Event Description

Time

(hh:mm)

Spontaneous SGTR 00:00

Reactor scram 00:03

Turbine stop valves close 00:03

Steam dump valves open and modulate Not accomplished*

Steam dump valves close (RCS temperature < 547 °F) Not accomplished*

HHSI initiated (3 pumps) 00:03

First AFW delivery 00:03

Operators take control of AFW as the SGs are overfilling 00:10**

AFW delivery to faulted steam generator secured 00:12***

1 of 3 HHSI pumps secured 00:15

Faulted steam generator PORV 1st lifts 00:32

Faulted steam generator flooded 00:42

TDAFW fails (turbine floods) 00:42

Faulted steam generator isolated Not accomplished by 

operators

HHSI secured Not accomplished by 

operators

Leakage through faulted steam generator PORV stopped Not accomplished by 

operators

100 °F/hr cool-down initiated Not accomplished by 

operators

RWST exhausted (safety injection ends) 11:03

RCPs trip 18:22

Steam Generator C PORV fails open due to excessive cycling 31:00

First accumulator discharge 31:16

ECST exhausted (AFW delivery ends) 33:29

Steam Generator B PORV fails open due to excessive cycling 38:20

Core uncovering begins 43:48

First fission product gap release 45:46

* The automatic operation of steam dump valves was not represented in the Surry MELCOR model. The 

thermal-hydraulic signature in the subject calculation suggests that the valves might be active for the first 

6 min following scram but at no other time. Valve action would reduce RCS temperature by ~25 °F for the 

first few min and by a few °F in the next few min. The differences are thought to be inconsequential.

** Best-estimate timing for operators assuming manual control of AFW is 15 min. A discrepancy in the 

MELCOR input initiated level control of AFW at 10 min.

*** Best-estimate timing for operators securing AFW delivery to the faulted steam generator is 15 min. A 

discrepancy in the MELCOR input interrupted AFW to the steam generator at 12 min and 30 sec. AFW to the 

faulted steam generator is stopped on account of runaway high level in the generator
5-100

5.4.2.1 Thermal-hydraulic Response

Figure 5-96 through Figure 5-101 present the thermal hydraulic response predicted by the 

MELCOR for a spontaneous SGTR at Surry where reactor systems operate as designed but 

reactor operators fail to accomplish key actions per training and procedure. Specifically, the 

operators fail to depressurize and cool down the RCS. System pressure histories are shown in 

Figure 5-96. Figure 5-97 and Figure 5-98 show reactor pressure vessel level and fuel cladding 

and reactor vessel lower head temperatures, respectively. RWST and ECST inventories are 

tracked in Figure 5-99 while containment pressure is tracked in Figure 5-100. Steam generator 

level is tracked in Figure 5-101. 

Operator actions are remiss in this scenario with respect to training and procedure in that the 

operators:

• Fail to isolate the faulted steam generator

• Fail to depressurize and cool down the RCS

• Fail to extend the available duration of ECCS injection by refilling the RWST or cross 

connecting to the other Surry unit’s RWST.

The tube rupture quickly leads to a reactor scram, turbine stop valve closure, HHSI actuation, 

and AFW delivery. The flow of primary system coolant through the tube rupture into the 

secondary side of the faulted steam generator results in a fairly sustained leak to the environment 

beginning at 32 min when the PORV serving the generator first lifts.

Once level control of AFW is initiated in the MELCOR calculation (at 10 min simulation the 

operators taking control), HHSI stably removes core decay heat and reduces core temperatures 

up to the time when the useable inventory of the RWST is exhausted thereby ending injection. 

There is no heat removal by the intact steam generators during this time as evidenced in 

Figure 5-99 where no drawdown on ECST inventory shows to occur. Note that pressure in the 

intact steam generators during this period is low relative to the setpoint of the main steam line 

PORVs such that no venting of the generators occurs and hence no heat removal is accomplished 

by them. Coinciding flow out the tube rupture is governed by the head-versus-flow 

characteristics of the HHSI pumps, the flow resistance at the tube rupture, and the set point of the 

PORV in the faulted steam generator. The RCS heats to saturation over the course of 

several hours following the RWST being exhausted and an extended boil-off of the RCS 

inventory begins. The RCPs are stopped in the MELCOR calculation at the first occurrence of 

void in the RCS simulating the pumps tripping on their own or the operators shutting them down 

on account of erratic performance. As the RCS heats to saturation, the intact steam generators 

pressurize up to the setpoint on the main steam line PORVs and the generators function to 

remove heat from the RCS. This remains the case until the ECST is exhausted. (Note that once 

the RCS begins to void and the RCPs stop, heat rejection to the steam generators is by reflux 

cooling only.) Late in the course of the boil-off of the RCS inventory, the pressure history of the 

system is influenced strongly by failures of the PORVs serving the intact steam generators. The 

PORVs are modeled to fail open due to excessive cycling for a particular per-demand failure 

probability and a particular cumulative distribution function value according to the relation:

P(n) = 1-(1-Pd)n
5-101

where P(n) is the cumulative distribution function value, Pd is the per-demand failure 

probability, and n is the number of cycles. Given a per-demand failure probability of 0.0058 

(i.e., per Surry’s response to a SOARCA information request) and an assumed cumulative 

distribution function value of 0.5, the valves fail open after 119 cycles. Following failure of the 

PORVs serving the intact steam generators, dramatic reductions in system pressure result as the 

steam generators blow down. Note that over-cycle failure of the PORV serving the faulted steam 

generator was prevented in the subject MELCOR calculation. If this had not been done, 

inordinate cycling of the simplistically-modeled valve would have occurred as liquid entered it 

from the flooded generator and the valve would have failed open at 70 min. The sensitivity of 

the progression of the accident to the status of this PORV is addressed by the calculation of the 

next Section where the valve is failed open when liquid first flows through it. 

The MELCOR calculation predicts that core uncovering initiates at 43 hr and 48 min. The first 

release of fission products from a fuel/cladding gap is seen at 45 hr and 46 min. Traces in 

Figures 99 through 104 end at first gap release and figures of fission product tracking are not 

presented for the subject scenario on account of the great unlikelihood that operators would fail 

to depressurize and cool down the reactor system for some 43 hours.

Figure 5-96 SGTR with Failed Operator Action – System Pressures

 

0

2

4

6

8

10

12

14

16

18

0 4 8 12 16 20 24 28 32 36 40 44 48

Time (hr)

Pressure (MPa)

SG A (faulted SG)

SG B

SG C

Pressurizer

RWST exhausted - HHSI ends

SG C PORV fails open

SG B PORV fails open

1st gap release

SG C dry SG B dry

ECST empty
5-102

Figure 5-97 SGTR with Failed Operator Action – RPV Water Level

Figure 5-98 SGTR with Failed Operator Action – Maximum Cladding and Lower Head 

Temperatures

0

2

4

6

8

10

12

14

0 4 8 12 16 20 24 28 32 36 40 44 48

Level (m)

Time (hr)

Vessel top

Vessel bottom

BAF

TAF

RCP stop

SG C 

PORV 

fails open

SG C 

dryout SG B 

dryout

RWST empty 1st gap release

Accumulator 

injection

SG B 

PORV 

fails open

 

300

400

500

600

700

800

900

1000

1100

1200

0 4 8 12 16 20 24 28 32 36 40 44 48

Time (hr)

Temperature (K)

Cladding

Lower head outer

Lower head inner

Saturation

RWST exhausted

Core uncovery begins

1st gap release
5-103

Figure 5-99 SGTR with Failed Operator Action – RWST and ECST Inventories

Figure 5-100 SGTR with Failed Operator Action-Containment Pressure

 

0

50000

100000

150000

200000

250000

300000

350000

400000

0 4 8 12 16 20 24 28 32 36 40 44 48

Time (hr)

Inventory (gal)

RWST

ECST

Useable inventory exhausted 1st gap release
5-104

Figure 5-101 SGTR with Failed Operator Action- Steam Generator Level

5.4.2.2 Radionuclide Release

The radionuclide release analysis is not presented for this scenario due to the low likelihood that 

operators would fail to depressurize and cool down the reactor system for the 43 hr necessary for 

the reactor core to begin to uncover (i.e., predicted by MELCOR). Table-top exercises performed 

with operators during site visits substantiate the low likelihood that operators would fail to 

depressurize and cool down the reactor system especially given the ample time for TSC and EOF 

intervention. 

5.4.3 Unmitigated - Spontaneous SGTR with Failed Operator Action and Faulted Steam 

Generator SORV

Table 5-13 summarizes the timing of the key events in the SGTR with failed operator actions, 

and as a resultant SORV on the faulted steam generator (unmitigated). As described in 

Section 3.3.1, the accident scenario initiates with a spontaneous double-ended-guillotine rupture

of one steam generator tube. After about three minutes, the reactor successfully trips and the 

containment isolates. The full complement of systems at Surry Power Station is considered 

functional in this scenario including all systems associated with engineered safeguards and 

instrumentation and control as well as all auxiliary and emergency systems. However, operator 

actions are not successful at isolating the faulted steam generator or at cooling the reactor system 

(i.e., to permit operation of the RHR system) and the PORV on the faulted steam generator fails 

open when liquid first flows through it. Eventually the RWST is exhausted (i.e., after ~8.75 hr) 

leading to core uncover (i.e., beginning at ~22.75 hr) and core damage. Due to the long amount 

of time to core damage, it was judged unlikely that the operators would not correct missed 

actions (i.e., failure to isolate the faulted SG, failure to cool down and depressurize, and failure 

to refill the RWST or connect to the unaffected unit’s RWST). Section 5.4.3.1 summarizes the 

thermal-hydraulic response of the reactor and containment while Section 5.4.3.2 summarizes the 

associated radionuclide release from the fuel to the environment.
5-105

Table 5-13 The timing of key events for the Spontaneous SGTR with Failed Operator 

Action and Faulted Steam Generator SORV

Event Description

Time

(hh:mm)

Spontaneous SGTR 00:00

Reactor scram 00:03

Turbine stop valves close 00:03

Steam dump valves open and modulate Not accomplished*

Steam dump valves close (RCS temperature < 547 °F) Not accomplished*

HHSI initiated (3 pumps) 00:03

First AFW delivery 00:03

Operators take control of AFW (as the SGs are overfilling) 00:10**

AFW delivery to faulted steam generator secured 00:12***

1 of 3 HHSI pumps secured 00:15

Faulted steam generator PORV 1st lifts 00:32

Faulted steam generator flooded 00:42

TDAFW fails (turbine floods) 00:42

Faulted steam generator PORV fails open (1st liquid flow through valve) 00:44

Faulted steam generator isolated Not accomplished by 

operators

HHSI secured Not accomplished by 

operators

Leakage through faulted steam generator PORV stopped Not accomplished by 

operators

100 °F/hr cool-down initiated Not accomplished by 

operators

RWST exhausted (safety injection ends) 08:43

First accumulator discharge 08:53

RCPs trip 12:43

Core uncovering begins 22:48

First fission product gap release 26:44

* The automatic operation of steam dump valves was not represented in the Surry MELCOR model. The 

thermal-hydraulic signature in the subject calculation suggests that the valves might be active for the first 

6 min following scram but at no other time. Valve action would reduce RCS temperature by ~25 °F for the 

first few min and by a few °F in the next few min. The differences are thought to be inconsequential.

** Best-estimate timing for operators assuming manual control of AFW is 15 min. A discrepancy in the 

MELCOR modeling initiated level control of AFW at 10 min.

*** Best-estimate timing for operators securing AFW delivery to the faulted steam generator is 15 min. A 

discrepancy in the MELCOR modeling interrupted AFW to the steam generator at 12 min and 30 sec. AFW 

to the faulted steam generator is stopped on account of runaway high level in the generator.
5-106

5.4.3.1 Thermal-hydraulic Response

Figure 5-102 through Figure 5-107 present the thermal hydraulic response predicted by 

MELCOR for an accident with the same initiation event, system availabilities, and mitigative 

actions as the accident of the preceding Section with one distinction. The distinction is the 

inclusion of an additional mechanism for the PORV serving the faulted steam generator where 

the valve is failed open when liquid first flows through it. System pressure histories are shown 

in Figure 5-102. Figure 5-103 and Figure 5-104 show reactor pressure vessel level and fuel 

cladding and reactor vessel lower head temperatures, respectively. RWST and ECST inventories 

are tracked in Figure 5-105, while containment pressure is tracked in Figure 5-106. Steam 

generator level is tracked in Figure 5-107.

As in the accident of the preceding section, the SGTR quickly leads to a reactor scram, turbine 

stop valve closure, HHSI actuation, and AFW delivery. The flow of primary system coolant 

through the tube rupture into the secondary side of the faulted steam generator results in a fairly 

sustained leak to the environment beginning at 32 min when the PORV serving the generator 

first lifts. The faulted steam generator floods at 42 min and liquid first flows through the PORV 

serving the generator at 44.5 min. The PORV is failed open at this time in the MELCOR 

calculation instituting a sustained leak from the RCS to the environment. The assumption here is 

that the valve is not designed to pass violent critical flows of flashing liquid and would hence 

fail.

Once level control of AFW is initiated in the MELCOR calculation (at 10 min simulating the 

operators taking control), HHSI stably removes core decay heat and reduces core temperatures 

up to the time when the useable inventory of the RWST is exhausted thereby ending injection. 

There is no heat removal by the intact steam generators during this time as evidenced in 

Figure 5-105 where no drawdown on ECST inventory shows to occur. Note that pressure in the 

intact steam generators during this period is low relative to the setpoint of the main steam line 

PORVs such that no venting of the generators occurs and hence no heat removal is accomplished 

by them. Coinciding flow out the tube rupture is governed by the head-versus-flow 

characteristics of the HHSI pumps, the flow resistance at the tube rupture, and by either the set 

point of the PORV on the faulted steam generator (i.e., until the PORV fails open) or by the 

relief capacity of the PORV (i.e., once it fails open). The RCS heats to saturation over the course 

of a few hours following the RWST being exhausted and an extended boil-off of the RCS 

inventory begins. Note that the RWST empties ~2.25 hr earlier in this scenario (i.e., failed 

operator action and SORV) than in the previous scenario (i.e., failed operator action). This is 

simply because more water is pumped by the HHSI system through the RCS, out the SGTR, and 

out the relief valve given a stuck-open valve than given a cycling valve. The RCPs are stopped 

in the MELCOR calculation at the first occurrence of void in the RCS simulating the pumps 

tripping on their own or the operators shutting them down on account of erratic performance. 

Unlike in the calculation of the previous section, the intact steam generators pressurize up the 

setpoint of the main steam line PORVs very late in the calculation close to the time when the 

core starts uncovering. Consequently, the generators do not serve in any meaningful heat 

removal as the accident progresses to core damage. The core uncovers with nominal level in the 

steam generators and most of the original ECST inventory unused.
5-107

The MELCOR calculation predicts that core uncovering initiates at 22 hr and 48 min. The first 

release of fission products from a fuel/cladding gap is seen at 26 hr and 44 min. Traces in 

Figure 5-102 through Figure 5-107 end at first gap release and figures of fission product tracking 

are not presented for the subject scenario on account of the great unlikelihood that operators 

would fail to depressurize and cool down the reactor system for some 22 hours.

Figure 5-102 SGTR with Failed Operator Action and SORV – System Pressures

Figure 5-103 SGTR with Failed Operator Action and SORV – RPV Water Level

 

0

2

4

6

8

10

12

14

16

18

0 4 8 12 16 20 24 28

Time (hr)

Pressure (MPa)

SG A (faulted SG)

SG B

SG C

Pressurizer

RWST exhausted - HHSI ends

Faulted SG PORV fails 

Faulted SG dry

1st gap release

 

0

2

4

6

8

10

12

14

0 4 8 12 16 20 24 28

Time (hr)

Level (m)

Vessel top

Vessel bottom

BAF

TAF

RCP stop

RWST empty 1st gap release
5-108

Figure 5-104 SGTR with Failed Operator Action and SORV – Max Clad and Lower Head 

Temperature

Figure 5-105 SGTR with Failed Operator Action with SORV – RWST and ECST 

Inventories

 

300

500

700

900

1100

1300

1500

0 4 8 12 16 20 24 28

Time (hr)

Temperature (K)

Cladding

Lower head outer

Lower head inner

Saturation

RWST exhausted

Core uncovery begins

1st gap release

 

0

50000

100000

150000

200000

250000

300000

350000

400000

0 4 8 12 16 20 24 28

Time (hr)

Inventory (gal)

RWST

ECST

Useable inventory exhausted 1st gap release
5-109

Figure 5-106 SGTR with Failed Operator Action and SORV- Containment Pressure

Figure 5-107 SGTR with Failed Operator Action and SORV- Steam Generator Level
5-110

5.4.3.2 Radionuclide Release

The radionuclide release analysis is not presented for this scenario due to the low likelihood that 

operators would fail to depressurize and cool down the reactor system for the 22 hr necessary for 

the reactor core to begin to uncover (i.e., predicted by MELCOR). Table-top exercises 

performed with operators during site visits substantiate the low likelihood that operators would 

fail to depressurize and cool down the reactor system especially given the ample time for TSC 

and EOF intervention.

5.5 Interfacing System Loss of Coolant Accident

The ISLOCA scenario initiates with a common mode failure of both inboard isolation check 

valves in one leg of safety injection piping. The LHSI piping in the Safeguards outside of 

containment pressurizes to failure initiating a loss-of-coolant accident with containment bypass.

5.5.1 Unmitigated ISLOCA

The thermal hydraulic responses resulting for the unmitigated ISLOCA sequence are presented 

initially followed by the associated radionuclide migration. Table 5-14 summarizes the timing of 

the key events in the unmitigated ISLOCA.
5-111

Table 5-14 Sequence of Events for the Unmitigated ISLOCA

Event Description Time

(hh:mm:ss)

LHSI check valves fail 00:00:00

LHSI piping ruptures in Safeguards Area (outside Containment) 00:00:00+

Safeguards Area personnel door opens 00:00:16

SCRAM 00:00:22

ECCS initiates 00:00:26

Safeguards Area filtered exhaust ventilation system starts 00:00:26

Safeguards Area roof flashing tears 00:00:36

LHSI isolation valve MOV 1890C motor floods (valve inoperable) 00:02:41

RCP trip 00:03:11

MSVH/Aux. Bldg. pipe tunnel opens (penetration sealant dislodges) 00:04:13

Operators stop LHSI Pump A 00:06:17

Operators secure 1 of 3 HHSI pumps 00:15:00

Operators stop LHSI Pump B 00:15:44

Operators isolate LHSI pump suctions

(RWST spillage to Safeguards Area ends) 00:16:18

Accumulators begin discharging 00:28:27

Switchover to hot leg injection 00:45:00

Operators begin cooldown 01:00:00

Accumulators exhausted 01:12:00

Operators secure 2 of 3 HHSI pumps 01:45:00

RWST exhausted, HHSI ends 06:12:00

Water level at TAF 10:15:00

First fuel rod gap release 12:49:00

First hydrogen burn 13:29:00

Release of 1% of core inventory of iodine to environment 13:39:00

Safeguards roof fails grossly (from hydrogen burn) 13:54:00

Reactor lower head fails 18:34:00

The operator actions identified in this table are critical to delaying the onset of core damage, 

those actions being:

1. Stopping LHSI Pump A at 6 min and 17 sec

2. Stopping LHSI Pump B at 15 min and 44 sec

3. Isolating the LHSI pump suctions from the RWST at 16 min and 18 sec

4. Stopping two of the three total HHSI pumps to conserve RWST inventory

With these actions accomplished, MELCOR predicts the onset of core damage at 12 hr and 49

min. Without these actions accomplished, MELCOR predicts the onset of core damage much 

earlier.
5-112

The thermal hydraulic responses and releases of radionuclides predicted by MELCOR in the 

unmitigated ISLOCA calculation and related separate-effects (deposition) calculation are 

presented below.

5.5.1.1 Thermal Hydraulic Response

The thermal hydraulic response in the MELCOR ISLOCA calculation is illustrated in

Figure 5-108 through Figure 5-130. Presentation of these figures is grouped by RCS, 

Containment, Safeguards, and Safeguards ventilation response.

RCS Response:

Figure 5-108 shows the pressure response of the RCS and steam generators to the ISLOCA. 

Figure 5-109 shows the flow through the break expressed as an equivalent volumetric flow of 

cold water. The pressure response to the 2.57” diameter break is dramatic. RCS inventory loss 

is severe to the point where the primary and secondary systems decouple. Timing comparisons 

with Figure 5-113 and Figure 5-114 indicate that pressure approaches atmospheric once steam 

and hydrogen production in the reactor vessel subside and especially after the reactor vessel 

lower head fails.

Figure 5-110 shows ECCS flow rates. LHSI flows are to the Safeguards Area through the 

ISLOCA pipe break. The shutoffs of LHSI Pump A at 6 min and 17 sec and of LHSI Pump B at 

15 min and 44 sec are evident in the combined LHSI trace. HHSI is to the RCS cold legs 

initially. RWST drainage to the Safeguards Area ends with the operator action at 16 min and 18 

sec of isolating the LHSI pump inlets. The shutoff of one HHSI pump at 15 min is evident in the 

combined HHSI trace as is the manual transition from cold leg injection to hot leg injection at 45 

min. The step drop at 1 hr and 45 min reflects the manual shutoff of another HHSI pump 

(leaving just one of three total HHSI pumps running). HHSI ends at 6 hr and 12 min when the 

RWST is exhausted. RWST inventory is shown in Figure 5-111. The inflections in the trace in 

this figure reflect the various LHSI and HHSI pump shutoffs.

Accumulator discharge is illustrated in Figure 5-112. Discharge initiates when RCS pressure 

drops to 600 psig. The accumulators discharge at an approximately uniform rate until exhausted.

Reactor vessel water level is shown in Figure 5-113. With the exception of some early 

fluctuations accentuated by shutting off one HHSI pump at 15 min, level is maintained well 

above TAF until HHSI ends at 6 hr and 12 min. Level begins receding at this time. Hydrogen 

production from fuel cladding oxidation is shown in Figure 5-114. Hydrogen production flattens 

in Figure 5-114 when the reactor vessel lower head fails dropping the disintegrated reactor core 

to the reactor cavity in Containment. Maximum reactor core temperature, intact fuel and fuel 

debris considered, is shown in Figure 5-115. The temperatures of the inner and outer surfaces of 

the reactor lower head (in Radial Ring 1) are shown in Figure 5-116. The sharp decrease in this 

figure beginning at 18 hr and 34 min is coincident with the lower head failing.
5-113

Figure 5-108 ISLOCA Reactor System Pressure

Figure 5-109 ISLOCA Break Flow (RCS to Safeguards)

0

500

1000

1500

2000

2500

0 3 6 9 12 15 18 21 24

Pressure (psi)

Time (hr)

Pressurizer

Steam Generator A

Steam Generator B

Steam Generator C

0

1000

2000

3000

4000

5000

6000

0 1 2 3 4 5 6 7 8

Flow Rate (gpm cold water)

Time (hr)
5-114

Figure 5-110 ISLOCA ECCS Flow

Figure 5-111 ISLOCA RWST Inventory

0

1000

2000

3000

4000

5000

6000

7000

8000

9000

0 1 2 3 4 5 6 7 8

Flow rate (gpm)

Time (hr)

LHSI leakage to Safeguards

HHSI to cold legs

HHSI to hot legs

0

200

400

600

800

1000

1200

1400

1600

-1 0 1 2 3 4 5 6 7 8 Inventory (m3)

Time (hr)
5-115

Figure 5-112 ISLOCA Integral Accumulator Flow

Figure 5-113 ISLOCA Reactor Vessel Water Level

0

5000

10000

15000

20000

25000

0 1 2 3 4 5 6 7 8

Volume (gal)

Time (hr)

Integral flow

Initial inventory

0

1

2

3

4

5

6

7

8

9

0 6 12 18 24 30 36 42 48 54

LENGTH (m)

Time (hr)

Swollen level

Hot leg centerline

TAF

BAF

Vessel bottom
5-116

Figure 5-114 ISLOCA In-Vessel Hydrogen Production

Figure 5-115 ISLOCA Maximum Core Temperature

0

100

200

300

400

500

600

700

800

0 6 12 18 24 30 36 42 48 54

Mass (kg)

Time (hr)

0

500

1000

1500

2000

2500

3000

0 6 12 18 24 30 36 42 48 54

Temperature (K)

Time (hr)
5-117

Figure 5-116 ISLOCA Reactor Vessel Lower Head Temperature

Containment Response:

Containment pressure is shown in Figure 5-117. Pressure is sub atmospheric (i.e., by design) 

until reactor lower head failure at 18 hr and 34 min. Spikes in the figure are attributable to 

hydrogen and carbon monoxide burns. Temperatures in Containment are illustrated in 

Figure 5-118. 

0

500

1000

1500

2000

2500

0 6 12 18 24 30 36 42 48 54

Temperature (K)

Time (hr)

Outer surface

Inner surface
5-118

Figure 5-117 ISLOCA Containment Pressure

Figure 5-118 ISLOCA Containment Temperature

0

5

10

15

20

25

30

35

0 6 12 18 24 30 36 42 48 54

Pressure (psia)

Time (hr)

CV 55 (Containment dome)

Atmospheric pressure

0

500

1000

1500

2000

2500

3000

0 6 12 18 24 30 36 42 48 54

Temperature (K)

Time (hr)

Reactor cavity

Containment dome
5-119

Safeguards Response:

The initial pressure response of the Safeguards buildings to the blowdown of the RCS is 

illustrated in Figure 5-119. While pressure increases are highest in the lower regions of the 

Safeguards Area, these regions are below grade and therefore reinforced. With respect to 

building boundary integrity, the pressure increase seen in the level of the Safeguards Area above 

grade (CV 860) is most important. The pressure response of this level can be seen in 

Figure 5-119 relative to the pressure necessary to fail the roof, (i.e., relative to the pressure 

estimated as necessary to tear the flashing spanning from the edge of the steel roofing to the 

containment cylinder).

Figure 5-120 illustrates the temperature (i.e., atmospheric) response of the Safeguards buildings. 

Spikes in this figure reflect hydrogen and carbon monoxide burns.

The depth of the pool formed in the Safeguards Area in the short term is shown in Figure 5-121

relative to the elevation estimated to flood the LHSI pump motors and the motor of LHSI 

isolation valve MOV 1890C. Based on this figure, it was concluded that the Safeguards flooding 

caused by the postulated ISLOCA would not threaten the LHSI pump motors but would short out 

the motor on MOV 1890C making it inoperable at 2 min and 41 sec. The “Pipe penetration” 

elevations labeled in the figure pertains to the 21” high by 5’ wide piping penetration in the 

12” thick concrete wall between the Safeguards Area and the Containment Spray Pump Area. 

This penetration governs how deep flooding would get in the Safeguards Area. Long term 

flooding in the Safeguards buildings is presented in Figure 5-122. The continual drop in level 

from 6 or 7 hr to reactor lower head failure at 18 hr and 34 min is due primarily to leakage 

through the shaker space (FL 636) between the Recirculation Spray Pump B cubicle (CV 853) 

and the Containment Spray Pump Area (CV 861). The sudden level drop at lower head failure is 

because of water in the Safeguards pool being drawn back through the LHSI piping, through the 

breach in the lower head, and into containment. The driver for this flow is the initial sub 

atmospheric pressure of containment. With the exception of this water drawn into containment 

the eventual destination of all water lost through the ISLOCA pipe break is the Auxiliary 

Building.

The burning in the Safeguards buildings of hydrogen produced by fuel cladding oxidation and 

core-concrete interaction and of carbon monoxide produced by core-concrete interaction is 

illustrated in Figure 5-123. The associated energy release is shown in Figure 5-124 with 

annotations of the power related to two sections of the trace. Noteworthy is how continual the 

energy release from burning is during much of the calculation.
5-120

Figure 5-119 ISLOCA Safeguards Pressure

Figure 5-120 ISLOCA Safeguards Temperature

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

0 5 10 15 20 25 30 35 40 45

Pressure (psig)

Time (s)

CV 852

CV 855 (Break location)

CV 857

CV 860 (Safeguards Area above grade)

CV 861

CV 862

CV 863

CV 864

Roof design pressure

250

450

650

850

1050

1250

1450

0 6 12 18 24 30 36 42 48 54

Temperature (K)

Time (hr)

CV 852 (LHSI Pump B)

CV 855 (Break location)

CV 857

CV 860 (Safeguards Area above grade)

CV 861 (Low Spray Pump Area)

CV 862 (High Spray Pump Area)

CV 863 (Low MSVH)

CV 864 (High MSVH)
5-121

Figure 5-121 ISLOCA Safeguards Level Short Term

2

4

6

8

10

12

14

16

18

20

22

0 1 2 3 4 5

Elevation (ft)

Time (min)

CV 851 (LHSI Pump A)

CV 852 (LHSI Pump B)

CV 854 (MOV 1890C)

CV 855 (Break)

CV 856

CV 861 (Spray Pump Area)

CV 863 (MSVH pit)

Aux Bldg

LHSI pump flood (19'-4")

Pipe penetration top (18'-6")

Pipe penetration bot (16'-9")

MOV 1890C flood (15'-7")

Break extents (13'-9" cntrline)

10" pipe extents

Safeguards Area floor (12'-0")

MSVH-Aux Bldg tunnel extents
5-122

Figure 5-122 ISLOCA Safeguards Level Long Term

2

4

6

8

10

12

14

16

18

20

22

0 6 12 18 24 30 36 42 48 54

Elevation (ft)

Time (hr)

CV 851 (LHSI Pump A)

CV 852 (LHSI Pump B)

CV 854 (MOV 1890C)

CV 855 (Break)

CV 856

CV 861 (Spray Pump Area)

CV 863 (MSVH pit)

Aux Bldg

LHSI pump flood (19'-4")

Pipe penetration top (18'-6")

Pipe penetration bot (16'-9")

MOV 1890C flood (15'-7")

Break extents (13'-9" cntrline)

10" pipe extents

Safeguards Area floor (12'-0")

MSVH-Aux Bldg tunnel extents
5-123

Figure 5-123 ISLOCA Integral Mass of H2 & CO Burned in Safeguards Area

Figure 5-124 ISLOCA Energy Release from H2 and CO Burns in Safeguards Area

0

500

1000

1500

2000

2500

3000

0 6 12 18 24 30 36 42 48 54

Mass (kg)

Time (hr)

Hydrogen burned

Carbon monoxide burned

0

10000

20000

30000

40000

50000

60000

0 6 12 18 24 30 36 42 48 54

Energy (MJ)

Time (hr)

1

2

3

4

1 to 2 reflects 1.9 MW

3 to 4 reflects 0.7 MW
5-124

Safeguards Ventilation System Response:

The head developed by the safety-related exhaust ventilation fans serving Safeguards is shown in 

Figure 5-125. The fans started upon generation of a safety injection signal. The head increased 

as the HEPA filters upstream of the fans loaded with particulate. The line at 21 inches of H2O 

gauge in Figure 5-125, indicates the threshold that would result in the fans tripping in the 

calculation. The threshold in the calculation mimics the actual threshold at Surry. It could be 

reached given extensive filter loading.

Figure 5-126 shows the flow drawn by the exhaust ventilation fans, the portion of the flow drawn 

by the fans from Safeguards, and the “Central” portion of the flow drawn by the fans from other 

regions of the plant. Noteworthy is that the flow drawn by the fans is largely from areas of the 

plant other than Safeguards. Flow gradually decreases as the HEPA filters load.

Particulate loading on the HEPA filters is shown in Figure 5-127. Loading from the onset of 

core damage to the time of lower head failure is largely comprised of radionuclides released 

from the core. Later loading is made up of radionuclides and concrete from core-concrete 

interactions. Figure 5-128 shows the increasing differential pressure across the filter units and 

the HEPA filters load with aerosol. The traces in this figure include clean filter losses across the 

prefilters, HEPA filters, and charcoal filters.

Figure 5-129 shows temperatures along the Safeguard exhaust ventilation ducting relative to the 

maximum continuous service temperature rating for the HEPA filters (250 oF). While the 

temperature of the flow drawn from Safeguards continually exceeds the filter temperature rating, 

the temperature in the filter inlet plenum does not. This is because of the mixing in the plenum 

of the hot Safeguards flow with the much larger cold flows (68 °F) drawn from other areas of the 

plant. 

The power associated with the decay of fission products captured in the HEPA filters is 

presented in Figure 5-130. The airflow through the filters could easily remove the heat 

generation indicated.
5-125

Figure 5-125 ISLOCA Safeguards Exhaust Ventilation Fan Head

Figure 5-126 ISLOCA Safeguards Exhaust Ventilation Flows

0

5

10

15

20

25

0 6 12 18 24 30 36 42 48 54

Head (in H2O)

Time (hr)

Fan 1

Fan 2

Fan trip

0

10000

20000

30000

40000

50000

60000

70000

0 6 12 18 24 30 36 42 48 54

Flow rate (cfm)

Time (hr)

Fan 1

Fan 2

Safeguards

Central
5-126

Figure 5-127 ISLOCA Safeguards Exhaust Ventilation Filter Aerosol Loading

Figure 5-128 ISLOCA Safeguards Exhaust Ventilation Filter Unit Pressure Differential

0

5

10

15

20

25

30

35

40

45

0 6 12 18 24 30 36 42 48 54

Particulate loading (kg)

Time (hr)

0

1

2

3

4

5

6

7

8

9

0 6 12 18 24 30 36 42 48 54

Delta pressure (in H2O)

Time (hr)

Unit 1

Unit 2
5-127

Figure 5-129 ISLOCA Safeguards Exhaust Ventilation Duct and Filter Inlet Temperature

Figure 5-130 ISLOCA Decay Power of Fission Products Captured in Safeguards Exhaust 

Ventilation HEPA Filters

0

200

400

600

800

1000

1200

1400

0 6 12 18 24 30 36 42 48 54

Temperature (oF)

Time (hr)

CV 879 (Ducting)

CV 880 (Ducting)

CV 881 (Filter inlet plenum)

HEPA Filter max continuous service (250 degF)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0 6 12 18 24 30 36 42 48 54

Power (MW)

Time (hr)
5-128

5.5.1.2 Radionuclide Release

The migration of radionuclides released from the core in the MELCOR ISLOCA calculation is 

illustrated in Figure 5-131 through Figure 5-134. 

Figure 5-131 shows the fractional release to the environment of the original core inventory of the 

different radionuclide classes. The static conditions suggested in this figure after 18 hr are 

deceiving for some of the radionuclide classes because of the scale of the drawing and because of 

the time dependence of revaporization of radionuclide deposits in the LHSI piping not being 

captured in the overall ISLOCA calculation. The time dependence of this revaporization can 

only be seen in the separate-effects calculation. The results of the separate-effects calculation are 

presented in Section 5.5.2. Note that although the time dependence of revaporization of 

radionuclide deposits in the LHSI piping is missing from the overall ISLOCA results, the 

amounts of materials revaporized are accounted for. The migrations of the materials to the 

Safeguards Area do, however, come earlier in the overall calculation than they should.

The distribution of cesium throughout the MELCOR model is presented in Figure 5-132. The 

step change seen at 16 hr and 15 min is coincident with the reactor vessel lower plenum drying 

out. When a pool within a MELCOR control volume evaporates entirely, the hydrosols 

suspended within it are introduced as aerosols to the atmosphere in the same control volume. 

This happened in the control volume representing the lower plenum at the subject time and at 

least some of the aerosols introduced were swept into the LHSI piping where they promptly 

deposited by means of turbulent deposition and/or impaction. 

The distribution of iodine throughout the MELCOR model is presented in Figure 5-133. The 

sudden increase in iodine in containment and the sudden decrease in iodine in Safeguards, 

coincident with reactor lower head failure at 18 hr and 34 min, is due to water being drawn into 

containment from Safeguards. The driver for this was the initial sub atmospheric pressure in 

containment. The gradual reduction in the RCS portion of the distribution between 25 hr and 

31 hr is cesium iodide vapor evolving from aerosol deposits in the RCS. The vapor later 

condensed and the resulting aerosol was drawn into the HEPA filters of the Safeguards exhaust 

ventilation system. 

The releases of cesium iodide to the environment through the various release pathways 

represented in the MELCOR model are plotted in Figure 5-134. Each trace in Figure 5-134

shows the release through a particular pathway as a percentage of the current (i.e., instantaneous) 

total release through all pathways. Since release rates through the various pathways peak at 

different times, some of the traces in the figure do not increase continually. Pathways through 

which less than 0.5% of the total release ultimately passed have been excluded.

Of interest in the MELCOR calculation were the amounts of radionuclides scrubbed by the pool 

formed on the Safeguards floor as gas laden with radionuclides bubbled up from the ISLOCA 

pipe break through the pool. Amounts were trivial compared to amounts deposited in the LHSI 

piping or captured in the HEPA filters (e.g., 3.31 kg (pool) versus 550 kg (piping) and 42.45 kg

(HEPA filters), respectively).
5-129

Figure 5-131 ISLOCA Fission Product Release to the Environment

Figure 5-132 ISLOCA Cesium Distribution

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 6 12 18 24 30 36 42 48 54

Fractional release

Time (hr)

NG

I

Cs

Ba

Te

Ru

Mo

Ce

La

U

Cd

Sn

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 6 12 18 24 30 36 42 48 54

Fraction of inventory

Time (hr)

COR release

RCS

LHSI piping

Safeguards

HEPA filters

Aux Bldg

Containment

Environment
5-130

Figure 5-133 ISLOCA Iodine Distribution

Figure 5-134 ISLOCA CsI Releases by Available Pathways (% of Instantaneous Total)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 6 12 18 24 30 36 42 48 54

Fraction of inventory

Time (hr)

COR Release

RCS

LHSI piping

Safeguards

HEPA filters

Aux Bldg

Containment

Environment

0

10

20

30

40

50

60

70

0 6 12 18 24 30 36 42 48 54

Release (% of total CsI release to environment)

Time (hr)

FL 612 (Safeguards roof flashing)

FL 613 (Safeguards gross roof failure)

FL 626 (Spray Pump House abandoned exh fan)

FL 631 (MSVH exh fan)

FL 885 (Safeguards ventilation intake)

Pathwayswith CsI release < 0.5% not displayed
5-131

5.5.2 ISLOCA Separate-effects Calculation

The thermal hydraulic responses resulting from the ISLOCA separate-effects calculation are 

presented initially followed by the associated radionuclide migration.

5.5.2.1 Thermal Hydraulic Response

The thermal hydraulic response in the MELCOR ISLOCA separate-effects (i.e., deposition) 

calculation is illustrated in Figure 5-135 through Figure 5-138. In considering these figures, note 

that the separate-effects calculation was started 12 hr into the ISLOCA scenario somewhat 

before (i.e., 49 minutes before) the onset of core damage in the initial MELCOR overall 

ISLOCA calculation. This was sufficiently early for pipe wall temperatures to begin tracking the 

temperature of the vapor flowing through the LHSI piping before the first release of fission 

products from the core. The separate-effects calculation was carried out to 48 hr consistent with 

the duration of the initial and final overall ISLOCA calculations.

Figure 5-135 shows the pressure response along the LHSI piping. Pressure is higher upstream of 

the cavitating flow venturi from 12 to 17 hr suggesting choked flow conditions in the venturi.

Figure 5-136 and Figure 5-137 show vapor temperature and pipe wall inside surface temperature 

along the LHSI piping, respectively. Pipe wall temperatures exceed vapor temperatures due to 

the pipe walls absorbing gamma radiation emitted by fission product deposits accumulated in the 

piping. The LHSI piping in the Safeguards Area (i.e., the CV 301 trace in Figure 5-137), 

remains relatively cold for a number of hours because it is submerged in the water pool formed 

in the Safeguards Area. The eventual heat-up of the piping is in response to the piping 

uncovering as the pool recedes. The marked temperature decreases seen in these figures at 36 hr 

result when the low point in the cold leg piping dries out, i.e., when the loop seal clears. 

Clearing of the loop seal promotes cooling full-loop convective flow through the steam 

generator.

Vapor velocity along the LHSI piping is shown in Figure 5-138. Negative velocities exist for a 

period following reactor vessel lower head failure. This is because of the normally subatmospheric state of the Surry containments. Flow is from Safeguards to containment during this 

period of negative velocities. Velocity is greatest at the cavitating venturi as this is the minimum 

flow area in the LHSI piping. Velocities are relatively low in the 6” piping upstream of the 

venturi and relatively high in the 6” piping downstream of it. This is because of the higher 

pressures and hence higher densities upstream of the venturi. Velocity through the ISLOCA pipe 

break is lower than velocities in the 6” piping downstream of the venturi because the break in the 

10” piping is larger than the 6” piping. 
5-132

Figure 5-135 Pressure Along LHSI Piping in ISLOCA Deposition Calculation

Figure 5-136 Vapor Temperature Along LHSI Piping in ISLOCA Deposition Calculation

0

50

100

150

200

250

300

350

400

6 12 18 24 30 36 42 48 54

Pressure (psi)

Time (hr)

CV 101

CV 105

CV 201

CV 202

CV 301

300

400

500

600

700

800

900

1000

1100

1200

1300

6 12 18 24 30 36 42 48 54

Temperature (K)

Time (hr)

CV 101

CV 102

CV 103

CV 104

CV 105

CV 201

CV 202

CV 301
5-133

Figure 5-137 Pipe Wall Inside Surface Temperature Along LHSI Piping in ISLOCA 

Deposition Calculation

Figure 5-138 Vapor Velocity Along LHSI Piping in ISLOCA Deposition Calculation

300

400

500

600

700

800

900

1000

1100

1200

1300

6 12 18 24 30 36 42 48 54

Temperature (K)

Time (hr)

CV 101

CV 102

CV 103

CV 104

CV 105

CV 201

CV 202

CV 301

-200

-100

0

100

200

300

400

500

600

700

6 12 18 24 30 36 42 48 54

Velocity (m/s)

Time (hr)

CV 101

CV 105

FL 15 (Venturi)

CV 201

CV 202

CV 301

FL 3 (Break)

Largest LACE velocity
5-134

5.5.2.2 Radionuclide Response

The capture of aerosolized fission products by LHSI piping in the MELCOR ISLOCA separate

effects (deposition) calculation is illustrated in Figure 5-139 through Figure 5-144. These figures 

address only cesium iodide, cesium molybdate, barium, and tellurium, but illustrate well the 

competing influences of deposition (i.e., via turbulent deposition and impaction) and 

vaporization of radionuclide deposits.

Figure 5-139 identifies that most all of the cesium molybdate introduced to the LHSI piping 

deposited in the piping and remained there. At the same time, the figure identifies that more than 

half of the cesium iodide that deposited in the piping subsequently revaporized and left the 

piping. Figure 5-140 identifies analogous information for the barium and tellurium radionuclide 

classes. Note that the results for the barium class are for a metallic as opposed to an oxidic form. 

The step changes in these plots at 16 hr and 48 min are the result of the water pool in the reactor 

lower plenum disappearing. When an evaporating pool in a MELCOR control volume 

disappears, the hydrosols it suspended are introduced to the atmosphere of the control volume as 

aerosols. When the pool in the lower plenum disappeared in the initial overall ISLOCA 

MELCOR calculation, such introduction of aerosols occurred, and some of the aerosols were 

swept into the LHSI piping.

Figure 5-141 shows, for cesium iodide, the capture of aerosol along the LHSI piping through 

reactor lower head failure at 18 hr 34 min. Standing out in this figure is the relatively large 

amount of cesium iodide retained in the section of piping in the Safeguards Area. The large 

retention is attributable to the early relatively cold temperature of this piping associated with its 

being under water. Figure 5-142 shows the capture of cesium iodide for the entire duration of 

the anlaysis. Figure 5-143 shows information analogous to Figure 5-142 but for cesium 

molybdate. Clear in this figure is that cesium molybdate deposits in the LHSI piping are not 

susceptible to vaporization driven by fission product decay heating.

Illustrated in Figure 5-144 is the volume of radionuclide deposits relative to physical piping 

volume along the LHSI piping. MELCOR does not formulate a material volume associated with 

radionuclide deposits, so, Figure 5-144 presents volume estimates based on an average 

mass-weighted density determined from the masses of all radionuclide classes captured in the 

LHSI piping as a whole. The resulting average density was 6,043 kg/m3

. Additionally, 

Figure 5-144 is based on a packing factor of 0.5. Noteworthy is that radionuclide deposits 

exceed physical pipe volume in the section of piping closest to the RCS cold leg. This section of 

6” Schedule 160 piping is 18’-1” long. The overfilled condition in the MELCOR calculation is 

not perceived as an indication that the piping would become blocked to fluid flow, as the 

radionuclide deposits of this nature are not envisioned to have the structural rigidity necessary to 

cause such. Instead, the perception is that such deposits would break loose and carry further 

down the piping towards Safeguards in a form not readily aerosolized (i.e., the assumption has 

been made that such relocated radionuclide deposits would not release to the environment in 

dispersible form). To emphasize in context, MELCOR does not address aerosol resuspension. 

Aerosols deposited by turbulent deposition, impaction, or by other means in a MELCOR 

calculation, remain deposited indefinitely unless evolved by revaporization. Additionally 

noteworthy with respect to Figure 5-144 is that the capture of concrete dust evolved from core 

concrete interaction in containment is not accounted for in the separate-effects calculation. 
5-135

Consideration of concrete dust in the final overall MELCOR ISLOCA calculation suggests a 

0.0062-m3 capture of this material in the LHSI piping.

In considering the implications of volumetrically significant deposits of aerosols, it may be 

important to realize that feedback to the MELCOR thermal hydraulic solution from aerosol 

deposits is limited to energy addition from fission product decay. Energy associated with 

gamma radiation is deposited locally in the heat structure upon which the fission product 

aerosols reside and in the liquid water in the control volume to which the heat structure is 

coupled. The energy associated with beta particles is deposited in the heat structure and the fluid 

in the control volume. There is no other feedback to the thermal hydraulic solution (e.g., a 

reduction in available fluid flow area due to the buildup of aerosol deposits that is not taken ino 

account).

Table 5-15 presents the results of the MELCOR ISLOCA separate-effects calculation. The DFs 

identified in this table were placed on the flow path representing the LHSI piping in the final 

overall MELCOR ISLOCA calculation. Additionally, a DF of 10 was placed on the flow path 

associated with concrete dust. Concrete aerosol was not treated in the separate-effects 

calculation but was assumed fairly represented by the uranium, which had a resultant DF of 11.

Figure 5-139 LHSI Piping CsI and Cs2MoO4 Capture in ISLOCA Deposition Calculation

0

50

100

150

200

250

6 12 18 24 30 36 42 48 54

Mass (kg)

Time (hr)

CsI captured

CsI sourced

Cs2MoO4 captured

Cs2MoO4 sourced
5-136

Figure 5-140 LHSI Piping Ba and Te Capture in ISLOCA Deposition Calculation

Figure 5-141 CsI Capture along LHSI Piping in ISLOCA Deposition Calculation (Early)

0

5

10

15

20

25

30

35

6 12 18 24 30 36 42 48 54

Mass (kg)

Time (hr)

Ba captured

Ba sourced

Te captured

Te sourced

(Initial core CsI inventory 

of 34.821 kg)
5-137

Figure 5-142 CsI Capture along LHSI Piping in ISLOCA Deposition Calculation

Figure 5-143 Cs2MoO4 Capture along LHSI Piping in ISLOCA Deposition Calculation

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0

5

10

15

20

25

30

6 12 18 24 30 36 42 48 54

Fraction of initial core inventory

Mass (kg)

Time (hr)

CV 101

CV 102

CV 103

CV 104

CV 105

CV 201

CV 202

CV 301

Sum

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0

50

100

150

200

250

300

6 12 18 24 30 36 42 48 54

Fraction of initial core inventory

Mass (kg)

Time (hr)

CV 101

CV 102

CV 103

CV 104

CV 105

CV 201

CV 202

CV 301

Sum
5-138

Figure 5-144 Captured Particulate Volume versus Pipe Volume in ISLOCA 

Deposition Calculation

0

0.02

0.04

0.06

0.08

0.1

0.12

6 12 18 24 30 36 42 48 54

Volume (m3)

Time (hr)

CV 101 deposited vol

CV 102 deposited vol

CV 103 deposited vol

CV 104 deposited vol

CV 105 deposited vol

CV 201 deposited vol

CV 202 deposited vol

CV 301 deposited vol

CV 101 & 103 vol

CV 102 & 104 vol

CV 105 & 201 vol

CV 202 vol

CV 301 vol (0.6 m3)

(Assuming 6,043 kg/m3 and 0.5 Packing 

Factor, Excluding 0.0062 m3 of Concrete)
5-139

Table 5-15 LHSI Piping Aerosol Capture in the ISLOCA Separate-effects Calculation 

and Associated DFs

Radionuclide 

class

Aerosol mass 

entering LHSI 

piping (kg)

Aerosol and 

vapor mass 

exiting LHSI 

piping (kg)

Fractional 

capture

Associated 

DF

Initial core 

inventory 

(kg)

Fraction 

of initial 

core 

inventory 

retained 

in LHSI 

piping

Cesium1 6.2 0.1 0.98 50 11.7 0.52

Barium2 1.6 0.1 0.96 24 187.6 0.01

Tellurium 29.6 13.6 0.54 2.2 40.9 0.39

Rubidium 2.7 0.2 0.92 13 309.5 0.01

Molybdenum 10.8 1.3 0.88 8.4 243.3 0.04

Cerium 0.0 0.0 0.60 2.5 1226.0 0.00

Lanthanum 0.0 0.0 0.50 2.0 621.2 0.00

Uranium 76.8 6.9 0.91 11 66770.0 0.00

Cadmium 3.9 0.3 0.94 15 7.3 0.50

Tin 4.0 0.3 0.93 14 9.2 0.40

Cesium iodide 23.7 15.8 0.33 1.5 34.8 0.23

Cesium 

molybdate 195.9 8.0 0.96 24 302.5 0.62

Control rod 

silver 75.5 2.9 0.96 26 102.9 0.71

Control rod 

indium 13.5 0.5 0.96 25 18.1 0.72

Control rod 

cadmium 6.8 6.8 0.00 1.0 6.2 0.00

Structural tin3 91.8 8.3 0.91 11 177.4 0.47

1 Initial fuel-cladding gap inventory of cesium assumed by MELCOR 1.8.6 default to have the vapor 

pressure attributes of cesium molybdate. 2 In deriving the DF for the barium class, it was assumed that the form of the material would be oxidic as 

opposed to metallic (i.e., it was assumed that deposits would not be susceptible to revaporization). 3 The aerosol accounted for here is the tin that would be released by the melting of zircaloy fuel cladding. 

The zircaloy was assumed 1.45% tin by mass.

5.5.3 Mitigation of the ISLOCA

The unmitigated ISLOCA scenario analyzed in SOARCA is a catastrophic failure of both of the 

inboard isolation check valve disks within the LHSI piping together with failure to refill the 

RWST or to cross-connect to the unaffected unit’s RWST. Detailed analysis of this scenario 

using MELCOR provides insight into conservatisms in the PRA models and leads to 

identification of additional mitigation measures that are practical. 
5-140

The MELCOR analysis of the unmitigated ISLOCA scenario predicts that that the RWST 

becomes empty at 6 hrs and the fission product release begins at 13 hours. Assuming failure to 

refill or cross-connect RWSTs for 13 hours is a significant conservatism.

The unmitigated ISLOCA scenario analyzed in SOARCA includes operator actions to stop and 

isolate both LHSI pumps and to stop two of the three HHSI pumps to preserve RWST inventory. 

During a Surry site visit on October 26, 2011, NRC staff learned that the operators would stop 

the second HHSI pump earlier (i.e., at 1 hr instead of at 1.75 hrs) and throttle HHSI flow starting 

between 1.5 and 2 hrs. Based on these more realistic assumptions, our additional MELCOR 

analysis shows that the RWST becomes empty at 30 hours and 40 minutes. 

Based on the MELCOR analysis of the unmitigated ISLOCA scenario and a review of 

emergency procedures, it was identified that there was an additional procedure, which would 

establish core cooling and prevent core damage. It was concluded that core damage could be 

averted by starting the RHR system before the RWST becomes empty. By 2 hours and 40 

minutes, RHR entry conditions are satisfied. Thus, RHR initiation can establish adequate core 

cooling without the need for RCS inventory loss. This is discussed further in Section 5.5.3.1 

and Section 5.5.3.2.

For the unmitigated ISLOCA scenario, MELCOR sensitivity analysis showed that opening 

pressurizer PORVs and any other RCS-to-containment valves diverts some of the fission product 

release into the containment where it deposits on surfaces. This sensitivity analysis is described 

in Section 5.5.3.3.

Finally, a discussion with the utility provided further insights into expected operator actions 

associated with the HHSI pumps during an ISLOCA. To preserve RWST inventory, operators 

would secure (i.e., stop) two HHSI pumps and throttle the remaining running pump as necessary 

to maintain water level in the reactor. This is further discussed in Section 5.5.3.4.

5.5.3.1 Review of Mitigation Measures for the ISLOCA

The expected operator actions for the ISLOCA were confirmed over several visits, numerous 

conference calls, review of the applicable procedures, a full-scope simulator run for key event 

timings, and some independent analysis (i.e., see Section 3.4). The initial review of operator 

actions was performed in 2007. Then in 2011, the latest version of the procedures were 

obtained. Surry also performed a full-scope simulator exercise in 2011 using an operations crew 

to obtain the timings of some of the initial events.

The expected operator actions and timing obtained from the activities cited above allowed 

specification of the events for the unmitigated scenario previously described in Section 5.5.1. 

The unmitigated scenario arises from failure of the operators to perform the following actions:

(a) failure to successfully establish RHR, (b) refill the RWST, and (c) cross connect to the 

unaffected unit’s RWST. The purpose of this section is to review the specific actions cited in the 

emergency procedures that would lead to mitigation of the ISLOCA without the assumed 

operator failures.
5-141

Upon identification of the reactor trip and ECCS actuation, the operators quickly transition to 

Emergency Procedure E-0, “Reactor Trip or Safety Injection.” Both a reactor trip and safety 

injection initiate normally in the ISLOCA. The procedure methodically goes through the 

annunciator alarms and plant signals to identify the confirm operation of safety systems and then 

identify the cause of the transient. At a specific step, a LOCA into the containment is ruled out 

because of normal containment radiation, pressure, and sump conditions. In a following step, 

abnormal conditions would be identified in the control areas outside containment (i.e., sump 

annunciator signals in the Safeguards Building). The Emergency Procedure E-0 directs the 

operators to Emergency Procedure ECA-1.2, “LOCA Outside of the Containment.” 

In Emergency Procedure ECA-1.2, the critical valve alignments are confirmed. Next, the 

operator attempts to identify the location of the leak. First, it would be determined that the RCS 

pressure could not be stabilized through the identified valve isolation actions. Next, 

confirmation of sump annunciator signals in the safeguards and auxiliary building leads to the 

following actions: (a) local inspection of the piping in the auxiliary and safeguards buildings, (b) 

notification to the TSC of water in those locations, and (c) directions to proceed to Emergency 

Procedure ECA-1.1, “Loss of Emergency Coolant Recirculation.”

Emergency Procedure ECA-1.1 includes the critical steps for mitigation of the ISLOCA. The 

purposes of the procedure are defined as follows:

“To provide guidance to restore emergency coolant recirculation capability, to delay 

RWST depletion by adding makeup and reducing outflow, and to depressurize the RCS 

to minimize break flow.”

The highlights of the key steps in ECA-1.1 that would lead to modeling actions in the MELCOR 

simulation include:

• Start make-up of the RWST

• Initiate RCS cooldown at no more than 100 °F/hr

• Raise RCS make-up flow to maintain RVLIS at >63% (i.e., top of the active fuel)

• Depressurize RCS using one pressurizer PORV. If RCS subcooling is <30 o

F, raise RCS 

make-up flow to restore subcooling

• Check if RHR system can be placed in service

• If RWST < 3%, use RWST crosstie, establish charging pump crosstie as needed

• Depressurize all intact SGs to atmospheric

• Use RHR system if in service
5-142

Consequently, as evidenced above, the three operator failures assumed in the unmitigated 

ISLOCA are identified as procedural steps in ECA-1.1 (i.e., failure to refill the RWST, failure to 

establish a RWST crosstie to the other unit, and RHR operation, respectively).

5.5.3.2 Analysis of RHR Operation

Mitigation of the ISLOCA would be achieved through actuation of the RHR system. The RHR 

system removes water out of the Loop A hot leg and returns it into the Loop B and C cold legs 

after cooling it in a heat exchanger. The RHR system is a relatively high flow system (i.e., 

~3700 gpm/pump versus 550 gpm/pump HHSI at runout conditions), which quickly subcools the 

core and leads to complete depressurization of the primary system. The entry conditions for the 

RHR include a RCS pressure below 450 psig and Tave below 350 oF. The water level must be at 

least mid-height in the hot leg to prevent vortexing vapor into the pump. Following Surry

procedures and confirmed by the EOF and TSC, the operators would attempt to establish RHR as 

soon as possible to mitigate the accident.

Figure 5-145 shows the RCS pressure and Tave with the RHR pressure and temperature entry 

criteria. After 2 hours and 40 minutes, the RHR pressure and temperature entry criteria are 

established. The vessel swollen level is shown in Figure 5-146. The swollen vessel level is 

above the hot leg centerline after 2 hours and 40 minutes. Hence, after 2 hours and 40 minutes, 

the operators could establish RHR based on the entry criteria.

There are several challenges to be considered. First, water will need to be added to the RCS until 

the break flow is terminated. The RHR system will fully depressurize the RCS, which 

terminates the break flow. The break flow from the unmitigated ISLOCA is shown in

Figure 5-147. At 2 hours and 40 minutes, the break flow is less than 40 kg/s and steady due to 

the successful operator actions to reduce HHSI injection to only one pump. There is sufficient 

inventory in the RWST to maintain HHSI injection for more than 3 hours (see Figure 5-148). 

Previous experience in modeling RHR operation for a similar, mitigated ISLOCA shows that the 

RCS will depressurize to atmospheric conditions in about 30 minutes after the start of RHR. 

Since the LHSI piping with the ISLOCA connects to the top of the cold leg piping and 

immediately rises vertically 4 ft upon leaving the cold leg,34 the break flow will terminate once 

the RCS pressure drops below the head of the piping rise (~1.7 psig). 

A second challenge is maintenance of the HHSI flow while the RHR is depressurizing the RCS. 

The HHSI pumps are vulnerable to flooding due to drainage of the effluent from the ISLOCA 

into the basement of the Auxiliary Building. Surry has done a flooding analysis of the Auxiliary 

Building and concluded the HHSI pumps will begin to flood when the water volume exceeds 

530,000 gallons in the basement. As shown in Figure 5-149, the water volume in the basement 

of the auxiliary building is ~235,000 gal at 2 hours and 40 minutes. Therefore, there is 

considerable margin before the HHSI pumps are flooded. As illustrated in the unmitigated 

ISLOCA Auxiliary Building water volume response (see Figure 5-149), there is insufficient 

water to flood the HHSI pumps. Consequently, they have some margin to use additional water 

from a RWST crosstie (i.e., ECA-1.1) or by refilling the RWST (i.e., ECA-1.1).

 34 Normal, forward flow for the LHSI piping is into the cold leg. Leaving the top of the cold leg in this context is 

the direction of the ISLOCA break flow, which is backwards through the LHSI piping.
5-143

The final challenge is maintenance of available Net Positive Suction Head (NPSH) of the RHR 

pumps. The NPSH is the pressure head difference between the actual pressure of a liquid in a 

pipeline and the liquid's vapor pressure at the liquid temperature. If the liquid pressure drops 

below the vapor pressure, then vapor bubbles will form. The vapor bubbles cause cavitation and 

reduce the efficiency of the pumps. The collapse of the cavitation bubble in the pump impeller 

creates a pressure wave that can destroy the internal pump components (usually the leading edge 

of the impeller). Additionally, the inevitable increase in vibration can cause other mechanical 

faults in the pump and associated equipment [67].

There are two aspects of the NPSH: the available NPSH and the required NPSH. The available 

NPSH is the suction pressure at the pump inlet. The available NPSH is defined as follows:

Available NPSH = p0 − pv

ρg

+ ∆z − hL

where:

hL is the head loss from the hot leg to the pump inlet, [m]

p0 is the pressure in the hot leg, [Pa]

pv(Tliq) is the saturation pressure at the temperature of the liquid in the piping, [Pa]

Δz is the difference in height, [m]

ρ is density of the liquid, [kg/m3]

g is gravitational acceleration, [m/s2]
5-144

The required NPSH is experimentally determined by the pump manufacturer. It represents the 

head needed before the pump’s total differential head performance is reduced by 3% due to 

cavitation. Cavitation occurs at suction pressure levels below the required NPSH and pump 

damage can occur from cavitation even though the pump may continue to provide injection flow.

At 3 hours in the Surry ISLOCA sequence, pv(Tliq) is near saturation conditions. Setting 

Tliq = Tsat in the previous equation, the available NPSH reduces to the following expression:

Available NPSH = p0 − pv

ρg

+ ∆z − hL

Available NPSH = ∆z − hL

The difference in height between the hot leg and the RHR pump (Δz) was determined from plant 

drawings to be 25-ft. The head loss hL is a function of the flow rate, fluid state, and the losses in 

the piping. Surry provide the results from engineering RHR NPSH calculations. The RHR 

NPSH calculations that indicated there was adequate (no less than 7” and no more than 20”) 

NPSH margin (i.e., NPSHmargin = NPSHavailable – NPSHrequired) at saturated conditions with 8” of 

water in the hot leg with one RHR pump running. The Surry NPSH calculations imply that the 

RHR must be started gradually (i.e., only one of the two pumps) until additional subcooling was 

achieved.35 In addition, as indicated in ECA-1.1, additional make-up flow should be initiated to 

increase the subcooling. Since only HHSI pump was running in the results presented in 

Figure 5-145 through Figure 5-147 there are additional resources and procedural direction to 

increase the RCS water level and the subcooling, which will further reduce the potential for 

inadequate NPSH margin.

In summary, RHR could be established after 2 hours and 40 minutes to provide closed loop 

cooling of the RCS and terminate the break flow. There is adequate water in the RWST for 

3 hours to complete this action. The available NPSH is adequate to start the pumps, especially if 

the operators followed procedural steps to increase subcooling (i.e., starting additional HHSI 

pumps). Once RHR is established, the core will become subcooled. This allows the RCS 

pressure to decrease to atmospheric conditions, which will terminate the ISLOCA break flow.

 35 Following discussions with Surry operations, there is plant experience in cavitating the RHR pumps during midloop operations. There are identifiable signals in the control room of the condition. The operator would monitor 

the flow rate of the RHR system and secure one or both RHR pumps if inadequate flow is developed. The plant 

operations staff shared that there are ample control room signals to identify when two pumps could be operated. 

The plant experience demonstrates that the pumps were robust enough to tolerate cavitation during RHR startup.

0
5-145

Figure 5-145 ISLOCA RCS Pressure and Temperature Relative to the RHR Entry 

Pressure and Temperature Criteria for the Unmitigated ISLOCA

Figure 5-146 ISLOCA Vessel Swollen Water Level for the Unmitigated ISLOCA

200

300

400

500

600

700

0

500

1000

1500

2000

2500

0 1 2 3 4 5 6 7 8

Temperature (°F)

Pressure (psig)

Time (hr)

ISLOCA RCS Pressure and Temperature Relative to RHR Entry 

Conditions

RCS pressure

RHR entry pressure (450 psig)

RCS saturation press

RCS Tave

RHR entry temp (350 degF)

3

4

5

6

7

8

9

0 1 2 3 4 5 6 7 8

Elevation (m)

Time (hr)

Swollen Water Level in the Vessel (Inside Core Barrel)

(0 m = Vessel Bottom)

Vessel Level

Top of the Cold Leg

Cold leg centerline

Bottom of the Cold Leg

TAF

BAF
5-146

Figure 5-147 ISLOCA Break Flow for the Unmitigated ISLOCA

Figure 5-148 RWST Volume for the Unmitigated ISLOCA

0

50

100

150

200

250

0 1 2 3 4 5 6 7 8

Mass Flow Rate (kg/s)

Time (hr)

Break Flow in the Unmitiigated ISLOCA

0

132,100

264,200

396,300

0

500

1000

1500

0 1 2 3 4 5 6 7 8

Inventory (m

Inventory (gal)

3)

Time (hr)

RWST Inventory
5-147

Figure 5-149 Auxiliary Building Water Volume for the Unmitigated ISLOCA

5.5.3.3 Analysis of Mitigation Using Pressurizer PORVs

Mitigation of the ISLOCA also could be achieved by opening valves between the RCS and 

containment, potentially diverting flow of fission products into the containment where fission 

products could deposit and the eventual release to the environment could be delayed. A review 

of the Surry plant design showed that paths between the RCS and containment could be opened 

by opening the pressurizer PORVs to reduce RCS subcooling (i.e., ECA-1.1). This section 

demonstrates the potential effects of reducing fission product released to the safeguards area 

should conditions permit PORV actuation. Similarly, the RCS vent and drain system could 

allow venting to containment. 

Opening the pressurizer PORVs allows steam, hydrogen, and fission products to reach the 

pressurizer relief tank. Once the pressurizer relief tank pressure reaches 100 psi above 

containment pressure, the PRT rupture disks will open providing a flow path from the pressurizer 

relief tank to the containment. This section investigates the effectiveness of opening pressurizer 

PORVs in reducing the offsite release. 

As noted above, a pressure difference of 100 psi is needed to open the PRT rupture disks. 

Figure 5-108 and Figure 5-117 shows the pressure in the RCS and containment, respectively. 

Pressure in the RCS appears sufficient to open the PRT rupture disks at two different phases: 

The first couple of hours of RCS depressurization and during the in-vessel core degradation. 

Several calculations were performed to estimate the time interval, post scram, when the PORVs 

must be opened to burst the rupture disks prior to in-vessel release. The variations were 

performed at half hour intervals. The longest delay in pressurizer PORV actuation that resulted 

0

100000

200000

300000

400000

500000

600000

0 1 2 3 4 5 6 7 8

Volume (gal)

Time (hr)

Auxiliary Building Water Volume

HHSI pump flood (530,000 gal)

RWST Volume (387,066 gal)

Aux. Bldg. water volume
5-148

in PRT rupture disk opening prior to in-vessel release was 2 hrs. PORV actuation following 2.5 

hrs resulted in PRT rupture disk opening during core degradation.

Two sensitivities were performed to investigate the effectiveness of PORV actuation in 

mitigating fission product release to the environment: (1) PORV actuation at 0.5 hrs after scram,

and (2) PORV actuation at the start of the in-vessel core degradation (i.e., when water level is at 

core mid-plane). Each sensitivity is a variation of the unmitigated ISLOCA, described in

Section 5.5.1. 

Facility pressure response to the actuation of the pressurizer PORVs at 0.5 hrs is presented in 

Figure 5-150. The PRT rupture disks opened at 0.74 hrs. The pressurizer level increases in 

response to PORV actuation, as shown in Figure 5-151. Initial flow is predominantly vapor, 

Figure 5-152. Once the level reaches the base of the PORVs at 1.1 hrs, the flow transitions to 

two-phase. At 6.6 hrs the flow returns to single phase vapor for the duration of the in-vessel 

release. 

Given the rupture disks are open, the differential pressure between the RCS and the containment 

atmosphere, established during the boildown, sustains vapor flow through the PORVs during the 

in-vessel release phase of the accident. Therefore, vapor released from the RCS is partitioned 

between the safeguards area and containment as well as the transported fission products, shown 

in Figure 5-153. The flow through the PORVs is diverting a significant portion of fission 

products into containment. Using the containment volume during the in-vessel release reduced 

the cesium and iodine released to the environment, presented in Figure 5-154 and Figure 5-155, 

respectively. Flow through the pressurizer PORVs terminates due to lower head failure. 

 

The sensitivity characterizing late actuation of the pressurizer PORVs was performed coinciding 

with core degradation at 12.8 hrs. It is postulated that given PORV actuation occurs between the 

initiation of core degradation and 2.5 hrs, post scram, the result will be late PRT rupture disk 

opening. Upon PORV actuation, flow through the PORVs occurs for a short time but terminates 

prior to fission product release; subsequently, fission products released evacuate the RCS 

through the LHSI piping. In Figure 5-156, the pressure excursion at 14.5 hrs resulted from the 

failure of the flow distributor element. This failure provides a pathway for core debris to 

relocate into the pooled water below the distributor, generating an increase in steam production 

and a corresponding pressure rise. The pressure rises above the rupture disks set point which 

opens the rupture disks. This decreases flow through the LHSI line as release paths are now 

established to the containment as well as the safeguards area. CsI released in-vessel is overlaid 

with the LHSI and PORV flow rates in Figure 5-157. The resulting reductions in cesium and 

iodine released to the environment are presented in Figure 5-158 and Figure 5-159, respectively. 

From the results presented, reductions to the environmental release were achieved through 

PORV actuation. PORV actuation resulted in PRT pressurization and eventually the PRT 

rupture disks to burst. Once the rupture disks were open, steam, vapor, and fission products 

produced in the core were partially diverted to containment. Observations from the two 

sensitivities demonstrated a strong dependency between the PRT rupture disk actuation with the 

back-pressure of the RCS. Two time periods were indentified when RCS pressure was great 

enough to induce rupture disk actuation. The first period was during early depressurization and 
5-149

the second was post depressurization and prior to core degradation. Early actuation established 

continual flow from the RCS to containment during the in-vessel release, until lower head 

failure. Actuation of the PORVs after depressurization resulted in PRT rupture disks remaining 

intact for the majority of the in-vessel release. 

The percent of iodine released to the environment was reduced from 15.8% to 11.3% and 14.4% 

and the percent of cesium released to the environment was reduced from 2.03% to 1.37% and 

1.97%, for PORV actuation at 0.5 hrs after scram and at the beginning of core degradation, 

respectively. 

Figure 5-150 Facility pressure response to PORV actuation at 0.5hrs

10

100

1000

0 5 10 15 20

Pressure [psi]

time [hr]

Pressure - PORVs Open at 0.5 hr

Pressurizer

Containment

Pressurizer Relief Tank
5-150

Figure 5-151 Pressurizer liquid level response to PORV actuation at 0.5hrs

Figure 5-152 Flow Rate to Pressurizer Relief Tank – PORVs Open at 0.5 hrs

0

5

10

15

20

25

0 2 4 6 8

LENGTH [m]

time [hr]

Pressurizer Liquid Level - PORVs Open at 0.5hrs

Pressurizer Liquid Level

Pressurizer Boundaries

Porv Opening Bottom

0

5

10

15

20

25

30

35

40

0 5 10 15 20

Mass Flow Rate [kg/s]

time [hr]

Flow Rate to Pressurizer Relief Tank - PORVs Open at 0.5hrs

Pool

Vapor
5-151

Figure 5-153 Flow Rates with CsI Invessel Release – PORVs Open at 0.5 hrs

Figure 5-154 Cesium Release to the Environment – PORVs Open at 0.5 hrs

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0

0.5

1

1.5

2

2.5

3

3.5

4

4.5

5

10 11 12 13 14 15 16

Invessel Release Fraction (-)

Mass Flow Rate [kg/s]

time [hr]

Flow Rates with CsI Invessel Release - PORVs Open at 0.5 hrs

LHSI Flow

Combined PORV Flow

Invessel Release Fraction CsI

0

0.005

0.01

0.015

0.02

0.025

0.03

0.035

0.04

0.045

0.05

0 6 12 18 24 30 36 42 48

Fraction release (-)

Time (hr)

Cesium Release to the Environment

PORVs Open at 0.5 hrs

Best Estimate
5-152

Figure 5-155 Iodine Release to the Environment – PORVs Open at 0.5 hrs

Figure 5-156 Facility pressure response to PORV actuation at Core Degradation

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

0 6 12 18 24 30 36 42 48

Fraction release (-)

Time (hr)

Iodine Release to the Environment

PORVs Open at 0.5 hrs

Best Estimate

10

100

1000

0 5 10 15 20

Pressure [psi]

time [hr]

Pressure - PORVs Open at Core Degradation

Pressurizer

Containment

Pressurizer Relief Tank
5-153

Figure 5-157 Flow Rates with CsI Invessel Release – PORVs Open at Core Degradation

Figure 5-158 Cesium Release to the Environment – PORVs Open at Core Degradation

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0

0.5

1

1.5

2

2.5

3

3.5

4

4.5

5

10 11 12 13 14 15 16

Invessel Release Fraction (-)

Mass Flow Rate [kg/s]

time [hr]

Flow Rates with CsI Invessel Release - PORVs Open at Core Degradation

LHSI Flow

Combined PORV Flow

Invessel Release Fraction CsI

0

0.005

0.01

0.015

0.02

0.025

0.03

0.035

0.04

0.045

0.05

0 6 12 18 24 30 36 42 48

Fraction release (-)

Time (hr)

Cesium Release to the Environment

PORVs Open at Core Degradation

Best Estimate
5-154

Figure 5-159 Iodine Release to the Environment – PORVs Open at Core Degradation

5.5.3.4 Effect of Throttling HHSI Flow

Discussion with the utility provided further insights into expected operator actions associated 

with the HHSI pumps during an ISLOCA. To preserve RWST inventory, operators would secure 

(i.e., stop) two HHSI pumps and throttle the remaining running pump as necessary to maintain 

water level in the reactor above TAF.

A timeline provided by the utility demonstrated that one HHSI pump would be secured 15 min 

after scram and another 1 hr after scram. Throttling of the third pump would commence at 1 hr 

and 45 min. Core water level would be maintained above TAF. A quasi-steady balance would be 

achieved between HHSI flow rate and the rate of steam production in the core.

The timeline described above was instituted in a MELCOR ISLOCA sensitivity calculation. 1 hr 

and 45 min into the calculation, when throttling of the remaining running HHSI pump began, 

roughly 190,000 gallons or 49% of the original 388,000 gallons in the RWST remained. This can 

be seen in Figure 5-160. Fission product decay power at this time had reduced to 32.5 MW as 

seen in Figure 5-161. By 4 hr, the reactor had depressurized to atmospheric pressure and the 

calculation had become uneventful. Figure 5-162, Figure 5-163, and Figure 5-164 show reactor 

pressure, water level, and maximum fuel cladding temperature in the calculation, respectively. 

Since the MELCOR calculation had become uneventful and was time consuming, a side 

calculation was performed to estimate how long it would take decay power to consume the 

remaining inventory in the RWST. Consuming the inventory was taken to involve heating the 

water to saturation and transforming it to steam. The decreasing trend of decay power was 

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

0 6 12 18 24 30 36 42 48

Fraction release (-)

Time (hr)

Iodine Release to the Environment 

PORVs Open at Core Degradation

Best Estimate
5-155

accounted for. The side calculation predicted that the RWST would be exhausted 30 hr and 40 

min after the onset of the ISLOCA.

Figure 5-160 ISLOCA HHSI Throttling Sensitivity - RWST Inventory

150000

200000

250000

300000

350000

400000

0 2 4 6 8 10

Volume (gal)

Time (hr)
5-156

Figure 5-161 ISLOCA HHSI Throttling Sensitivity - Fission Product Decay Power

Figure 5-162 ISLOCA HHSI Throttling Sensitivity - RCS Pressure

0

20

40

60

80

100

120

140

160

180

200

0 10 20 30 40

Power (MW)

Time (hr)

0

2

4

6

8

10

12

14

16

0 2 4 6 8 10

Pressure (MPa)

Time (hr)
5-157

Figure 5-163 ISLOCA HHSI Throttling Sensitivity - Core Water Level

Figure 5-164 ISLOCA HHSI Throttling Sensitivity - Maximum Core Temperature

2

3

4

5

6

7

8

9

0 2 4 6 8 10

Level (m above vessel bottom)

Time (hr)

Core exit

swollen level

Core Swollen

Water Level

Hot leg

centerline

TAF

BAF

0

100

200

300

400

500

600

700

800

900

1000

0 2 4 6 8 10

Temperature (K)

Time (hr)
5-158

5.6 Other Sensitivity Studies

During the peer review of the MELCOR calculations, several other more generic issues were 

identified relative to the ones already discussed in Sections 5.2.3 and 5.3.3. They include 

uncertainties in the chemical form of iodine, iodine spiking, uncertainties of the impact of air 

ingression into the vessel, and uncertainties in the aerosol deposition rate in containment.

5.6.1 Chemical Form of Iodine

The chemical forms and quantities of gaseous iodine are an active research topic as new 

information is still being evaluated in existing and planned tests. The SOARCA calculations did 

not include gaseous iodine. All iodine was assumed to be combined with cesium to form cesium 

iodide and remain in that chemical form. New data from Phebus suggests some iodine is 

released in elemental form yet undergoes complex chemical reactions in the containment to form 

organic compounds unless liberated by the chemosorption process. MELCOR does not include a 

model for surface chemistry with paint. Furthermore, MELCOR’s ex-vessel iodine pool model 

is very slow running and not fully validated. Consequently, all the iodine was modeled as a

cesium iodide compound and the pool iodine model was not used. It should be noted that the 

uncertainty study will investigate the influence of different fixed amounts of gaseous iodine 

(i.e., elemental and organic forms). 

Relative to the current results, it is worth making some simple evaluations using recent 

interpretations of Phebus data [33]. Phebus Test FTP-1 shows that the concentration of iodine 

reaches a steady state in the containment that is independent of the pool pH and condensing or 

evaporating conditions. In particular, the prototypical Phebus configuration shows a steady state 

exchange between the painted surfaces where the iodine is absorbed and released to maintain a 

steady concentration. 

Two evaluations were performed to assess the impact of gaseous iodine on the source term using 

Phebus FTP1 data. In the first evaluation using the STSBO, a range of gaseous iodine 

concentrations were considered with the calculated containment leakage rate to estimate the 

additional iodine source term. The measured Phebus gaseous iodine containment concentrations 

are shown in Figure 5-165 with the conversion to an iodine release fractions based on the 

containment release rate. The calculated iodine release magnitude was 0.65% in the unmitigated 

STSBO at 48 hours (see Section 5.2.1). Assuming gaseous iodine concentrations of 0.05%, 

0.10%, and 0.15%, the additional source term would be less than 0.10%. Given the small 

absolute and relative magnitude of the iodine release, the impact of a 0.10% additional gaseous 

iodine release was judged as not significant.

The second evaluation examined the additional source term to the environment through the failed 

steam generator tube, which occurred earlier in the accident progression. The measured Phebus 

gaseous iodine containment concentrations are shown in Figure 5-166. The higher short-term 

values were used to estimate the additional gaseous release to the environment. Using the noble 

gas leakage rate into the environment through the steam generator secondary and the early, 
5-159

higher concentrations from Phebus containment, the gaseous iodine leak rate was calculated.

36

The calculated iodine release rate was 0.6% in the first 24 hours when the dominant releases 

through the TI-SGTR occurred (see Section 5.3.1.2). Assuming gaseous iodine concentrations of 

0.10%, 0.15%, and 0.20%, the additional source term would be <<0.10%, respectively. Given 

the small absolute and relative magnitude of the iodine release, the impact of gaseous iodine on 

the source term was also judged small. Following vessel failure, any remaining gaseous iodine 

in the reactor vessel was discharged into the containment. All further releases through the failed 

TI-SGTR were diluted by the volume of the containment.

In summary, gaseous forms of iodine have the potential to increase the severity of the 

environmental source term because they do not settle like other aerosol radionuclides. Gaseous 

iodine remains an important source term issue, especially with respect to long-term containment 

performance issues after the comparatively much larger airborne radioactivity has settled from 

the atmosphere. The mechanistic modeling treatment for gaseous iodine behavior is a 

technology still under development with important international research programs underway to 

determine the dynamic behavior of iodine chemistry with respect to paints, wetted surfaces, 

buffered and unbuffered water pools undergoing radiolysis, and gas phase chemistry. In 

SOARCA, gaseous iodine was not specifically included except as a revaporization gas from 

deposited cesium iodide that had chemisorbed onto stainless steel surfaces. The magnitude of 

the revaporization release was generally very small. A review of Phebus data suggests that 

gaseous iodine is released from the fuel and complex surface and pool reactions take place in the 

containment. Based on the analysis of the early and the long-term behavior from Phebus for two 

of the SOARCA calculations, it does not significantly change the magnitude or the timing of the 

overall iodine release. It was judged that the small additional source term of gaseous in the other 

sequences would also not change the overall conclusions of the study. Nevertheless, it should be 

noted that on-going and future NRC research is dedicated to better understanding iodine 

behavior.

Gaseous iodine remains an important source term issue, especially with respect to long-term 

containment performance issues after the comparatively much larger airborne radioactivity has 

settled from the atmosphere. The mechanistic modeling treatment for gaseous iodine behavior is 

a technology still under development with important international research programs underway 

to determine the dynamic behavior of iodine chemistry with respect to paints, wetted surfaces, 

buffered and unbuffered water pools undergoing radiolysis, and gas phase chemistry. The base 

case treatment under the best practices recommendation are sufficient for the mean effects 

addressed in SOARCA.

 36 The gaseous iodine release was estimated by examining the noble gas release from the fuel and the fraction 

subsequently transported through the failed steam generator tube. Since the release of the noble gases occurred 

roughly coincidental with the expected gaseous iodine release (i.e., both considered highly volatile), the 

subsequent accumulation and transport of the non-condensable noble gas would be a good surrogate for gaseous 

iodine. In particular, the portion transported out the failed tube versus the amount discharged in the containment 

could be assessed. As described above, the magnitude of the gaseous iodine release from the fuel was scaled 

based on the short-term estimates from Phebus FPT1.
5-160

Figure 5-165 The additional gaseous iodine source term using Phebus data is compared to 

the iodine source term for the unmitigated short-term station blackout

Figure 5-166 The additional gaseous iodine source term using Phebus data is compared to 

the iodine source term for the unmitigated short-term station blackout with a 

thermally-induced steam generator tube rupture

Long-term Iodine Leakage as a Function of the 

Containment Steady State Gaseous Iodine Concentration

0.000

0.001

0.002

0.003

0.004

0.005

0.006

0.007

0.008

0.009

0.010

0 12 24 36 48

Time (hr)

Fraction of Initial Inventory (-)

STSBO Iodine Release

0.15%

0.10%

0.05%

0.65%

0.03 – 0.1%

Estimated Iodine Gas Environmental Release Using Phebus FPT1

STSBO + 100% TI-SGTR - No Mitigation

0.000

0.001

0.002

0.003

0.004

0.005

0.006

0.007

0.008

0.009

0.010

0 6 12 18 24

Time (hr)

Fraction of Inventory (-)

0.10%

0.15%

0.20%

Base Case

Estimated 

gaseous iodine 

release from fuel 

to environment
5-161

5.6.2 Additional Source Term from Iodine Spiking

Iodine spiking was identified by one of the review committee members as a possible alternate

source of iodine to the environment for an early release in the spontaneous SGTR. Using the 

water leakage from the unmitigated SGTR, the maximum recorded iodine spike (18 μCi/g), and 

the recommended partition factor from Regulatory Guide 1.83 (i.e., 100)37, the fractional iodine 

release was 10-6 [35]. While an iodine spike might be an operational concern, it is not significant 

relative to the magnitude of release fractions from the other considered severe accidents (see 

Table 5-16).

Table 5-16 Comparison of Iodine Spike Source Term to Iodine Source Terms from the 

Other Unmitigated Accidents

Unmitigated Scenario

Core fraction of iodine 

released to environment

Long-term SBO 0.003

Short-term SBO 0.006

Short-term SBO with thermally 

induced SGTR 0.009

ISLOCA 0.158

Spontaneous SGTR

(Iodine Spike) 10-6

5.6.3 Air Ingression into the Vessel

Air ingression into the vessel was identified by one of the peer review committee members as an 

important concern for enhanced air oxidation of metals and enhanced ruthenium releases. There 

are two events in the Surry SOARCA sequences where air ingression into the reactor could 

occur.38 First, while the fuel is degrading, natural circulation into the RCS could lead to creep 

rupture of the hot leg. Since there was a large decay heat source in the reactor vessel, all cases 

showed a slight pressurization of the reactor coolant system relative to the containment (or 

auxiliary building) that maintained a steady flow outward of the pipe breaks. Consequently, 

inward flow of air during this time was not expected.

Later in the accident progression, the fuel will collapse onto the lower head and fail the reactor 

vessel. Following failure of reactor vessel, the hot contents in the lower plenum poured into the 

reactor cavity. In the progression of events calculated in the unmitigated scenarios, all injection 

 37 Although Regulatory Guide 1.83 was withdrawn on November 12, 2009 (Federal Register Volume 74, Number 

217, 58324), the use of the recommended partition factor remains appropriate for the purposes of this analysis. 38 A pipe break in the reactor coolant system for a loss-of-coolant accident (LOCA) could be another event. The 

ISLOCA was the only LOCA examined for Surry. It is incredible to imagine any significant air ingression 

upstream through 30 m of LHSI piping against sonic two-phase flow in the ISLOCA. However, if a large break 

LOCA of the RCS piping had been considered , there is a potential for air ingression.
5-162

had terminated and the entire core had degraded and collapsed prior to vessel failure. 

Consequently, all the debris relocated to the cavity prior to any significant air-ingression 

(e.g., see Figure 5-167). 

Finally, MELCOR includes models for both steam and air oxidation of metals in the core 

package. Consequently, air oxidation is considered if any air ingression occurs. However, there 

are no models for automatically changing the ruthenium release model in an air-oxidizing

environment. Consequently, each calculation must be reviewed for the presence of high air 

concentration conditions, which was done.

Figure 5-167 Comparison of the oxygen concentration and UO2 mass in the vessel during 

the unmitigated STSBO

0

0.2

0.4

0.6

0.8

1

0 4 8 12 16

Gas Fraction (-) and UO2 Fraction (-)

Time (hr)

Comparison of Oxygen Concentration versus UO2 Mass in the Vessel

Surry Unmitigated STSBO

Oxygen Concentration in the core region

Vessel UO2 Fraction

Vessel Failure
5-163

5.6.4 Aerosol Settling Rate in the Containment

A peer review committee member thought the aerosol settling rate in the containment looked 

high (e.g., Figure 5-36) for the STSBO scenario without B.5.b mitigation. To address this issue, 

two time phases were investigated. The first time phase occurred after the hot leg failure. 

Following hot leg failure, co-dispersing and flashing water from accumulator injection with the 

aerosols immediately led to a very high mass median diameter of the airborne aerosols (>10 μm), 

which caused them to settle very quickly. The condensing steam helped increase agglomeration 

enhanced deposition. Within one hour after hot leg failure, over 50% of the airborne aerosols 

had settled. 

The second phase occurred with the releases following prior to and at vessel failure. To analyze 

the settling rate, the mass of airborne aerosols in the STSBO at vessel failure were normalized to 

one (see Figure 5-168). Following vessel failure, the airborne aerosol concentration decreased 

steadily. The airborne decay constant, λ, was calculated and compared to Phebus FTP0 

data [34]. The calculated decay constant is a strong function of the mass-median diameter of the 

airborne aerosols. However, the results in Figure 5-168 show the settling rate was comparable to 

the test data and actually slightly slower.

Figure 5-168 Unmitigated STSBO airborne aerosol mass in the containment following 

vessel failure

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

6 12 18 24

Fraction Airborne at Vessel Failure (-)

Time (hr)

Surry Unmitigated STSBO Airborne Aerosol Mass in the Containment

λ = -0.7 /hr

Containment failure

λ = -1.0 

Phebus FTP0 aerosol deposition in the 

containment (with code comparisons with 

6-1

6. EMERGENCY RESPONSE

Advancements in consequence modeling now allow more detailed and more realistic treatment 

of emergency response when performing consequence analyses. This includes the ability to 

model protective action decisions from offsite response organizations (OROs) and the 

implementation of protective actions by individual population segments. To best utilize these 

advancements, detailed information was obtained from local sources and OROs. Through a user 

interface added to the consequence model, this detailed information was input to account for 

differences in the implementation of protective actions by individual population segments. 

These advancements are significant because they facilitate more realistic modeling of response

activities, timing of decisions, and implementation of protective actions across different 

population segments.

Emergency response programs for nuclear power plants (NPPs) are designed to protect public 

health and safety in the event of a radiological accident. These emergency response programs 

are developed, tested, and evaluated and are in place as defense in depth to respond in the 

unlikely event of an accident. To support a state of the art approach and integrate realism in the 

analyses, the modeling of the emergency response was based on the site-specific emergency 

planning documentation and on research of public response to non-nuclear emergencies. The 

information developed in this Emergency Response Section was used to support the MACCS2 

consequence analyses for the accident scenarios. These analyses are conducted for the 

unmitigated accident sequences only. Many of these response actions would be similar for the 

mitigated case because response officials initiate protective actions upon notification, which as 

described herein occurs very early in the incident, even before mitigation actions have been 

implemented. This is because emergency planning is designed to be proactive to remove the 

public prior to plume arrival when possible. For each accident scenario, evacuation of the plume 

exposure pathway emergency planning zone (EPZ) was assessed. This included consideration of 

a shadow evacuation to a distance of 20 miles from the plant. Including a shadow evacuation, 

which occurs when members of the public evacuate from areas that are not under official 

evacuation orders, provides realism because these are observed in large-scale evacuations [44]

and have the potential to slow down the evacuation from the affected area. Also, for each 

scenario, members of the public were modeled as being relocated from any area where doses are 

projected, based on the consequence analysis model, to exceed established criteria. Figure 6-1

identifies the location of the Surry plant and radial distances of 10 and 20 miles from the plant.

Sensitivity analyses were completed for one of the accident scenarios to evaluate evacuation 

distances of 16 miles and 20 miles from the plant. The sensitivity analysis of an evacuation to 20 

miles is different than the shadow evacuation to 20 miles described above, because the 

sensitivity analysis evaluates the conditions under which residents of the entire 20 mile area are 

notified to evacuate and leave the area. A sensitivity analysis was performed to assess the effect 

of a delay in the implementation of protective actions, as suggested by the peer review 

committee. An analysis was also conducted that included consideration of the effects on 

infrastructure, emergency response, and response of the public due to a seismic event. 
6-2

Figure 6-1 Surry 10 and 20 Mile Areas

As required by 10 CFR 50, OROs develop emergency response plans for implementation in the 

event of an NPP accident. These plans are regularly drilled and are inspected biennially, through 

a demonstration exercise performed in conjunction with the licensee. In biennial exercises, ORO 

personnel demonstrate timely decision making and the ability to implement public protective 

actions. Emergency plans escalate response activities in accordance with a classification scheme 

based on emergency action levels (EALs). Preplanned actions are implemented at each 

classification level including Unusual Event, Alert, Site Area Emergency (SAE), and General 

Emergency (GE). Public protective actions are required at the GE level, but ORO plans 

commonly include precautionary protective actions at the SAE level and sometimes at an Alert.

The plume exposure pathway EPZ is identified in NUREG-0654 / FEMA – REP-1, Rev. 1, [23]

as the area around an NPP of about 10 miles. Within the EPZ, detailed emergency plans are in 

place to reduce the risk of public health consequences in the unlikely event of an accident. 

Emergency planning within the EPZ provides a substantial basis for expansion of response 

efforts should it be necessary. ORO personnel have repeatedly demonstrated the ability to 

implement protective actions within the EPZ during inspected biennial exercises. Modeling of 

expected protective action response is described in this section. Analyses were conducted for 

accident scenarios identified in Table 6-1. 
6-3

Table 6-1 Scenarios Assessed for Emergency Response

Report Section

Number Scenario

6.3.1 Unmitigated STSBO 

6.3.2 Unmitigated STSBO with TI-SGTR 

6.3.3 Mitigated STSBO with TI-SGTR 

6.3.4 Unmitigated LTSBO 

6.3.5 Unmitigated ISLOCA 

6.4.1 Sensitivity 1 Unmitigated ISLOCA and evacuation to 16 miles

6.4.2 Sensitivity 2 Unmitigated ISLOCA and evacuation to 20 miles

6.4.3 Sensitivity 3 Unmitigated ISLOCA with a Delay in Implementation 

of Protective Actions

6.5.6 Seismic Analysis - Unmitigated STSBO with TI-SGTR

6.1 Population Attributes

The population near the Surry plant was modeled using six cohorts. A cohort is a population 

group that mobilizes or moves differently from other population groups. Modeling includes 

members of the public who evacuate early, evacuate late, those who refuse to evacuate, and 

those who evacuate from areas not under an evacuation order (e.g., the shadow evacuation). The 

consequence model does not constrain the number of cohorts but there is no benefit to defining 

an excessive number of cohorts with little difference in response characteristics. The following 

cohorts were established for SOARCA analyses: 

Cohort 1: 0 to 10 Public:

This cohort includes the public residing within the EPZ.

Cohort 2: 10 to 20 Shadow:

This cohort includes the shadow evacuation from the 10 to 20 mile area beyond the EPZ. A 

shadow evacuation occurs when members of the public evacuate from areas that are not under 

official evacuation orders and generally begins when a large scale evacuation is ordered [44].

Shadow evacuations are often reported and observed, although there is little quantitative data 

available. In a national telephone survey of residents of EPZs, more than 20 percent of people 

that had been asked to evacuate for emergencies such as hurricanes, had also evacuated for

similar emergencies in which they were asked not to evacuate (e.g., they were shadow 

evacuees) [44]. Additional information used to develop a value for use in the SOARCA analysis 

included a review of more than 20 NPP evacuation time estimate (ETE) studies. Although not 

currently required, most of these ETE studies included an analysis of a shadow evacuation. 

Typically, a shadow evacuation of 30 percent of the public outside the EPZ to a distance of 15 

miles was included in the analysis and often sensitivity analyses were provided that varied the 

shadow evacuation percentage to values as high as 60 percent. Review of the ETE values
6-4

showed that increasing the percentage of public participating in the shadow evacuation did not 

affect the ETE in 11 of 16 studies that included a sensitivity analysis. Using the above 

information, combined with the early decision in the SOARCA project to consider effects 

beyond the EPZ to a distance of 20 miles, a shadow evacuation of 20 percent of the public from 

the area 10 to 20 miles from the plant was modeled. The shadow evacuation was determined to 

have no effect on the evacuation of the Surry EPZ because much of the area beyond the EPZ is 

lightly populated. 

Cohort 3: 0 to 10 Schools:

This cohort includes elementary, middle and high school student populations within the EPZ. 

Schools receive early and direct warning from OROs and have response plans in place to support 

busing of students out of the EPZ, but sirens are not sounded as SAE.

Cohort 4: 0 to 10 Special Facilities:

The Special Facilities population includes residents of hospitals, nursing homes, assisted living 

communities and prisons. Special facility residents are assumed to reside in robust facilities such 

as hospitals, nursing homes, or similar structures, which provide additional shielding. Shielding 

factors for this population group consider this fact. In an emergency, Special Facilities would be 

evacuated individually over a period of time based upon available transportation. 

Cohort 5: 0 to 10 Tail:

The 0 to 10 Tail is defined as the last 10 percent of the public to evacuate from the EPZ. The 

approach to modeling the Tail is an analysis simplification to support inclusion of this population 

group. The Tail takes longer to evacuate for many valid reasons such as the need to return home 

from work to evacuate with the family, pick up children, shut down farming or manufacturing 

operations or performing other actions prior to evacuating as well as those who may miss the 

initial notification. 

Cohort 6: Non-evacuating public:

This cohort represents a portion of the public from 0 to 10 miles who may refuse to evacuate and 

is assumed to be 0.5 percent of the population. The cohort is modeled as though they are 

performing normal activities. Research of large scale evacuations has shown that a small 

percentage of the public refuse to evacuate [44] and this cohort accounts for this potential group. 

It is important to note that emergency planning is in place to support evacuation of 100 percent 

of the public.

6.1.1 Population Distribution

The Surry 2001 ETE was used to develop the population fractions for the cohorts within the 

EPZ. The populations provided in the Surry 2001 ETE present a detailed estimate of the 

population of the 0 to 10 mile region. 

A separate estimate was developed for the permanent residents and special facilities population 

beyond the EPZ to support development of the shadow population cohort and sensitivity 

analyses. SECPOP2000 [49] was used to estimate the population within 20 miles of the plant. 

The population was projected to 2005 using a multiplier of 1.0533 obtained from Census Bureau 

information. The population of the 10 to 20 mile area outside the EPZ was then calculated as the 
6-5

difference between the total estimated population within 20 miles and the 10 mile EPZ 

population. School children are not a separate cohort in the 10 to 20 mile area because it is 

assumed there is ample time for schools to close and children to go home and evacuate with

families; therefore they are included in the 10 to 20 public. Special facilities data for hospitals, 

nursing homes, and detention facilities in the 10 to 20 mile area was developed by researching 

available public information.

To establish the population distributions, the Shadow population was assessed first and defined 

as 20 percent of the total population within 10 to 20 miles from the plant. This value for the 

Shadow was then combined with non-evacuee, special facilities, and schools and then subtracted 

from the remaining total to establish the general public. Ten percent of the general public 

defines the evacuation tail, and the remainder was used as the total for the general public. The 

non-evacuating population is 0.5 percent of the total population in each region. Cohort 

populations are provided in Table 6-2.

Table 6-2 Surry Cohort Population Values

Cohort Description Population

1 0 to 10 Public 88,590

2 10 to 20 Shadow 63,171

3 0 to 10 Schools 23,262

4 0 to 10 Special Facilities 844

5 0 to 10 Tail 7,232

6 0 to 10 Non-evacuating 603

6.1.2 Evacuation Time Estimates

As provided in 10 CFR 50.47 Appendix E, each licensee is required to estimate the time to 

evacuate the EPZ. Appendix 4 of NUREG-0654/FEMA-REP-1, Revision 1 [23] provides 

information on the requirements of ETEs, and NUREG/CR-6863, “Development of Evacuation 

Time Estimate Studies for Nuclear Power Plants,” [24] provides detailed guidance on 

development of ETEs. A typical ETE includes many scenarios to help identify the combination 

of events for normal and off-normal conditions and provides emergency planners with estimates 

of the time to evacuate the EPZ under varying conditions [24]. The ETE study provides 

information regarding population characteristics, mobilization of the public, special facilities, 

transportation infrastructure, and other information used to estimate the time to evacuate the 

EPZ.

The SOARCA project used a normal weather weekday scenario that includes schools in session. 

This scenario was selected because it presents several challenges to timely protective action 

implementation including evacuating while residents are at work and mobilizing buses to 

evacuate children at school. The Surry 2001 ETE report provides the following regarding 

evacuation of the general public:
6-6

East of the James River- The densely populated area.

• 100 percent evacuation: 13 hours; and

• 90 percent evacuation: 10 hours and 50 minutes (rounded to 11 hours).

West of the James River- The rural low population area.

• 100 percent ETE: 3 hours 10 minutes; and 

• 90 percent ETE: Not provided. 

These values were used to develop the speeds for the cohorts used in the analysis. The Surry 

study describes the ETE scenario used as a ‘worst case’ because it includes the high number of 

transients in the area and schools in session. This scenario can be considered the bounding ETE 

case for the analysis and alternative seasonal evaluations and time of day are not necessary.

For the evacuation scenarios, a speed is input into the consequence model. The evacuation 

speed is developed from the ETE and is primarily influenced by population density and roadway 

capacity. When using ETE information, it is important to understand the components of the 

time estimate. The ETE includes mobilization activities the public undertakes upon receiving 

the initial notification of the incident [23], [24]. These actions include receiving the warning, 

verifying information, gathering children, pets, belongings, etc., packing, securing the home, 

and other evacuation preparations. Thus, a 13 hour ETE does not indicate that all of the vehicles 

are en route for 13 hours, but is the end of a 13 hour period in which the public mobilizes and 

evacuates the area. An evacuating population does not enter the roadway system at once. 

Rather an ideal model would include a “road loading function” that represents the expected 

movement. Most ETE studies use such a model. However, MACCS2 does not currently have 

the capability to move populations in this manner. This being the case, cohorts are modeled to 

begin moving together at a specific time after notification. To represent this movement of the 

cohort evacuating together, a single linear value based on distance divided by time (i.e., the 

ETE) was used for the speed. This distance over ETE ratio provides a slightly slower average 

speed than would be expected in an evacuation and adds some conservatism to the analysis. 

Evacuations can therefore be represented as a curve that is relatively steep at the beginning and 

tends to flatten as the last members of the public exit the area. Through review of more than 20 

existing ETE studies, the point at which the curve tends to flatten occurs where approximately 

90 percent of the population has evacuated. This is consistent with research that has shown that 

a small portion of the population that takes a longer time to evacuate than the rest of the general 

public and is the last to leave the evacuation area [23]. This last 10 percent of the population is 

identified as the evacuation tail. A goal of emergency preparedness is to protect the public 

health and safety of the public. To best achieve this goal, new guidance from the NRC [65]

suggests the 90 percent ETE be used when making protective action recommendations and 

decisions; therefore, for the analyses in this study, the 90 percent ETE value was used to 

develop evacuation speeds.
6-7

6.2 WinMACCS 

WinMACCS is a user interface for the MACCS2 code and was used to generate input for 

MACCS2 model runs. WinMACCS has the ability to integrate the information described above 

into the consequence analysis. The entire evacuation area was mapped onto a radial sector grid 

network around the plant. The roadway network was reviewed against site-specific evacuation 

plans to determine likely evacuation direction in each grid element. The results of the ETE were 

reviewed to determine localized areas of congestion and areas where no congestion would be 

expected. Using this information, speed adjustment factors ranging from one to three were 

applied at the grid element level to speed up vehicles in the rural uncongested areas and to slow 

down vehicles in more urban settings. 

6.2.1 Hotspot and Normal Relocation and Habitability

In the unlikely case of a severe accident and radiological release, protective actions in addition to

evacuation may be implemented. For instance, residents would be relocated if their potential 

dose exceeds protective action criteria. OROs would base this determination on dose projections 

using state, utility, and Federal agency computer models as well as measurements taken in the 

field. Hotspot relocation and normal relocation models are included in the MACCS2 code to 

treat this contingency. Total dose commitment pathways for the relocation models are 

cloudshine, groundshine, direct inhalation, and resuspension inhalation. Relocated individuals 

are removed from the calculation for the remainder of the emergency phase and receive no 

additional dose during that phase. The dose criteria are applied after plume arrival at the affected 

area.

Hotspot relocation of individuals beyond ten miles occurs 24 hours after plume arrival if the total 

lifetime dose commitment for the weeklong emergency phase exceeds 0.05 Sv (5 rem). Normal 

relocation of individuals occurs 36 hours after plume arrival if the total lifetime dose 

commitment exceeds 0.01 Sv (1 rem). The relocation times of 24 hours for hotspot and 36 hours 

for normal relocation were established based on review of the emergency response timelines, 

which suggest that because of the high population density in some areas off the EPZ, OROs 

would not likely be available earlier to assist with relocation due to higher priority tasks in the 

evacuation area. Relocation is a process that requires identification of the affected areas and 

notification of residents within those areas. The time values represent the average time expected 

to implement each action. 

The hotspot value used in NUREG-1150 [2] was 0.5 Sv (50 rem) and the relocation value was 

0.25 Sv (25 rem). The long term habitability criteria used in NUREG-1150 was 0.04 Sv (4 rem) 

over a 5 year period. The NUREG-1150 long term habitability criterion is the same as the site 

specific value used for the Surry analysis. It should be noted that the non-evacuating cohort is 

still subject to the Hotspot and Normal Relocation criterion. It is assumed these individuals will 

evacuate when they understand a release has in fact occurred and they are informed they are 

located in high dose areas.

6.2.2 Shielding Factors

Shielding factors vary by geographical region across the United States, and those used in the 

Surry analysis are shown in Table 6-3. The factors represent the fraction of dose that a person 
6-8

would be exposed to when performing normal activities, evacuating, or staying in a shelter in 

comparison to a person outside with full exposure and are applied to all cohorts. Special 

Facilities are typically larger and more robust structures than housing stock and therefore have 

better shielding factors as identified in the table. Special facilities have the same factor for 

normal and shelter indicating these individuals are all indoors. 

Table 6-3 Surry Shielding Factors

Ground Shine Cloud Shine Inhalation/Skin

Normal Evac. Shelter Normal Evac. Shelter Normal Evac. Shelter

Cohorts 0.26 0.50 0.20 0.68 1.00 0.60 0.46 0.98 0.33

Special 

Facilities 0.05 0.50 0.05 0.31 1.00 0.31 0.33 0.98 0.33

The shielding factors provided in Table 6-3 were obtained from a variety of sources. Where 

appropriate, site specific values for sheltering were obtained from NUREG-1150 [2]. An 

updated inhalation/skin evacuation shielding factor was obtained from NUREG/CR-6953, 

Vol. 1, [45]. The normal activity shielding factors have been adjusted to account for the 

understanding that people do not spend a great deal of time outdoors. The normal activity values 

are all weighted averages of indoor and outdoor values based on being indoors 81 percent of the 

time and outdoors 19 percent of the time [66]. Indoor values are assumed to be the same as 

sheltering.

6.2.3 Potassium Iodide

The State of Virginia implements a potassium iodide (KI) program. The Virginia Department of 

Health provides potassium iodide to people who live, work or visit within 10 miles of the Surry 

NPP. Potassium iodide also is available to the public for purchase without a prescription at 

pharmacies and from manufacturers. 

The purpose of the KI is to saturate the thyroid gland with stable iodine so that further uptake of 

radioactive iodine by the thyroid is diminished. If taken at the right time and in the appropriate 

dosage, KI can nearly eliminate doses to the thyroid gland from inhaled radioiodine. Factors that 

contribute to the effectiveness of KI include the availability (i.e., whether residents can find 

their KI), the timing of ingestion, and the degree of pre-existing stable iodine saturation of the 

thyroid gland. It is considered that some residents will not remember where they have placed 

their KI or may not have it available and will therefore not take KI. It is also assumed some 

residents will not take their KI when directed (i.e., they may take it early or late which reduces 

the efficacy). To account for this, KI was turned on in the model for approximately 50 percent of 

the public, and the efficacy of the KI was set at 70 percent. 

6.2.4 Adverse Weather

Adverse weather is typically defined as rain, ice, or snow that affects the response of the public 

during an emergency. The affect of adverse weather on the mobilization of the public is not 

directly considered in establishing emergency planning parameters for this project because such 

a consideration approximates a worst case evacuation scenario. However, adverse weather was 
6-9

addressed in the movement of cohorts within the analysis. The evacuation speed multiplier

(ESPMUL) parameter in WinMACCS is used to reduce travel speed when precipitation is 

occurring as indicated from the meteorological weather file. The ESPMUL factor was set at 0.7, 

which effectively slows down the evacuating public to 70 percent of the established travel speed 

when precipitation exists.

6.2.5 Modeling using Evacuation Time Estimates

The purpose of using the ETE as a parameter in consequence modeling is to better approximate 

the real time actions expected of the public. Although consequence modeling has evolved to 

allow use of many cohorts and can address many individual aspects of each cohort, the approach 

to modeling evacuations is not direct. As stated earlier, evacuations include mobilizing and 

evacuating the public over a period of time, which is best modeled as a distribution. To use 

WinMACCS, this distribution of data must be converted into discrete events. For instance, upon 

the sounding of the sirens and issuance of the Emergency Alert System messaging, it is assumed 

all members of the public shelter and one hour later all members of the public enter the roadway 

network together at the same time and begin to evacuate. In research of existing evacuations for 

technological hazards, it is shown that most members of the public would enter the roadway 

network over a period of about an hour. It is not realistic that all vehicles would load 

simultaneously; however, this treatment within the model is necessary due to the current 

modeling abilities of WinMACCS. 

Although WinMACCS can accommodate more cohorts, expert judgment was used to balance the 

number of cohorts with model run time. The speeds for each cohort are developed from the ETE, 

and the elements that factor into the speeds include:

• Time to receive notification and prepare to evacuate (i.e., mobilization time);

• Time to evacuate; and

• Distance of travel.

The time to receive notification requires assurance that sirens sound when needed. In review of 

the Reactor Oversight Program data regarding sirens for Surry, the average siren performance 

indicator was 99.9 percent, indicating that sirens will perform when needed. With few 

exceptions, travel speeds were established as whole numbers. A simple ratio of distance to time 

would show that evacuation of the 0 to 10 public from the 10 mile EPZ at Surry, which has an 

ETE of 10 hours 50 minutes, would provide a speed of 0.92 mph. However, as indicated above, 

notification and preparation to evacuate are included in the ETE.

For the general public, a one hour delay to shelter is assigned to reflect the mobilization time 

when residents receive the warning and prepare to evacuate. If the one hour mobilization time is 

subtracted from the ETE (10:50 – 1 hour) there remains 9 hours and 50 minutes to travel a 

maximum of 10 miles. As observed in actual evacuations due to technological or other hazards, 

people perform these mobilization activities at varying times with some residents ready to 

evacuate quickly while others can take up to an hour or longer. While this cohort is sheltered, a 

greater shielding factor is applied, and while en route during the evacuation, a lower shielding 

factor is applied.
6-10

During the evacuation, roadway congestion occurs rather quickly and traffic exiting the EPZ 

begins to slow. In review of over 20 ETE studies, this congestion typically occurs in 1 to 2 hours 

depending upon the population density and roadway capacity of the EPZ. In the SOARCA 

analysis, the 0-10 public is sheltered and preparing to evacuate for one hour. The public is then 

loaded onto the roadway and congestion is assumed to occur within 15 minutes. This total time 

of 1 hour 15 minutes for congestion to occur was established to be consistent with ETE studies. 

The calculation of the speed of evacuees includes the first 15 minutes to the point when 

congestion occurs. For this first 15 minutes, evacuees are assumed to travel at 5 mph. The slow 

initial speed was set to account for the model loading all members of a cohort at one time. In the 

first 15 minutes at 5 mph, a distance of 1.25 miles has been traveled. At that time congestion is 

heavy and speeds slow for the next 8.75 miles. 

The ETE is 10 hours 50 minutes for this cohort. Having sheltered and prepared to evacuate for 

1 hour and then traveled the first 15 minutes at 5 mph, the remaining time is 9 hours and 

35 minutes (10:50 - 1 hour shelter - 15 minutes at 5 mph). To determine the speed of travel for 

the remaining 8.75 miles, the distance is divided by the time (8.75 miles / 9 hours and 

35 minutes) which provides a speed of 0.9 mph. The calculated speed used in the analysis for 

this cohort was rounded to 1 mph for this cohort. This approach applies to evacuees travelling 10 

miles. Most evacuees would travel less distance and some travel a greater distance because the 

roadways are not radial away from the plant. Adjustments were made to the speeds based on 

review of population densities and aerial photos of roadways. The baseline speed of 1 mph was 

assigned to the general public cohort Because this represents the slowest speed expected, the 

speed was increased using a speed adjustment factor of 3 in the rural areas where no congestion 

is expected. 

6.2.6 Establishing the Initial Cohort in the Calculation

The WinMACCS parameters for the cohorts are stored in multi-dimensional arrays, and the 

dimensions of the arrays are defined by geographical area for the analysis. WinMACCS requires 

the dimensions be established with the first cohort. All subsequent cohorts must be defined 

within these array dimensions, meaning they can extend from the origin to any distance equal to 

or less than the maximum distance established with the first cohort. 

Cohort 1 was defined as the 0 to 10 mile public and has the same response characteristics as 

Cohort 2. The cohort that extends the greatest distance and defines the limits of the array is the 

Shadow Evacuation, which is Cohort 2. Thus, in the WinMACCS model, Cohorts 1 and 2 had to 

be redefined to meet the above requirement. The model input parameters for Cohort 1 were 

extended from the plant out to the maximum array distance of 20 miles, and Cohort 2 extends 

from the plant out to 10 miles. Cohort 1 is input as 20 percent of the population from 0 to 

20 miles. This captures the 20 percent of the population between 10 and 20 miles involved in the 

shadow evacuation beyond the EPZ. As noted earlier, for this site the shadow evacuation has no 

effect on the evacuation of the residents of the EPZ. The combination of Cohorts 1 and 2 from 0 

to 10 miles in the WinMACCS model represent the Public (0 – 10) Cohort defined above. For 

the remaining cohorts, application of parameters in the WinMACCS model is direct, and the 

population fractions directly correspond to the cohort descriptions.
6-11

6.3 Accident Scenarios

An emergency response timeline was developed for each accident scenario using information 

from the MELCOR analyses, expected timing of Emergency Classification declarations, and 

information from the ETE. The timeline identifies points at which cohorts would receive 

instruction from OROs to implement protective actions. In practice, initial evacuation orders are 

based on the severity of the accident and in Virginia would likely include an evacuation of the 

2 mile zone and a 5 mile downwind keyhole consistent with the guidance in Supplement 3 to

NUREG-0654/FEMA-REP-1, Revision 1 [23]. The distance would be expanded to 10 miles 

based on dose projections. Because WinMACCS does not readily support modeling a keyhole 

area, the SOARCA project modeled evacuation of the full EPZ and a shadow evacuation from 

the 10 to 20 mile area. 

6.3.1 Unmitigated LTSBO 

The emergency response timeline for the unmitigated LTSBO scenario is shown in Figure 6-2. 

The timing of emergency classification declarations was based on the EALs contained in site 

emergency plan implementing procedures. Protective actions were assumed to be recommended 

by OROs in accordance with approved emergency plans and procedures. Discussions were held 

with site representatives to help ensure proper understanding of EALs for each accident scenario 

and emergency response practices. Discussions with OROs confirmed that sirens are only 

sounded for a GE at Surry. Siren systems are tested routinely within all EPZs and the results of 

the Response Oversight Program indicate a 99.9 percent performance rating for sirens at Surry. 

Therefore, it is assumed that sirens do not fail and in the event one or two do fail, societal 

notification and route alerting by OROs would alert residents in these areas within the same 

mobilization time period as estimated for the EPZ. 

For this scenario, EAL SS1.1 specifies that if all offsite power and all onsite AC power is lost for 

greater than 15 minutes an SAE is declared. If restoration of power is not likely within 4 hours, 

EAL SG1.1 establishes that a GE be declared. It is expected the SAE is declared in about 

15 minutes, and plant operators would recognize rather soon that restoration of power within 

4 hours is unlikely. A 2 hour period from loss of power was selected as a reasonable time for 

declaration of a GE. It is expected that notification to OROs is timely and the sounding of sirens 

and broadcast of EAS messages occurs approximately 45 minutes after declaration of GE. From 

the MELCOR analysis, the first fission product gap release occurs 16 hours into the event with a

significant radioactive release to the environment occurring 45.5 hours into the event. The 

duration of specific protective actions for each cohort are summarized in Figure 6-3.
6-12

Figure 6-2 Unmitigated LTSBO Emergency Response Timeline

Figure 6-3 Duration of Protective Actions - Unmitigated LTSBO

The Virginia Department of Emergency Management will directly communicate with schools 

upon receiving the declaration of SAE. This allows for the preparation and early response of 

schools, but the public is largely unaware at this time. It could be noted that there would be a 

societal communication process as members of the public become aware of the school 

evacuation. Sirens are sounded at an SAE in many states, but Virginia only sounds sirens in 

response to declaration of a GE. Although there could potentially be some shadow evacuation 

due to societal communication, it is assumed that there would be no significant movement of the 

general public. The initiating event for the Unmitigated STSBO, Unmitigated STSBO with 
6-13

TI-SGTR, Mitigated STSBO with TI-SGTR, and Unmitigated LTSBO scenarios is a station 

blackout and EAL SS1.1 and SG1.1 are reached. Therefore, the cohort actions are identical for 

each of these scenarios:

Cohort 1: 0 to 10 Public:

Following declaration of a GE, sirens are sounded and an evacuation order would be issued via 

an EAS message for affected areas within the EPZ. Cohort 1 is assumed to shelter when the 

sirens sound. The time to receive the warning and prepare to mobilize is assumed to be 1 hour 

after the siren. One hour is based on evacuation research, which shows the public mobilizes over 

a period of time with some members of the public moving soon after hearing the sirens, while 

most take some time to prepare and then evacuate. One hour was selected as a reasonable 

centroid of an evacuation curve for this cohort, which is consistent with empirical data from 

previous large scale evacuations [26].

Cohort 2: 10 to 20 Shadow: 

This cohort is assumed to begin movement at the same time as the 0 to 10 Public after sirens 

have sounded within the EPZ and when widespread media broadcasts are underway. Residents 

in the 10 to 20 area begin seeing large numbers of people evacuating and initiate a shadow 

evacuation. There is no warning or notification for the public residing in this area, which is not 

under an evacuation order.

Cohort 3: 0 to 10 Schools: 

Schools are the first to take action. Upon receipt of the declaration of SAE by the site, the 

Virginia Department of Emergency Management would notify the schools in accordance with 

the offsite emergency response plan. It is assumed schools are notified at SAE and begin 

sheltering in about 15 minutes. Buses would be mobilized, and it is assumed schools begin 

evacuating 1:15 after the start of the incident. At this time in the event, roads are uncongested 

and school buses are able to exit the EPZ quickly.

Cohort 4: 0 to 10 Special Facilities: 

Special Facilities can take longer to evacuate than the general public because transportation 

resources, some of which are very specialized, must be mobilized. Special Facilities would be 

evacuated individually over a period of time based upon available transportation and the number 

of return trips needed. Special Facilities provide better shielding for the residents, thus while 

residents are in the facility, they are better protected than when they are evacuating. It was 

determined that the best representation of this cohort in the modeling is to evacuate with the tail 

and apply shielding factors consistent with the types of structures within which these residents 

reside. The Special Facilities cohort is assumed to depart at the same time as the evacuation tail, 

although it is recognized this cohort would begin mobilization about the same time as the 

schools.

Cohort 5: 0 to 10 Tail:

Using the evacuation data provided in the Surry ETE study, 90 percent of the evacuation of the 

EPZ is complete at approximately 11 hours into the evacuation, and this corresponds to the 

departure time for the 0 to 10 Tail. 
6-14

Cohort 6: Non-evacuating public:

This cohort group represents the portion of the public who may refuse to evacuate and is 

assumed to be 0.5 percent of the population. Any member of the public who does not evacuate is 

still subject to the Hotspot and Normal Relocation criterion discussed earlier.

Selected input parameters for WinMACCS are provided to support detailed use of this study. 

More detailed information regarding modeling parameters is available in the MACCS2 User’s 

Guide [41]. A brief description of the parameters is provided below:

• Delay to Shelter (DLTSHL) represents a delay from the time of the start of the accident 

until cohorts enter the shelter. 

• Delay to Evacuation (DLTEVA) represents the length of the sheltering period from the 

time a cohort enters the shelter until the point at which they begin to evacuate. 

• The speed (ESPEED) is assigned for each of the three phases used in WinMACCS 

including Early, Middle, and Late. Average evacuation speeds were developed from the 

Surry 2001 ETE report. Speed adjustment factors were then utilized in the WinMACCS 

application to represent free flow in rural areas and congested flow in urban areas.

• Duration of Beginning phase (DURBEG) is the duration assigned to the beginning phase 

of the evacuation and may be assigned uniquely for each cohort. 

• Duration of Middle phase (DURMID) is the duration assigned to the middle phase of the 

evacuation and may also be assigned uniquely for each cohort.

For the 0 to 10 Public and the 0 to 10 Tail cohorts, by definition the sum of the DLTEVA, 

DURBEG and DURMID is equal to the ETE. This is because the ETE does not include shelter 

time. Table 6-4 provides a summary of the evacuation timing actions for each cohort. The 

MACCS2 variable names are included at the top of each column. 
6-15

Table 6-4 Unmitigated LTSBO cohort timing

† Values represent speeds east of the James River. Speeds west of the river are increased through use 

of multipliers in the WinMACCS model.

Departure speeds and durations of the beginning and middle periods for the WinMACCS runs 

were developed from the Surry ETE study. Adjustments were made to individual elements of 

the WinMACCS grid to reflect differences in vehicle direction and speed of travel through the 

network. The timeline identifies the point at which it is assumed that cohorts begin to take action. 

The actions taken by each cohort last for a given period as indicated in Table 6-4.

6.3.2 Unmitigated STSBO

The emergency response timeline for the unmitigated STSBO scenario is shown in Figure 6-4. 

The timing of emergency classification declarations was based on the EALs contained in site 

emergency plan implementing procedures. Protective actions were expected to be recommended 

by OROs in accordance with approved emergency plans and procedures. Discussions were held 

with site representatives to help ensure proper understanding of EALs for each accident scenario 

and emergency response practices. Discussions with OROs confirmed that sirens are only 

sounded for a GE at Surry. Siren systems are tested routinely within all EPZs and the results of 

the Response Oversight Program indicate a 99.9 percent performance rating for sirens at Surry. 

Therefore, it is assumed that sirens do not fail and in the event one or two do fail, societal 

notification and route alerting by OROs would alert residents in these areas within the same 

mobilization time period as estimated for the EPZ. Figure 6-5 summarizes the duration of 

specific protective actions for each cohort.

For this scenario, EAL SS1.1 specifies that if all offsite power and all onsite AC power is lost for 

greater than 15 minutes an SAE is declared. If restoration of power is not likely within 4 hours, 

EAL SG1.1 establishes that a GE be declared. It is expected the SAE is declared in about 

15 minutes as shown in Figure 6-4 and that plant operators would recognize rather soon that 

restoration of power within 4 hours is unlikely. A 2 hour period from loss of power was selected 

as a reasonable time for declaration of a GE. It is expected that notification to OROs is timely 

Cohort

Delay to 

Shelter

DLTSHL 

(hr)

Delay to 

Evacuation

DLTEVA 

(hr)

Duration of 

beginning 

phase

DURBEG

(hr)

Duration of 

middle phase

DURMID

(hr)

Speed of 

early phase

ESPEED†

(early)

mph

Speed of 

middle phase

ESPEED†

(mid)

mph

0 to 10

Public 2.75 1 0.25 9.75 5 1

10 to 20 

Shadow 2.75 1 0.25 9.75 5 1

0 to 10

Schools 0.25 1 0.25 0.5 10 10

0 to 10

Special 

Facilities

2.75 11 1 1 1 10

0 to 10

Tail 2.75 11 1 1 1 10

Non-Evac NA NA NA NA 0 0
6-16

and the sounding of sirens and broadcasting the EAS message occurs approximately 45 minutes

after declaration of GE. From the MELCOR analysis, the first fission product gap release occurs

about three hours into the event with a significant radioactive release occurring 25.5 hours into 

the event.

Figure 6-4 Unmitigated STSBO emergency response timeline

Figure 6-5 Duration of Protective Actions - Unmitigated STSBO
6-17

The Virginia Department of Emergency Management will directly communicate with schools 

upon receiving the declaration of SAE. This allows for the preparation and early response of 

schools, but the public is largely unaware at this time. It could be noted that there would be a 

societal communication process as members of the public become aware of the school 

evacuation. Sirens are sounded at an SAE in many states, but Virginia only sounds sirens in 

response to declaration of a GE. Although there could potentially be some shadow evacuation 

due to societal communication, it is assumed that there would be no significant movement of the 

general public. The initiating event for the Unmitigated STSBO, Unmitigated STSBO with 

TI-SGTR, Mitigated STSBO with TI-SGTR, and Unmitigated LTSBO scenarios is a station 

blackout and EAL SS1.1 and SG1.1 are reached. Therefore, the cohort actions are identical for 

each of these scenarios. The cohorts are expected to respond to the EAS messages as described 

below:

Cohort 1: 0 to 10 Public: 

Following declaration of a GE, sirens are sounded and an evacuation order would be issued via 

an EAS message for affected areas within the EPZ. Cohort 1 is assumed to shelter when the 

sirens sound. The time to receive the warning and prepare to mobilize is assumed to be 1 hour

after the siren. One hour is based on evacuation research, which shows the public mobilizes over 

a period of time with some members of the public moving soon after hearing the sirens, while 

most take some time to prepare and then evacuate. One hour was selected as a reasonable 

centroid of an evacuation curve for this cohort, which is consistent with empirical data from 

previous large scale evacuations [26].

Cohort 2: 10 to 20 Shadow:

This cohort is assumed to begin movement at the same time as the 0 to 10 Public after sirens 

have sounded within the EPZ and when widespread media broadcasts are underway. Residents 

in the 10 to 20 area begin seeing large numbers of people evacuating and initiate a shadow 

evacuation. There is no warning or notification for the public residing in this area, which is not 

under an evacuation order.

Cohort 3: 0 to 10 Schools: 

Schools are the first to take action. Upon receipt of the declaration of SAE by the site, the 

Virginia Department of Emergency Management would notify the schools in accordance with 

the offsite emergency response plan. It is assumed schools are notified at SAE and begin 

sheltering in about 15 minutes. Buses would be mobilized, and it is assumed schools begin 

evacuating 1:15 after the start of the incident. At this time in the event, roads are uncongested 

and school buses are able to exit the EPZ quickly.

Cohort 4: 0 to 10 Special Facilities:

Special Facilities can take longer to evacuate than the general public because transportation 

resources, some of which are very specialized, must be mobilized. Special Facilities would be 

evacuated individually over a period of time based upon available transportation and the number 

of return trips needed. Special Facilities provide better shielding for the residents, thus while 

residents are in the facility, they are better protected than when they are evacuating. It was 

determined that the best representation of this cohort in the modeling is to evacuate with the tail 

and apply shielding factors consistent with the types of structures within which these residents 
6-18

reside. The Special Facilities cohort is assumed to depart at the same time as the evacuation tail, 

although it is recognized this cohort would begin mobilization about the same time as the 

schools.

Cohort 5: 0 to 10 Tail: 

Using the evacuation data provided in the Surry ETE study, 90 percent of the evacuation of the 

EPZ is complete at approximately 11 hours into the evacuation, and this corresponds to the 

departure time for the 0 to 10 Tail. 

Cohort 6: Non-evacuating public:

This cohort group represents the portion of the public who may refuse to evacuate and is 

assumed to be 0.5 percent of the population. Any member of the public who does not evacuate is 

still subject to the Hotspot and Normal Relocation criterion discussed earlier. 

The evacuation timing for each cohort is presented in Table 6-5. Selected input parameters for 

WinMACCS are provided to support detailed use of this study. More detailed information 

regarding modeling parameters is available in the MACCS2 User’s Guide [41]. A brief 

description of the parameters is provided below:

• Delay to Shelter (DLTSHL) represents a delay from the time of the start of the accident 

until cohorts enter the shelter.

• Delay to Evacuation (DLTEVA) represents the length of the sheltering period from the 

time a cohort enters the shelter until the point at which they begin to evacuate. 

• The speed (ESPEED) is assigned for each of the three phases used in WinMACCS 

including Early, Middle, and Late. Average evacuation speeds were developed from the 

Surry 2001 ETE report. Speed adjustment factors were then utilized in the WinMACCS 

application to represent free flow in rural areas and congested flow in urban areas.

• Duration of Beginning phase (DURBEG) is the duration assigned to the beginning phase 

of the evacuation and may be assigned uniquely for each cohort. 

• Duration of Middle phase (DURMID) is the duration assigned to the middle phase of the 

evacuation and may also be assigned uniquely for each cohort.

For the 0 to 10 Public and the 0 to 10 Tail cohorts, by definition the sum of the DLTEVA, 

DURBEG and DURMID is equal to the ETE. This is because the ETE does not include shelter 

time.
6-19

Table 6-5 Unmitigated STSBO cohort timing

Cohort

Delay to 

Shelter

DLTSHL 

(hr)

Delay to 

Evacuation

DLTEVA 

(hr)

Duration of 

beginning 

phase

DURBEG

(hr)

Duration of 

middle 

phase

DURMID

(hr)

Speed of 

early 

phase

ESPEED†

(early)

mph

Speed of 

middle 

phase

ESPEED†

(mid)

mph

0 to 10

Public 2.75 1 0.25 9.75 5 1

10 to 20 Shadow 2.75 1 0.25 9.75 5 1

0 to 10

Schools 0.25 1 0.25 0.5 10 10

0 to 10 Special 

Facilities 2.75 11 1 1 1 10

0 to 10 Tail 2.75 11 1 1 1 10

Non-Evac NA NA NA NA 0 0

† Values represent speeds east of the James River. Speeds west of the river are increased through use of 

multipliers in the WinMACCS model.

6.3.3 Unmitigated STSBO with TI-SGTR 

The emergency response timeline for the unmitigated STSBO with TI-SGTR scenario is shown 

in Figure 6-6. For this scenario, EAL SS1.1 specifies that if all offsite power and all onsite AC 

power is lost for greater than 15 minutes an SAE is declared. If restoration of power is not likely 

within 4 hours, EAL SG1.1 establishes that a GE be declared. It is expected the SAE is declared 

in about 15 minutes, and plant operators would recognize rather soon that restoration of power 

within 4 hours is unlikely. A 2 hour period from loss of power was selected as a reasonable time

for declaration of a GE. It is expected that notification to OROs is timely and the sounding of

sirens and broadcast of an EAS message occurs approximately 45 minutes after declaration of 

GE. From the MELCOR analysis, the first fission product gap release occurs about three hours 

into the event with the SGTR occurring 30 minutes later. The duration of specific protective 

actions for each cohort in this scenario is described in Figure 6-7.
6-20

Figure 6-6 Unmitigated STSBO with TI-SGTR emergency response timeline

Figure 6-7 Duration of Protective Actions - Unmitigated STSBO with TI-SGTR

The initiating event for the Unmitigated STSBO, Unmitigated STSBO with TI-SGTR, Mitigated 

STSBO with TI-SGTR, and Unmitigated LTSBO scenarios is a station blackout and EAL SS1.1 

and SG1.1 are reached. Therefore, the cohort actions are identical for each of these scenarios as 

described below:

Cohort 1: 0 to 10 Public:

Following declaration of a GE, sirens are sounded and an evacuation order would be issued via 

an EAS message for affected areas within the EPZ. Cohort 1 is assumed to shelter when the 

sirens sound. The time to receive the warning and prepare to mobilize is assumed to be 1 hour 
6-21

after the siren. One hour is based on evacuation research, which shows the public mobilizes over 

a period of time with some members of the public moving soon after hearing the sirens, while 

most take some time to prepare and then evacuate. One hour was selected as a reasonable 

centroid of an evacuation curve for this cohort, which is consistent with empirical data from 

previous large scale evacuations [26].

Cohort 2: 10 to 20 Shadow: 

This cohort is assumed to begin movement at the same time as the 0 to 10 Public after sirens 

have sounded within the EPZ and when widespread media broadcasts are underway. Residents 

in the 10 to 20 area begin seeing large numbers of people evacuating and initiate a shadow 

evacuation. There is no warning or notification for the public residing in this area, which is not 

under an evacuation order.

Cohort 3: 0 to 10 Schools: 

Schools are the first to take action. Upon receipt of the declaration of SAE by the site, the 

Virginia Department of Emergency Management would notify the schools in accordance with 

the offsite emergency response plan. It is assumed schools are notified at SAE and begin 

sheltering in about 15 minutes. Buses would be mobilized, and it is assumed schools begin 

evacuating 1:15 after the start of the incident. At this time in the event, roads are uncongested 

and school buses are able to exit the EPZ quickly.

Cohort 4: 0 to 10 Special Facilities: 

Special Facilities can take longer to evacuate than the general public because transportation 

resources, some of which are very specialized, must be mobilized. Special Facilities would be 

evacuated individually over a period of time based upon available transportation and the number 

of return trips needed. Special Facilities provide better shielding for the residents, thus while 

residents are in the facility, they are better protected than when they are evacuating. It was 

determined that the best representation of this cohort in the modeling is to evacuate with the tail 

and apply shielding factors consistent with the types of structures within which these residents 

reside. Special Facilities are assumed to depart at the same time as the evacuation tail, although 

it is recognized this cohort would begin mobilization about the same time as the schools.

Cohort 5: 0 to 10 Tail: 

Using the evacuation data provided in the Surry ETE study, 90 percent of the evacuation of the 

EPZ is complete at approximately 11 hours into the evacuation, and this corresponds to the 

departure time for the 0 to 10 Tail. 

Cohort 6: Non-evacuating public:

This cohort group represents the portion of the public who may refuse to evacuate and is 

assumed to be 0.5 percent of the population. Any member of the public who does not evacuate is 

still subject to the Hotspot and Normal Relocation criterion discussed earlier. 

Table 6-6 provides a summary of the evacuation timing activities for each cohort.
6-22

Table 6-6 Unmitigated STSBO with TI-SGTR cohort timing

Cohort

Delay to 

Shelter

DLTSHL 

(hr)

Delay to 

Evacuation

DLTEVA 

(hr)

Duration of 

beginning 

phase

DURBEG

(hr)

Duration of 

middle 

phase

DURMID

(hr)

Speed of 

early phase

ESPEED†

(early)

mph

Speed of 

middle phase

ESPEED†

(mid)

mph

0 to 10

Public 2.75 1 0.25 9.75 5 1

10 to 20 

Shadow 2.75 1 0.25 9.75 5 1

0 to 10

Schools 0.25 1 0.25 0.5 10 10

0 to 10

Special 

Facilities

2.75 11 1 1 1 10

0 to 10

Tail 2.75 11 1 1 1 10

Non-Evac NA NA NA NA 0 0

† Values represent speeds east of the James River. Speeds west of the river are increased through use 

of multipliers in the WinMACCS model.

6.3.4 Mitigated STSBO with TI-SGTR 

The accident scenario timeline for the Mitigated STSBO with TI-SGTR is identical to the 

unmitigated STSBO with TI-SGTR as shown in Figure 6-8 and Figure 6-9. The values identified

in Table 6-6 were used to support the consequence analyses for the Mitigated STSBO with 

TI-SGTR. 

Figure 6-8 Mitigated STSBO with TI-SGTR emergency response timeline
6-23

Figure 6-9 Duration of Protective Actions - Mitigated STSBO with TI-SGTR

The initiating event for the Unmitigated STSBO, Unmitigated STSBO with TI-SGTR, Mitigated 

STSBO with TI-SGTR, and Unmitigated LTSBO scenarios is a station blackout and EAL SS1.1 

and SG1.1 are reached. Therefore, the cohort actions are identical for each of these scenarios. 

Cohort 1: 0 to 10 Public: 

Following declaration of a GE, sirens are sounded and an evacuation order would be issued via 

an EAS message for affected areas within the EPZ. Cohort 1 is assumed to shelter when the 

sirens sound. The time to receive the warning and prepare to mobilize is assumed to be 1 hour 

after the siren. One hour is based on evacuation research, which shows the public mobilizes over 

a period of time with some members of the public moving soon after hearing the sirens, while 

most take some time to prepare and then evacuate. One hour was selected as a reasonable 

centroid of an evacuation curve for this cohort, which is consistent with empirical data from 

previous large scale evacuations [26].

Cohort 2: 10 to 20 Shadow:

This cohort is assumed to begin movement at the same time as the 0 to 10 Public after sirens 

have sounded within the EPZ and when widespread media broadcasts are underway. Residents 

in the 10 to 20 area begin seeing large numbers of people evacuating and initiate a shadow 

evacuation. There is no warning or notification for the public residing in this area, which is not 

under an evacuation order.

Cohort 3: 0 to 10 Schools:

Schools are the first to take action. Upon receipt of the declaration of SAE by the site, the 

Virginia Department of Emergency Management would notify the schools in accordance with 

the offsite emergency response plan. It is assumed schools are notified at SAE and begin 

sheltering in about 15 minutes. Buses would be mobilized, and it is assumed schools begin 

evacuating 1:15 hour after the start of the incident. At this time in the event, roads are 

uncongested and school buses are able to exit the EPZ quickly.
6-24

Cohort 4: 0 to 10 Special Facilities:

Special Facilities can take longer to evacuate than the general public because transportation 

resources, some of which are very specialized, must be mobilized. Special Facilities would be 

evacuated individually over a period of time based upon available transportation and the number 

of return trips needed. Special Facilities provide better shielding for the residents, thus while 

residents are in the facility, they are better protected than when they are evacuating. It was 

determined that the best representation of this cohort in the modeling is to evacuate with the tail 

and apply shielding factors consistent with the types of structures within which these residents 

reside. The Special Facilities cohort is assumed to depart at the same time as the evacuation tail, 

although it is recognized this cohort would begin mobilization about the same time as the 

schools.

Cohort 5: 0 to 10 Tail:

Using the evacuation data provided in the Surry ETE study, 90 percent of the evacuation of the 

EPZ is complete at approximately 11 hours into the evacuation, and this corresponds to the 

departure time for the 0 to 10 Tail. 

Cohort 6: Non-evacuating public:

This cohort group represents the portion of the public who may refuse to evacuate and is 

assumed to be 0.5 percent of the population. Any member of the public who does not evacuate is 

still subject to the Hotspot and Normal Relocation criterion discussed earlier. 

Table 6-7 provides a summary of the evacuation timing activities for each cohort.

Table 6-7 Unmitigated STSBO with TI-SGTR cohort timing

Cohort

Delay to 

Shelter

DLTSHL 

(hr)

Delay to 

Evacuation

DLTEVA 

(hr)

Duration of 

beginning 

phase

DURBEG

(hr)

Duration of 

middle 

phase

DURMID

(hr)

Speed of 

early phase

ESPEED†

(early)

mph

Speed of 

middle phase

ESPEED†

(mid)

mph

0 to 10

Public 2.75 1 0.25 9.75 5 1

10 to 20 

Shadow 2.75 1 0.25 9.75 5 1

0 to 10

Schools 0.25 1 0.25 0.5 10 10

0 to 10

Special 

Facilities

2.75 11 1 1 1 10

0 to 10

Tail 2.75 11 1 1 1 10

Non-Evac NA NA NA NA 0 0

† Values represent speeds east of the James River. Speeds west of the river are increased through use 

of multipliers in the WinMACCS model.
6-25

6.3.5 Unmitigated ISLOCA 

The emergency response timeline for the unmitigated ISLOCA scenario is shown in Figure 6-10. 

As shown in the figure, SAE is declared 15 minutes after the initiating event based on EAL 

FS1.1. A GE is declared 2 hours after the start of the event, the basis of which is discussed 

below. Sirens sound and EAS messages are broadcast approximately 45 minutes after declaration 

of GE. The durations of specific protective actions for each cohort are summarized in 

Figure 6-11.

Establishing the response timing for the Surry ISLOCA scenario required review of affected 

systems and expected operator actions. For this scenario, procedures would direct plant 

personnel to add water to the RWST. If water injection to the RWST is not possible, procedures 

identify use of a cross connect to the RWST of the unaffected unit. However, the boundary 

conditions for this accident scenario include failure to refill the RWST or crossconnect to the 

unaffected units RWST.

Within the bounds of the above constraint, staff discussed the response timing of this scenario 

with cognizant licensee technical staff. Following procedures, the licensee stated declaration of 

SAE would occur within about 15 minutes. Direct interpretation of response procedures for the 

declaration of a GE was not as straightforward.

The ISLOCA affects multiple plant systems, which require evaluation to determine whether the 

change in status requires declaration of a GE. For instance, MELCOR results indicate the 

reactor vessel water level approaches TAF in about 15 minutes, however, the water level is 

quickly recovered. The procedure calls for a GE if RVLIS indicates 46 percent (TAF), so a GE 

may not be warranted due to the brief uncovery of TAF. Another path to a GE would be 

radiation dose. Because the core is not exposed, radiation monitors may not show dose rates high 

enough to warrant declaration of a GE. Following the orange path of plant procedures, a GE 

would be declared when there is no injection or when core thermocouples indicate potential loss 

of fuel clad integrity. The orange path would be entered when the RWST is empty and water is 

no longer available for core injection. A GE would be declared based on a potential loss of fuel 

clad. As shown in Figure 5-111, ISLOCA RWST Inventory, the RWST empties in just over 

6 hours. 

From a system status perspective, the GE would be declared at 6 hours. However, emergency 

plan implementing procedures instruct operators to declare an emergency classification when it 

is determined that the emergency action level will be reached. In this case, when operators 

become aware that there will be no source of injection after the tank is empty, they would likely 

declare a GE. The timing of the declaration would be dependent upon the situation and the 

efforts to refill the tank or provide another source. Following declaration of an SAE, all licensee, 

offsite response organization and federal emergency response organizations and centers would 

be activated. These organizations would provide response support to the control room and 

execution of mitigation efforts. In particular, TSC management would provide onsite 

engineering and management resources to determine the likelihood of successful mitigative 

action. When those efforts are determined to be unsuccessful with the RWST level falling below 

50 percent capacity and continuing to fall, emergency response management would be expected 

to declare the GE. The licensee would use the ETE in support of a protective action 
6-26

recommendation. It was determined that a GE would be declared in about 2 hours. A response 

timeline was therefore developed assuming a GE would be declared at 2 hours. 

Figure 6-10 Unmitigated ISLOCA Emergency Response Timeline

Figure 6-11 Duration of Protective Actions - Unmitigated ISLOCA
6-27

The timing of these EALs and cohort actions are different than the scenarios where the initiating 

event is a station blackout. 

Cohort 1: 0 to 10 Public:

Cohort 1 is assumed to shelter when the sirens sound, and the time to receive the warning and 

prepare to mobilize is assumed to be 1 hour after the siren at which time this cohort begins to 

evacuate.

Cohort 2: 10 to 20 Shadow: 

This cohort is assumed to begin movement at the same time as the 0 to 10 Public after sirens 

have sounded within the EPZ and when widespread media broadcasts are underway. Residents 

in the 10 to 20 area begin seeing large numbers of people evacuating and initiate a shadow 

evacuation. There is no warning or notification for the public residing in this area, which is not 

under an evacuation order.

Cohort 3: 0 to 10 Schools:

Upon notification of the declaration of SAE by the site, the Virginia Department of Emergency 

Management would notify the schools promptly in accordance with the emergency response 

plan. This is a slowly developing accident and although it is assumed that notifications are 

prompt, extra time is included for drivers and buses to be mobilized. It is assumed schools begin 

evacuating 2 hours after start of the accident. 

Cohort 4: 0 to 10 Special Facilities:

The Special Facilities cohort is assumed to depart at the same time as the evacuation tail.

Cohort 5: 0 to 10 Tail:

The Tail evacuates 11 hours after notification to evacuate. 

Cohort 6: Non-evacuating public:

This cohort group represents the portion of the public who may refuse to evacuate and is 

assumed to be 0.5 percent of the population.

Table 6-8 provides a summary of the evacuation timing actions for each cohort.
6-28

Table 6-8 Unmitigated ISLOCA cohort timing

† Values represent speeds east of the James River. Speeds west of the river are increased through 

use of multipliers in the WinMACCS model.

Departure speeds and durations of the beginning and middle periods for the WinMACCS runs 

were developed from the Surry ETE study. Adjustments made to individual elements of the 

WinMACCS grid to reflect differences in vehicle direction and speed of travel through the 

network remained the same for this scenario. The timeline identifies the point at which it is 

assumed cohorts begin to take action. The actions taken by each cohort last for a given period as 

indicated in the timing table. 

6.4 Sensitivity Studies

Analysis of emergency preparedness and response parameters such as demographics, 

infrastructure, timing, etc., provide opportunity for further evaluation through sensitivity studies. 

The project team selected three additional calculations to assess variations in the implementation 

of protective actions. Each of the sensitivity studies was conducted using the ISLOCA accident 

scenario.

Sensitivity 1 - Evacuation of a 16 mile area including a shadow evacuation from within the 16 to 

20 mile area. 

Sensitivity 2 – Evacuation of the 0 to 20 mile area. 

Sensitivity 3 – Delay in implementation of protective actions for the public within the EPZ.

Sensitivities 1 and 2 assessed the effects of expanding the initial protective actions to distances 

of 16 and 20 miles respectively. The objective of this sensitivity analysis was to determine 

whether consequences might be reduced if the initial evacuation area was larger. Twenty miles 

was selected because it is twice the distance of the EPZ. A middle distance was also desired and 

Cohort

Delay to 

Shelter

DLTSHL 

(hr)

Delay to 

Evacuation

DLTEVA 

(hr)

Duration of 

beginning 

phase

DURBEG

(hr)

Duration of 

middle 

phase

DURMID

(hr)

Speed of 

early phase

ESPEED†

(early)

mph

Speed of 

middle phase

ESPEED†

(mid)

mph

0 to 10

Public 2.75 1 0.25 9.75 5 1

10 to 20 

Shadow 2.75 1 0.25 9.75 5 1

0 to 10

Schools 0.25 1.75 0.25 0.5 10 10

0 to 10

Special 

Facilities

2.75 11 1 1 1 10

0 to 10

Tail 2.75 11 1 1 1 10

Non-Evac NA NA NA NA 0 0
6-29

16 miles was selected because a ‘ring’ had been established in the underlying nodalization 

network in WinMACCS. A ring was not available at 15 miles.

A full scale evacuation model was developed to assess the sensitivity of consequences to changes 

in protective action strategies. Although the modeling of the area beyond the EPZ includes a full 

scale evacuation for the sensitivity analysis, this does not reflect likely protective action 

recommendations. To support the assessment of implementing protective actions beyond the 

EPZ, data was obtained for the 10 to 20 mile area around the NPP. Evacuation speeds for the 

area 10 to 20 miles beyond the EPZ were developed using OREMS Version 2.6. OREMS is a 

Windows-based application used to simulate traffic flow and was designed specifically for 

emergency evacuation modeling [25]. The main features of OREMS utilized in the analysis 

include:

• Determining the length of time associated with complete or partial evacuation of the 

population at risk within an emergency zone, or for specific sections of highway network 

or sub-zones;

• Determining potential congestion areas in terms of traffic operations within the 

emergency zone.

The OREMS model considers special conditions which may be imposed during an emergency 

evacuation. For example, intersections that normally have pre-timed controllers are assumed to 

be manned by emergency personnel to facilitate traffic flow [25]. This function is consistent 

with the emergency response actions that may be expected during an evacuation. Detail for 

roadway networks was obtained from aerial mapping and was input into OREMS using the 

standard intersection functions available in the model. Judgment and experience were necessary 

in determining the number of nodes that are established for the model. OREMS can manage 

hundreds of nodes, but there is a point at which the addition of nodes and links provides little 

change in the results. The nodal network established for the Surry plant is a moderately 

populated network for this code because about half of the EPZ is rural, southwest of the James 

River, and half is more urban, northeast of the river. With fewer nodes needed in rural areas, the 

mix of rural and urban areas results in a moderately populated nodal network.

The population values for the 10 to 20 mile area were developed using SECPOP 2000. A total of

171,182 vehicles were loaded onto 47 nodes distributed over 5 one-hour time periods. Vehicle 

data from the Surry 2001 ETE was also loaded onto the 10 to 20 mile area evacuation network 

consistent with the Surry ETE. The following evacuation times were produced from the 

OREMS calculation:

• 100 percent evacuation to a distance of 20 miles: 17 hours and 30 minutes; and

• 90 percent evacuation to a distance of 20 miles: 13 hours and 15 minutes. 

These times were used to develop the evacuation speeds for input into the WinMACCS model. 

The evacuation modeling conducted for the Surry plant was developed consistent with the 

characteristics observed in prior evacuations conducted for non-nuclear incidents. Most notably, 

the analysis includes the common phenomenon of evacuations in which travelers who depart the 
6-30

threat zone the earliest experience lower amounts of delay. This occurs because the routes have 

yet to become fully utilized during the emergency and the traffic volume and corresponding 

route congestion is generally lower. Evacuees who depart during the middle part of the 

evacuation, when the greatest number of people are seeking to depart, generally experience the 

highest amount of congestion and delay. This is because the demand on the roadway network is 

at its greatest, exceeding the available capacity in many areas. Evacuees who depart the hazard 

zone later, while potentially putting themselves at greater risk, enter the transportation network 

as the demand is near or even less than the roadway capacity. This means that this group, the 

tail, generally avoids the delays associated with the peak evacuation demand period. The 

OREMS output evacuation curve for the 20 mile evacuation is provided in Figure 6-12.

Figure 6-12 Evacuation Timeline from Surry for the 10 to 20 Mile Region

The initial accident scenarios were evaluated for protective actions within the EPZ. Expanding 

the protective actions to distances beyond the EPZ is not readily accommodated using the 

modeling approach selected for these analyses. For instance, although OROs may request the 10 

to 20 population shelter, this population group is treated within the modeling as performing 

normal activities throughout the emergency. The normal activity shielding factors are weighted 

averages of indoor and outdoor values based on being indoors 81 percent of the time and 

outdoors 19 percent of the time [66]. The hotspot and normal relocation model within MACCS2 

will move affected individuals out of the area if the dose criteria apply.

Surry

0%

10%

20%

30%

40%

50%

60%

70%

80%

90%

100%

0:00 2:24 4:48 7:12 9:36 12:00 14:24 16:48

Elapsed Time

% Evacuated

13:15

17:30
6-31

6.4.1 Sensitivity 1 ISLOCA 

For sensitivity 1, evacuation of a 16 mile area around the NPP is assessed. In addition, a shadow 

evacuation occurs from within the 16 to 20 mile area, and the remaining members of the public 

in the 16 to 20 mile area were assumed to shelter. Figure 6-13, Figure 6-14, and Table 6-9

summarize the cohort timing for sensitivity 1.

Figure 6-13 Unmitigated ISLOCA Timeline for Sensitivity 1

Figure 6-14 Duration of Protective Actions - Unmitigated ISLOCA Sensitivity 1
6-32

Cohort 1: 0 to 16 Public: 

Following declaration of a GE, sirens are sounded and an EAS message is broadcast that 

includes an evacuation order for affected areas within the EPZ. The public is assumed to shelter 

upon receipt of the EAS message, and the time to receive the warning and prepare to mobilize is 

assumed to be 1 hour. An assumption in this sensitivity analysis is the 10 to 16 public would be 

notified at the same time as the EPZ via EAS messaging and route alerting. The ETE for the 

public was estimated as a linear projection between the Surry 2001 ETE Study and the 10 to 20 

mile ETE developed for the sensitivity 2 analysis. 

Cohort 2: 16 to 20 Shadow:

This cohort is assumed to begin movement at the same time as the 0 to 16 Public after sirens 

have sounded within the EPZ, EAS messages are broadcast, and when widespread media 

broadcasts are underway. Residents in the 16 to 20 area begin seeing large numbers of people 

evacuating and initiate a shadow evacuation.

Cohort 3: 0 to 16 Schools: 

Schools are the first to take action. Upon receipt of the declaration of SAE by the site, the 

Virginia Department of Emergency Management would notify the schools within the EPZ in 

accordance with the emergency response plan. For this scenario, it is assumed schools within the 

10 mile to 16 mile area would evacuate beginning 2 hours after start of the accident. For this 

sensitivity study, it is assumed schools beyond the EPZ would decide, based upon media 

information, that it is prudent to evacuate or close schools immediately.

Cohort 4: 0 to 16 Special Facilities: 

For this sensitivity study, it is assumed Special Facilities beyond the EPZ would decide, based 

upon media information that it is prudent to evacuate. Special Facilities can take longer to 

evacuate than the general public because transportation resources must be mobilized, some of 

which are very specialized; therefore, the Special Facilities cohort is assumed to depart at the 

same time as the evacuation tail. 

Cohort 5: 0 to 16 Tail: 

An estimate of the departure for the evacuation tail is established as a linear projection between 

the Surry 2001 ETE Study and the OREMS 10 to 20 mile ETE developed for evacuation to a 

distance of 20 miles from the plant. 

Cohort 6: Non-evacuating public: 

This cohort group represents the portion of the 0 to 16 mile public who may refuse to evacuate 

and is assumed to be 0.5 percent of the population.
6-33

Table 6-9 Sensitivity Case 1 cohort timing

Cohort

Delay to 

Shelter

DLTSHL 

(hr)

Delay to 

Evacuation

DLTEVA 

(hr)

Duration of 

beginning 

phase

DURBEG

(hr)

Duration of 

middle 

phase

DURMID

(hr)

Speed of 

early phase

ESPEED†

(early)

mph

Speed of 

middle phase

ESPEED†

(mid)

mph

0 to 16

Public 2.75 1 0.25 10.75 5 1.5

16 to 20 

Shadow 2.75 1 0.25 10.75 5 1.5

0 to 16

Schools 0.25 1.75 0.25 0.5 10 10

0 to 16

Special 

Facilities

2.75 12 2 2 1.5 10

0 to 16

Tail 2.75 12 2 2 1.5 10

Non-Evac NA NA NA NA 0 0

† Values represent speeds east of the James River. Speeds west of the river are increased through use 

of multipliers in the WinMACCS model.

The timeline identifies the point at which it is assumed that cohorts begin to take action. The

actions taken by each cohort last for a given period as indicated in the timing table.

6.4.2 Sensitivity 2 ISLOCA

For Sensitivity Case 2, evacuation of a 20 mile area around the NPP is assessed. It is not 

expected that evacuation would be required beyond the EPZ; however, this sensitivity analysis

considers the possibility. Because the limit of the evacuation in this sensitivity analysis extends 

a considerable distance away from the plant, it was determined that adding a shadow evacuation 

beyond 20 miles would not be realistic. Therefore, no shadow evacuation is assumed in this

calculation. Figure 6-15, Figure 6-16, and Table 6-10 summarize the cohort timing for 

sensitivity 2. 

The WinMACCS model structure requires the first cohort to extend to the limits of the 

calculation, and for earlier calculations this was the limit of the shadow evacuation. In this 

sensitivity analysis, because there is no shadow evacuation, the limit of the first cohort is 

20 miles and represents the 0 to 20 public. 
6-34

Figure 6-15 Unmitigated ISLOCA Timeline for Sensitivity 2

Figure 6-16 Protective Action Durations - Unmitigated ISLOCA Sensitivity 2

Cohort 1: 0 to 10 Public: 

Cohort 1 is assumed to shelter when the sirens sound and the EAS message is broadcast. The 

time to receive the warning and prepare to mobilize is assumed to be 1 hour after the siren at 

which time this cohort begins to evacuate. 

Cohort 2: 10 to 20 Public: 

There are no sirens in the 10 to 20 mile area and no preplanned EAS messages; therefore, 

notification is assumed to be media broadcasts to residents in this area. The time to receive the 

warning and prepare to mobilize is still assumed to be 1 hour after the initial notification. The 

ETE for the 10 to 20 public was estimated using OREMS.
6-35

Cohort 3: 0 to 20 Schools: 

Upon receipt of the declaration of SAE by the site, the Virginia Department of Emergency 

Management would notify the schools within the EPZ in accordance with the emergency 

response plan. For this sensitivity study, it is assumed schools beyond the EPZ would decide, 

based upon media information, that it is prudent to evacuate or close schools. 

Cohort 4: 0 to 20 Special Facilities: 

For this sensitivity study, it is assumed Special Facilities beyond the EPZ would decide, based 

upon media information that it is prudent to evacuate. Special Facilities can take longer to 

evacuate than the general public because transportation resources must be mobilized, some of 

which are very specialized; therefore, the Special Facilities cohort is assumed to depart at the 

same time as the evacuation tail.

Cohort 5: 0 to 20 Tail: 

The ETE for the evacuation tail was estimated based on the OREMS analysis. 

Cohort 6: Non-evacuating public:

This cohort group represents the portion of the 0 to 20 mile public who may refuse to evacuate 

and is assumed to be 0.5 percent of the population.

Table 6-10 Sensitivity Case 2 cohort timing

Cohort

Delay to 

Shelter

DLTSHL 

(hr)

Delay to 

Evacuation

DLTEVA 

(hr)

Duration of 

beginning 

phase

DURBEG

(hr)

Duration of 

middle 

phase

DURMID

(hr)

Speed of 

early phase

ESPEED†

(early)

mph

Speed of 

middle phase

ESPEED†

(mid)

mph

0 to 10 

Public 2.75 1 0.25 12 5 1.6

10 to 20 

Public 2.75 1 0.25 12 5 1.6

0 to 20

Schools 0.25 1.75 0.25 0.5 10 10

0 to 20

Special 

Facilities

2.75 13.5 2 2 1.6 10

0 to 20

Tail 2.75 13.5 2 2 1.6 10

Non-Evac NA NA NA NA 0 0

† Values represent speeds east of the James River. Speeds west of the river are increased through 

use of multipliers in the WinMACCS model.

The timeline identifies the point at which it is assumed that cohorts begin to take action. The 

actions taken by each cohort last for a given period as indicated in the timing table.
6-36

6.4.3 Sensitivity 3 ISLOCA with Delay of Implementation of Protective Actions

There is a high level of confidence regarding the actions expected from control room operators in 

the event of accident scenarios identified for analysis in the SOARCA project. The initiating 

conditions provide clear indication to these operators and the response actions of the control 

room are prescribed. The Peer Review committee suggested that analysis consider a delay of the 

implementation of protective actions. Such a delay could be due to delay in control room 

declaration of an incident, delay in the decision process of OROs, delay in communication to the 

public regarding implementation of protective actions or other reasons. To address the potential 

for delay, an additional protective action timeline has been developed for the ISLOCA. This 

timeline reflects a delay in the implementation of protective actions by the public within the 

EPZ. Because protocols and procedures are in place, exercised and tested frequently, it is 

assumed that a delay of 30 minutes is adequate for this sensitivity study. Figure 6-17,

Figure 6-18, and Table 6-11 summarize cohort timing for sensitivity 3.

Figure 6-17 Unmitigated ISLOCA Timing for Sensitivity 3
6-37

Figure 6-18 Protective Action Durations - Unmitigated ISLOCA Sensitivity 3

The timing of these EALs and cohort actions are different than the scenarios where the initiating 

event is a station blackout. This sensitivity study includes a delay of 30 minutes in the 

implementation of protective actions. This delay is accounted for by delaying declaration of 

SAE by 30 minutes. 

Cohort 1: 0 to 10 Public:

Cohort 1 is assumed to shelter when the sirens sound and the initial EAS message is broadcast. 

The time to receive the warning and prepare to mobilize is assumed to be 1 hour after the receipt 

of the EAS message at which time this cohort begins to evacuate. 

Cohort 2: 10 to 20 Shadow:

This cohort is assumed to begin movement at the same time as the 0 to 10 Public after sirens 

have sounded within the EPZ and when widespread media broadcasts are underway. Residents 

in the 10 to 20 area begin seeing large numbers of people evacuating and initiate a shadow 

evacuation. There is no warning or notification for the public residing in this area which is not 

under an evacuation order.

Cohort 3: 0 to 10 Schools: 

Upon receipt of the declaration of SAE by the site, the Virginia Department of Emergency 

Management would notify the schools in accordance with the emergency response plan. Buses 

would be mobilized, and it is assumed schools begin evacuating 2.5 hours after start of the 

accident. 

Cohort 4: 0 to 10 Special Facilities:

The Special Facilities cohort is assumed to depart at the same time as the evacuation tail. 

Cohort 5: 0 to 10 Tail:

The Tail evacuates 11 hours after the public has been notified to evacuate. 

Cohort 6: Non-evacuating public:

This cohort group represents the portion of the public who may refuse to evacuate and is 

assumed to be 0.5 percent of the population.
6-38

Table 6-11 Sensitivity Case 3 cohort timing

Cohort

Delay to 

Shelter

DLTSHL 

(hr)

Delay to 

Evacuation

DLTEVA 

(hr)

Duration of 

beginning 

phase

DURBEG

(hr)

Duration of 

middle 

phase

DURMID

(hr)

Speed of 

early phase

ESPEED†

(early)

mph

Speed of 

middle phase

ESPEED†

(mid)

mph

0 to 10 

Public 3.25 1 0.25 9.75 5 1

10 to 20 

Public 3.25 1 0.25 9.75 5 1

0 to 10

Schools 0.75 1.75 0.25 0.5 10 10

0 to 10

Special 

Facilities

3.25 11 1 1 1 10

0 to 10

Tail 3.25 11 1 1 1 10

Non-Evac NA NA NA NA 0 0

† Values represent speeds east of the James River. Speeds west of the river are increased through use 

of multipliers in the WinMACCS model.

6.5 Analysis of Earthquake Impact

A seismic analysis was developed to assess the potential effects on local infrastructure, 

communications, and emergency response in the event of a large scale earthquake. The accident 

used in the earthquake analysis is the STSBO with TI-SGTR. Integrating the effects of the 

earthquake into the analysis required assessing the damage potential of the earthquake, 

identification of parameters that would be affected, and determining the adjusted values for 

affected parameters. 

The potential for an earthquake is largely identified by the occurrence of previous earthquakes in 

the region. Understanding of where earthquake faults exist in the eastern United States is not 

robust; whereas, in the west geological fault lines can be identified on the surface. Faults in the 

east are usually buried below layers of soil and rock and are not identifiable making prediction of 

earthquake location and magnitude difficult. The earthquakes hypothesized in SOARCA are 

assumed to be close to the plant site, and it may be assumed that severe damage is generally 

localized. Housing stock would generally survive the earthquake with some damage. The local 

electrical grid is assumed out of service due to the failure of lines, switch yard equipment, or 

other impacts. There is a backup power system for the sirens at Surry, and it is expected sirens 

would function. Under these postulated conditions, the potential for such an earthquake to affect 

emergency response and public evacuation is considered. 

6.5.1 Soils Review 

To approximate the extent of damage, an evaluation of the potential failure of infrastructure was 

conducted by NRC seismic experts to determine which, if any, roadways or bridges may fail 

under the postulated earthquake conditions. The assessment was performed using readily 

available information and professional judgment. Existing information on basic bedrock geology 

of the region was developed from reports and papers from the United States Geological Service. 
6-39

Soils of this region are formed from unconsolidated sediments deposited when the ocean level 

was much higher than at present. As sea levels lowered, many of these deposits were reworked 

by meandering rivers and streams that originated in the western part of the state and flowed to 

the east. In general, the closer to the coast, the nearer the water table is to the soil surface. Soils 

in the coastal plain are acidic, infertile, highly weathered, and vary from sandy textures to very 

clayey textures. Soil types are mostly silts and sands deposited in low energy environments, 

which make them potentially susceptible to liquefaction in an earthquake; however, site specific 

liquefaction potential is highly variable. 

Most landscapes are nearly level to gently sloping and because of this feature, the soils are not as 

susceptible to erosion. 

6.5.2 Infrastructure Analysis

The seismic evaluation of the potential failure of roadway infrastructure identified 40 bridges 

and roadway segments that could fail under the postulated conditions. The major areas where 

problems occur are in and around the urban area of Williamsburg, Virginia. Figure 6-19 shows 

an example of a bridge that could potentially fail under the earthquake conditions. Figure 6-20

shows the transportation network and the locations of the affected roadway segments and 

bridges. Table 6-12 provides a description of each of the roadway segments and bridges that 

could fail.

Figure 6-19 Highway 199 Over Highway 321

(Bridge 13 on Figure 6-20)
6-40

Figure 6-20 Roadway Network Identifying Potentially Affected Roadways and Bridges

The earthquake may cause structural damage in other areas of the EPZ. The structures within the 

EPZ are primarily light commercial and residential housing, both of which would largely be 

expected to stay intact. However, there are areas of larger commercial facilities and theme parks, 

which could sustain damage. The urban setting is also likely to experience localized fires caused 

by ruptured gas lines. 
6-41

Table 6-12 Description of the Potential Evacuation Failure Locations

Location Description

1 Small bridge on Highway 10 across pond

2 Overpass on Highway 10 near Smithfield

3 Bridge on Highway 10 near Smithfield

4 Stream culvert and boggy area under Highway 10

5 Road 617 (Whitemarsh Rd) culvert and bog beneath roadway

6 Highway 31 south of Highway 10- Bridge over pond

7 Road 621 (Burwell Bay Rd) pond and bog

8 Bridge over Route 5 (John Tyler Highway)

9 Bridge on Route 31 (Jamestown Rd)

10 Bridge over small lake on Route 31

11 Route 31 across small dam

12 Bridge over river on Highway 199-2 bridges in this area

13 Highway 199-overpass over Route 321 (Monticello Ave)

14 Highway 199-overpass over Longhill road 

15 Highway 199-overpass Old Towne Rd

16 Route 321 (Monticello Ave) adjacent to Lake Powhaton, Potential slope failure/slumping

17 Route 321 (Monticello Ave) bridge over bog

18 Highway 60 in Williamsburg-Overpass over rail tracks

19 Overpass US 199 over US 60 in Williamsburg

20 US 199/I-64 Interchange in Williamsburg

21 I-64/US 60 Interchange in Williamsburg

22 I-64 overpass near Williamsburg

23 I-64 Bridges over Colonial Parkway, edge of Williamsburg

24 I-64 Bridge over river and swampy area

25 I-64/Route 143 overpass

26 Route 143 ridge over river and swampy area

27 US 60-Dike over lake, southwest of Williamsburg

28 US 60 and SR 105 interchange

29 Small dam above US 60

30 I-64 overpass with SR 143 (Jefferson Ave)

31 I-64 overpass - Bland Ave.

32 I-64 Overpass, SR 173 (Denbigh Ave)

33 I-64 bridges over Industrial Park Drive

34 I-64 bridge over lake

35 I-64 and SR 238 (Old Williamsburg Rd) overpass

36 SR 105 (Ft. Eustis Blvd) bridge over lake

37 Colonial Historic Parkway- road on dam

38 Colonial Historic Parkway- bridge over creek

39 Highway 17 bridge across York River

40 Highway 17/258 across James River
6-42

6.5.3 Electrical and Communications 

There are many high voltage power lines traversing the Surry EPZ. It is assumed transformers 

and switchgear fail; however, power lines and related structures are assumed to not fail to a 

degree that they affect the emergency response (i.e., it is assumed that power lines do not fall 

across roadways potentially affecting evacuation routes). The siren system at Surry includes 

backup batteries, which would be sufficient to sound the sirens upon declaration of a GE. 

Loss of power limits the potential for residents to receive instructions via the Emergency Alert 

System (EAS) messaging. Televisions, household radios, and some telephones will not operate; 

although battery operated radios and car radios will operate. It is expected that the public will 

utilize these means of communication as well as societal forms, such as neighbor to neighbor,

propagating the EAS message throughout the EPZ. The alert and notification would be 

supplemented by route alerting conducted by OROs. This is a planned back up form of 

communication for the EPZ, and research shows this is effective and can be conducted in a 

timely manner. As observed in other large scale disasters, calls can inundate the emergency 

telecommunications systems with questions and requests for help, many of which will require 

emergency responder support [44]. This may cause cell phone service to be overloaded and may 

delay communications. However, for a localized event such as an earthquake, it is assumed that 

cell phone service is restored quickly.

The loss of power will affect traffic signalization within the EPZ. Typically, traffic signals 

default to flashing red in both directions in a power outage requiring all directions to stop prior to 

entering an intersection. This effectively turns signalized intersections into four-way stop signs. 

Four-way stop, as an intersection control, is less effective signalization for moving large 

numbers of vehicles, particularly when traffic is present on multiple approaches [46].

6.5.4 Emergency Response 

The assumption on the event timing is a mid-week winter day in which the public is at work and 

children are at school. The primary shift of emergency responders would be on duty and 

immediately available at the time of the incident. These initial priorities for emergency response 

personnel may delay implementation of traffic control to support an evacuation. It is expected 

that alerting the public would not be appreciably delayed, because Surry has backup batteries for 

the siren system. Route alerting may supplement alert and notification in localized areas.

During large scale emergencies, OROs routinely supplement staff with on-call and off-duty 

personnel. Although communications are assumed to be initially limited, radios are available to 

contact needed staff, and off-duty responders are expected to report for duty during such 

emergencies. By the time an evacuation is ordered, it is expected that OROs would have been 

augmented with additional staff; however, the effect on the infrastructure within the EPZ will 

require that OROs initially support activities that are protective of health and safety. 
6-43

6.5.4.1 Evacuation Time Estimate

Evacuation times are affected by bridges and roadways that fail, traffic signalization, and EAS 

messaging is not disseminated timely to inform evacuees of protective actions and evacuation 

routes. There are 40 locations identified as potential failures of infrastructure. Although 

evacuations are planned and conducted to move the public radially away from the NPP, 

evacuation following this postulated earthquake will be constrained to the few unobstructed

access routes out of the EPZ. West of the James River, the population is sparse and

infrastructure failures are relatively isolated. The result is a negligible affect on the evacuation

time for this area. However, the infrastructure failures north and east of the river will have a 

pronounced affect on the movement of vehicles and requires that an ETE for this area be 

developed to reflect the conditions. 

In review of the Surry 2001 ETE report, no significant traffic congestion was noted in the areas 

west of the James River. The ETE analysis does indicate significant traffic congestion in all 

zones that are north and east of the James River. The James River varies in width from about 1.5 

to 5.5 miles with only three crossings along a 50-mile stretch between Hopewell and Portsmouth. 

For this reason, the evacuation network north and east of the James River is essentially 

disconnected from the network west of the river. 

The major road system on the east side of the river, as discussed in the Surry 2001 ETE report, is 

oriented northwesterly and southeasterly, parallel with the peninsula. Primary evacuation routes 

east of the river include U.S. Highway 17, Interstate 64, U.S. Highway 60, and Route 143 and 

the Surry ETE report identifies all of these roadways as congested under evacuation conditions. 

These roadways are further affected by the seismic event. 

As indicated in Figure 6-20, 32 of the 40 affected roadways and bridges are located east of the 

James River and are clustered in and around Williamsburg, Virginia. Of particular importance is 

the major affect on Interstate 64 through this section of the EPZ. Most of the bridges and 

overpasses on this interstate are assumed to fail which causes some very difficult issues with an 

assessment of the evacuation time. To truly understand the effect of such damage on the ETE, 

the roadway system should be modeled. A basic vehicle/capacity approach was applied to 

develop a reasonable ETE and associated evacuation speeds. This assessment considers the 

timing and activities of residents, but may not fully account for factors such as driver confusion 

over which routes are accessible. 

This scenario is a mid-day mid-week event where the interstate can be assumed to be moderately 

traveled. A priority for emergency response personnel will be assisting those who are in life 

threatening conditions such as occupants of vehicles that are stranded on or under the sections of 

Interstate 64 between the failed bridge segments. The assistance in removing the vehicles has a

two fold effect of tying up emergency response personnel and creating additional congestion 

around the roadways leading to the interstate. 

Even before an SAE has been declared, the failure of the interstate bridges will affect traffic and 

cause a gridlock within the area. The interstate is unusable, and the underpasses to the interstate 

also become unusable which are major impediments to an evacuation. The significant failure of 

infrastructure causes the limiting factor of the ETE to be the queuing and loading of the 
6-44

evacuating vehicles at the points at which evacuation routes are available near the edge of the 

EPZ. At approximately the point at which Interstate 64 crosses the 10 mile EPZ boundary, there 

is no further damage to the Interstate. This is true at both the north and south ends of the EPZ,

and at this point Interstate 64 is available to support the evacuation with 3 lanes northbound and 

5 lanes southbound. 

Theoretical lane capacity for the Interstate would approximate 2,400 passenger cars per lane 

per hour; however, studies of lane capacity during planned evacuations, such as the evacuation in 

response to Hurricane Katrina, concluded that observed peak interstate flows were between 

1,350 and 1,500 passenger cars per lane per hour [43]. There are an estimated 90,000 evacuating 

vehicles, as developed from the Surry ETE report. Applying a rate of 1,500 passenger cars per 

lane per hour, the 8 available lanes would support all of the vehicles in approximately 7.5 hours; 

however, vehicles cannot simply access the interstate, so this simple Interstate capacity analysis 

only confirms that once on the Interstate, traffic will be in a free-flow state. 

The controlling point in the evacuation is the access capacity to Interstate 64. As vehicles travel 

to the available Interstate onramps, the lines of traffic, or queue, saturate the roadway network. 

To evaluate onramp capacity, the segment-flow density function is applied for oversaturated 

conditions following the procedure in Chapter 22 of the Highway Capacity Manual [46]. Using 

a saturated traffic density estimated at 160 passenger cars per mile per lane, which is appropriate 

for gridlock conditions, a corresponding onramp capacity of 500 passenger cars per hour per lane 

is obtained. Using aerial mapping, there are an estimated 10 onramps to Interstate 64 within a 

short distance of the EPZ boundary. Ten onramps with 500 passenger cars per hour per lane can 

load 90,000 in 18 hours. Because outbound capacity once on the Interstate is not a limiting 

factor and the loading points are beyond the limits of the EPZ, it is assumed that the time to load 

the vehicles is effectively the ETE.

This is a simplified approach to evacuation under the seismic conditions identified. However, 

for purposes of understanding the effects of protective actions, the estimate of 18 hours to 

complete the evacuation of areas east of the James River appears reasonable. Residents west of 

the river evacuate more quickly because there are few roadway impacts in this area, and speed 

adjustments were made in modeling this area. 

6.5.5 Development of WinMACCS parameters 

Traffic movement was approximated in each grid element by assigning a direction and speed for 

the vehicles within the grid. To account for the potential loss of bridges and roadway sections, 

the routing patterns in the WinMACCS model were adjusted to divert traffic around the locations 

identified. 

6.5.5.1 Relocation Outside the Evacuation Area

In the event of a significant release, the population in the region outside the evacuation area 

would be moved if their potential dose exceeds protective action criteria based on field 

measurements. The MACCS2 code uses hotspot and normal relocation, which is a dose based 

rather than distance based protective action. The values used in the earthquake analysis are the 

same as those used in the baseline analysis.
6-45

For hotspot relocation, individuals beyond twenty miles are relocated 24 hours after plume 

arrival if the total lifetime dose commitment for the weeklong emergency phase exceeds 0.05 Sv 

(5 rem). For the normal relocation, individuals are relocated 36 hours after plume arrival if the 

total lifetime dose commitment exceeds 0.01 Sv (1 rem). Review of the accident sequence 

timelines suggest that OROs would not be available earlier to assist with relocation due to higher 

priority tasks in the evacuation area. 

6.5.5.2 Shielding Factors

Shielding factors are the same as those used in the baseline analyses. It may be expected that the 

damage to structures caused by an earthquake of this magnitude would include broken windows 

and some structural damage. However, because residents within the seismic area are assumed to 

shelter only a short period of time, no adjustments in the modeling were made.

6.5.6 Seismic Analysis STSBO with TI-SGTR

The emergency response timeline for the unmitigated STSBO with TI-SGTR seismic scenario is 

shown in Figure 6-21. For this scenario, EAL SS1.1 specifies that if all offsite power and all 

onsite AC power is lost for greater than 15 minutes an SAE is declared. If restoration of power 

is not likely within 4 hours, EAL SG1.1 establishes that a GE be declared. It is expected the 

SAE is declared in about 15 minutes, and plant operators would recognize rather soon that 

restoration of power within 4 hours is unlikely. A 2 hour period from loss of power was selected 

as a reasonable time for declaration of a GE. It is expected that notification to OROs is timely 

and sounding of sirens occurs approximately 45 minutes after declaration of GE. From the 

MELCOR analysis, the first fission product gap release about three hours into the event with the 

SGTR occurring 30 minutes later. The duration of specific protective actions for each cohort is 

summarized in Figure 6-22.
6-46

Figure 6-21 STSBO with TI-SGTR Emergency Response Timeline (Seismic Scenario)

Figure 6-22 Protective Action Durations - STSBO with TI-SGTR (Seismic Scenario)

The timeline identifies points at which cohorts would receive instruction from OROs to 

implement protective actions. While protective actions within the EPZ can be modeled in 

accordance with procedures, assumptions were made that reasonably approximate those actions 

that could be taken due to the effects of the earthquake; however, the actual decisions made by 

OROs could differ.
6-47

The emergency response procedures for Surry provide for the sounding of sirens for a 

declaration of General Emergency. Backup batteries are available to support the sounding of 

sirens. It is assumed the large earthquake will be felt by everyone within the EPZ, and 

individuals will begin to prepare for an evacuation prior to receiving official notice. 

Cohort 1: 0 to 10 Public: 

The 0 to 10 Public is assumed to begin evacuating when the sirens sound. OROs would prepare 

and broadcast an EAS message, but the loss of power and infrastructure may limit the range of 

the broadcast. It is assumed that the effects of the earthquake are severe such that members of 

the public, knowing they live within an EPZ, begin preparations for evacuation shortly after the 

earthquake and are ready to leave when sirens sound.

Cohort 2: 10 to 20 Shadow: 

This cohort is assumed to begin movement at the same time as the 0 to 10 Public after sirens 

have sounded within the EPZ and when widespread media broadcasts are underway. Residents 

in the 10 to 20 area begin seeing large numbers of people evacuating and initiate a shadow 

evacuation. There is no warning or notification for the public residing in this area which is not 

under an evacuation order. It is assumed that the shadow population increases to 30 percent of 

the public in the area beyond the EPZ. 

Cohort 3: 0 to 10 Schools:

It is assumed schools take the initiative to prepare to evacuate prior to notification from the 

Virginia Department of Emergency Management. Buses would be mobilized, and it is assumed 

schools begin evacuating 30 minutes after the start of the incident; however, traffic congestion 

resulting from infrastructure failure causes a very slow evacuation speed. The analysis also 

considers that some drivers may not report due to an inability to get to the bus depot or need to 

address other immediate concerns. It is also assumed that given the magnitude of the 

earthquake, parents in the vicinity of the schools will pick up their children, reducing the need 

for a full complement of buses. In addition, it is expected schools will respond as needed and 

make do with the resources that arrive to evacuate the children in a single wave. This may 

include placing more than the normal 50 to 70 children on a bus and / or using school 

administrator’s and teacher’s vehicles to augment transportation needs. 

Cohort 4: 0 to 10 Shadow:

This cohort is assumed to begin movement first. They experience the earthquake and quickly 

begin to evacuate avoiding the congestion.

Cohort 5: 0 to 10 Special Facilities: 

The Special Facilities cohort is assumed to depart at the same time as the evacuation tail. Special 

Facilities need to have transportation resources mobilized, some of which are very specialized. 

Inbound lanes on roadways will be useable for emergency support vehicles, but traffic

congestion will delay the arrival of specialized vehicles. Special Facilities are assumed to leave 

at the same time as the evacuation tail; however, as discussed earlier, this is a simplification of 

the analysis because Special Facilities would realistically evacuate individually as resources are 

available.
6-48

Cohort 6: 0 to 10 Tail: 

The Tail takes longer to evacuate for many valid reasons such as the need to return home from 

work to evacuate with the family; the need to shut down farming or manufacturing operations 

prior to evacuating; and for the earthquake, the need to move rubble or other items prior to 

evacuating. However, with the extent of damage within the Surry EPZ, the tail simply becomes 

a continuous extension of the evacuating public. 

Cohort 7: Non-evacuating public: 

This cohort group represents the portion of the 0 to 10 public who may refuse to evacuate and is 

assumed to be 0.5 percent of the population.

Table 6-13 provides a summary of the evacuation timing for each cohort. The values in the table 

represent the minimum evacuation speeds corresponding to the area north and east of the site. In 

general, the cohorts in the seismic study have faster mobilization times but significantly slower 

evacuation speeds. The delay to shelter represents a delay before people get to the shelter, and 

delay to evacuation represents the length of the sheltering period prior to evacuation. These 

delays correspond to the different shielding factors that were applied to each cohort during these 

timeframes. Speed values are adjusted within each grid element of the WinMACCS model. 

Table 6-13 Cohort Timing STSBO with TI-SGTR Including Speeds

Cohort

Delay to 

Shelter

DLTSHL 

(hr)

Delay to 

Evacuation

DLTEVA 

(hr)

DURBEG

(hr)

DURMID

(hr)

ESPEED 

early

(mph)

ESPEED

middle

(mph)

E W E W

0-10 Public 2.75 0 0.25 15.75 5 15 0.7 2.1

10 to 20 

Shadow 2.75 0 0.25 15.75 5 15 0.7 2.1

0-10 

Schools 0.50 0 1.00 10.00 5 15 0.7 2.1

0 to 10 

Shadow 0.50 0 0.25 10.00 10.0 30 0.7 2.1

0-10 

Special 

Facilities

2.75 16 1.00 1.00 0.7 2.1 10.0 30.0

0-10 Tail 2.75 16 1.00 1.00 0.7 2.1 10.0 30.0

Non-evac NA NA NA NA NA NA NA NA

As indicated in Table 6-13, the evacuation speeds east of the James River are significantly

slower than the speeds west of the river. The actions taken by each cohort last for a given period 

as indicated in the timing table. To model the difference in speeds between the east and west 

sides of the river, the speeds were developed for the east side and the WinMACCS multipliers 

were applied to the west side. 
6-49

6.6 Accident Response and Mitigation of Source Terms

The Surry SOARCA study has concluded that scenarios can be mitigated by the licensee through 

the use of safety and security enhancements, including SAMGs and 10 CFR 50.54(hh) mitigation 

measures. Analyses were conducted of the consequences that may result if the onsite emergency 

response organization (ERO) is unable to prevent core damage and radiological release to the 

environment. It is expected that mitigative actions would be attempted. However, a human 

reliability assessment and a detailed seismic damage assessment were not performed for 

implementation of mitigative measures. The staff believed it appropriate to perform the 

consequence analyses to develop an understanding of core melt sequences, source term 

evolution, and offsite response dynamics to compare to previous studies. Perhaps the most 

important objective of these analyses is to quantify the benefit of mitigation enhancements. The 

following analysis describes an expected national level response to a severe nuclear power plant 

accident and provides a basis for truncating the release no later than 48 hours after the accident 

begins. Note that past studies, including PRAs such as NUREG-1150, typically truncated 

releases after 24 hours.

Mitigative actions taken during an accident are intended to:

• prevent the accident from progressing;

• terminate core damage if it begins;

• maintain the integrity of the containment as long as possible; and

• minimize the effects of offsite releases.

Response to a General Emergency would begin with the onsite ERO and would expand as 

needed to include utility corporate resources, State and local resources, and resources available 

from the Federal government, should these be necessary. If plant personnel efforts were 

unsuccessful, the national level response would provide resources to support mitigation of the 

source term. The discussion below presents a timeline for bringing resources onto the Surry site 

in order to flood containment to a level above a hypothetical hole in the reactor vessel to 

minimize the effects of an offsite release. Flooding containment would jeopardize vital 

equipment and monitoring capabilities and is considered a last resort option. Such an option 

would not begin until all other efforts have been exhausted. The approach described herein also 

discusses, to the extent practical at this time, the NRC Task Force review of the response to the 

nuclear power plant accidents at the Fukushima Dia-ichi facility in Japan.

On March 11, 2011, the Great East Japan Earthquake caused a large tsunami estimated to have 

exceeded 14 meters (45 feet) in height at the Fukushima Daiichi nuclear power plant site. The 

earthquake and tsunami produced widespread devastation across northeastern Japan, resulting in 

approximately 25,000 people dead or missing, displacing tens of thousands of people, and 

significantly impacting the infrastructure and industry in the northeastern coastal areas of Japan. 

The earthquake and tsunami caused accidents at Units 1, 2 and 3 of the Fukushima Daiichi 

nuclear power plant facility and caused concern for the remaining units and spent fuel pools at 

the site. Amid the vast devastation and competing health and safety priorities in the region, 

onsite and offsite response agencies worked diligently to bring the accidents under control. 

Shortly after the nuclear accident, the NRC established a task force to conduct a methodical and 

systematic review of the NRC’s processes and regulations to determine whether the agency 
6-50

should make additional improvements to its regulatory system. The Task Force report, 

“Recommendations for Enhancing Reactor Safety in the 21st Century,” [57] identified that 

prolonged SBO and multiunit events present challenges to Emergency Preparedness facilities 

that were not considered when the NRC issued NUREG – 0696, “Functional Criteria for 

Emergency Response Facilities,” in 1981. The Task Force report also states that an overarching 

lesson is that major damage to infrastructure in the area surrounding the plant might challenge an 

effective emergency response. A number of recommendations are presented in the report that 

address physical, administrative, and regulatory enhancements to further reduce the risk of 

similar challenges occurring among the US fleet of NPPs. As a state-of-art analyses, the 

SOARCA project included a degree of depth in the analyses beyond the scope of many previous 

studies. In this regard, some of the recommendations of the Task Force report had already been 

considered in SOARCA.

For instance, SOARCA investigated the challenges of potential damage to infrastructure within 

and beyond the EPZ as a result of an earthquake. The site specific analysis showed that 

40 bridges and/or roadway segments may fail. Review of the locations of the affected roadways 

showed that there would be a significant impact on the evacuation of the public from the area 

east of the James River, while the effects on infrastructure west of the river were shown to be 

relatively minor. The seismic analysis described in Section 6.5 quantified the offsite effects of 

challenges to both infrastructure and resources within the Surry EPZ. The Task Force report also 

recommends further enhancement of current capabilities for onsite emergency actions by 

requiring licensee’s modify the EOP technical guidelines to include SAMGs. This would 

enhance current capabilities, but would not change the manner in which SAMGs have already 

been considered in SOARCA.

With regard to the emergency response, the Surry analyses applied the site specific EALs. For 

Surry, an SAE would be declared within 15 minutes for all of the accident scenarios. A GE 

would be declared about 1 hour 45 minutes hours after SAE for all scenarios. Licensees are 

required by regulation to notify OROs within 15 minutes of declaring an emergency, and the 

OROs then initiate a planned response by offsite agencies who are able to direct necessary 

resources upon request, such as fire trucks, to support mitigation of the accident. The declaration 

by the licensee is not only a notification; this declaration initiates an ongoing communication 

between the control room, licensee staff, OROs, NRC, and other response agencies. 

As supported by the SOARCA analyses, it is shown that the accidents evaluated could be 

mitigated through the actions of the onsite and offsite response agencies. The evaluation of the 

mitigation of source term and truncation of the accident at 48 hours further expands upon the 

response resources through identification of corporate, State, local and Federal offsite resources. 

The responsibilities and resources of each of these organizations are described in onsite and 

offsite emergency response plans. These response organizations would mobilize upon request 

and as needed to support a severe nuclear power plant accident. These resources are in addition 

to the mitigative actions by the licensee through the use of safety and security enhancements, 

including SAMGs and 10 CFR 50.54(hh) mitigation measures. 

Although the response to the Fukushima Daiichi has taken much time and challenges still 

remain, it is expected that the regulatory structure, protocols, and resources available to support a 
6-51

response in the US are sufficient to mitigate the release of the accident scenarios identified in this 

study within 48 hours. Implementation of recommendations from the Task Force report would 

further improve the capabilities to mitigate the accident itself in a timely manner, to help prevent 

a release from occurring.

The NRC has onsite inspectors that are available to provide firsthand knowledge of accident 

conditions. NRC would be notified, following plant procedures, within one hour of the 

declaration; although, for a seismic event, it is likely that the NRC would be informed sooner 

than one hour by the onsite NRC resident inspectors, if not the licensee. Upon receipt of 

notification of the emergency classification, the NRC would activate the Headquarters 

Operations Center (HOC). NRC response teams reporting to the HOC include the Reactor 

Safety Team, Protective Measures Team, Executive Team, and other teams that support response 

related activities. Plant drawings and procedures are available in the HOC. Data that is typically 

communicated to the HOC via the Emergency Response Data System (ERDS), the Emergency 

Notification System bridge line, and other communication bridges, would likely be 

communicated via the resident inspectors satellite phones and via site battery operated systems. 

Licensees are required to provide guaranteed power to the emergency communications 

equipment per NRC Bulletin 80-15, “Possible Loss of Emergency Notification System (ENS) 

with Loss of Offsite Power.” The NRC region office would activate with similar response teams 

and would deploy a Site Team to the licensee’s EOF to support the response. A Site Team 

would include reactor safety engineers and protective measures specialists to review actions 

taken to mitigate the accident and to review protective action decisions that will be 

recommended to the public to assure the most appropriate actions are taken. Arrival of the Site 

Team may take several hours. The HOC, Regional Operations Center, and Site Team include 

liaisons to support coordination of resources when requested by the licensee.

Surry is part of the Dominion fleet which includes a remote EOF that would be activated and has 

access to fleet wide emergency response personnel and equipment, including equipment from 

sister plants following 10 CFR 50.54(hh) reactor security requirements to mitigate the effects of 

large fires and explosions. Significant resources would be made available to the site to mitigate 

the accident. The equipment would support multiple response needs and include such items as 

generators, pumps, compressed gas, etc. In addition to those directly involved in the incident 

and those agencies that fully test and exercise response plans, the Institute for Nuclear Power 

Operations and the Nuclear Energy Institute would activate their emergency response centers to 

assist the site. These efforts would provide knowledgeable personnel and an extensive array of 

equipment would be available and as such are considered in the decision to truncate the release at 

48 hours.

Concurrently with the NRC and industry response, the National Response Framework (NRF) 

establishes a coordinated response of national assets and would be implemented for an accident 

that progresses to a General Emergency. As described in the Nuclear/Radiological Incident 

Annex to the NRF, NRC is the Coordinating Agency for incidents occurring at NRC-licensed 

facilities. As Coordinating Agency, NRC has technical leadership for the Federal Government’s 

response to the incident. Under established agreement, if the severity of an emergency rises to 

the level of General Emergency, overall coordination of the incident would be conducted by the 

DHS. In this case, NRC retains the Federal technical leadership role but does not coordinate 
6-52

overall Federal response. Some of the other agencies cooperating in an incident include the 

EPA, FEMA, Health and Human Services, and any other Federal agency that may be needed. 

The assets of the Department of Energy (DOE) would be activated and brought to bear on the 

accident. Every licensee participates with many of these organizations in a full onsite and offsite 

exercise biennially. The NRC has an extensive, well-trained and exercised emergency response 

capability that would support, and under unusual circumstances, direct licensee efforts. 

Communications systems require battery backup in accordance with 10 CFR 50.54 Appendix E, 

and multiple communication bridge lines would be established to facilitate structured 

communication among the various response teams. Satellite phones, cell phones, radios, and 

other means are available for those instances where communications have been affected.

The above organizations would be developing onsite and offsite mitigation strategies with 

different objectives. These strategies would be implemented concurrently. An onsite mitigation 

strategy relies upon onsite resources and is expected to be immediate in order to prevent core 

melt and radiological release. Offsite mitigation strategies would bring national resources to the 

site and take more time to develop than onsite measures.

6.6.1 External Resources

The primary focus of the site and utility ERO would be mitigating core damage, and State and 

local resources would focus on the public evacuation. However, it is typical, as demonstrated in 

drills and exercises, for EROs to develop contingency plans in case initial onsite mitigative 

actions are not successful. The NRC ERO would focus on protection of the public and methods 

to reduce consequences reviewing the licensee and ORO information, actions, and decisions 

while performing independent analyses. If the site ERO is not successful with the onsite 

mitigative actions, as the unmitigated cases assume, various EROs would be considering in 

parallel the availability of portable power and pumping capacity from offsite locations. Virginia 

has a statewide mutual aid agreement for assistance from every fire department in the 

commonwealth. 

The Surry volunteer fire department and Rushmere volunteer fire department are both 15 miles 

from the site. The fire departments are in opposite directions, but they share a six mile stretch of 

road to the power plant. Figure 6-23 shows Surry and Rushmere Volunteer Fire Departments in 

relation to Surry Nuclear Power Plant. 
6-53

Figure 6-23 Surry and Rushmere Fire Departments

The Surry volunteer fire department practices annual response drills with Surry Nuclear Power 

Plant. This fire department has five pumper trucks, each with a capability to pump 500 to 

1000 gpm and has the ability to draft water directly from the river. A fire truck typically 

provides a pressure of up to 125 psi when limited by the firefighter controlling the nozzle. Fire 

hoses have a test pressure of at least 300 psi and a burst pressure of between 600 and 1000 psi. 

To obtain an Underwriter’s Laboratory (U.L.) Certification, a fire pump inside the truck must 

meet the following rated capacity shown in Table 6-14.

Table 6-14 Rated Flowrate of Fire Pumps

Additional resources are available through the State of Virginia Office of Emergency 

Management (OEM) which has an emergency services contract for the delivery of significant 

amounts of equipment within eight hours. OEM has access to high-pressure, high capacity 

pumps with 8 and 12 inch flanges that could readily move the water necessary for containment 

flooding. Such equipment is available for immediate State emergency use from many locations 

including Norfolk, Elizabeth City, and Raleigh.

The initiating event for the reactor accident is an earthquake in close proximity to the plant. This 

event causes significant ground motion and damage to certain types of structures. The impact to 

Pressure 

(psi)

Rated 

Flowrate

150 100%

165 100%

200 70%

250 50%
6-54

infrastructure, including the loss of some highway bridges within the EPZ, is described in 

Section 6.5. However, the loss of these bridges would not affect the ability to truck equipment 

into the site, because roads west of the James River were shown to be largely unaffected.

The EPZ environment is generally rural west of the river, and the roadway system is substantial. 

There are small bridges and culvert crossings on the plant access route that could be damaged, 

but these would not prevent passage by heavy trucks. A Virginia department of transportation 

area headquarters is 3 miles from the Surry fire department and has six dump trucks, two 

front-end loaders, a road grader, and other heavy equipment. Staff at the area headquarters have 

the capability to open roads and could immediately respond. Even if the culverts were to 

collapse, personnel at the transportation facility believe they could fix the road so that equipment 

could pass in less than 10 hours. In addition, Norfolk Naval Station has airlift resources that 

could be used to bring equipment into the plant. Fire trucks, which can weigh 35,000 lbs, are too 

heavy for air lift with MH-53 Helicopters which have a lift capacity of about 13,000 pounds. 

Airlift is not expected to be necessary, but if it was, the EROs would work together to identify 

appropriate pump and electrical equipment resources. The State has access to fire pumps and 

diesel generators that range from 15 kW to 15 MW and electrical generators are also available 

from commercial sources. These fire pumps and generators could be on site within about 

10 hours.

In addition, there are six fireboats in the Newport News/Norfolk area each with the pumping 

capacity of 1500 to 3000 gpm, and a marine fire fighting capability exists for three more boats 

equipped for firefighting, if needed. These fireboats have tremendous pumping capacity but may 

need to use long lengths of hose to support the site. It is likely that fire trucks and other portable 

equipment can be trucked onto the site more rapidly than the fire boats can be deployed, but they 

are a viable option should they be needed. EROs may pursue several options to ensure success. 

The timing of the hypothetical radiological release is scenario dependent, but when it occurs the 

site would be contaminated and working conditions more difficult. However, plant staff are 

trained in radiological work and the full staff of health physics technicians would be available. 

Within about 12 hours, staff from Dominion fleet plants could be at the site as well as staff from 

the DOE. Additional radiological technicians could likely be obtained from neighboring plants 

or perhaps from the Norfolk Naval Shipyard approximately 40 miles away should they be 

needed.

6.6.2 Mitigation Strategies

Strategies initially focus on injection of water into the reactor vessel. If all efforts in this regard 

fail, flooding of containment might be attempted to reduce the release potential through the 

cooling and scrubbing action of water. The site, utility and NRC EROs would identify various 

water injection methods corresponding to the damage circumstances. An effective option may 

be to inject water via the containment spray system. The two containment spray systems at 

Surry each have the capacity to inject more than 3,000 gpm into containment. This would have 

an immediate effect of lowering pressure and scrubbing radionuclides from the air and would 

suppress the release. The containment spray system has 8 inch flanges which can accept a milled 

flange and fire hose arrangement. The containment spray system has two sections: the 

containment spray (CS) subsystem and the recirculation spray (RS) subsystem. The recirculation 
6-55

spray subsystem also has two sections, inside and outside. Both the inside recirculation spray 

(ISRS) and the outside recirculation spray (OSRS) are designed for 100 percent capacity. 

There are several water sources but earthquake based scenarios assume that most water supply 

tanks fail. If nothing else is available, the intake or discharge canals would provide a water 

source. The site ERO would perform a damage survey to determine tank status and either find

usable tanks or make repairs. The utility would connect a water tank to a containment spray 

connection and make up to the water tank from the river if necessary. Local fire departments are 

experienced in drafting water directly from the river if needed. The availability of portable 

electrical power sources was discussed above, and portable power would likely be available at 

about the same time tank repairs and pumping capacity would be arranged. 

Mitigation of the ISLOCA scenario is different from the SBO scenarios in this volume because it 

is not seismically induced. There is no widespread onsite damage to prevent the use of normal 

systems and water sources. Actually, most ECCSs would be available to prevent core damage or 

truncate the release. As described in the site procedure key steps in Section 5.5.3, the utility has 

established at lease three proceduralized steps to align different water sources to one of many 

injection systems for truncating a release. 

In addition to the proceduralized steps described in Section 5.5.3, Surry has a portable 

low-pressure pump with a capacity of ~2000 gpm, and two portable high-pressure diesel-driven 

(Kerr) pumps. There is also a diesel-driven fire pump available at the Surry site. The utility has 

procedures to align installed pumps, portable pumps or the fire pump to the containment spray 

system and would use this strategy for the LTSBO and STSBO accident sequences. ISLOCA 

and STSBO TI-SGTR on the other hand are bypass events, so this strategy would not stop the 

release of radioactive material. However, connecting the fire pump to the containment spray 

system could be effective for any sequence after the molten core falls into the lower plenum or

breaches the vessel.

A high-pressure pump would be more effective because of the elevation of the containment spray 

nozzles and containment pressure. A high-pressure, high capacity fire pump (i.e., available from 

the State) connected to either the CS or the OSRS subsystems would immediately start to lower 

the pressure, scrub the volatiles, and eventually cover the molten core on the drywell floor. 

Figure 6-24 shows the analyzed pressure of the unmitigated STSBO accident.
6-56

Figure 6-24 Containment Pressure for Unmitigated STSBO

The ERO could also use a fire truck for pumping capacity. The utility has guidelines for using a 

fire truck to feed plant systems such as auxiliary feed water. However, to obtain a desirable flow 

rate in containment spray it would be necessary to gang fire trucks, and high capacity 

high-pressure portable pumps may be a better solution. 

Another method for release truncation would be to inject water directly into the primary system. 

This may be preferred because there is little elevation head to overcome and water would flow 

directly to the breach in the containment vessel and cover the molten core on the drywell floor 

scrubbing volatiles from the release. Containment pressure would continue to increase due to 

decay heat and the site ERO would need a means to remove heat (e.g., air coolers or heat 

exchangers via containment spray system). As necessary, all EROs would work together to 

identify other measures for mitigation interacting with plant personnel who know the plant well 

and may identify innovative solutions to inject water into containment. 

In Unit-2 the water cannot overflow into the reactor cavity until it reaches a level 25 feet from 

the bottom of the sump. In Unit-2 about 1.75 million gallons of water, as extrapolated from

Figure 6-25, would be necessary. Table 6-15 provides the data used to extrapolate the level from 

the bottom of the sump with respect to cumulative water volume.
6-57

Figure 6-25 Containment Water Level vs. Volume

Table 6-15 Cumulative Water Volume vs. Elevation for Unit 1

ELEVATION 

(sea level)

Distance from 

bottom of sump 

(ft)

Water 

volume 

(gal)

Water 

volume 

(ft3)

-29' 4 7/8" 0 0 0

-29' 4" 0.07 52 7

-29' 3" 0.16 108 14.4

-29' 3" 0.24 164 21.9

-29' 3" 0.41 277 37

-28' 7" 0.82 558 74.6

-28' 2" 1.24 965 129

-27' 10" 1.57 1,831 244.8

-27' 7" 1.82 9,113 1,218.2

-26' 7" 2.82 87,990 11,762.6

-25' 7" 3.82 166,867 22,306.9

-24' 7" 4.82 245,744 32,851.3

-23' 7" 5.82 324,620 43,395.5

-22' 7" 6.82 404,997 54,140.4

-21' 10 1/2" 7.53 461,933 61,751.6

-20' 1" 9.32 600,688 80,300.5

-10' ** 19.41 1,381,588 184,692.0

** Last data point extrapolated
6-58

The location of radiological release from containment may be expected in the vicinity of the 

equipment hatch, but the location of the increased leakage is not expected to depressurize the 

containment immediately. Therefore, in the unlikely event that the increased leakage is at or 

below the level of the water, the containment would still be capable of retaining water. In any 

case, if water is injected into containment through containment spray or primary injection it 

would suppress the release.

Figure 6-26 shows the Surry containment system elevations. Figure 6-27 shows the time it 

would take to pump 2 million gallons of water at various pumping capacities. At 3,000 gpm, it 

would take about 11 hours and at 6,000 gpm it would take about 5.5 hours.

Figure 6-26 Surry Containment System
6-59

Figure 6-27 Pumping Capacity

6.6.3 Truncation Summary

The types of resources needed are multi-use and would begin being acquired or established early 

in the response to support restoration of cooling, depressurizing the RCS, injecting into the RCS, 

etc. Based on the above, flooding of containment would need to begin about 40 hours after the 

start of the accident. This would allow 8 hours to fill the containment to an appropriate depth, 

which would require a pumping rate of about 4,000 gpm. It is reasonable to assume that within 

40 hours response personnel would be prepared to make a decision to flood the containment. 

The approach described provides a supporting basis for truncating the release at 48 hours. It is 

expected that onsite mitigative actions would limit core damage and reduce the release 

magnitude; however, this analysis is not definitive. Some of the actions identified are described 

within emergency response plans and some are ad hoc. The availability of the equipment is 

likely but not certain. If core damage is not prevented from the onsite mitigative measures, this 

truncation analysis demonstrates the offsite mitigative measures would likely take no more than 

48 hours from accident initiation to truncate the radiological release. Based on the approach 

provided, which considers in detail the timing of many activities underway during an event, it is 

reasonable to bound the truncation of the accidents at 48 hours. 

6.7 Emergency Preparedness Summary and Conclusions 

Advancements in consequence modeling provide an opportunity to integrate more detail and

realism in the application of protective action decisions applied for discrete population segments 

to represent implementation of protective actions. To best utilize these advancements, detailed 

information was developed and / or obtained from response of OROs to assure the quality of 

input data. Consequence modeling now provides for analysis of individual population segments 

and a user interface has been added to the consequence model to facilitate input detailed 

information that incorporates differences in the response to protective actions by various 

population segments. These advancements are significant because they now allow more detailed
6-60

modeling of response activities, timing of decisions, and implementation of protective actions 

across a wide range of population segments. 

Licensees develop ETEs to support emergency planning and help assure the most appropriate 

protective actions are implemented in an emergency. These ETEs provide detailed information 

regarding the evacuation of the general public, schools, special facilities and the evacuation tail. 

The improvements to consequence modeling and improved understanding of implementation of 

protective actions now allows use of this detailed information when modeling potential 

consequences of reactor accidents. For the first time, consequence modeling can represent the 

actions of OROs and the timing of public response to an emergency with a clearer and more 

defensible basis provided for the timing of these actions.

In this analysis, six cohorts were modeled for each of the accident scenarios and a seventh cohort 

was added for the seismic analysis. Protective action factors were applied to each specific 

cohort.

• For the general public, shielding factors appropriate for the region were applied during 

normal, sheltering, and evacuation and response times and speeds were developed from 

the Surry 2001 ETE.

• Schools are notified directly in accordance with offsite emergency response plans and 

buses are mobilized to support expedited evacuation of schoolchildren. Mobilizing 

school resources early allows the evacuation of schools to occur first, prior to roadways 

becoming congested from evacuation of the general public. Therefore, the speed of the 

school cohort was established based on relatively little traffic on the roadways at the 

time.

• Special facilities are also notified early, but respond quite differently than schools. 

Transportation resources for special facilities are quite specialized, can be limited, and 

typically take extra time to mobilize. Evacuation of these facilities starts later than 

schools and continues longer than the evacuation of the general public. This is because 

transportation resources take longer to mobilize and make return trips to evacuate each 

facility independently. A benefit of special facilities is the robust nature of the structures 

of these nursing homes, hospitals, etc. The shielding protection values are increased for 

these facilities. For this analysis, this cohort is sheltered until the point at which

evacuation begins which for calculation purposes was set at the same time as the 

evacuation tail begins.

• The evacuation tail was treated as a separate cohort and includes those members of the 

public who take longer to evacuate and are the last to leave the area. Indoor shielding 

values were applied to this cohort and it was evacuated late in the emergency moving at 

faster speeds because of the lower volume of traffic on the roadways at this time. The 

timing of the evacuation tail was developed from the ETE as the time when the last 

10 percent of the public begin to evacuate.
6-61

• A shadow evacuation, which occurs when people evacuate from areas that are not under 

an evacuation order, was represented in the area beyond the EPZ in the analysis to 

account for additional vehicles on the roadway network. Including a shadow evacuation 

adds realism to the analysis and allows consideration of the impact of the shadow 

evacuation vehicles on the evacuation of the EPZ.

• Consistent with NUREG-1150, evacuees received a dose until they traveled to a point 

10 miles beyond the analysis area, which for SOARCA was 30 miles.

• For the seismic analysis, it may be expected that a shadow evacuation of residents from 

within the EPZ may occur prior to the issuance of an evacuation order. This additional 

shadow evacuation was included in the analysis.

• A non-evacuating cohort was also included in the analysis assuming that a small 

percentage of the public may refuse to evacuate. Normal activity shielding values were 

applied to this cohort.

The Surry EALs were obtained for each of the accident scenarios modeled to reflect the timing 

of the declaration of SAE and GE. 

The following accident scenarios were modeled:

1. Unmitigated STSBO

2. Unmitigated STSBO with TI-SGTR

3. Mitigated STSBO with TI-SGTR

4. Unmitigated LTSBO

5. Unmitigated ISLOCA

6. Sensitivity 1, Unmitigated ISLOCA with evacuation to 16 miles

7. Sensitivity 2, Unmitigated ISLOCA with evacuation to 20 miles

8. Sensitivity 3, Unmitigated ISLOCA with a delay in implementation of protective actions

9. Seismic Unmitigated STSBO with TI-SGTR

For each of these accident scenarios, the specific EAL information and cohort movement was 

applied and the WinMACCS files were complied for the consequence analysis.

The sensitivity analyses were performed to identify differences when varying selected 

parameters. This included expanding the limits of the evacuation and adjusting the timing of 

implementation of protective actions. In the first sensitivity analysis, the limits of the evacuation 

were extended to 16 miles. In order to evaluate evacuation to this distance, an ETE was 

developed using OREMS. A second sensitivity analysis was conducted modeling evacuation to 

a distance of 20 miles from the plant. An ETE was developed using OREMS, and for this case, 

there was no shadow evacuation assumed due to the extreme distance from the point of the 

accident. 

An additional sensitivity analysis was conducted to assess the sensitivity of the timing of ORO 

decisions to consequences. This required increasing the delay times for cohorts to take action.
6-62

As supported by the SOARCA analyses, it is expected that the accidents evaluated would be 

mitigated through the actions of the onsite and offsite response agencies. The evaluation of the 

mitigation of source term and truncation of the accident at 48 hours further expands upon the 

response resources through identification of corporate, local, State, and Federal offsite resources. 

The responsibilities and resources of each of these organizations are described in onsite and 

offsite emergency response plans. These response organizations would mobilize upon request

and as needed to support a severe nuclear power plant accident. These resources are in addition 

to the mitigative actions by the licensee through the use of safety and security enhancements, 

including SAMGs and 10 CFR 50.54(hh) mitigation measures.

The parameters developed in this section provide input to the MACCS2 consequence model 

presented in Section 7.
7-1

7. OFF-SITE CONSEQUENCES

7.1 Introduction

The MACCS2 consequence model (Version 2.5) was used to calculate offsite doses and their 

effect on members of the public. MACCS2 was developed at Sandia National Laboratories for 

the NRC for use in PRAs for commercial nuclear reactors to simulate the impact of accidental 

atmospheric releases of radiological materials on humans and on the surrounding environment. 

The principal phenomena considered in MACCS2 are atmospheric transport using a straight-line 

Gaussian plume model, short-term and long-term dose accumulation through several pathways 

including cloudshine, groundshine, inhalation, deposition onto the skin, and food and water 

ingestion. The ingestion pathway was not treated in the analyses reported here because 

uncontaminated food and water supplies are abundant within the United States and it is unlikely 

that the public would consume radioactively contaminated food or water. The following doses 

are included in the reported risk metrics:

∗ Cloudshine during plume passage

∗ Groundshine during the emergency and long-term phases from deposited aerosols

∗ Inhalation during plume passage and following plume passage from resuspension of 

deposited aerosols. Resuspension is treated during both the emergency and long-term 

phases.

The SOARCA project made additional enhancements to MACCS2. In general, these 

enhancements reflect recommendations obtained during the SOARCA external review and also 

reflect needs identified by the broader consequence analysis community. The code 

enhancements done for SOARCA were primarily to improve fidelity and code performance, and

to enhance existing functionality. These enhancements are anticipated to have a significant effect 

on the fidelity of the analyses performed under the SOARCA project.

MACCS2 previously allowed up to three emergency-phase cohorts. Each emergency-phase 

cohort represents a fraction of the population who behave in a similar manner, although response 

times can be a function of radius. For example, a cohort might represent a fraction of the 

population who rapidly evacuate after officials instruct them to do so. To create a high fidelity 

model for SOARCA, the number of emergency-phase cohorts was increased, as described in

Section 6 of this report. This allowed significantly more variations in emergency response, 

(e.g., variations in preparation time prior to evacuation to more accurately reflect the movement 

of the public during an emergency). In a similar way, modeling evacuation routes using the 

network-evacuation model added a greater degree of realism than in previous analyses that used 

the simpler radial evacuation model.

7.2 Surry Source Terms

Brief descriptions of the source terms for the Surry accident scenarios are provided in Table 7-1. 

For comparison, the largest source term (SST1) from the 1982 NUREG/CR-2239, “Technical 
7-2

Guidance for Siting Criteria Development” (referred to hereafter as the ‘The 1982 Siting Study’ 

or just ‘The Siting Study’) [42] is also shown. Of the Surry source terms shown in the table, the 

unmitigated interfacing-systems loss-of-coolant accident (ISLOCA) is the largest in terms of 

release magnitude, but the release begins at 12.8 hours after accident initiation. Release begins 

earliest for the two thermally induced steam generator tube rupture scenarios (TISGTR), only 

3.6 hours after accident initiation, but the magnitudes are very small. The unmitigated STSBO 

and LTSBO scenarios begin very late in time and have very small release magnitudes. 

In comparison, the SST1 source term is significantly larger in magnitude, especially for the 

cesium class, than any of the Surry source terms. Moreover, it begins only 1.5 hours after 

accident initiation, about 2 hours earlier than the fastest release of the set of Surry source terms. 

The current understanding of accident progression has lead to a very different characterization of 

release signatures than was assumed for the 1982 Siting Study.

Table 7-1 Brief Source-Term Description for Unmitigated Surry Accident Scenarios

and the SST1 Source Term from the 1982 Siting Study

Scenario

Integral Release Fractions by Chemical Class

Atmospheric 

Release

Timing

CDF 

(Events/yr) Xe Cs Ba I

* Te Ru Mo Ce La

Start

(hr)

End

(hr)

Surry STSBO 2x10-6 0.518 0.001 0.000 0.006 0.006 0.000 0.000 0.000 0.000 25.5 48.0

Surry STSBO 

w/ TISGTR 4x10-7 0.592 0.004 0.000 0.009 0.007 0.000 0.001 0.000 0.000 3.6 48.0

Surry Mitigated 

STSBO w/

TISGTR

4x10-7 0.085 0.004 0.000 0.005 0.004 0.000 0.001 0.000 0.000 3.6 48.0

Surry LTSBO 2x10-5 0.537 0.000 0.000 0.003 0.006 0.000 0.000 0.000 0.000 45.3 72.0

Surry ISLOCA 3x10-8 0.983 0.020 0.000 0.154 0.132 0.000 0.003 0.000 0.000 12.8 48.0

SST1 1x10-5 1.000 0.670 0.070 0.450 0.640 0.050 0.050 0.009 0.009 1.5 3.5

∗ The Iodine release fraction for the ISLOCA is less than that presented in Section 5.5 (0.158) due to the plume 

segment limitations of MACCS2. A MACCS2 analysis is limited to a maximum of 200 plume segments. In 

order to fully capture the ISLOCA Iodine release fraction and retain the one-hour plume segment durations, it 

would have taken ~300 plume segments in the analysis. Thus, the 200 plume segments that represented the 

majority of the release were used.

For comparison purposes, a consequence analysis using the old SST1 source term is presented in 

this chapter. This allows a direct comparison, using the same modeling options and result 

metrics, of the SST1 source term and the current, best-estimate source terms. An attempt to 

replicate the results of the 1982 Siting Study and comparison with SOARCA results is also 

presented. 
7-3

7.3 Consequence Analyses

The results of the consequence analyses are presented in terms of risk to the public for each of 

the five accident scenarios identified for Surry. Both conditional and absolute risks are 

tabulated. The conditional risks are per reactor event. The absolute risks are likelihood of 

receiving a fatal cancer or early fatality for an average individual living within a specified radius 

of the plant per year of reactor operation from a potential plant accident.

The risk metrics are latent-cancer-fatality and early-fatality risks to residents in circular regions 

surrounding the plant. Population and economic data used in these analyses are projected for the 

year 2005. They are averaged over the entire residential population within the circular region. 

The risk values represent the predicted number of fatalities divided by the population for three

choices of dose-truncation level. These risk metrics account for the distribution of the 

population within the circular region and for the interplay between the population distribution 

and the wind rose probabilities. 

LCF risk results are presented for three dose-response assumptions: linear no threshold (LNT); 

US average natural background dose rate combined with average annual medical exposure as a 

dose truncation level (US BGR), which is 620 mrem/yr; and a dose truncation level based on the 

Health Physics Society’s Position [50] that there is a dose below which, due to uncertainties, a 

quantified risk should not be assigned (HPS), which is 5 rem/yr with a lifetime dose limit of 10 

rem. A 10 mrem/yr dose truncation level was investigated, but it produced results that were just 

slightly lower than with the LNT assumption and thus these results were not included in the final 

version of this document.

In addition to the base-case mitigated and unmitigated accident scenarios, several sensitivity 

analyses are reported in this chapter. A sensitivity analysis for the unmitigated STSBO scenario 

shows the influence of the size of the evacuation zone or a delay in evacuation on predicted risk. 

Another sensitivity analysis considers the effect of seismic activity on emergency response. This 

sensitivity is considered because the base case results account for the effect of the seismic event 

on the plant but do not account for its effects on evacuation. This sensitivity analysis takes the 

latter effects into account as well. A separate analysis of the SST1 source term (shown in

Table 7-1) allows older source-term assumptions to be compared with the current state-of-the-art 

methods for source term evaluation using otherwise equivalent assumptions and models. This 

analysis does not try to reproduce the 1982 Siting Study results; it merely overlays the older 

source term onto what are otherwise SOARCA assumptions for dose-response modeling, 

emergency response, and other factors. The final analyses show the relative contributions of 

each of the chemical classes included in the MELCOR source term analysis.

In this section, the risk tables represent rounded values obtained from the full data sets. The 

plots were developed from the full data sets and slight differences may be noticed due to this 

rounding. 

7.3.1 Unmitigated Long-Term Station Blackout 

The unmitigated LTSBO scenario is similar to the STSBO scenario except that cooling of the 

primary system is maintained until the batteries die, so degradation of the fuel and subsequent 
7-4

failure of the containment are delayed in the LTSBO scenario. As a result, the source term is 

later and smaller for this scenario than for the STSBO. In fact, the source term for this scenario 

begins more than 45 hr after accident initiation. This source term is also unique in that it was 

truncated at 72 hr rather than 48 hr, the time that was used for all of the other source terms. 

Table 7-2 displays the conditional, mean, latent-cancer-fatality (LCF) risks to residents within a 

set of concentric circular areas centered at the Surry site for the unmitigated LTSBO scenario. 

Three values of dose-truncation level are shown in the table: Linear, no threshold (LNT), i.e., a 

dose-truncation level of zero; the average, annual, US-background radiation (including average 

medical radiation) of 620 mrem/yr; and a dose-truncation level based on the Health Physics 

Society (HPS) Position [50], i.e., 5 rem/yr with a lifetime limit of 10 rem. 

The truncation based on the HPS Position is more complex than the others because it involves 

both annual and lifetime limits. According to the recommendation, annual doses below the 5 

rem truncation level do not need to be counted toward health effects; however, if the lifetime 

dose exceeds 10 rem, all annual doses, no matter how small, count toward health effects. 

Because of the 10 rem lifetime limit, risks predicted with the truncation based on the HPS 

Position can sometimes exceed those using the background radiation level for dose truncation.

Table 7-3 is analogous to Table 7-2, but displays absolute rather than conditional risks. In the 

case of the Surry LTSBO scenario, the mean CDF is 2x10-5 pry, a frequency that is based on the 

assumption that B.5.b mitigation does not succeed.

The values in Figure 7-3 are shown in Figure 7-1. The plot shows that for all dose-truncation 

levels, the risk is greatest for those closest to the plant and diminishes monotonically as distance 

increases. The trends shown in this figure are the same as those shown below in Figure 7-3 for 

the unmitigated STSBO. The NRC Safety Goal for latent cancer fatality risk from nuclear power 

plant operation (i.e., 2x10-6 or two in one million) which is shown in Figure 7-1and subsequent

figures is set 1,000 times lower than the sum of cancer fatality risks resulting from all other 

causes (i.e., 2x10-3 or two in one thousand). The calculated cancer fatality risks from the 

selected, important scenarios analyzed in SOARCA are thousands of times lower than the NRC 

Safety Goal and millions of times lower than the general U.S. cancer fatality risk.

Comparisons of SOARCA’s calculated LCF risks to the NRC Safety Goal and the average 

annual US cancer fatality risk from all causes are provided to give context that may help the 

reader to understand the contribution to cancer risks from these nuclear power plant accident 

scenarios. However, such comparisons have limitations for which the reader should be aware. 

Relative to the safety goal comparison, the safety goal is intended to encompass all accident 

scenarios. SOARCA does not examine all scenarios typically considered in a PRA, even though 

it includes the important scenarios. In fact, any analytical technique, including PRAs, will have 

inherent limitations of scope and method. As a result, comparison of SOARCA’s scenariospecific calculated LCF risks to the NRC Safety Goal is necessarily incomplete. However, it is 

intended to show that adding multiple scenarios’ low risk results in the ~ 10-10 range to 

approximate a summary risk from all scenarios, would yield a summary result that is also below 

the NRC Safety Goal of 2x10-6 or two in one million.
7-5

Table 7-2 Mean, Individual LCF Risk per Event (Dimensionless) for Residents within 

the Specified Radii of the Surry Site for the Unmitigated LTSBO for a Mean 

CDF of 2x10-5 pry

Radius of

Circular Area 

(mi)

LNT US BGR HPS

10 4.7E-05 4.0E-07 1.5E-09

20 2.6E-05 1.4E-07 4.1E-10

30 1.7E-05 7.8E-08 2.3E-10

40 1.1E-05 4.0E-08 1.2E-10

50 8.1E-06 2.7E-08 7.9E-11

Table 7-3 Mean, Individual LCF Risk per Reactor Year for Residents within the 

Specified Radii of the Surry Site for the Unmitigated LTSBO Scenario for a 

Mean CDF of 2x10-5 pry

Radius of

Circular Area 

(mi)

LNT US BGR HPS

10 7.1E-10 6.0E-12 2.2E-14

20 3.8E-10 2.1E-12 6.2E-15

30 2.6E-10 1.2E-12 3.4E-15

40 1.6E-10 6.0E-13 1.8E-15

50 1.2E-10 4.1E-13 1.2E-15

Figure 7-1 Mean, individual LCF risk per Reactor Year from the Surry unmitigated 

LTSBO scenario for residents within a circular area of specified radius from 

the plant for three values of dose-truncation level.

0.0E+0

1.0E-10

2.0E-10

3.0E-10

4.0E-10

5.0E-10

6.0E-10

7.0E-10

8.0E-10

10 20 30 40 50

Individual LCF Risk per Reactor Year 

(pry)

Radius of Circular Area (mi)

LNT

US BGR

HPS

Note: The NRC Safety Goal for LCF Risk is 2.0 x 10-6 pry
7-6

Figure 7-2 shows absolute LNT risks for the Surry unmitigated LTSBO for the emergency and 

long-term phases. The entire height of each column shows the combined (total) risk for the two 

phases. The figure shows that the emergency response is very effective within the EPZ and that 

the long-term phase dominates the overall risks. The habitability (i.e., return) criterion, which is 

implemented as a limit of 4 rem in the first 5 years after returning to live in a residential area, 

controls the overall risk to the public for this accident scenario. The trends shown in this figure 

are the same as those shown in Figure 7-4 for the unmitigated STSBO. In both cases, evacuees 

have ample time to evacuate before release begins.

All of the emergency-phase risk within the 10-mile EPZ is for the non-evacuating cohort. This is 

because all of the other cohorts avoid exposure to the plume. Thus, for this accident scenario, 

the residents within the EPZ who comply with the request to evacuate have no increased risk 

prior to the long-term phase. The peak emergency-phase risk is at 20 miles, which is the first 

location in the plot outside of the evacuation zone.

The prompt-fatality risks are identically zero for this accident scenario. This is because the 

release fractions (shown in Table 7-1) are too small to produce doses large enough to exceed the 

dose thresholds for early fatalities, even for the 0.5% of the population that does not evacuate.

Estimated risks below 1 × 10-7 per reactor year should be viewed with caution because of the 

potential impact of events not studied in the analyses and the inherent uncertainty in very small 

calculated numbers.

Figure 7-2 Mean, individual LNT, LCF risk per Reactor Year from the Surry 

unmitigated LTSBO scenario for residents within a circular area of specified 

radius from the plant for the emergency and long-term phases

0.0E+0

1.0E-10

2.0E-10

3.0E-10

4.0E-10

5.0E-10

6.0E-10

7.0E-10

8.0E-10

10 20 30 40 50

Individual LNT LCF Risk per Reactor 

Year (pry)

Radius of Circular Area (mi)

Long-Term Phase

Emergency Phase

Note: The NRC Safety Goal for LCF Risk is 2.0 x 10-6 pry
7-7

7.3.2 Unmitigated Short-Term Station Blackout 

Table 7-4 displays the conditional, mean, latent-cancer-fatality risks to residents within a set of 

concentric circular areas centered at the Surry site for the unmitigated STSBO scenario. 

Table 7-4 Mean, Individual, LCF Risk per Event (Dimensionless) for Residents within 

the Specified Radii of the Surry Site for the Unmitigated STSBO Scenario, 

for a Mean CDF of 1.5x10-6

/pry

Radius of

Circular Area 

(mi)

LNT US BGR HPS

10 9.4E-05 3.4E-06 1.4E-08

20 4.8E-05 1.5E-06 4.9E-09

30 3.2E-05 8.4E-07 2.7E-09

40 2.0E-05 4.3E-07 1.4E-09

50 1.5E-05 2.9E-07 9.4E-10

Table 7-5 is analogous to Table 7-4 but shows absolute rather than conditional risks. In the case 

of the Surry unmitigated STSBO, the mean CDF of 1.5x10-6 per year is used, a frequency that is 

based on the assumption that B.5.b mitigation does not succeed. These risks are shown 

graphically in Figure 7-3. The plot shows that for all dose-truncation levels, the risk is greatest 

for those closest to the plant and diminishes monotonically with distance.

Table 7-5 Mean, Individual, LCF Risk per Reactor Year for Residents within the 

Specified Radii of the Surry Site for the Unmitigated STSBO Scenario for a 

Mean CDF of 1.5x10-6 pry

Radius of

Circular Area 

(mi)

LNT US BGR HPS

10 1.4E-10 5.1E-12 2.1E-14

20 7.2E-11 2.3E-12 7.3E-15

30 4.8E-11 1.3E-12 4.0E-15

40 2.9E-11 6.5E-13 2.1E-15

50 2.2E-11 4.4E-13 1.4E-15
7-8

Figure 7-3 Mean, Individual, LCF Risk per Reactor Year from the Surry unmitigated 

STSBO scenario for residents within a circular area of specified radius from 

the plant for three values of dose-truncation level

Figure 7-4 shows the mean, individual, LNT, LCF risk for the Surry unmitigated STSBO for the 

emergency and the long-term phases. The height of each column indicates the combined (total) 

risk for the two phases. The figure shows that the emergency response is very effective within 

the EPZ, so those risks are very small and entirely represent the 0.5 percent of the population that 

does not evacuate. Thus, for this accident scenario, the residents within the EPZ who comply 

with the request to evacuate have no increased risk prior to the long-term phase. The peak in the 

emergency-phase risk is at 20 miles, which is the first location outside of the evacuation zone.

The long-term phase dominates the overall risks under the LNT assumption. The habitability 

(i.e., return) criterion, which is implemented as a limit of 4 rem in the first 5 years after returning 

to live in a residential area, controls the overall risk to the public. 

0.0E+0

2.0E-11

4.0E-11

6.0E-11

8.0E-11

1.0E-10

1.2E-10

1.4E-10

1.6E-10

10 20 30 40 50

Individual LCF Risk per Reactor Year 

(pry)

Radius of Circular Area (mi)

LNT

US BGR

HPS

Note: The NRC Safety Goal for LCF Risk is 2.0 x 10-6 pry
7-9

Figure 7-4 Mean, Individual, LNT, LCF Risk per Reactor Year from the Surry 

unmitigated STSBO scenario for residents within a circular area of specified 

radius from the plant for the emergency and long-term phases

The prompt-fatality risks are identically zero for this accident scenario. This is because the 

release fractions (shown in Table 7-1) are too small to produce doses large enough to exceed the 

dose thresholds for early fatalities, even for the 0.5% of the population that does not evacuate.

7.3.3 Unmitigated Short-Term Station Blackout with TI-SGTR

Table 7-6 displays the conditional, mean, latent-cancer-fatality risks to residents within a set of 

concentric circular areas centered at the Surry site for the unmitigated STSBO initiated,

thermally induced, steam-generator-tube rupture (TISGTR) scenario.

Table 7-7 is analogous to Table 7-6, but shows absolute rather than the conditional risks. In the 

case of the Surry unmitigated STSBO with TISGTR scenario, the mean CDF of 3.75x10-7 pry is 

used39. This frequency is based on the assumption that B.5.b mitigation does not succeed.

 39 The frequency of the Surry short-term station blackout is 1x10-6 to 2x10-6

/yr. The conditional probability of a 

thermally induced steam generator tube rupture is 0.1 to 0.4. The mean core damage frequency of 3.75x10-7

/yr

represents the product of the mid points of these two ranges, (i.e., 0.25*1.5x10-6

/yr).

0.0E+0

2.0E-11

4.0E-11

6.0E-11

8.0E-11

1.0E-10

1.2E-10

1.4E-10

1.6E-10

10 20 30 40 50

Individual LNT LCF Risk per Reactor 

Year (pry)

Radius of Circular Area (mi)

Long-Term Phase

Emergency Phase

Note: The NRC Safety Goal for LCF Risk is 2.0 x 10-6 pry
7-10

Table 7-6 Mean, Individual, LCF Risk per Event (Dimensionless) for Residents within 

the Specified Radii of the Surry Site for the Unmitigated STSBO with 

TISGTR With a Mean CDF of 4x10-7 pry

Radius of

Circular Area 

(mi)

LNT US BGR HPS

10 3.2E-04 7.4E-05 1.3E-05

20 1.9E-04 4.0E-05 4.5E-06

30 1.3E-04 2.5E-05 2.5E-06

40 8.4E-05 1.4E-05 1.3E-06

50 6.5E-05 9.9E-06 8.6E-07

Table 7-7 Mean, Individual, LCF Risk per Reactor Year for Residents within the 

Specified Radii of the Surry Site for the Unmitigated STSBO with TISGTR

With a Mean CDF of 4x10-7 pry

Radius of

Circular Area 

(mi)

LNT US BGR HPS

10 1.2E-10 2.8E-11 5.0E-12

20 7.2E-11 1.5E-11 1.7E-12

30 4.9E-11 9.5E-12 9.2E-13

40 3.2E-11 5.3E-12 4.7E-13

50 2.4E-11 3.7E-12 3.2E-13

The values in Figure 7-7 are shown in Figure 7-5. The plot shows that for all dose-truncation 

levels, the risk is greatest for those closest to the plant and diminishes monotonically with 

distance. The general trends in this figure are very similar to those shown in Figure 7-3 in the 

previous subsection. 

Figure 7-6 shows the LNT latent-cancer-fatality risks for the Surry unmitigated STSBO with 

TISGTR for the emergency and long-term phases. The figure shows that the emergency 

response does not entirely eliminate doses within the EPZ, but nonetheless, the risks are very 

small compared with the long-term phase risks. The doses received during the early phase stem 

from the relatively early release, which begins 3.6 hr after accident initiation (as shown in

Table 7-1). The habitability (i.e., return) criterion, which is implemented as a limit of 4 rem in 

the first 5 years after returning home, controls the overall risk to the public for this accident 

scenario. The general trends in this figure are very similar to those shown in Figure 7-4 in the 

previous subsection, with one exception. The risk from exposure during the emergency phase 

within a 10-mile radius is very small compared with the other distances shown in Figure 7-4; it is 

larger within a 10-mile radius than it is for the larger radii in Figure 7-6. The difference is 

directly related to the source-term characteristics, particularly the initiation of release, as 

discussed in the following paragraph.

Most of the emergency-phase risk within the 10-mile EPZ is for the evacuees. This is because 

release begins at 3.6 hr after accident initiation; the public begins to evacuate at 3.75 hr. Thus, 
7-11

some of the evacuees travel through the plume. By comparison, release begins 25.5 hours after 

accident initiation in the unmitigated STSBO discussed in the previous subsection while 

evacuation begins at the same time for both sequences.

The nonevacuating cohort represents 1.4% of the overall emergency-phase risk using the LNT 

hypothesis. This is a larger percentage of the overall risk than the population fraction 

represented by this cohort, which is 0.5%. This is expected, i.e., that the nonevacuating cohort 

should represent a greater risk than the cohorts that evacuate.

Figure 7-5 Mean, Individual, LCF Risk per Reactor Year from the Surry, unmitigated 

STSBO with TISGTR scenario for residents within a circular area of 

specified radius from the plant for three choices of dose-truncation level

0.0E+0

2.0E-11

4.0E-11

6.0E-11

8.0E-11

1.0E-10

1.2E-10

1.4E-10

10 20 30 40 50

Individual LCF Risk per Reactor Year 

(pry)

Radius of Circular Area (mi)

LNT

US BGR

HPS

Note: The NRC Safety Goal for LCF Risk is 2.0 x 10-6 pry
7-12

Figure 7-6 Mean, Individual, LNT, LCF Risk per Reactor Year from the Surry,

unmitigated, STSBO with TISGTR scenario for residents within a circular 

area of specified radius from the plant for the emergency and long-term 

phases.

The prompt-fatality risks are zero for this accident scenario. This is because the release fractions 

(shown in Table 7-1) are too small to produce doses large enough to exceed the dose thresholds 

for early fatalities, even for the 0.5% of the population that does not evacuate.

7.3.4 Mitigated Short-Term Station Blackout with TI-SGTR Rupture

Table 7-8 displays the conditional, mean, latent-cancer-fatality risks to residents within a set of 

concentric circular areas centered at the Surry site for the mitigated STSBO with TISGTR 

scenario. This scenario is similar to the one in the previous Section except that it is mitigated by 

operator actions to restore containment sprays. Because of the restored containment sprays, the 

risks are slightly lower than those shown in the previous subsection.

Table 7-9 is analogous to Table 7-8, but displays absolute rather than the conditional risk. In the 

case of the Surry mitigated STSBO with TISGTR scenario, the mean core damage frequency is 

4x10-7 pry a frequency that is based on the assumption that B.5.b mitigation does not succeed.

The values in this table are plotted in Figure 7-7. The trends are identical to those shown in the 

previous subsection for the unmitigated STSBO with TISGTR.

0.0E+0

2.0E-11

4.0E-11

6.0E-11

8.0E-11

1.0E-10

1.2E-10

1.4E-10

10 20 30 40 50

Individual LNT LCF Risk per Reactor 

Year (pry)

Radius of Circular Area (mi)

Long-Term Phase

Emergency Phase

Note: The NRC Safety Goal for LCF Risk is 2.0 x 10-6 pry
7-13

Table 7-8 Mean, Individual, LCF Risk per Event (Dimensionless) for Residents within 

the Specified Radii of the Surry Site for the Mitigated STSBO with TISGTR 

Scenario with a CDF of 4x10-7 pry

Radius of

Circular Area 

(mi)

LNT US BGR HPS

10 2.8E-04 7.1E-05 1.4E-05

20 1.7E-04 3.8E-05 4.5E-06

30 1.1E-04 2.4E-05 2.5E-06

40 7.3E-05 1.3E-05 1.3E-06

50 5.6E-05 9.3E-06 8.7E-07

Table 7-9 Mean, Individual, LCF Risk per Reactor-Year for Residents within the 

Specified Radii of the Surry Site for the Mitigated STSBO with TISGTR 

Scenario with a CDF of 4x10-7 pry

Radius of

Circular Area 

(mi)

LNT US BGR HPS

10 1.0E-10 2.7E-11 5.1E-12

20 6.2E-11 1.4E-11 1.7E-12

30 4.3E-11 8.9E-12 9.3E-13

40 2.7E-11 4.9E-12 4.8E-13

50 2.1E-11 3.5E-12 3.3E-13
7-14

Figure 7-7 Mean, Individual, LCF Risk per Reactor Year from the Surry, mitigated,

STSBO with TISGTR scenario for residents within a circular area of 

specified radius from the plant for three choices of dose-truncation level

Figure 7-8 shows the LCF risk for the Surry mitigated STSBO with TISGTR scenario for the 

emergency and long-term phases. The figure shows that the emergency response does not

entirely eliminate doses within the EPZ, but nonetheless, the risks are very small compared with 

the long-term risks. The doses received during the emergency phase stem from the relatively 

early release, which begins just 3.6 hr after accident initiation (cf., shown in Table 7-1). The 

habitability (i.e., return) criterion, which is implemented as a limit of 4 rem in the first 5 years 

after returning home, controls the overall risk to the public for this accident scenario. The trends

shown in Figure 7-8 are identical to those shown in Figure 7-6 for the unmitigated scenario, but 

the magnitudes are slightly smaller. 

0.0E+0

2.0E-11

4.0E-11

6.0E-11

8.0E-11

1.0E-10

1.2E-10

10 20 30 40 50

Individual LCF Risk per Reactor Year 

(pry) 

Radius of Circular Area (mi)

LNT

US BGR

HPS

Note: The NRC Safety Goal for LCF Risk is 2.0 x 10-6 pry
7-15

Figure 7-8 Mean, Individual, LNT, LCF Risk per Reactor Year from the Surry 

mitigated STSBO with TISGTR scenario for residents within a circular area 

of specified radius from the plant for the emergency and long-term phases

The prompt-fatality risks are zero for this accident scenario. This is because the release fractions 

(shown in Table 7-1) are too low to produce doses large enough to exceed the dose thresholds for 

early fatalities, even for the 0.5% of the population that does not evacuate.

Most of the emergency-phase risk within the 10-mile EPZ is for the evacuees. This is because 

release begins at 3.6 hr and, as a result, most of the evacuees are unable to avoid exposure to the 

plume. The nonevacuating cohort represents 1.1% of the overall emergency-phase risk using the 

LNT hypothesis. This is clearly a larger percentage of the overall risk than the population 

fraction represented by this cohort, which is 0.5%. It is expected that the nonevacuating cohort 

should have a greater risk than the cohorts who evacuate. 

7.3.5 Unmitigated ISLOCA

The unmitigated interfacing systems loss of coolant accident (ISLOCA) has the largest predicted 

release and it begins earlier than for the SBO scenarios without TISGTR. The release for this 

scenario begins at 12.8 hours after accident initiation. Emergency response is very effective and 

essentially no early fatalities are predicted to occur. However, predicted latent cancer fatality 

risk is larger than it is for the scenarios described in the previous subsections. 

Table 7-10 displays the conditional, mean, latent cancer fatality risks to residents within a set of 

concentric circular areas centered at the Surry site for the unmitigated ISLOCA scenario. 

0.0E+0

2.0E-11

4.0E-11

6.0E-11

8.0E-11

1.0E-10

1.2E-10

10 20 30 40 50

Individual LNT LCF Risk per Reactor 

Year (pry)

Radius of Circular Area (mi)

Long-Term Phase

Emergency Phase

Note: The NRC Safety Goal for LCF Risk is 2.0 x 10-6 pry
7-16

Table 7-11 is analogous to Table 7-10, but displays absolute rather than the conditional risks. In 

the case of the Surry unmitigated ISLOCA scenario, the mean CDF is 3x10-8/yr. This frequency 

is used to multiply the results in Table 7-10, as described above.

The values in Figure 7-11 are plotted in Figure 7-9. The plot shows that for all dose truncation 

levels, the risk is greatest at the 20-mile radius. This trend is due to evacuation in the 10-mile 

EPZ and only a partial shadow evacuation between 10 and 20 miles. Predicted LCF risk is larger 

in this scenario than it is in the scenarios described in the preceding subsections. 

Figure 7-10 shows the latent cancer fatality risks as a function of the radius from the plant for the 

emergency phase (EARLY), the long-term phase (CHRONC), and the combined phases (sum of 

the two). The figure shows that the emergency response does not entirely eliminate risks within 

the EPZ. This is because release begins at 12.8 hours after accident initiation, which is before 

evacuation is complete. Figure 6-10 shows that the public evacuates from 3 hr to 13 hr after 

accident initiation and the Special Facilities and Tail Cohorts only begin to evacuate at 13 hr 

after accident initiation. Therefore, there is a potential for exposure to the plume during 

evacuation for this accident scenario. This accounts for the emergency-phase risk within the 

10-mile EPZ shown in Figure 7-10.

Nonetheless, the long-term phase dominates the overall risks, even within the EPZ. The 

habitability (i.e., return) criterion, which is implemented as a limit of 4 rem in the first 5 years 

after returning home, controls the overall risk to the public for this accident scenario.

Most of the emergency-phase risk within the 10-mile EPZ is for the evacuees. The 

nonevacuating cohort represents 29.8% of the overall emergency-phase risk using the LNT 

hypothesis. This is more than an order-of-magnitude larger percentage of the overall risk than 

the population fraction represented by this cohort, which is 0.5%. This result is expected, that is, 

that the individual risk for the cohort that does not evacuate should be greater than for the 

cohorts who do evacuate.

Table 7-10 Mean, Individual LCF Risk per Event (Dimensionless) for Residents within 

the Specified Radii of the Surry Site for the Unmitigated ISLOCA Scenario

with a CDF of 3x10-8 pry

Radius of

Circular Area 

(mi)

LNT US BGR HPS

10 3.0E-04 6.5E-05 3.3E-05

20 3.4E-04 1.7E-04 1.5E-04

30 2.7E-04 1.4E-04 1.1E-04

40 2.0E-04 9.4E-05 6.6E-05

50 1.6E-04 7.5E-05 4.9E-05
7-17

Table 7-11 Mean, Individual, LCF Risk per Reactor-Year for Residents within the 

Specified Radii of the Surry Site for the Unmitigated ISLOCA Scenario with 

a CDF of 3x10-8 pry

Radius of

Circular Area 

(mi)

LNT US BGR HPS

10 9.1E-12 1.9E-12 1.0E-12

20 1.0E-11 5.2E-12 4.5E-12

30 8.1E-12 4.1E-12 3.2E-12

40 5.9E-12 2.8E-12 2.0E-12

50 4.9E-12 2.2E-12 1.5E-12

Figure 7-9 Mean, Individual, LCF Risk per Reactor Year from the Surry, unmitigated,

ISLOCA scenario for residents within a circular area of specified radius 

from the plant for three choices of dose truncation level

0.0E+0

2.0E-12

4.0E-12

6.0E-12

8.0E-12

1.0E-11

1.2E-11

10 20 30 40 50

Individual LCF Risk per Reactor Year 

(pry)

Radius of Circular Area (mi)

LNT

USBGR

HPS

Note: The NRC Safety Goal for LCF Risk is 2.0 x 10-6 pry
7-18

Figure 7-10 Mean, Individual, LNT, LCF Risk per Reactor Year from the Surry,

unmitigated, ISLOCA scenario for residents within a circular area of 

specified radius from the plant for the emergency and long-term phases

The prompt fatality risks are essentially zero for this accident scenario. The releases provide 

doses that are just above the threshold for early fatalities to the small fraction of the population 

that does not evacuate. There is no prompt fatality risk for the cohorts that evacuate. Conditional 

prompt fatality risks are shown in Table 7-12 as a function of distance from the plant. Absolute

prompt fatality risks are shown in Table 7-13. 

Table 7-12 Mean, Individual, Prompt Fatality Risk per Event (Dimensionless) for 

Residents within the Specified Radii of the Surry Site for the ISLOCA 

Scenario with a Mean CDF of 3x10-8 pry

Radius of

Circular Area 

(mi)

Prompt

Fatality Risk

1.3 1.5E-06

2.0 6.4E-07

2.5 4.0E-07

0.0E+0

2.0E-12

4.0E-12

6.0E-12

8.0E-12

1.0E-11

1.2E-11

10 20 30 40 50

Individual LNT LCF Risk per Reactor 

Year (pry) 

Radius of Circular Area (mi)

Long-Term Phase

Emergency Phase

Note: The NRC Safety Goal for LCF Risk is 2.0 x 10-6 pry
7-19

Table 7-13 Mean, Individual Prompt-Fatality Risk per Reactor-Year for Residents 

within the Specified Radii of the Surry Site for the ISLOCA Scenario with a 

Mean CDF of 3x10-8 pry

Radius of

Circular Area 

(mi)

Prompt

Fatality Risk

1.3 4.4E-14

2.0 1.9E-14

2.5 1.2E-14

The NRC quantitative health objective (QHO) for prompt fatalities is generally interpreted as the 

absolute risk within 1 mile of the exclusion area boundary. For Surry, the exclusion area 

boundary is 0.35 miles from the reactor building from which release occurs, so the outer 

boundary of this 1-mile zone is at 1.35 miles. The closest MACCS2 grid boundary to 1.35 miles 

used in this set of calculations is at 1.3 miles. Using the risk within 1.3 miles should reasonably 

approximate the risk within 1 mile of the exclusion area boundary. The absolute risk of a prompt 

fatality to an individual for this source term is approximately 4.4x10-14 per reactor year, which is 

well below the QHO. In fact, this risk is so low that for practical purposes it is zero.

7.3.6 Sensitivity Analyses on Size of the Evacuation Zone and Start Time

The baseline analyses included evacuation of the 10-mile EPZ, a partial shadow evacuation 

between 10 and 20 miles, and relocations of the remaining members of the public. For the 

unmitigated ISLOCA scenario, three additional calculations were performed to assess variations 

in the protective actions.

Sensitivity 1: Evacuation of a 16-Mile Circular Area:

In this calculation, the evacuation zone is expanded to 16 miles. Shadow evacuation occurs from 

within the 16- to 20-mile area. 

Sensitivity 2: Evacuation of a 20-Mile Circular Area:

In this calculation, the evacuation zone is expanded to 20 miles. No shadow evacuation beyond 

the evacuation zone is considered. 

Sensitivity 3: Delayed Evacuation of a 10-Mile Circular Area:

This calculation is identical to the baseline case described above except that implementation of 

protective action is delayed by 30 minutes.

The results of all three sensitivity analyses are compared with the base case in Table 7-14, which

shows that very little benefit results from increasing the size of the evacuation zone beyond the 

standard 10 miles. A delay in evacuation timing also has little impact on the results.
7-20

Table 7-14 Effect of Size of Evacuation Zone on Mean, LNT, LCF Risks for Residents 

within the Specified Radii of the Surry Site for the Unmitigated 

ISLOCA Scenario

Radius of

Circular Area 

(mi)

Base Case

10-Mile 

Evacuation

Sensitivity 1

16-Mile

Evacuation

Sensitivity 2

20-Mile 

Evacuation

Sensitivity 3

10-Mile Delayed

Protective Action

10 3.0E-04 3.9E-04 4.1E-04 3.2E-04

20 3.4E-04 2.7E-04 2.5E-04 3.5E-04

30 2.7E-04 2.3E-04 2.2E-04 2.7E-04

40 2.0E-04 1.8E-04 1.7E-04 2.0E-04

50 1.6E-04 1.5E-04 1.4E-04 1.6E-04

7.3.7 Evaluation of the Effect of the Seismic Activity on Emergency Response

The effects of seismic activity on emergency response are evaluated in this subsection for the 

unmitigated TISGTR scenario. Several impacts of the seismic activity are considered. One of 

these is the effect of collapsed bridges and impassible roadways on the evacuation itself, which is 

expected to increase risk. Another effect is on the size of the shadow evacuation, which is 

expected to decrease risk. The overall impact of the seismic activity on emergency response at 

the Surry site is insignificant, as shown in Table 7-15. Prompt fatality risk remains zero for this 

scenario. 

Table 7-15 Mean, Individual, LNT, LCF Risk per Event (Dimensionless) for Residents 

within the Specified Radii of the Surry Site for the unmitigated TISGTR 

Scenario and Comparing the Unmodified Emergency Response (ER) and ER 

Adjusted to Account for the Effect of Seismic Activity on Evacuation Routes 

and Human Response

Radius of

Circular Area 

(mi)

Unmodified 

ER

ER Adjusted 

for

Seismic Effects

10 3.2x10-4 3.3x10-4

20 1.9x10-4 1.9x10-4

30 1.3x10-4 1.3x10-4

40 8.4x10-5 8.4x10-5

50 6.5x10-5 6.5x10-5

7.3.8 Evaluation of SST1 Source Term

One of the differences between the SOARCA study and the 1982 Siting Study is the character of 

the radiological releases in terms of magnitude and timing. Because of this difference, it is 

useful to characterize and compare the risk to the public that derives from these releases. 

The approach used in this section is to substitute the SST1 source term for the SOARCA source 

term into the MACCS2 input files for the unmitigated STSBO with TISGTR scenario. These 
7-21

sensitivity analyses show the impact of the improvements made in the source term methods and 

practices on the consequence results.

The characteristics of the SST1 source term are described in the 1982 Siting Study report as 

follows:

• Severe core damage

• Essentially involves loss of all installed safety features

• Severe direct breach of containment

An exact scenario and containment failure mechanism (e.g., hydrogen detonation, direct 

containment heating, or alpha-mode failure) are not specified in the report. 

Notification time (i.e., sounding a siren to notify the public that a GE has been declared) for the 

unmitigated STSBO with TISGTR occurs at 2.75 hr, as shown in Figure 6-4. Declaration of a 

general emergency occurs at 2 hr and it takes an additional 45 min to notify the public. 

Notification of the public is thus after the beginning of release for the SST1 source term (cf.,

Table 7-1), which occurs 1.5 hr after accident initiation. Evacuation of the general public begins 

one hour after notification, or 3.75 hr after accident initiation. The start of evacuation here for 

this scenario is slightly earlier, but comparable to that for the largest segment of the population in 

the 1982 Siting Study, which occurred 4 hr after accident initiation. 

While the 1982 Siting Study treated emergency response very simplistically, a major emphasis of 

the SOARCA project is to treat all aspects of the consequence analysis as realistically as 

possible. No attempt was made in this sensitivity analysis to reproduce the treatment of 

emergency response used in the 1982 Siting Study.

Table 7-16 shows the latent-cancer-fatality risks for a release corresponding to the SST1 source 

term occurring at Surry based on the unmitigated STSBO with TISGTR ER. Table 7-17

compares the LNT risks with those for the unmitigated ISLOCA and the unmitigated STSBO 

with TISGTR scenarios discussed in preceding subsections. The LNT risk within 10 miles is 

about a factor of 30 higher than for the largest Surry source term considered in this study, which 

is for the ISLOCA. The 10-mile risk using a 620 mrem/yr dose truncation criterion is a factor of 

150 higher (cf., Table 7-10). At 50 miles the LNT risk is about a factor of 10 higher.

Table 7-16 Mean, Individual, LCF Risk per Event (Dimensionless) for Residents within 

the Specified Radii of the Surry Site for the SST1 Source Term from the 1982

Siting Study with All Parameters Other than for Source Terms Are Taken 

from the Unmitigated STSBO with TISGTR Scenario

Radius of

Circular Area 

(mi)

LNT US BGR HPS

10 9.1E-03 9.1E-03 9.0E-03

20 4.8E-03 4.8E-03 4.8E-03

30 3.1E-03 3.1E-03 3.1E-03

40 1.9E-03 1.9E-03 1.8E-03

50 1.4E-03 1.4E-03 1.3E-03
7-22

Table 7-17 Mean, Individual, LNT, LCF Risk per Event (Dimensionless) for Residents 

within the Specified Radii of the Surry Site for the SST1 Source Term from 

the 1982 Siting Study Using Emergency Response Parameters from the 

STSBO with TISGTR Scenario. Results are Compared with the Unmitigated 

ISLOCA and the Unmitigated STSBO with TISGTR Scenarios

Radius of

Circular Area 

(mi)

SST1 Using 

Unmitigated 

STSBO with 

TISGTR ER

Unmitigated 

ISLOCA

Unmitigated 

STSBO with 

TISGTR

10 9.1E-03 3.0E-04 3.2E-04

20 4.8E-03 3.4E-04 1.9E-04

30 3.1E-03 2.7E-04 1.3E-04

40 1.9E-03 2.0E-04 8.4E-05

50 1.4E-03 1.6E-04 6.5E-05

The maximum risk for the SST1 source term is within 10 miles, which is partially due to the fact 

that emergency response is not rapid enough to prevent exposures within the EPZ during the 

emergency phase. This is expected since release begins before notification of the public and, 

therefore, before evacuation begins. 

A notable feature of the risks presented in Table 7-16 is that the choice of dose truncation 

criterion has a minor influence on risk. This is very different from the SOARCA accident 

scenarios discussed in preceding subsections. Figure 7-11 provides some insights into this 

behavior. Figure 7-11 shows the absolute risk per reactor year for the population living near the 

Surry site, which accounts for the 10-5 pry frequency assigned to the SST1 source term. For the 

SST1 source term, nearly all of the risk, especially at short distances from the plant, is from 

exposures that occur during the emergency phase. Because a significant fraction of these doses 

are received over a short period of time and the doses are large due to the large source term, the 

level used for the dose truncation criterion has little influence on predicted risks. Again, this is a 

very different trend than is observed for the current, state-of-the-art source terms.

Table 7-18 shows the risk of prompt fatalities for several circular areas of specified radii 

centered at the plant. Unlike the source terms presented above, the predicted prompt-fatality 

risks are significantly greater than zero. Furthermore, the maximum distance at which prompt 

fatalities occur is more than 10 miles for this calculation. The SST1 release fractions are more 

than large enough to induce prompt fatalities for members of the public who live close to the 

plant and who do not evacuate quickly.
7-23

Figure 7-11 Mean, LNT, latent cancer fatality risks from the SST1 source term for 

residents within a circular area of specified radius from the Surry plant for 

the emergency and long-term phases.

Table 7-18 Mean, Individual, Prompt-Fatality Risk per Event (Dimensionless) for 

Residents within the Specified Radii of the Surry Site for the SST1 Source 

Term from the 1982 Siting Study Using Emergency Response Parameters 

from the Unmitigated STSBO with TISGTR Scenario

Radius of 

Circular Area 

(mi)

Prompt-Fatality Risk per 

Event Using STSBO with 

TISGTR ER 

(dimensionless)

1.3 1.3E-02

2.0 1.5E-02

2.5 1.1E-02

3.0 8.4E-03

3.5 5.4E-03

5.0 3.7E-04

7.0 5.0E-05

10.0 1.5E-05

The NRC QHO for prompt fatality risk is generally interpreted as the absolute risk within 1 mile 

of the exclusion area boundary. For Surry, the exclusion area boundary is 0.35 miles from the 

reactor building from which release occurs, so the outer boundary of this 1-mile zone is at 1.35 

miles. The closest MACCS2 grid boundary to 1.35 miles used in this set of calculations is at 1.3 

miles. Using the risk within 1.3 miles should reasonably approximate the risk within 1 mile of 
7-24

the exclusion area boundary. The frequency stated for the SST1 source term in the 1982 Siting 

Study is 10-5 pry, so the absolute risk of a prompt fatality to an individual for this source term is 

approximately 1.3x10-7 pry, which is well below the QHO. The actual risk of a prompt fatality 

(cf., Table 7-13), using current best-estimate practices for calculating source terms, is about five 

orders of magnitude lower than using the SST1 source term would imply (cf., Table 7-13 and 

Table 7-18). 

The acute-fatality risks presented in Table 7-18 are lower than the risks that would have been 

calculated in the 1982 Siting Study. There are two primary reasons for this difference. One is 

that 30% of the population within the EPZ is assumed to evacuate a full 6 hr after accident 

initiation in the 1982 Siting Study; here, 97.4% of the population within the EPZ begin to 

evacuate at least by 3.75 hr after accident initiation. A second reason is that the coefficients in 

the equations for acute health effects have been updated since the 1982 Siting Study based on 

more recent expert data [34]. The updated coefficients result in lower predicted acute fatalities 

across most of the exposure range for which these health effects can occur.

7.3.9 Surface Roughness

All of the SOARCA analyses presented above use a surface roughness length that represents a 

typical value for the US, which is 10 cm. This value was used in the 1982 Siting Study and in 

NUREG-1150 and has become a de facto default for most if not all license-related consequence 

analyses (e.g., severe accident mitigation alternatives (SAMA) analyses for license extension). 

However, this value of surface roughness is not necessarily the best choice for all regions of the 

country. In this section, we consider a site-specific value of surface roughness. 

The effect of increased surface roughness is twofold: It increases vertical mixing of the plume 

and it increases deposition velocities for all aerosol sizes. Both effects are treated in this 

sensitivity analysis and are discussed in the subsequent paragraphs.

Surry is located on the James River and the area surrounding the plant may be characterized by a 

mosaic of surface water, forests, farmland, suburban areas, and urban areas. The James River 

covers a large fraction of the EPZ. Land-use types correspond to a typical surface roughness or a 

range of surface roughness, some of which are shown in Table 7-19 [62].

Table 7-19 Surface roughness for various land-use categories for the area surrounding 

the Surry site

Land-Use Category Surface Roughness (cm)

Farmland recently plowed 1

Farmland with mature corn 10

Suburban housing 5 to 20

Suburban institutional buildings 70

Woodland forests 20 to 100

Urban Areas 100 to 600

Determining the best choice of surface roughness to represent the range of land-use categories is 

not a simple task. The value of 10 cm used in the base case is representative of the farmland and 
7-25

river that make up a significant fraction of the countryside surrounding the plant. Woodland 

forests and urban areas also make up a large fraction of the area and have a mean surface 

roughness of about 100 cm. An intermediate choice representing the average between cornfields 

and woodland forest, about 50 to 60 cm, might also have been a reasonable choice for this area. 

But the proximity of the surface features must also be considered. Because the river dominates 

the landscape and the urban areas are across the river, which is 2 to 3 miles wide in some areas, 

the lower value of 10 cm was chosen for this site. Additional discussion regarding surface 

roughness, including a sensitivity analyses, are provided in NUREG/CR-7110, Volume 1.

The effect on vertical mixing has traditionally been modeled by means of a multiplicative factor 

on vertical dispersion. The empirical expression for this factor is the ratio of surface roughness 

at the site in question to a standard value of surface roughness to the 1/5th power. Most of the 

data upon which empirical dispersion models have been based were taken at a site characterized 

by prairie grass [63], which was estimated to have a surface roughness of 3 cm. Thus, the 

empirical equation used to scale vertical dispersion uses the actual surface roughness divided by 

3 cm to the 1/5th power. The standard multiplicative factor corresponding to a 10-cm surface 

roughness is (10 / 3)0.2 = 1.27, which is the value used in all of the calculations presented above. 

A surface roughness length of 60 cm corresponds to a multiplicative factor of 1.82.

The effect of surface roughness on deposition velocity has been characterized by Bixler et al.,

and is based on expert elicitation data [64]. Bixler et al. provides a set of correlations for 

estimating deposition velocity as a function of aerosol diameter, wind speed, surface roughness, 

and percentile representing degree of belief by the experts. Here, we use the 50th percentile from 

the experts to get a best estimate deposition velocity. The 50th percentile correlation is as 

follows:

ln(vd) = −3.112 + 0.992 ∙ ln�dp� + 0.190 ∙ [ln�dp�]

2 − 0.072 ∙ [ln�dp�]3 + 5.922 ∙ z0

− 6.314 ∙ z0

2 + 0.169 ∙ v

where

vd = deposition velocity (cm/s)

dp = aerosol diameter (μm)

z0 = surface roughness (m)

v = mean wind speed (m/s)

Table 7-20 shows the aerosol deposition velocities calculated with the above equation that were 

used in this study for each aerosol bin in the MELCOR model. A mean wind speed of 2.2 m/s 

was used to obtain the results in the table. The column of deposition velocities corresponding to 

a surface roughness of 10 cm were used for all of the results shown in the preceding subsections. 

Increasing surface roughness from 10 to 60 cm roughly doubles the deposition velocity. The 

effect of surface roughness on deposition velocity was also investigated in the sensitivity study in 

NUREG/CR-7110, Vol. 1. 
7-26

Table 7-20 Deposition Velocities Used for the Base Case Calculations and for the 

Surface Roughness Sensitivity Study for Each of the Ten Aerosol Bins in the 

MELCOR Model

Mass Median 

Aerosol Diameter 

(μm)

Deposition Velocity (cm/s) for 

Specified Surface Roughness

10 cm 60 cm

0.15 0.053 0.11

0.29 0.049 0.10

0.53 0.064 0.14

0.99 0.11 0.23

1.8 0.21 0.45

3.4 0.43 0.92

6.4 0.84 1.8

11.9 1.4 2.9

22.1 1.7 3.7

41.2 1.7 3.7

7.3.10 Importance of Chemical Classes

Each isotope present in the core of a nuclear reactor contributes to the overall risk of an accident; 

however, the release of some isotopes contributes to risk much more than others. There are three 

reasons some isotopes are more important than others:

• Abundance of an isotope in the inventory in the core at the beginning of an accident,

• Release fraction of an isotope into the atmosphere, and

• The dose conversion factors for an isotope, which depends on the type and energy of the 

radiation produced, the half life of the isotope, and for internal pathways, the biokinetics 

of the isotope.

There are 69 isotopes in the treatment of consequences considered in the MACCS2 analysis, as 

described in Appendix B. These isotopes are grouped into a set of 9 chemical classes in the 

MELCOR analyses that generated the source terms used in the SOARCA analyses. Since 

release fractions are calculated by MELCOR at the level of chemical classes, it is both 

reasonable and useful to examine how these same chemical classes influence the evaluation of 

risk.

One approach to estimate the relative importance of each chemical class on risk is to release one 

chemical class at a time and evaluate the fraction of the overall risk that results, where overall 

risk is evaluated by releasing all chemical classes simultaneously. The problem with this 

approach is that the contributions from the individual chemical classes add up to more than the 

overall risk. The difference results from the amount of remedial action that is taken to reduce 

doses to the public. For example, much less remedial action is taken when doses are small, 
7-27

which may be the case when only one chemical class is released at a time, than when doses are 

large. Because less remedial action is taken, the contribution of an individual chemical group to 

risk is greater when it is released on its own than when it is part of a larger release. To make the 

fractional contributions from individual chemical classes add to unity, the contribution from a 

single chemical class must be normalized by the sum of the individual contributions of the 

chemical classes rather than the risk calculated for the combined effect of all chemical classes. 

This inherent nonlinearity tends to diminish the effect of the major contributors and exaggerate 

the effect of the minor contributors.

To minimize the effect of the nonlinearities described in the previous paragraph, an alternative 

approach is adopted here. That is to evaluate the contribution of a chemical class by performing 

calculations with all but that one chemical class. The effect of that chemical class is then 

calculated by taking the difference between the risk when all chemical classes are included and 

the risk for all but that one chemical class (i.e., setting the release fractions for that chemical 

class to zero).

The relative importance of each chemical class was evaluated for the unmitigated ISLOCA 

accident sequence, for each dose truncation level: LNT, US BGR, and HPS. The results for the 

population within 10 miles are shown in Figure 7-12, Figure 7-13, and Figure 7-14. Results at 

longer distances are shown in subsequent figures.

The first of these, Figure 7-12, is for LNT for the population within 10 miles. It shows the 

importance of each chemical group on total risk, on just the emergency-phase risk, and on just 

the long-term-phase risk. The cesium group dominates the total risk and the long-term phase 

risk, but contributes only a few percent to the emergency-phase risk owing to the relatively long 

half lives of the cesium isotopes (e.g., 137Cs has a half life of 30 yrs). Tellurium and iodine 

contribute most of the emergency-phase risk owing to the short half-lives of the isotopes 

represented by these chemical classes. However, the emergency phase contributes very little to 

the total risk because 99.5% of the population within 10 miles evacuate and do not receive any 

dose during the emergency phase.
7-28

Figure 7-12 Percentage contribution to total, emergency-phase, and long-term-phase, 

mean, individual risk for the population within 10 miles by chemical class for 

the Surry unmitigated ISLOCA based on the LNT hypothesis

Figure 7-13 and Figure 7-14 show the total risk contributions of each chemical class for the 

unmitigated ISLOCA using US BGR dose truncation and truncation based on the HPS Position

[50], respectively. These plots also show risk to the population living within 10 miles of the 

plant. They only show the total risk contribution because annual doses in the first year are 

combinations of emergency- and long-term-phase doses. Because of the overlapping 

contributions to the first year, the individual contributions of the two phases cannot be easily 

deconvolved from the whole. These figures show that the tellurium, cesium, and iodine 

chemical classes contribute most of the risk for these dose truncation criteria, with the same 

order of importance in the two figures. Isotopes with relatively short half-lives tend to be more 

dominant than those with longer half-lives because most of the risk is from doses received during 

the first year for the US BGR and truncation based on the HPS Position. Longer-term annual 

doses are limited by the habitability criterion to values below the dose truncation levels. 

0%

10%

20%

30%

40%

50%

60%

70%

80%

90%

100%

Xe Cs Ba/Sr I Te Ru Mo Ce La

Percentage of Risk

Radionuclide Group

Total

Emergency Phase

Long-Term Phase
7-29

Figure 7-13 Percentage contribution to total, mean, individual risk for the population 

within 10 miles by chemical class for the Surry unmitigated ISLOCA based 

on US BGR dose truncation

Figure 7-14 Percentage contribution to total, mean, individual risk for the population 

within 10 miles by chemical class for the Surry unmitigated ISLOCA based 

on a truncation level reflecting the HPS Position for quantifying health 

effects

0%

10%

20%

30%

40%

50%

60%

Xe Cs Ba/Sr I Te Ru Mo Ce La

Percentage of Risk Contribution

Radionuclide Group

0%

10%

20%

30%

40%

50%

Xe Cs Ba/Sr I Te Ru Mo Ce La

Percentage of Risk Contribution

Radionuclide Group
7-30

Figure 7-15 through Figure 7-20 are analogous to those above but show the relative importance 

of the chemical classes for the population within 20 and 50 miles. The trends are similar, but the 

emergency phase plays a larger role because significant portions of the population do not 

evacuate before the plume arrives and, thereby, receive a dose during the emergency phase. The 

most important set of chemical classes using the LNT hypothesis is cesium, tellurium, and iodine 

in that order. For the two dose truncation criteria, cesium is less important because of the 

relatively long half-lives of the dominant isotopes. 

Figure 7-15 Percentage contribution to total, emergency-phase, and long-term-phase, 

mean, individual risk for the population within 20 miles by chemical class for 

the Surry unmitigated ISLOCA based on LNT hypothesis

0%

10%

20%

30%

40%

50%

60%

70%

80%

90%

100%

Xe Cs Ba/Sr I Te Ru Mo Ce La

Percentage of Risk

Radionuclide Group

Total

Emergency Phase

Long-Term Phase
7-31

Figure 7-16 Percentage contribution to total, mean, individual risk for the population 

within 20 miles by chemical class for the Surry unmitigated ISLOCA based 

on US BGR dose truncation

Figure 7-17 Percentage of contribution to total, mean, individual risk for the population 

within 20 miles by chemical class for the Surry unmitigated ISLOCA based 

on a truncation level reflecting the HPS Position for quantifying health 

effects

0%

10%

20%

30%

40%

50%

60%

70%

80%

Xe Cs Ba/Sr I Te Ru Mo Ce La

Percentage of Risk Contribution

Radionuclide Group

0%

10%

20%

30%

40%

50%

60%

70%

80%

Xe Cs Ba/Sr I Te Ru Mo Ce La

Percentage of Risk Contribution

Radionuclide Group
7-32

Figure 7-18 Percentage contribution to total, emergency-phase, and long-term-phase, 

mean, individual risk for the population within 50 miles by chemical class for 

the Surry unmitigated ISLOCA based on the LNT hypothesis

Figure 7-19 Percentage contribution to total, mean, individual risk for the population 

within 50 miles by chemical class for the Surry unmitigated ISLOCA based 

on US BGR dose truncation

0%

10%

20%

30%

40%

50%

60%

70%

80%

90%

100%

Xe Cs Ba/Sr I Te Ru Mo Ce La

Percentage of Risk

Radionuclide Group

Total

Emergency Phase

Long-Term Phase

0%

10%

20%

30%

40%

50%

60%

70%

Xe Cs Ba/Sr I Te Ru Mo Ce La

Percentage of Risk Contribution

Radionuclide Group
7-33

Figure 7-20 Percentage contribution to total, mean, individual risk for the population 

within 50 miles by chemical class for the Surry unmitigated ISLOCA based 

on a truncation level reflecting the HPS Position for quantifying health 

effects

0%

10%

20%

30%

40%

50%

60%

70%

Xe Cs Ba/Sr I Te Ru Mo Ce La

Percentage of Risk Contribution

Radionuclide Group

8-1

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8-2

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8-3

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Model Test,” NUREG/CR–5671, 1998.
8-4

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D.C.: NRC. January 2005. 

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Action Recommendations for Severe Accidents." Washington D.C.: NRC. December 

2007. 

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September 30,2004

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SAND2008-4195P. "Review of NUREG-0654, Supplement 3, “ Criteria for Protective 

Action Recommendations for Severe Accidents: Focus Groups and Telephone Survey" 

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8-5

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Licensing Evaluations, ORNL/TM-2005/39, Version 5.1, Vols. I-III, November 2006.

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Release Frequency Probabilistic Risk Assessment for Nuclear Power Plant Applications,” 

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– The Near-Term Task Force Review of Insights From the Fukushima Dai-Ichi 

Accident,” U.S. Nuclear Regulatory Commission, Washington D.C., July 2011.

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Scinta, Inc. 2421 Homestead Drive, Silver Spring Md. 20902, 1992.

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1996.

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Coolant System Under Severe Accident Conditions, NUREG/CR-6131, December 1998.

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NP-6094-D, November 1988.

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III, Report AFCRC-TR-58-235, Air Force Cambridge Research Center. 1959.
8-6

[64] Harper, F. T., et al., “Probabilistic Accident Consequence Uncertainty Analysis, 

Dispersion and Deposition Uncertainty Analysis,” NUREG/CR-6244, U.S. Nuclear 

Regulatory Commission: Washington DC. 1994.

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Development of Evacuation Time Estimate Studies.” Washington D.C. NRC. (NRC, 

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4th edition. 2012.
APPENDIX A

SURRY CONTAINMENT PERFORMANCE

A-1

APPENDIX A: SURRY CONTAINMENT PERFORMANCE

1.0 INTRODUCTION

Containment performance at beyond design basis accident internal pressure and temperature is 

required as an input for determining the offsite consequences and accident progression of a 

nuclear power plant during a severe accident. This appendix documents the analysis and 

assessment of Surry Nuclear Power Plant (SNPP) containment at beyond design basis internal 

pressures and temperatures developed during a severe accident. The design-specific SNPP 

containment failure pressure, leakage area, and leakage location as documented is used as an 

input for the State of-the-Art Reactor Consequences Analysis (SOARCA) of the SNPP. 

2.0 APPROACH

Extensive research and scale model testing of reinforced and prestressed concrete containments 

to determine behavior at beyond design basis accident pressure has been performed in the last 

25 years at SNL [2] and CEGB [3]. Concrete containments start to leak before a complete 

rupture or failure. It is extremely difficult to accurately predict the location and leakage rate of 

the concrete containment due to beyond design basis internal pressure and temperature. 

Hessheimer and Dameron [2], and Dameron, Rashid, and Tang [4] provide guidance for 

predicting leak area and leak rate in containments. Hessheimer and Dameron [2] recommend a 

non-linear finite element analysis of the concrete containment to predict containment 

performance and leakage. In the past, reactor severe accident progression analysis has often 

assumed that the concrete containment starts to leak through a small hole as soon as containment 

is pressurized. The area of the small hole is calculated based on nominal design leakage rate of 

0.10 to 0.20 percent of containment free volume mass per day at the design internal pressure. 

The area of the hole is assumed to remain constant until containment failure in the accident 

progression analysis. Results of concrete containment model tests [5, 6] indicate that leakage 

area increases appreciably with internal pressure. In addition, if the rate of pressurization is 

gradual and does not exceed the leakage rate, catastrophic failure of the concrete containment is 

not possible. 

2.1 Concrete Containment Performance under Internal Pressure

A 1:6 scale model of a representative PWR concrete containment was tested at the Sandia 

National Laboratories (SNL) in July 1987 [5]. Prior to performing the test, 10 international 

organizations performed an independent and separate (round robin) pretest analyses of the 

containment [7] to predict containment behavior. A summary of the round robin analyses and 

test results is presented in Table A-1 . 

Hessheimer and Dameron [2] have concluded that global, free field strain of 1.5 to 2.0% for 

reinforced and 0.5 to 1.0% for prestressed concrete can be achieved before failure or rupture. In 

addition, leakage in concrete containment increases appreciably after the rebars and liner plate 

yield. Furthermore, under gradual increase in internal pressure, containment leakage continues 

to grow without failure and rupture. Using these criteria, failure and yield pressure for the 

1:6 scale model concrete containment were calculated using the following equations.
A-2

The results of these simple calculations, as shown in Table A-1 , are quite consistent with 

detailed finite element analyses using state of the art computer codes and test data. 

Pfail = (Ahoop * Yrebar@2% + Aliner * Yliner@2% ) / R

Pyield = (Ahoop * Yrebar + Aliner * Yliner) / R

where:

Pfail = Containment failure pressure 

Pyield = Containment pressure at which hoop rebars and liner plate yield 

Ahoop = Area of the hoop rebars

Aliner = Area of the liner plate

Yrebar = Yield stress of the rebar

Yliner = Yield stress of the liner plate

Yrebar@2% = Stress in the rebar at 2% strain

Yliner@2% = Stress in the rebar at 2% strain

R = Radius of the containment
A-3

Table A-1 Internal Pressure in 1:6-Scale Reinforced Concrete Containment

Source Hoop Rebar and Liner Plate Yield

MPa (psig)

Containment Failure

MPa (psig)

Round Robin 

Analyses 

(Maximum)

0.951 (138) 1.276 (185)

Round Robin 

Analyses 

(Minimum)

0.827 (120) 0.883 (128)

Round Robin 

Analyses 

(Average)

0.869 (126) 1.076 (156)

Test Data 0.820 (119) 1.00 (145)

Proposed 

Simplified 

Analysis

0.876 (127) 0.986 (143)

The simplified analysis approach was then used to determine the behavior of three existing 

reinforced concrete PWR containments. The comparison of results using the simplified 

approach with information provided by the three plant licensees in their Individual Plant 

Examination (IPE) reports is presented in Table A-2. A review of this table indicates that failure 

pressure predicted in the IPE reports for all three containments is 10 to 25 percent higher than the 

one obtained by simplified approach. Similarly, the pressure at which rebars in the three 

containments yield, as reported in the IPE reports, varies from the simplified analysis by 4 to 

40 percent. These differences in the predictions are similar to the ones reported by the round 

robin analysts for the 1:6 scale containment, and are essentially due to the use of different criteria 

for postulated failure. For instance, the licensees have used strains greater than 2 percent to 

determine failure pressure reported in the IPEs. 

Table A-2 Internal Pressure at Yield and Failure in Reinforced Concrete PWR 

Containments

Item Containment 

#1

Containment 

#2

Containment 

#3

Internal Pressure at Rebar Yield from 

IPE Report MPa (psig) 0.758 (110) 1.000 (145) 1.248 (181)

Internal Pressure at Failure from IPE 

Report MPa (psig) 1.062 (154) 1.048 (152) 1.489 (216)*

Internal Pressure at Rebar Yield from 

Simplified Analysis MPa (psig) 0.779 (113) 0.848 (123) 1.062 (154)

Internal Pressure at from the Proposed 

Simplified Analysis MPa (psig) 0.855 (124) 0.958 (139) 1.200 (174)

* IPE confirmatory analysis determined the failure pressure as 158 psi at 1% strain. 

To further confirm the validity of the simplified analysis approach, it was applied to the 1⁄4 scale 

model of a PWR prestressed concrete containment that was tested at Sandia National 
A-4

Laboratories in 2000 [6]. Prior to performing the test, a round robin pretest analysis of the 

containment [8] was performed by 17 international organizations to predict containment 

behavior. A summary of the round robin analysis, test results, and results of simplified analysis 

based on free field strain of 1.0% for failure is presented in Table A-3. The simplified approach 

for prestressed containment is similar to the one described above for reinforced concrete except 

that effect of prestressing steel is included in the calculations.

Table A-3 Internal Pressure at Yield and Failure in the 1:4-Scale Prestressed 

Concrete Containment

Source Hoop Reinforcement Yield

MPa(psig)

Containment Failure

MPa (psig)

(Leakage >100%)

Round Robin 

Analyses (Maximum) 1.248 (181) 1.979 (287)

Round Robin 

Analyses (Minimum) 0.855 (124) 0.814 (118)

Round Robin 

Analyses (Average) 1.034 (150) 1.413 (205)

Test Data 1.055 (153) 1.296 (188)

Proposed Simplified 

Analysis 1.062 (154) 1.331(193)

A review of the Table A-3 indicates that there is a wide variation in predicted pressures by 

17 organizations. The maximum predicted failure pressure is 2.4 times more than the minimum 

predicted failure pressure. However, the average round robin and the proposed simplified 

analysis predicted pressures are quite close to the pressures recorded during the test. 

The simplified approach described above was also applied to the 1:10 Scale Sizewell B model. 

The results of proposed simplified analysis are compared with the 3-D finite element analysis 

and pressure test data in Table A-4. The proposed simplified approach results closely match 

with detailed 3-dimensional non-linear finite element analysis and test data.

Table A-4 Internal Pressure at Yield and Failure in the 1:10-Scale Prestressed 

Concrete Containment

Item

Test Result 

[2]

Proposed 

Simplified 

Approach

3-D Analysis 

[2]

Internal Pressure at Rebar 

Yield

MPa (psig)

0.586 (85) 0.683 (99) 0.662 (96)

Internal Pressure at Failure 

MPa (psig) (Leakage > 100%) 0.772 (112) 0.738 (107) 0.738 (107)
A-5

Based on the above discussion, simplified analysis approach provides good agreement with the 

more detailed finite analysis and test data for the concrete containment performance under 

internal pressure, and was to determine Surry containment behavior.

2.2 Containment Leakage

The containment performance criteria used for severe accident analysis require prediction of 

leakage rate as a function of internal pressure and temperature. There is lack of experimental 

data for containment leakage beyond design pressure. Rizkalla et al. [9], Dameron et al. [4], and 

others have attempted to quantify leakage through concrete sections. This guidance cannot be 

used directly to determine leakage thru concrete containments in which steel liner plate is 

designed to act as a leakage barrier. Detailed 3-dimensional nonlinear analysis of the 

containments with equipment hatch and other penetrations can determine the local strains in the 

liner plate and concrete. The results of the 3-dimensional nonlinear analysis can be used to 

determine airflow through the liner plate and containment concrete. All these complicated 

analyses will lead to leak rate predictions with large uncertainties due to variation in the 

properties of the materials, quality and porosity of welds, and concrete placement. 

The relationship between containment leakage and internal pressures for reinforced concrete and 

prestressed concrete containment model tests from References 4 and 5 is shown in Figure A-1

and Figure A-2, respectively. A review of these figures indicates that the concrete containments 

start to leak appreciably once the liner plate yields. The rate of leakage when the liner plate 

yields is about 10 times more than normal leakage of 0.10 percent of containment air mass 

per day at the containment design pressure. The leakage rate increases appreciably with further 

increase in test pressure. Once the rebars yield, the leakage rate is about 10-15 percent. 

Thereafter, the leakage rate continues to increase and reaches to about 60-65 percent when the 

strain in the rebars is about 1-2 percent. Containment pressure does not increase significantly 

after leakage rate exceeds 60-65 percent of the containment air mass per day. The liner welds 

and concrete crack after rebars and liner plate yields to create a path for leakage. The leakage 

occurs in areas such as equipment hatch, personnel airlocks and penetrations where local strains 

are substantially higher than the global strains. 

The results in Figure A-1 and Figure A-2 are for scale model tests of two concrete containments. 

Rebar and concrete crack spacing, and aggregate size can affect the leakage rates in full size 

containments. However, based on the results of testing and analyses presented above, it is 

reasonable to conclude that all concrete containments start to leak once the rebars and liner plate 

yield. In addition, leakage becomes excessive once the strains in the reinforced and prestressed 

concrete containments reach about 2 and 1 percent respectively. Based on information of the 

containment model test results and analyses data presented in Figure A-1 and Figure A-2, it is 

reasonable to assume that containment leakage is about one percent of the containment mass 

per day when the liner plate yields, this increases to 13 percent of containment mass per day 

when rebars yield. Similarly, leakage rate of 62 percent can be used in severe accident analysis 

when the containment global strains are 1-2 percent. Uncertainty in the leakage rate can be 

accounted for by conservatively reducing the yield and failure pressure calculated by simplified 
A-6

analysis to 85 percent of the calculated value.

Figure A-1 1:6 Scale Prestressed Concrete Containment Test Pressure versus Leakage 

Rate

1:6 Scale Containment Reinforced Concrete Containment Test Pressure Vs. Leakage Rate

0

20

40

60

80

100

120

140

160

0.1 1 10 100 1000

Leakage Rate (% Mass/day)

Internal Pressure (psig)

Test Pressure 

Pressure @ Liner Yield 

Pressure @ Rebar Yield 

Pressure @ 2.00% Strain
A-7

Figure A-2 1:4 Scale Prestressed Concrete Containment Test

The location of the leakage can have a significant effect on the results of the severe accident 

analysis and dose rates. For instance, if the containment leakage occurs thru penetrations that are 

located inside adjoining plant buildings, the fission product release into atmosphere would be 

significantly less as compared to direct leakage to the environment. Previously, some of the 

severe accident analyses were based on the assumption that the leakage takes place at the top of 

the containment dome. A more realistic approach is to consider leakage to occur at equipment 

hatch and other penetrations as demonstrated by tests data, and non-linear analyses. 

1:4 Scale Prestressed Concrete Containment Test Pressure Vs. Leakage Rate

0

50

100

150

200

250

0.10 1.00 10.00 100.00 1000.00

Leakage Rate (% Mass/day)

Internal Pressure (psig)

Test Pressure 

Pressure @ Liner Yield

Pressure @ Rebar Yield 

Pressure @ 1.00% Strain
A-8

3.0 ANALYSIS

3.1 Containment Internal Pressure at Liner Plate Yield

The liner plate material yield strength is less than yield strength of rebars. However, modulus of 

elasticity of carbon steel liner plate and rebars is about the same. Therefore, the liner plate is 

likely to yield first under internal pressure. When the liner plate yields, the stress in the rebar 

and liner plate would be the same and equal to yield strength of the liner plate. Using this 

approach, the internal pressure at which liner plate will yield Plineryield was calculated as follows:

Plineryield = (Ahoop + Aliner ) * Yliner / R

where:

Plineryield = Containment pressure at which liner plate yield 

Ahoop = Area of the hoop rebars = 18.777 in2

/ft (Reference 10) 

Aliner = Area of the 3/8” thick liner plate = 4.5 in2

/ft (Reference 11) 

Yrebar = Yield stress of the rebar = 50,000 psi (Reference 12) 

Yliner = Yield stress of the liner plate = 32,000 psi (Reference 11)

R = Radius of the containment = 63 feet (Reference 11)

Using the above listed values:

Plineryield = 82.10 psi

To account for uncertainties in material properties and other simplifying assumptions, this 

pressure at liner plate yield was reduced to 85 percent of the calculated value for the MELCOR

analysis. 

Therefore: 

P@lineryield = 69.79 psi
A-9

3.2 Containment Internal Pressure at Rebar Yield

Pyield = (Ahoop * Yrebar + Aliner * Yliner) / R

Using this equation:

Pyield = 119.36 psig

To account for uncertainties in material properties and other simplifying assumptions, this 

pressure at yield will be reduced to 85 percent of the calculated value for the MELCOR analysis.

Therefore:

P@yield = 101.46 psi

3.3 Containment Internal Pressure at 2% Strain

Pfail = (Ahoop * Yrebar@2% + Aliner * Yliner@2% ) / R

where:

Pfail = Containment failure pressure at 2% strain

Yrebar@2% = Stress in the rebar at 2% strain = 53,000 psi

Yliner@2% = Stress in the liner at 2% strain = 34,300 psi

Using the above listed values:

Pfail = 126.71 psi

To account for uncertainties in material properties and other simplifying assumptions, this 

pressure at failure will be reduced to 85 percent of the calculated value for the MELCOR

analysis. 

Therefore: 

P@fail = 107.70 psi
A-10

3.4 Containment Leakage

Surry minimum containment free volume per Table 5.4-24 of Reference 13 = 1, 730,000 ft3

Density of air at containment pressure of 119.36 psi and 200oF (rebar yield):

ρ = 0.55 lb/ft3

, (Page A-10 of Reference 14)

Mass of air inside containment at Pyield

MassPyield = ρ Vcontainment

MassPyield = 9.515 x 105 lb

Mass leak rate of the containment at Pyield:

MassleakratePyield = 13%/day

Massleakperday = MassPyield MassleakratePyield

Massleakperday = 1.237 x 105 lb/day

Density of air at 70o F and atmospheric pressure, ρa

ρa = 0.075 lb/ft3, (Reference 14)

Leakage flow “Q” calculation

Q = Massleakperday / ρa

Q = 1.649 x 106 lb/day

Therefore, leakage volume

Volleak% = Q/Vc

Volleak% = 95.33%/day

Table A-5 provides a summary of these results for containment air temperature of 200o F.
A-11

Table A-5 Recommended Leakage Rates and Areas for the Surry Analyses

Containment 

Pressure

(psig)

Containment 

Temperature

(

o

F)

Containment 

Leakage 

(% Mass/day)

45.00 70 0.1

69.79 200 1.0

101.46 200 13

107.70 200 62

123.20 200 352

4.0 CONCLUSION

Simplified analysis of concrete containment provides good agreement with the more detailed 

finite analysis and test data for the concrete containment performance under internal pressure, 

and has been used in this report to determine Surry containment. 

Extensive research and scale model testing of reinforced and prestressed concrete containments 

indicate that concrete containments start to leak before a complete rupture or failure. Unless the 

rate of pressurization is extremely rapid, concrete containments are not likely to have a 

catastrophic failure. There is some uncertainty about containment leakage rate; however, 

concrete containments start to leak significantly after the liner plate and rebar yield. Leakage 

rate becomes excessive after the global strains in the liner plate and rebar reach 2%. Surry 

containment leakage rates under different internal pressures have been determined using the 

results of previous tests performed on different scale models of the concrete containments. The 

leakage rates are presented in Table A-5. Most of the leakage is likely to occur at equipment 

hatch and other penetrations as demonstrated by tests data, and non-linear analyses. 
A-12

5.0 REFERENCES

1. U. S. Nuclear Regulatory Commission, NUREG/CR-2239, "Technical Guidance for Siting 

Criteria Development," Prepared by Sandia National Laboratories, December 1982.

2. Hessheimer, M. F. and Dameron, R. A., “Containment Integrity Research at Sandia National 

Laboratories,” NUREG/CR–6906, 2006.

3. Dameron, R. A., Rashid, Y. R., and Sullaway, M. F., “Pretest Prediction Analysis and 

Posttest Correlation of the Sizewell-B 1;10 Scale Prestressed Concrete Containment Model 

Test,” NUREG/CR–5671, 1998.

4. Dameron, R. A., Rashid, Y. R., and Tang, H. T., “Leak Area and Leakage Rate Prediction for 

Probabilistic Risk Assessment of Concrete Containments under Severe Core Conditions,” 

Nuclear Engineering and Design, 1995.

5. Horschell, D. S., “Experimental Results From Pressure Testing a 1:6 – Scale Nuclear Power 

Plant Containment,” NUREG/CR–5121, 1992.

6. Hessheimer, M. F., Klamerus, E. W., Rightly, G. S., Lambert, L. D., and Dameron, R. A., 

Overpressurization Test of a 1:4–Scale Prestressed Concrete Containment Vessel Model, 

NUREG/CR-6810, 2003.

7. Clauss, D. B., Round Robin Pretest Analyses of a 1:6-Scale Reinforced Concrete 

Containment Model Subject to Static Internal Overpressurization, NUREG/CR-4913, April 

1987. 

8. Luk, V. K., “Pretest Round Robin Analysis of a Prestressed Concrete Containment Vessel 

Model,” NUREG/CR-6678, August 2000.

9. Rizkalla, S. H., Lau, B. L., and Simmonds, S. H., “Air leakage Characteristics in Reinforced 

Concrete, Journal of Structural Engineering, American Society of Civil Engineers, Vol. 110, 

No. 5, 1984.

10. Surry Power Station, Drawing No. 11448-FC-15C

11. Surry Power Station, Calculation 11448-EA-53

12. Surry Power Station, Drawing No. 11448-FC-15A 

13. Surry Power Station Updated Final Safety Analysis Report (UFSAR)

14. Crane Company Technical Paper 410, “Flow of Fluids Through Valves, Fittings, and Pipe.”
APPENDIX B

SURRY RADIONUCLIDE INVENTORY

B-1

APPENDIX B: SURRY RADIONUCLIDE INVENTORY

The following tables summarize the radionuclide core inventory for the Surry plant at the time of 

shutdown. This isotopic inventory was used in each of the accident progression scenarios 

considered in this report.

Table B-1 Surry radionuclide core inventory and class definition

Radionuclide Class Name Representative

Element

Member 

Elements

Radioactive

Mass (kg)

Noble Gas Xe He, Ne, Ar, Kr, 

Xe, Rn, H, N 448.2

Alkali Metals Cs Li, Na, K, Rb, Cs, 

Fr, Cu 251.7

Alkaline Earths Ba Be, Mg, Ca, Sr, 

Ba, Ra, Es, Fm 187.6

Halogens I F, Cl, Br, I, At 17.0

Chalcogens Te O, S, Se, Te, Po 40.9

Platinoids Ru Ru, Rh, Pd, Re, 

Os, Ir, Pt, Au, Ni 309.5

Early Transition Elements Mo

V, Cr, Fe, Co, 

Mn, Nb, Mo, Tc, 

Ta, W

323.5

Tetravalent Ce Ti, Zr, Hf, Ce, Th, 

Pa, Np, Pu, C 1226.0

Trivalents La

Gd, Tb, Dy, Ho, 

Er, Tm, Yb, Lu, 

Am, Cm, Bk, Cf

621.2

Uranium U U 66770.0

More Volatile Main Group Cd Cd, Hg, Zn, As, 

Sb, Pb, Tl, Bi 7.26

Less Volatile Main Group Sn Ga, Ge, In, Sn, Ag 9.19

Table B-2 Surry noble gas radionuclide class specific isotopic activity at the time of 

reactor shutdown

Isotope Activity (Bq)

Kr-85 2.94E+16

Kr-85m 8.07E+17

Kr-87 1.60E+18

Kr-88 2.14E+18

Xe-133 6.07E+18

Xe-135 1.80E+18

Xe-135m 1.29E+18
B-2

Table B-3 Surry alkali metals radionuclide class specific isotopic activity at the time of 

reactor shutdown

Isotope Activity (Bq)

Cs-134 4.32E+17

Cs-136 1.57E+17

Cs-137 3.05E+17

Rb-86 5.36E+15

Rb-88 2.16E+18

Table B-4 Surry alkali earths radionuclide class specific isotopic activity at the time of 

reactor shutdown

Isotope Activity (Bq)

Ba-139 5.54E+18

Ba-140 5.37E+18

Sr-89 2.98E+18

Sr-90 2.27E+17

Sr-91 3.75E+18

Sr-92 4.00E+18

Ba-137m 2.92E+17

Table B-5 Surry halogen radionuclide class specific isotopic activity at the time of 

reactor shutdown

Isotope Activity (Bq)

I-131 2.78E+18

I-132 4.08E+18

I-133 5.76E+18

I-134 6.48E+18

I-135 5.49E+18

Table B-6 Surry chalcogen radionuclide class specific isotopic activity at the time of 

reactor shutdown

Isotope Activity (Bq)

Te-127 2.60E+17

Te-127m 4.22E+16

Te-129 7.79E+17

Te-129m 1.49E+17

Te-131m 5.71E+17

Te-132 4.29E+18

Te-131 2.55E+18
B-3

Table B-7 Surry platinoid radionuclide class specific isotopic activity at the time of 

reactor shutdown

Isotope Activity (Bq)

Rh-105 2.90E+18

Ru-103 4.61E+18

Ru-105 3.14E+18

Ru-106 1.40E+18

Rh-103m 4.61E+18

Rh-106 1.56E+18

Table B-8 Surry early transition element radionuclide class specific isotopic activity at 

the time of reactor shutdown

Isotope Activity (Bq)

Nb-95 5.18E+18

Co-58 4.79E+13

Co-60 2.65E+14

Mo-99 5.68E+18

Tc-99m 5.03E+18

Nb-97 5.24E+18

Nb-97m 4.95E+18

Table B-9 Surry tetravalent radionuclide class specific isotopic activity at the time of 

reactor shutdown

Isotope Activity (Bq)

Ce-141 4.87E+18

Ce-143 4.55E+18

Ce-144 3.42E+18

Np-239 5.67E+19

Pu-238 8.31E+15

Pu-239 9.56E+14

Pu-240 1.17E+15

Pu-241 3.39E+17

Zr-95 4.96E+18

Zr-97 5.00E+18
B-4

Table B-10 Surry trivalent radionuclide class specific isotopic activity at the time of 

reactor shutdown

Isotope Activity (Bq)

Am-241 3.43E+14

Cm-242 1.14E+17

Cm-244 1.13E+16

La-140 5.67E+18

La-141 5.10E+18

La-142 4.92E+18

Nd-147 2.04E+18

Pr-143 4.65E+18

Y-90 2.39E+17

Y-91 3.93E+18

Y-92 4.11E+18

Y-93 4.62E+18

Y-91m 2.20E+18

Pr-144 3.63E+18

Pr-144m 5.06E+16
APPENDIX C

INPUT PARAMETERS 

FOR CONSEQUENCE ANALYSIS 

C-1

APPENDIX C: INPUT PARAMETERS FOR CONSEQUENCE ANALYSIS 

The input parameters used for the LTSBO, STSBO, ISLOCA, TI-SGTR Mitigated and 

Unmitigated Scenarios are shown in this appendix in tabular form. Table C-1 contains the more 

general ATMOS input parameters used for these three scenarios. Table C-2 through Table C-4 

contains specific inputs related to the source terms that were extracted from MELCOR results via 

the MELMACCS code. Table C-5 contains general EARLY input parameters. Table C-6 and

Table C-7 contain parameters associated with the network evacuation model that was used to 

treat emergency response. Table C-8 contains the evacuation direction parameters. Table C-9

contains the CHRONC input parameters.

Table C-1 ATMOS Input Parameters Used in the Surry LTSBO, STSBO, ISLOCA, 

TISGTR Mitigated and Unmitigated scenarios

Variable Description LTSBO STSBO ISLOCA TISGTR 

Mitigated

TISGTR 

Unmitigated

APLFRC

Method of 

Applying Release 

Fraction

PARENT PARENT PARENT PARENT PARENT

ATNAM1

Title Describing 

the ATMOS 

Assumptions

SOARCA Surry 

Unmitigated 

LTSBO

SOARCA Surry 

Unmitigated 

STSBO

SOARCA Surry 

Unmitigated 

STSBO

SOARCA Surry 

Mitigated 

TISGTR

SOARCA Surry 

Unmitigated 

TISGTR

ATNAM2 Title Describing 

the Source Term

Surry source 

term for 

unmitigated 

long-term station 

blackout.

Surry source term 

for unmitigated 

short-term station 

blackout.

Surry source term 

for unmitigated 

short-term station 

blackout.

Surry source term 

for mitigated, 

thermallyinduced, steamgenerator-tube 

rupture

Surry source 

term for 

unmitigated 

short-term 

station blackout.

BNDMXH

Boundary 

Weather Mixing 

Layer Height

1000 1000 1000 1000 1000

BNDRAN

Boundary 

Weather Rain 

Rate

5 5 5 5 5

BNDWND Boundary Wind 

Speed 2.2 2.2 2.2 2.2 2.2

BRKPNT

Breakpoint Time 

for Plume 

Meander

3600 3600 3600 3600 3600

BUILDH

Building Height 

for all Plume 

Segments

50 50 50 50 50

CORINV

Isotopic 

Inventory at 

Time of Reactor 

Shutdown

from 

MELMACCS 

(see Appendix B)

from 

MELMACCS 

(see Appendix B)

from 

MELMACCS (see 

Appendix B)

from 

MELMACCS 

(see Appendix B)

from 

MELMACCS 

(see Appendix B)

CORSCA

Linear Scaling 

Factor on Core 

Inventory

1 1 1 1 1

CWASH1

Linear 

Coefficient for 

Washout

1.89E-05 1.89E-05 1.89E-05 1.89E-05 1.89E-05
C

-2

Variable Description LTSBO STSBO ISLOCA TISGTR 

Mitigated

TISGTR 

Unmitigated

CWASH2

Exponential 

Term for 

Washout

0.664 0.664 0.664 0.664 0.664

CYSIGA

Linear 

Coefficient for 

sigma-y

Stability Class A 0.7507 0.7507 0.7507 0.7507 0.7507

Stability Class B 0.7507 0.7507 0.7507 0.7507 0.7507

Stability Class C 0.4063 0.4063 0.4063 0.4063 0.4063

Stability Class D 0.2779 0.2779 0.2779 0.2779 0.2779

Stability Class E 0.2158 0.2158 0.2158 0.2158 0.2158

Stability Class F 0.2158 0.2158 0.2158 0.2158 0.2158

CYSIGB

Exponential 

Term

for sigma-y

Stability Class A 0.866 0.866 0.866 0.866 0.866

Stability Class B 0.866 0.866 0.866 0.866 0.866

Stability Class C 0.865 0.865 0.865 0.865 0.865

Stability Class D 0.881 0.881 0.881 0.881 0.881

Stability Class E 0.866 0.866 0.866 0.866 0.866

Stability Class F 0.866 0.866 0.866 0.866 0.866

CZSIGA

Linear 

Coefficient for 

sigma-z

Stability Class A 0.0361 0.0361 0.0361 0.0361 0.0361

Stability Class B 0.0361 0.0361 0.0361 0.0361 0.0361

Stability Class C 0.2036 0.2036 0.2036 0.2036 0.2036

Stability Class D 0.2636 0.2636 0.2636 0.2636 0.2636

Stability Class E 0.2463 0.2463 0.2463 0.2463 0.2463

Stability Class F 0.2463 0.2463 0.2463 0.2463 0.2463

CZSIGB

Exponential 

Term

for sigma-z

Stability Class A 1.277 1.277 1.277 1.277 1.277
C-3

Variable Description LTSBO STSBO ISLOCA TISGTR 

Mitigated

TISGTR 

Unmitigated

Stability Class 

B 1.277 1.277 1.277 1.277 1.277

CZSIGB Stability Class 

C 0.859 0.859 0.859 0.859 0.859

Stability Class 

D 0.751 0.751 0.751 0.751 0.751

Stability Class 

E 0.619 0.619 0.619 0.619 0.619

Stability Class 

F 0.619 0.619 0.619 0.619 0.619

DISPMD Dispersion 

Model Flag LRDIST LRDIST LRDIST LRDIST LRDIST

DRYDEP Dry Deposition 

Flag

Xe = .FALSE.

Other Groups = 

.TRUE.

Xe = .FALSE.

Other Groups = 

.TRUE.

Xe = .FALSE.

Other Groups = 

.TRUE.

Xe = .FALSE.

Other Groups = 

.TRUE.

Xe = .FALSE.

Other Groups = 

.TRUE.

ENDAT1

Control flag 

indicating only 

ATMOS is to be 

run

.FALSE. .FALSE. .FALSE. .FALSE. .FALSE.

GRPNAM

Names of the 

Chemical Classes 

(Used by 

WinMACCS)

Chemical 

Class 1 Xe Xe Xe Xe Xe

Chemical 

Class 2 Cs Cs Cs Cs Cs

Chemical 

Class 3 Ba Ba Ba Ba Ba

Chemical 

Class 4 I I I I I

Chemical 

Class 5 Te Te Te Te Te

Chemical 

Class 6 Ru Ru Ru Ru Ru

Chemical 

Class 7 Mo Mo Mo Mo Mo

Chemical 

Class 8 Ce Ce Ce Ce Ce

Chemical 

Class 9 La La La La La

IBDSTB

Boundary 

Weather Stability 

Class Index

4 4 4 4 4

IDEBUG

Debug Switch for 

Extra Debugging 

Print

0 0 0 0 0

IGROUP

Definition of 

Radionuclide 

Group Numbers

1 = Xe 1 = Xe 1 = Xe 1 = Xe 1 = Xe

2 = Cs 2 = Cs 2 = Cs 2 = Cs 2 = Cs

3 = Ba 3 = Ba 3 = Ba 3 = Ba 3 = Ba
C-4

Variable Description LTSBO STSBO ISLOCA TISGTR 

Mitigated

TISGTR 

Unmitigated

4 = I 4 = I 4 = I 4 = I 4 = I

5 = Te 5 = Te 5 = Te 5 = Te 5 = Te

6 = Ru 6 = Ru 6 = Ru 6 = Ru 6 = Ru

IGROUP 7 = Mo 7 = Mo 7 = Mo 7 = Mo 7 = Mo

8 = Ce 8 = Ce 8 = Ce 8 = Ce 8 = Ce

9 = La 9 = La 9 = La 9 = La 9 = La

INWGHT

Number of 

Samples for Each 

Bin Used for 

Nonuniform 

Weather Bin 

Sampling

Bin 1 12 12 12 12 12

Bin 2 14 14 14 14 14

Bin 3 12 12 12 12 12

Bin 4 39 39 39 39 39

Bin 5 85 85 85 85 85

Bin 6 94 94 94 94 94

Bin 7 16 16 16 16 16

Bin 8 12 12 12 12 12

Bin 9 27 27 27 27 27

Bin 10 134 134 134 134 134

Bin 11 119 119 119 119 119

Bin 12 74 74 74 74 74

Bin 13 92 92 92 92 92

Bin 14 43 43 43 43 43

Bin 15 12 12 12 12 12

Bin 16 12 12 12 12 12

Bin 17 37 37 37 37 37

Bin 18 12 12 12 12 12

Bin 19 13 13 13 13 13

Bin 20 17 17 17 17 17

Bin 21 14 14 14 14 14

Bin 22 12 12 12 12 12

Bin 23 6 6 6 6 6

Bin 24 6 6 6 6 6

Bin 25 12 12 12 12 12

Bin 26 12 12 12 12 12

Bin 27 12 12 12 12 12

Bin 28 1 1 1 1 1

Bin 29 5 5 5 5 5

Bin 30 6 6 6 6 6

Bin 31 5 5 5 5 5

Bin 32 12 12 12 12 12

Bin 33 5 5 5 5 5

Bin 34 12 12 12 12 12

Bin 35 12 12 12 12 12

Bin 36 12 12 12 12 12
C-5

Variable Description LTSBO STSBO ISLOCA TISGTR 

Mitigated

TISGTR 

Unmitigated

IRSEED

Seed for Random 

Number 

Generator

79 79 79 79 79

LATITU Latitude of 

Power Plant 37 ̊ 9' 56" 37 ̊ 9' 56" 37 ̊ 9' 56" 37 ̊ 9' 56" 37 ̊ 9' 56"

LIMSPA

Last Interval for 

Measured 

Weather

25 25 25 25 25

LONGIT Longitude of 

Power Plant 76 ̊ 41' 54" 76 ̊ 41' 54" 76 ̊ 41' 54" 76 ̊ 41' 54" 76 ̊ 41' 54"

MAXGRP

Number of 

Radionuclide 

Groups

9 9 9 9 9

MAXHGT Flag for Mixing 

Height

DAY_AND_NI

GHT

DAY_AND_NIG

HT

DAY_AND_NIG

HT

DAY_AND_NIG

HT

DAY_AND_NI

GHT

MAXRIS Selection of Risk 

Dominant Plume 3 3 1 1 1

METCOD

Meteorological 

Sampling Option 

Code

2 2 2 2 2

MNDMOD Plume Meander 

Model Flag OFF OFF OFF OFF OFF

NAMSTB List of Pseudo

stable Nuclides

Isotope 1 I-129 I-129 I-129 I-129 I-129

Isotope 2 Xe-131m Xe-131m Xe-131m Xe-131m Xe-131m

Isotope 3 Xe-133m Xe-133m Xe-133m Xe-133m Xe-133m

Isotope 4 Cs-135 Cs-135 Cs-135 Cs-135 Cs-135

Isotope 5 Sm-147 Sm-147 Sm-147 Sm-147 Sm-147

Isotope 6 U-234 U-234 U-234 U-234 U-234

Isotope 7 U-235 U-235 U-235 U-235 U-235

Isotope 8 U-236 U-236 U-236 U-236 U-236

Isotope 9 U-237 U-237 U-237 U-237 U-237

Isotope 10 Np-237 Np-237 Np-237 Np-237 Np-237

Isotope 11 Rb-87 Rb-87 Rb-87 Rb-87 Rb-87

Isotope 12 Zr-93 Zr-93 Zr-93 Zr-93 Zr-93

Isotope 13 Nb-93m Nb-93m Nb-93m Nb-93m Nb-93m

Isotope 14 Nb-95m Nb-95m Nb-95m Nb-95m Nb-95m

Isotope 15 Tc-99 Tc-99 Tc-99 Tc-99 Tc-99

Isotope 16 Pm-147 Pm-147 Pm-147 Pm-147 Pm-147

NPSGRP

Number of 

Particle Size 

Groups

10 10 10 10 10

NRINTN

Number of Rain 

Intensity 

Breakpoints

3 3 3 3 3

NRNINT

Number of Rain 

Distance 

Intervals

5 5 5 5 5

NSBINS Number of 36 36 36 36 36
C-6

Variable Description LTSBO STSBO ISLOCA TISGTR 

Mitigated

TISGTR 

Unmitigated

Weather Bins to 

Sample

NUCNAM Radionuclide 

Names See Appendix B See Appendix B See Appendix B See Appendix B See Appendix B

NUCOUT

Radionuclide 

Used in 

Dispersion Print

Cs-137 Cs-137 Cs-137 Cs-137 Cs-137

NUMCOR

Number of 

Compass Sectors 

in the Grid

64 64 64 64 64

NUMISO Number of 

Radionuclides 69 69 69 69 69

NUMRAD

Number of 

Radial Spatial 

Intervals

26 26 26 26 26

NUMREL

Number of 

Released Plume 

Segments

28 24 107 49 115

NUMSTB

Number of 

Defined Pseudo

stable 

Radionuclides

16 16 16 16 16

OALARM

Time to Reach 

General 

Emergency 

Conditions

0 0 0 0 0

PDELAY Plume Release 

Times

MELMACCS 

Data (See Table 

C-2)

MELMACCS 

Data (See Table 

C-2)

MELMACCS Data 

(See Table C-2)

MELMACCS 

Data (See Table 

C-2)

MELMACCS 

Data (See Table 

C-2)

PLHEAT Plume Heat 

Contents

MELMACCS 

Data (See Table 

C-2)

MELMACCS 

Data (See Table 

C-2)

MELMACCS Data 

(See Table C-2)

MELMACCS 

Data (See Table 

C-2)

MELMACCS 

Data (See Table 

C-2)

PLHITE Plume Release 

Heights

MELMACCS 

Data (See Table 

C-2)

MELMACCS 

Data (See Table 

C-2)

MELMACCS Data 

(See Table C-2)

MELMACCS 

Data (See Table 

C-2)

MELMACCS 

Data (See Table 

C-2)

PLMDEN Plume Mass 

Density

MELMACCS 

Data (See Table 

C-2)

MELMACCS 

Data (See Table 

C-2)

MELMACCS Data 

(See Table C-2)

MELMACCS 

Data (See Table 

C-2)

MELMACCS 

Data (See Table 

C-2)

PLMFLA Plume Mass 

Flow Rate

MELMACCS 

Data (See Table 

C-2)

MELMACCS 

Data (See Table 

C-2)

MELMACCS Data 

(See Table C-2)

MELMACCS 

Data (See Table 

C-2)

MELMACCS 

Data (See Table 

C-2)

PLMMOD

Flag for Plume 

Rise Input 

Option

DENSITY DENSITY DENSITY DENSITY DENSITY

PLUDUR Plume Segment 

Durations

MELMACCS 

Data (See Table 

C-2)

MELMACCS 

Data (See Table 

C-2)

MELMACCS Data 

(See Table C-2)

MELMACCS 

Data (See Table 

C-2)

MELMACCS 

Data (See Table 

C-2)

PSDIST

Particle Size 

Distribution by 

Group

MELMACCS 

Data (See Table 

C-3)

MELMACCS 

Data (See Table 

C-3)

MELMACCS Data 

(See Table C-3)

MELMACCS 

Data (See Table 

C-3)

MELMACCS 

Data (See Table 

C-3)

REFTIM Plume Reference 

Time Point

0. for first

0.5 for 

0. for first

0.5 for 

0. for first

0.5 for subsequent

0. for first

0.5 for 

0. for first

0.5 for 
C-7

Variable Description LTSBO STSBO ISLOCA TISGTR 

Mitigated

TISGTR 

Unmitigated

subsequent subsequent subsequent subsequent

RELFRC

Release Fractions 

of the Source 

Term

MELMACCS 

Data (See Table 

C-4)

MELMACCS 

Data (See Table 

C-4)

MELMACCS Data 

(See Table C-4)

MELMACCS 

Data (See Table 

C-4)

MELMACCS 

Data (See Table 

C-4)

RNDSTS

Endpoints of 

Rain Distance 

Intervals

Interval 1 3.22 3.22 3.22 3.22 3.22

Interval 2 5.63 5.63 5.63 5.63 5.63

Interval 3 11.27 11.27 11.27 11.27 11.27

Interval 4 20.92 20.92 20.92 20.92 20.92

Interval 5 32.19 32.19 32.19 32.19 32.19

RNRATE

Rain Intensity 

Breakpoints for 

Weather Binning

Intensity 1 2 2 2 2 2

Intensity 2 4 4 4 4 4

Intensity 3 6 6 6 6 6

SCLADP

Scaling Factor 

for A-D Plume 

Rise

1.0 1.0 1.0 1.0 1.0

SCLCRW

Scaling Factor 

for Critical Wind 

Speed

1.0 1.0 1.0 1.0 1.0

SCLEFP

Scaling Factor 

for E-F Plume 

Rise

1.0 1.0 1.0 1.0 1.0

SIGYINIT

Initial Sigma-y 

for All Plume 

Segments

9.3 9.3 9.3 9.3 9.3

SIGZINIT

Initial Sigma-z 

for All Plume 

Segments

23.3 23.3 23.3 23.3 23.3

SPAEND

Radial distances 

for grid 

boundaries

Ring 1 0.16 0.16 0.16 0.16 0.16

Ring 2 0.52 0.52 0.52 0.52 0.52

Ring 3 1.21 1.21 1.21 1.21 1.21

Ring 4 1.61 1.61 1.61 1.61 1.61

Ring 5 2.13 2.13 2.13 2.13 2.13

Ring 6 3.22 3.22 3.22 3.22 3.22

Ring 7 4.02 4.02 4.02 4.02 4.02

Ring 8 4.83 4.83 4.83 4.83 4.83

Ring 9 5.63 5.63 5.63 5.63 5.63

Ring 10 8.05 8.05 8.05 8.05 8.05

Ring 11 11.27 11.27 11.27 11.27 11.27

Ring 12 16.09 16.09 16.09 16.09 16.09

Ring 13 20.92 20.92 20.92 20.92 20.92
C-8

Variable Description LTSBO STSBO ISLOCA TISGTR 

Mitigated

TISGTR 

Unmitigated

Ring 14 25.75 25.75 25.75 25.75 25.75

Ring 15 32.19 32.19 32.19 32.19 32.19

Ring 16 40.23 40.23 40.23 40.23 40.23

Ring 17 48.28 48.28 48.28 48.28 48.28

SPAEND Ring 18 64.37 64.37 64.37 64.37 64.37

Ring 19 80.47 80.47 80.47 80.47 80.47

Ring 20 112.65 112.65 112.65 112.65 112.65

Ring 21 160.93 160.93 160.93 160.93 160.93

Ring 22 241.14 241.14 241.14 241.14 241.14

Ring 23 321.87 321.87 321.87 321.87 321.87

Ring 24 563.27 563.27 563.27 563.27 563.27

Ring 25 804.67 804.67 804.67 804.67 804.67

Ring 26 1609.34 1609.34 1609.34 1609.34 1609.34

TIMBAS

Time Base for 

Plume Expansion

Factor

600 600 600 600 600

VDEPOS Dry Deposition 

Velocities

Aerosol Bin 1 5.35E-04 5.35E-04 5.35E-04 5.35E-04 5.35E-04

Aerosol Bin 2 4.91E-04 4.91E-04 4.91E-04 4.91E-04 4.91E-04

Aerosol Bin 3 6.43E-04 6.43E-04 6.43E-04 6.43E-04 6.43E-04

Aerosol Bin 4 1.08E-03 1.08E-03 1.08E-03 1.08E-03 1.08E-03

Aerosol Bin 5 2.12E-03 2.12E-03 2.12E-03 2.12E-03 2.12E-03

Aerosol Bin 6 4.34E-03 4.34E-03 4.34E-03 4.34E-03 4.34E-03

Aerosol Bin 7 8.37E-03 8.37E-03 8.37E-03 8.37E-03 8.37E-03

Aerosol Bin 8 1.37E-02 1.37E-02 1.37E-02 1.37E-02 1.37E-02

Aerosol Bin 9 1.70E-02 1.70E-02 1.70E-02 1.70E-02 1.70E-02

Aerosol Bin 10 1.70E-02 1.70E-02 1.70E-02 1.70E-02 1.70E-02

WETDEP Wet Deposition 

Flag

Xe = .FALSE.

Other groups = 

.TRUE.

Xe = .FALSE.

Other groups =

.TRUE.

Xe = .FALSE.

Other groups = 

.TRUE.

Xe = .FALSE.

Other groups = 

.TRUE.

Xe = .FALSE.

Other groups = 

.TRUE.

XPFAC1

Base Time for 

Meander 

Expansion Factor

0.2 0.2 0.2 0.2 0.2

XPFAC2

Breakpoint for 

Expansion Factor 

Model

0.25 0.25 0.25 0.25 0.25

YSCALE

Scale Factor for 

Horizontal 

Dispersion

1 1 1 1 1

ZSCALE

Scale Factor for 

Vertical 

Dispersion

1.27 1.27 1.27 1.27 1.27
C-9

Table C-2 Plume Parameters Used in the Surry LTSBO, Unmitigated STSBO, TISGTR Mitigated and Unmitigated Scenarios

Surry LTSBO

Plume Segment PDELAY PLHEAT PLHITE PLMDEN PLMFLA PLUDUR

1 1.63E+05 5.11E+04 8.40E+00 1.77E-01 4.85E-01 3.60E+03

2 1.67E+05 1.14E+05 8.40E+00 3.89E-01 4.83E-01 3.60E+03

3 1.70E+05 2.07E+05 8.40E+00 7.01E-01 4.81E-01 3.60E+03

4 1.74E+05 3.05E+05 8.40E+00 1.02E+00 4.79E-01 3.60E+03

5 1.77E+05 3.81E+05 8.40E+00 1.27E+00 4.76E-01 3.72E+03

6 1.81E+05 4.44E+05 8.40E+00 1.46E+00 4.74E-01 3.48E+03

7 1.85E+05 4.82E+05 8.40E+00 1.57E+00 4.72E-01 3.60E+03

8 1.88E+05 5.07E+05 8.40E+00 1.63E+00 4.70E-01 3.60E+03

9 1.92E+05 5.25E+05 8.40E+00 1.67E+00 4.69E-01 3.60E+03

10 1.95E+05 5.38E+05 8.40E+00 1.70E+00 4.67E-01 3.60E+03

11 1.99E+05 5.48E+05 8.40E+00 1.72E+00 4.65E-01 3.60E+03

12 2.03E+05 5.59E+05 8.40E+00 1.73E+00 4.63E-01 3.60E+03

13 2.06E+05 5.71E+05 8.40E+00 1.76E+00 4.62E-01 3.60E+03

14 2.10E+05 5.75E+05 8.40E+00 1.75E+00 4.60E-01 3.60E+03

15 2.13E+05 5.79E+05 8.40E+00 1.75E+00 4.58E-01 3.60E+03

16 2.17E+05 5.85E+05 8.40E+00 1.75E+00 4.57E-01 3.60E+03

17 2.21E+05 5.94E+05 8.40E+00 1.77E+00 4.55E-01 3.60E+03

18 2.24E+05 6.02E+05 8.40E+00 1.78E+00 4.54E-01 3.60E+03

19 2.28E+05 6.11E+05 8.40E+00 1.79E+00 4.52E-01 3.60E+03

20 2.31E+05 6.17E+05 8.40E+00 1.80E+00 4.51E-01 3.60E+03

21 2.35E+05 6.22E+05 8.40E+00 1.80E+00 4.50E-01 3.60E+03

22 2.39E+05 6.33E+05 8.40E+00 1.81E+00 4.48E-01 3.60E+03

23 2.42E+05 6.56E+05 8.40E+00 1.81E+00 4.42E-01 3.60E+03

24 2.46E+05 6.52E+05 8.40E+00 1.76E+00 4.39E-01 3.60E+03

25 2.49E+05 6.46E+05 8.40E+00 1.71E+00 4.36E-01 3.60E+03

26 2.53E+05 6.37E+05 8.40E+00 1.66E+00 4.34E-01 3.72E+03

27 2.57E+05 6.30E+05 8.40E+00 1.62E+00 4.32E-01 2.40E+03

28 2.59E+05 6.27E+05 8.40E+00 1.60E+00 4.31E-01 1.20E+02
C-10

Surry Unmitigated STSBO

Plume 

Segment PDELAY PLHEAT PLHITE PLMDEN PLMFLA PLUDUR

1 9.19E+04 5.78E+04 8.40E+00 4.84E-01 1.97E-01 3.72E+03

2 9.56E+04 1.49E+05 8.40E+00 4.81E-01 5.03E-01 3.48E+03

3 9.91E+04 2.84E+05 8.40E+00 4.78E-01 9.44E-01 3.72E+03

4 1.03E+05 4.10E+05 8.40E+00 4.76E-01 1.35E+00 3.48E+03

5 1.06E+05 5.07E+05 8.40E+00 4.73E-01 1.64E+00 3.60E+03

6 1.10E+05 5.66E+05 8.40E+00 4.71E-01 1.81E+00 3.60E+03

7 1.14E+05 6.01E+05 8.40E+00 4.68E-01 1.90E+00 3.72E+03

8 1.17E+05 6.23E+05 8.40E+00 4.66E-01 1.95E+00 3.60E+03

9 1.21E+05 6.39E+05 8.40E+00 4.64E-01 1.97E+00 3.60E+03

10 1.24E+05 6.53E+05 8.40E+00 4.62E-01 2.00E+00 3.60E+03

11 1.28E+05 6.63E+05 8.40E+00 4.60E-01 2.01E+00 3.48E+03

12 1.32E+05 6.71E+05 8.40E+00 4.58E-01 2.01E+00 3.60E+03

13 1.35E+05 6.78E+05 8.40E+00 4.56E-01 2.01E+00 3.72E+03

14 1.39E+05 6.84E+05 8.40E+00 4.54E-01 2.01E+00 3.48E+03

15 1.42E+05 6.90E+05 8.40E+00 4.53E-01 2.01E+00 3.60E+03

16 1.46E+05 6.96E+05 8.40E+00 4.51E-01 2.01E+00 3.72E+03

17 1.50E+05 7.04E+05 8.40E+00 4.49E-01 2.02E+00 3.60E+03

18 1.53E+05 7.09E+05 8.40E+00 4.48E-01 2.02E+00 3.48E+03

19 1.57E+05 7.26E+05 8.40E+00 4.44E-01 2.01E+00 3.60E+03

20 1.60E+05 7.34E+05 8.40E+00 4.38E-01 1.96E+00 3.60E+03

21 1.64E+05 7.26E+05 8.40E+00 4.35E-01 1.90E+00 3.72E+03

22 1.68E+05 7.16E+05 8.40E+00 4.32E-01 1.84E+00 3.48E+03

23 1.71E+05 7.08E+05 8.40E+00 4.30E-01 1.80E+00 1.68E+03

24 1.73E+05 7.05E+05 8.40E+00 4.30E-01 1.78E+00 1.20E+02
C-11

Surry ISLOCA

Plume Segment PDELAY PLHEAT PLHITE PLMDEN PLMFLA PLUDUR

1 4.59E+04 3.65E+05 2.17E+01 1.07E+00 1.83E+01 4.20E+03

2 4.59E+04 3.65E+05 2.17E+01 1.07E+00 1.83E+01 4.20E+03

3 4.83E+04 7.72E+04 7.90E+00 8.94E-01 1.05E+00 4.20E+03

4 4.83E+04 2.66E+05 5.64E+00 5.77E-01 6.65E-01 4.20E+03

5 4.86E+04 5.34E+03 6.39E-01 4.54E-01 1.13E-02 3.90E+03

6 4.86E+04 7.64E+05 4.91E+00 5.51E-01 1.75E+00 3.90E+03

7 5.01E+04 5.28E+05 4.91E+00 5.10E-01 1.66E+00 3.00E+02

8 5.01E+04 6.68E+05 2.17E+01 1.00E+00 1.68E+01 3.30E+03

9 5.01E+04 6.68E+05 2.17E+01 1.00E+00 1.68E+01 3.30E+03

10 5.04E+04 4.32E+05 4.91E+00 5.09E-01 1.14E+00 3.00E+03

11 5.25E+04 4.90E+05 4.91E+00 6.22E-01 2.13E+00 3.60E+03

12 5.25E+04 6.64E+04 7.90E+00 9.09E-01 1.23E+00 3.60E+03

13 5.25E+04 1.92E+05 5.64E+00 5.99E-01 8.21E-01 3.60E+03

14 5.34E+04 4.39E+05 4.91E+00 3.72E-01 1.11E+00 3.60E+03

15 5.34E+04 4.47E+05 2.17E+01 1.04E+00 1.76E+01 3.90E+03

16 5.34E+04 4.47E+05 2.17E+01 1.04E+00 1.76E+01 3.90E+03

17 5.61E+04 4.13E+05 4.91E+00 6.57E-01 2.03E+00 3.60E+03

18 5.61E+04 6.61E+04 7.90E+00 9.11E-01 1.21E+00 3.60E+03

19 5.61E+04 1.56E+05 5.64E+00 6.86E-01 8.55E-01 3.60E+03

20 5.70E+04 5.58E+04 4.91E+00 4.25E-01 1.47E-01 3.60E+03

21 5.73E+04 2.91E+05 2.17E+01 1.08E+00 1.75E+01 3.30E+03

22 5.73E+04 2.91E+05 2.17E+01 1.08E+00 1.75E+01 3.30E+03

23 5.97E+04 1.67E+04 4.91E+00 1.07E+00 8.97E-01 3.60E+03

24 5.97E+04 2.33E+04 7.90E+00 1.03E+00 8.08E-01 3.60E+03

25 5.97E+04 1.34E+04 5.64E+00 1.06E+00 6.77E-01 3.60E+03

26 6.06E+04 1.58E+04 2.17E+01 1.15E+00 1.83E+01 3.90E+03

27 6.06E+04 1.58E+04 2.17E+01 1.15E+00 1.83E+01 3.90E+03

28* 6.45E+04 -2.35E+04 2.17E+01 1.16E+00 1.82E+01 3.30E+03

29* 6.45E+04 -2.35E+04 2.17E+01 1.16E+00 1.82E+01 3.30E+03

30 7.02E+04 3.67E+02 4.91E+00 1.14E+00 5.20E-02 3.60E+03

31 7.02E+04 4.58E+03 7.90E+00 1.12E+00 3.38E-01 3.60E+03

32 7.02E+04 4.68E+02 5.64E+00 1.16E+00 2.11E-01 3.60E+03

33 7.17E+04 8.31E+00 1.20E+00 8.27E-01 5.62E-05 3.60E+03

34 7.17E+04 6.12E+04 3.38E+00 1.12E+00 7.11E+00 3.30E+03

35 7.50E+04 5.92E+04 3.38E+00 1.12E+00 7.12E+00 3.90E+03

36 7.53E+04 1.03E+01 1.20E+00 8.56E-01 1.19E-04 3.60E+03

37 7.89E+04 1.30E+01 1.20E+00 8.51E-01 1.39E-04 3.60E+03

38 7.89E+04 5.68E+04 3.38E+00 1.12E+00 7.13E+00 3.60E+03

39 8.10E+04 1.35E+03 4.91E+00 1.12E+00 1.80E-01 3.60E+03

40 8.10E+04 5.49E+02 5.64E+00 1.14E+00 1.38E-01 3.60E+03

* This plume segment is for the Safeguards building exhaust ventilation system and experienced flow 

oscillations during this one-hour time segment. This positive/negative flow rate and heat output ‘averaged’ to a 

negative plume heat contribution (PLHEAT). Since the MACCS2 analysis did not consider plume heat as an 

input (i.e., PLMMOD = DENSITY), the variable PLHEAT was not used in the consequence analysis.
C-12

Surry ISLOCA

Plume Segment PDELAY PLHEAT PLHITE PLMDEN PLMFLA PLUDUR

41 8.22E+04 4.76E+02 4.91E+00 9.18E-01 1.47E-01 3.60E+03

42 8.25E+04 1.63E+01 1.20E+00 8.39E-01 1.50E-04 3.60E+03

43 8.25E+04 5.54E+04 3.38E+00 1.12E+00 7.14E+00 3.30E+03

44 8.46E+04 1.51E+03 4.91E+00 1.12E+00 1.60E-01 3.60E+03

45 8.46E+04 5.91E+02 5.64E+00 1.15E+00 1.52E-01 3.60E+03

46 8.58E+04 5.41E+04 3.38E+00 1.12E+00 7.65E+00 4.20E+03

47 9.00E+04 5.16E+04 3.38E+00 1.12E+00 7.15E+00 3.60E+03

48 9.18E+04 3.12E+03 4.91E+00 1.14E+00 2.60E-01 4.35E+03

49 9.36E+04 2.07E+03 4.91E+00 1.04E+00 6.92E-02 3.60E+03

50 9.36E+04 5.03E+04 3.38E+00 1.12E+00 7.16E+00 3.60E+03

51 9.62E+04 3.43E+03 4.91E+00 1.12E+00 4.74E-01 3.75E+03

52 9.62E+04 1.16E+03 5.64E+00 1.13E+00 1.93E-01 3.75E+03

53 9.72E+04 4.88E+04 3.38E+00 1.12E+00 7.17E+00 3.60E+03

54 1.01E+05 4.70E+04 3.38E+00 1.13E+00 7.17E+00 3.60E+03

55 1.03E+05 3.37E+03 4.91E+00 1.12E+00 3.66E-01 4.50E+03

56 1.04E+05 4.53E+04 3.38E+00 1.13E+00 7.18E+00 3.60E+03

57 1.07E+05 5.20E+03 4.91E+00 1.10E+00 5.49E-01 2.70E+03

58 1.08E+05 4.48E+04 3.38E+00 1.13E+00 7.18E+00 3.60E+03

59 1.10E+05 6.56E+03 4.91E+00 1.11E+00 6.23E-01 4.50E+03

60 1.12E+05 4.27E+04 3.38E+00 1.13E+00 7.19E+00 3.60E+03

61 1.14E+05 8.29E+03 4.91E+00 1.13E+00 6.96E-01 3.60E+03

62 1.15E+05 4.04E+04 3.38E+00 1.13E+00 7.20E+00 3.60E+03

63 1.18E+05 1.41E+04 4.91E+00 1.12E+00 7.79E-01 3.60E+03

64 1.18E+05 4.19E+03 5.64E+00 1.12E+00 2.50E-01 3.60E+03

65 1.19E+05 3.78E+04 3.38E+00 1.13E+00 7.21E+00 3.60E+03

66 1.22E+05 2.59E+04 4.91E+00 1.07E+00 9.14E-01 3.60E+03

67 1.22E+05 7.68E+03 5.64E+00 1.07E+00 3.05E-01 3.60E+03

68 1.22E+05 6.05E+03 4.91E+00 9.36E-01 2.44E-01 3.60E+03

69 1.25E+05 2.95E+04 4.91E+00 1.04E+00 9.47E-01 2.70E+03

70 1.25E+05 8.85E+03 5.64E+00 1.02E+00 3.20E-01 2.70E+03

71 1.26E+05 9.16E+03 4.91E+00 8.37E-01 2.63E-01 3.60E+03

72 1.28E+05 3.05E+04 4.91E+00 1.08E+00 9.52E-01 4.50E+03

73 1.28E+05 5.90E+03 7.90E+00 1.10E+00 3.51E-01 4.50E+03

74 1.28E+05 9.19E+03 5.64E+00 1.06E+00 3.22E-01 4.50E+03

75 1.30E+05 1.78E+04 4.91E+00 8.29E-01 3.61E-01 3.60E+03

76 1.30E+05 3.40E+04 3.38E+00 1.14E+00 7.23E+00 3.60E+03

77 1.32E+05 4.56E+04 4.91E+00 1.02E+00 1.06E+00 2.70E+03

78 1.32E+05 8.92E+03 7.90E+00 1.09E+00 4.36E-01 2.70E+03

79 1.33E+05 3.27E+04 4.91E+00 8.54E-01 6.85E-01 3.60E+03

80 1.35E+05 4.78E+04 4.91E+00 1.06E+00 1.10E+00 3.60E+03

81 1.35E+05 1.02E+04 7.90E+00 1.08E+00 4.68E-01 3.60E+03

82 1.35E+05 1.51E+04 5.64E+00 1.01E+00 3.96E-01 3.60E+03

83 1.37E+05 3.49E+04 4.91E+00 8.44E-01 7.07E-01 3.60E+03
C-13

Surry ISLOCA

Plume Segment PDELAY PLHEAT PLHITE PLMDEN PLMFLA PLUDUR

84 1.39E+05 4.88E+04 4.91E+00 1.09E+00 1.11E+00 3.60E+03

85 1.39E+05 1.56E+04 5.64E+00 1.07E+00 4.05E-01 3.60E+03

86 1.40E+05 3.80E+04 4.91E+00 8.90E-01 7.59E-01 3.60E+03

87 1.42E+05 5.31E+04 4.91E+00 1.05E+00 1.16E+00 4.50E+03

88 1.42E+05 1.72E+04 5.64E+00 1.02E+00 4.22E-01 4.50E+03

89 1.44E+05 4.67E+04 4.91E+00 8.54E-01 9.27E-01 3.60E+03

90 1.47E+05 5.61E+04 4.91E+00 1.01E+00 1.19E+00 2.70E+03

91 1.48E+05 5.53E+04 4.91E+00 8.60E-01 1.09E+00 3.60E+03

92 1.49E+05 5.79E+04 4.91E+00 1.06E+00 1.22E+00 3.60E+03

93 1.51E+05 5.47E+04 4.91E+00 8.62E-01 1.08E+00 3.60E+03

94 1.53E+05 5.64E+04 4.91E+00 1.08E+00 1.21E+00 3.60E+03

95 1.55E+05 5.22E+04 4.91E+00 8.62E-01 1.04E+00 3.60E+03

96 1.57E+05 5.42E+04 4.91E+00 1.10E+00 1.19E+00 3.60E+03

97 1.58E+05 5.03E+04 4.91E+00 8.64E-01 1.02E+00 3.60E+03

98 1.60E+05 4.98E+04 4.91E+00 1.05E+00 1.14E+00 4.50E+03

99 1.60E+05 1.27E+04 7.90E+00 1.08E+00 5.15E-01 4.50E+03

100 1.60E+05 1.65E+04 5.64E+00 1.03E+00 4.25E-01 4.50E+03

101 1.62E+05 3.94E+04 4.91E+00 8.64E-01 8.23E-01 3.60E+03

102 1.62E+05 2.78E+04 3.38E+00 1.14E+00 7.27E+00 3.60E+03

103 1.65E+05 3.89E+04 4.91E+00 1.08E+00 9.95E-01 2.70E+03

104 1.65E+05 1.06E+04 7.90E+00 1.08E+00 4.70E-01 2.70E+03

105 1.66E+05 2.81E+04 4.91E+00 8.65E-01 6.22E-01 3.60E+03

106 1.67E+05 3.70E+04 4.91E+00 1.11E+00 9.62E-01 3.60E+03

107 1.69E+05 2.69E+04 4.91E+00 8.69E-01 6.18E-01 3.60E+03
C-14

Surry Mitigated TISGTR

Plume 

Segment PDELAY PLHEAT PLHITE PLMDEN PLMFLA PLUDUR

1 10600 4.32E+01 8.4 0.73919 4.11E-04 3619.6

2 12800 8.88E+05 24.643 0.29498 1.0716 3639.9

3 14220 8.97E+01 8.4 0.6666 7.40E-04 3600.4

4 16440 6.76E+04 24.643 0.3208 0.091864 3600.1

5 17820 1.12E+02 8.4 0.64171 8.13E-04 3599.9

6 20040 6.10E+04 24.643 0.2937 0.08302 3599.9

7 21420 1.28E+02 8.4 0.62148 8.23E-04 3600

8 23640 5.47E+04 24.643 0.28461 0.074318 3600

9 25020 1.48E+02 8.4 0.60385 8.97E-04 3599.9

10 27240 5.95E+04 24.643 0.33478 0.088723 3600.1

11 28620 1.21E+02 8.4 0.6199 8.26E-04 3660.1

12 30840 4.60E+04 24.643 0.33994 0.069402 3600.1

13 34440 2.31E+04 24.643 0.3649 0.039585 3599.9

14 38040 7.49E+03 24.643 0.40498 0.01627 3600

15 41640 3.83E+02 24.643 0.71955 0.002878 3600

16 52440 1.49E+03 24.643 0.71162 0.007903 3600

17 56040 1.24E+04 24.643 0.50849 0.036427 3599.9

18 59640 2.03E+04 24.643 0.42415 0.045113 3600.2

19 63240 2.50E+04 24.643 0.41032 0.05232 3599.9

20 66840 2.88E+04 24.643 0.40342 0.058182 3600

21 70440 3.23E+04 24.643 0.39798 0.063655 3600.1

22 74040 3.58E+04 24.643 0.39336 0.069014 3599.8

23 77640 3.92E+04 24.643 0.38949 0.074187 3600

24 81240 4.26E+04 24.643 0.38619 0.079248 3600.2

25 84840 4.41E+04 24.643 0.38476 0.081656 3599.9

26 88440 4.63E+04 24.643 0.3831 0.085028 3600

27 92040 4.86E+04 24.643 0.38153 0.088779 3599.9

28 95640 5.10E+04 24.643 0.38011 0.09261 3600

29 99240 5.34E+04 24.643 0.37894 0.096455 3600.1

30 1.03E+05 5.58E+04 24.643 0.37793 0.10035 3600.1

31 1.06E+05 5.83E+04 24.643 0.37704 0.10438 3600

32 1.10E+05 6.08E+04 24.643 0.3762 0.10847 3599.9

33 1.14E+05 6.33E+04 24.643 0.37533 0.11268 3600.1

34 1.17E+05 6.58E+04 24.643 0.37489 0.11697 3600.1

35 1.21E+05 6.83E+04 24.643 0.37457 0.12131 3599.8

36 1.24E+05 7.09E+04 24.643 0.37437 0.12571 3600

37 1.28E+05 7.34E+04 24.643 0.37426 0.13016 3600.1

38 1.32E+05 7.59E+04 24.643 0.37422 0.13468 3600

39 1.35E+05 7.84E+04 24.643 0.37428 0.13924 3600

40 1.39E+05 8.09E+04 24.643 0.37442 0.14384 3599.9

41 1.42E+05 8.33E+04 24.643 0.37464 0.14848 3600.2

42 1.46E+05 8.58E+04 24.643 0.37494 0.15318 3599.9

43 1.50E+05 8.82E+04 24.643 0.3753 0.15791 3600.1
C-15

Surry Mitigated TISGTR

Plume 

Segment PDELAY PLHEAT PLHITE PLMDEN PLMFLA PLUDUR

44 1.53E+05 9.06E+04 24.643 0.37573 0.16268 3599.9

45 1.57E+05 9.30E+04 24.643 0.37623 0.1675 3599.8

46 1.60E+05 9.53E+04 24.643 0.37687 0.17236 3600.3

47 1.64E+05 9.76E+04 24.643 0.37757 0.17724 3599.7

48 1.68E+05 9.98E+04 24.643 0.37832 0.18216 3600.1

49 1.71E+05 1.01E+05 24.643 0.37889 0.1857 1560.1
C-16

Surry Unmitigated TISGTR

Plume 

Segment PDELAY PLHEAT PLHITE PLMDEN PLMFLA PLUDUR

1 10600 4.32E+01 8.4 7.39E-01 4.11E-04 3619.6

2 12800 8.88E+05 24.643 2.95E-01 1.07E+00 3639.9

3 14220 8.97E+01 8.4 6.67E-01 7.40E-04 3600.4

4 16440 6.76E+04 24.643 3.21E-01 9.19E-02 3600.1

5 17820 1.12E+02 8.4 6.42E-01 8.13E-04 3599.9

6 20040 6.10E+04 24.643 2.94E-01 8.30E-02 3599.9

7 21420 1.28E+02 8.4 6.21E-01 8.23E-04 3600

8 23640 5.47E+04 24.643 2.85E-01 7.43E-02 3600

9 25020 1.48E+02 8.4 6.04E-01 8.97E-04 3599.9

10 27240 6.21E+04 24.643 3.45E-01 9.66E-02 3600

11 28620 2.00E+02 8.4 5.80E-01 1.08E-03 3660

12 30840 6.41E+04 24.643 3.64E-01 1.06E-01 3600

13 32280 2.26E+02 8.4 5.66E-01 1.12E-03 3600.2

14 34440 6.88E+04 24.643 3.62E-01 1.12E-01 3600.1

15 35880 2.45E+02 8.4 5.60E-01 1.18E-03 3600

16 38040 7.38E+04 24.643 3.58E-01 1.18E-01 3600.1

17 39480 2.63E+02 8.4 5.55E-01 1.24E-03 3600

18 41640 7.87E+04 24.643 3.55E-01 1.24E-01 3599.9

19 43080 2.85E+02 8.4 5.49E-01 1.31E-03 3600

20 45240 8.38E+04 24.643 3.53E-01 1.29E-01 3600.1

21 46680 3.07E+02 8.4 5.43E-01 1.37E-03 3600.1

22 48840 8.93E+04 24.643 3.55E-01 1.36E-01 3600.1

23 50280 3.30E+02 8.4 5.39E-01 1.43E-03 3600

24 52440 9.51E+04 24.643 3.52E-01 1.42E-01 3599.8

25 53880 3.54E+02 8.4 5.35E-01 1.51E-03 3600

26 56040 1.01E+05 24.643 3.48E-01 1.49E-01 3600

27 57480 3.80E+02 8.4 5.31E-01 1.60E-03 3599.9

28 59640 1.08E+05 24.643 3.45E-01 1.55E-01 3600

29 61080 4.06E+02 8.4 5.28E-01 1.66E-03 3600

30 63240 1.15E+05 24.643 3.42E-01 1.63E-01 3600.1

31 64680 4.33E+02 8.4 5.24E-01 1.74E-03 3599.8

32 66840 1.22E+05 24.643 3.39E-01 1.70E-01 3600

33 68280 4.62E+02 8.4 5.21E-01 1.83E-03 3600.3

34 70440 1.29E+05 24.643 3.36E-01 1.78E-01 3599.9

35 71880 4.91E+02 8.4 5.18E-01 1.92E-03 3599.9

36 74040 1.37E+05 24.643 3.33E-01 1.86E-01 3600.1

37 75480 5.21E+02 8.4 5.14E-01 2.01E-03 3600

38 77640 1.45E+05 24.643 3.31E-01 1.95E-01 3600

39 79080 5.53E+02 8.4 5.11E-01 2.10E-03 3600

40 81240 1.53E+05 24.643 3.29E-01 2.03E-01 3600.1

41 82680 5.84E+02 8.4 5.08E-01 2.18E-03 3599.8

42 84840 1.61E+05 24.643 3.27E-01 2.12E-01 3599.8

43 86280 6.16E+02 8.4 5.05E-01 2.27E-03 3600.3
C-17

Surry Unmitigated TISGTR

Plume 

Segment PDELAY PLHEAT PLHITE PLMDEN PLMFLA PLUDUR

44 88440 1.68E+05 24.643 3.26E-01 2.20E-01 3600.2

45 89880 6.49E+02 8.4 5.02E-01 2.36E-03 3600

46 92040 1.76E+05 24.643 3.24E-01 2.29E-01 3599.9

47 93480 6.82E+02 8.4 4.99E-01 2.45E-03 3599.8

48 95640 1.85E+05 24.643 3.23E-01 2.38E-01 3599.8

49 97080 7.16E+02 8.4 4.97E-01 2.55E-03 3600

50 99240 1.93E+05 24.643 3.22E-01 2.47E-01 3600.3

51 1.00E+05 4.91E+04 8.4 4.94E-01 1.72E-01 3720

52 1.01E+05 7.49E+02 8.4 4.94E-01 2.63E-03 3600

53 1.03E+05 2.00E+05 24.643 3.21E-01 2.55E-01 3600

54 1.04E+05 1.57E+05 8.4 4.91E-01 5.45E-01 3480.1

55 1.04E+05 7.79E+02 8.4 4.91E-01 2.71E-03 3600.1

56 1.06E+05 2.06E+05 24.643 3.20E-01 2.62E-01 3600

57 1.08E+05 2.85E+05 8.4 4.89E-01 9.81E-01 3720

58 1.08E+05 8.02E+02 8.4 4.89E-01 2.76E-03 3600.1

59 1.10E+05 2.11E+05 24.643 3.20E-01 2.67E-01 3600

60 1.11E+05 4.00E+05 8.4 4.87E-01 1.36E+00 3599.9

61 1.11E+05 8.20E+02 8.4 4.87E-01 2.79E-03 3600

62 1.14E+05 2.13E+05 24.643 3.19E-01 2.70E-01 3600

63 1.15E+05 4.77E+05 8.4 4.85E-01 1.61E+00 3480.3

64 1.15E+05 8.33E+02 8.4 4.85E-01 2.79E-03 3561.3

65 1.17E+05 2.15E+05 24.643 3.18E-01 2.71E-01 3600

66 1.18E+05 5.32E+05 8.4 4.83E-01 1.77E+00 3719.9

67 1.19E+05 8.45E+02 8.4 4.83E-01 2.86E-03 3638.5

68 1.21E+05 2.17E+05 24.643 3.18E-01 2.72E-01 3599.8

69 1.22E+05 5.65E+05 8.4 4.81E-01 1.86E+00 3599.8

70 1.22E+05 8.53E+02 8.4 4.81E-01 2.82E-03 3600.1

71 1.24E+05 2.18E+05 24.643 3.17E-01 2.73E-01 3600.2

72 1.26E+05 5.85E+05 8.4 4.79E-01 1.91E+00 3600.2

73 1.26E+05 8.61E+02 8.4 4.79E-01 2.81E-03 3600.1

74 1.28E+05 2.19E+05 24.643 3.17E-01 2.73E-01 3599.9

75 1.29E+05 5.98E+05 8.4 4.77E-01 1.94E+00 3480

76 1.29E+05 8.68E+02 8.4 4.77E-01 2.81E-03 3600

77 1.32E+05 2.19E+05 24.643 3.16E-01 2.74E-01 3600

78 1.33E+05 6.08E+05 8.4 4.75E-01 1.95E+00 3600.1

79 1.33E+05 8.74E+02 8.4 4.75E-01 2.81E-03 3599.7

80 1.35E+05 2.20E+05 24.643 3.16E-01 2.74E-01 3600.1

81 1.36E+05 6.17E+05 8.4 4.73E-01 1.97E+00 3599.9

82 1.37E+05 8.80E+02 8.4 4.73E-01 2.80E-03 3600.1

83 1.39E+05 2.20E+05 24.643 3.16E-01 2.74E-01 3600

84 1.40E+05 6.26E+05 8.4 4.72E-01 1.98E+00 3720.1

85 1.40E+05 8.87E+02 8.4 4.72E-01 2.80E-03 3600

86 1.42E+05 2.21E+05 24.643 3.15E-01 2.74E-01 3599.9
C-18

Surry Unmitigated TISGTR

Plume 

Segment PDELAY PLHEAT PLHITE PLMDEN PLMFLA PLUDUR

87 1.44E+05 6.60E+05 8.4 4.67E-01 2.03E+00 3479.8

88 1.44E+05 9.08E+02 8.4 4.67E-01 2.78E-03 3600.1

89 1.46E+05 2.19E+05 24.643 3.15E-01 2.71E-01 3599.8

90 1.47E+05 6.75E+05 8.4 4.62E-01 2.01E+00 3720.1

91 1.47E+05 9.18E+02 8.4 4.62E-01 2.73E-03 3600

92 1.50E+05 2.15E+05 24.643 3.15E-01 2.67E-01 3600.3

93 1.51E+05 6.74E+05 8.4 4.60E-01 1.97E+00 3479.9

94 1.51E+05 9.15E+02 8.4 4.60E-01 2.67E-03 3600

95 1.53E+05 2.12E+05 24.643 3.14E-01 2.62E-01 3599.7

96 1.54E+05 6.72E+05 8.4 4.57E-01 1.94E+00 3600

97 1.55E+05 9.11E+02 8.4 4.57E-01 2.62E-03 3600

98 1.57E+05 2.08E+05 24.643 3.14E-01 2.57E-01 3600.4

99 1.58E+05 6.69E+05 8.4 4.55E-01 1.90E+00 3600.2

100 1.58E+05 9.05E+02 8.4 4.55E-01 2.57E-03 3600.1

101 1.60E+05 2.05E+05 24.643 3.14E-01 2.52E-01 3599.6

102 1.62E+05 6.65E+05 8.4 4.53E-01 1.86E+00 3719.9

103 1.62E+05 8.98E+02 8.4 4.53E-01 2.51E-03 3599.8

104 1.64E+05 2.01E+05 24.643 3.14E-01 2.48E-01 3600

105 1.65E+05 6.60E+05 8.4 4.51E-01 1.83E+00 3480

106 1.65E+05 8.90E+02 8.4 4.51E-01 2.46E-03 3600.3

107 1.68E+05 1.98E+05 24.643 3.13E-01 2.43E-01 3600.3

108 1.69E+05 6.55E+05 8.4 4.49E-01 1.79E+00 3600

109 1.69E+05 8.82E+02 8.4 4.49E-01 2.41E-03 3599.8

110 1.71E+05 1.95E+05 24.643 3.13E-01 2.40E-01 1559.9

111 1.72E+05 6.52E+05 8.4 4.49E-01 1.77E+00 359.91

112 1.73E+05 8.77E+02 8.4 4.48E-01 2.38E-03 120

113 1.73E+05 6.52E+05 8.4 4.48E-01 1.77E+00 119.89

114 1.73E+05 8.78E+02 8.4 4.48E-01 2.38E-03 119.89

115 1.73E+05 1.95E+05 24.643 3.13E-01 2.38E-01 119.89
C-19

Table C-3 Plume Parameters Used in the Surry LTSBO, Unmitigated STSBO, TISGTR Mitigated and Unmitigated Scenarios

Surry LTSBO

Class Bin 1 Bin 2 Bin 3 Bin 4 Bin 5 Bin 6 Bin 7 Bin 8 Bin 9 Bin 10

Xe 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01

Cs 1.73E-03 1.26E-02 6.19E-02 2.00E-01 3.36E-01 2.52E-01 1.09E-01 2.37E-02 2.32E-03 1.05E-03

Ba 6.94E-03 3.56E-02 1.35E-01 3.69E-01 3.45E-01 8.98E-02 1.49E-02 2.50E-03 2.54E-04 3.58E-04

I 6.47E-03 3.22E-02 1.21E-01 3.28E-01 3.57E-01 1.32E-01 1.85E-02 1.95E-03 3.19E-04 1.44E-03

Te 7.53E-03 3.45E-02 1.31E-01 3.49E-01 3.40E-01 1.14E-01 1.86E-02 2.44E-03 2.44E-04 1.34E-03

Ru 8.80E-03 3.73E-02 1.35E-01 3.24E-01 3.13E-01 1.27E-01 2.71E-02 1.02E-02 3.38E-03 1.40E-02

Mo 2.39E-04 3.89E-03 2.78E-02 1.01E-01 2.67E-01 3.30E-01 1.98E-01 6.34E-02 8.46E-03 3.76E-04

Ce 7.52E-03 3.22E-02 1.15E-01 2.85E-01 3.34E-01 1.73E-01 3.59E-02 7.97E-03 1.65E-03 8.02E-03

La 4.89E-03 2.37E-02 9.23E-02 2.54E-01 3.44E-01 2.04E-01 6.15E-02 1.20E-02 1.34E-03 2.86E-03

Surry Unmitigated STSBO

Class Bin 1 Bin 2 Bin 3 Bin 4 Bin 5 Bin 6 Bin 7 Bin 8 Bin 9 Bin 10

Xe 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01

Cs 1.22E-03 1.01E-02 5.26E-02 1.73E-01 3.33E-01 2.86E-01 1.21E-01 2.05E-02 1.15E-03 1.19E-03

Ba 7.02E-03 3.61E-02 1.42E-01 3.43E-01 3.18E-01 1.19E-01 2.65E-02 3.58E-03 2.58E-04 4.40E-03

I 6.19E-03 2.94E-02 1.06E-01 2.66E-01 3.44E-01 2.03E-01 3.96E-02 2.54E-03 1.97E-04 2.36E-03

Te 4.03E-03 2.31E-02 9.45E-02 2.70E-01 3.62E-01 1.94E-01 4.57E-02 4.65E-03 2.33E-04 1.50E-03

Ru 5.15E-03 2.69E-02 1.07E-01 2.75E-01 3.43E-01 1.80E-01 4.05E-02 6.86E-03 1.67E-03 1.36E-02

Mo 2.51E-04 4.22E-03 3.10E-02 1.14E-01 2.91E-01 3.42E-01 1.78E-01 3.66E-02 2.37E-03 9.36E-05

Ce 5.06E-03 2.57E-02 9.95E-02 2.57E-01 3.41E-01 2.04E-01 4.92E-02 6.85E-03 1.30E-03 9.52E-03

La 3.14E-03 1.80E-02 7.61E-02 2.18E-01 3.44E-01 2.45E-01 8.08E-02 1.18E-02 7.95E-04 2.68E-03
C-20

Surry ISLOCA

Class Bin 1 Bin 2 Bin 3 Bin 4 Bin 5 Bin 6 Bin 7 Bin 8 Bin 9 Bin 10

Xe 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01

Cs 9.37E-03 2.56E-02 6.96E-02 2.73E-01 4.01E-01 1.78E-01 4.04E-02 2.62E-03 4.33E-05 5.48E-04

Ba 2.23E-02 4.72E-02 1.38E-01 2.92E-01 3.24E-01 1.41E-01 3.05E-02 2.46E-03 7.74E-05 2.58E-03

I 9.15E-03 2.41E-02 7.56E-02 2.81E-01 3.96E-01 1.72E-01 3.85E-02 2.53E-03 4.28E-05 6.21E-04

Te 1.28E-02 3.01E-02 8.65E-02 2.83E-01 3.82E-01 1.65E-01 3.72E-02 2.50E-03 4.58E-05 5.42E-04

Ru 1.13E-02 3.05E-02 9.69E-02 2.91E-01 3.75E-01 1.59E-01 3.34E-02 2.13E-03 3.54E-05 4.80E-04

Mo 1.02E-02 2.46E-02 6.13E-02 2.34E-01 3.80E-01 2.16E-01 6.41E-02 8.89E-03 7.41E-04 3.93E-04

Ce 7.90E-03 3.31E-02 2.20E-01 4.00E-01 1.92E-01 8.59E-02 4.09E-02 1.27E-02 1.26E-03 6.03E-03

La 2.07E-02 4.49E-02 2.46E-01 3.92E-01 1.74E-01 7.55E-02 3.43E-02 9.34E-03 8.08E-04 2.15E-03

.

Surry Mitigated TISGTR

Class Bin 1 Bin 2 Bin 3 Bin 4 Bin 5 Bin 6 Bin 7 Bin 8 Bin 9 Bin 10

Xe 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01

Cs 2.39E-02 3.93E-02 6.33E-02 2.06E-01 3.47E-01 2.16E-01 7.68E-02 2.30E-02 4.35E-03 6.63E-04

Ba 1.29E-02 1.87E-02 5.47E-02 2.44E-01 4.15E-01 1.96E-01 4.50E-02 1.17E-02 2.38E-03 3.65E-04

I 3.45E-02 4.99E-02 6.35E-02 2.09E-01 3.47E-01 2.03E-01 6.82E-02 2.00E-02 3.72E-03 5.92E-04

Te 1.50E-02 2.65E-02 5.69E-02 2.17E-01 3.69E-01 2.16E-01 7.29E-02 2.20E-02 4.16E-03 6.11E-04

Ru 1.60E-03 5.33E-03 3.94E-02 2.20E-01 4.24E-01 2.24E-01 6.31E-02 1.87E-02 3.45E-03 4.43E-04

Mo 1.61E-02 5.22E-02 6.90E-02 2.98E-01 4.12E-01 1.16E-01 2.79E-02 6.85E-03 1.28E-03 1.92E-04

Ce 1.27E-03 8.45E-03 7.31E-02 2.80E-01 4.20E-01 1.79E-01 2.96E-02 7.03E-03 2.00E-03 2.70E-04

La 8.71E-03 2.82E-02 8.14E-02 2.99E-01 4.07E-01 1.45E-01 2.36E-02 5.32E-03 1.43E-03 1.89E-04
C-21

Surry Unmitigated TISGTR

Class Bin 1 Bin 2 Bin 3 Bin 4 Bin 5 Bin 6 Bin 7 Bin 8 Bin 9 Bin 10

Xe 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01 1.00E-01

Cs 4.84E-03 1.55E-02 5.85E-02 1.91E-01 4.04E-01 2.69E-01 5.05E-02 5.54E-03 7.83E-04 1.17E-04

Ba 6.37E-03 2.58E-02 8.09E-02 2.35E-01 3.81E-01 2.25E-01 4.35E-02 2.94E-03 2.34E-04 4.00E-05

I 1.14E-02 2.75E-02 6.76E-02 2.02E-01 3.54E-01 2.49E-01 7.56E-02 1.12E-02 1.36E-03 2.20E-04

Te 8.56E-03 2.61E-02 7.12E-02 2.00E-01 3.56E-01 2.59E-01 7.19E-02 7.63E-03 6.94E-04 1.06E-04

Ru 1.80E-03 6.34E-03 4.20E-02 2.21E-01 4.19E-01 2.23E-01 6.39E-02 1.89E-02 3.48E-03 4.41E-04

Mo 3.95E-03 1.86E-02 6.60E-02 1.76E-01 3.29E-01 2.92E-01 1.04E-01 1.05E-02 3.31E-04 8.27E-05

Ce 4.18E-03 2.17E-02 9.15E-02 2.48E-01 3.59E-01 2.03E-01 5.65E-02 1.35E-02 2.38E-03 2.42E-04

La 2.27E-03 1.47E-02 6.50E-02 2.00E-01 3.99E-01 2.64E-01 5.12E-02 3.73E-03 2.51E-04 2.79E-05
C-22

Table C-4 Release Fraction Parameters Used in the Surry LTSBO, Unmitigated 

STSBO, TI-SGTR Mitigated and Unmitigated Scenarios

Surry LTSBO

Plume 

Segment Xe Cs Ba I Te Ru Mo Ce La

1 3.36E-03 3.50E-06 5.31E-07 3.62E-05 5.09E-05 7.77E-08 2.69E-07 1.93E-07 1.21E-08

2 7.20E-03 7.08E-06 1.03E-06 7.43E-05 1.04E-04 1.52E-07 5.22E-07 3.77E-07 2.42E-08

3 1.27E-02 1.17E-05 1.64E-06 1.24E-04 1.75E-04 2.44E-07 8.35E-07 6.06E-07 3.96E-08

4 1.80E-02 1.55E-05 2.17E-06 1.66E-04 2.36E-04 3.18E-07 1.09E-06 7.92E-07 5.28E-08

5 2.24E-02 1.79E-05 2.51E-06 1.94E-04 2.79E-04 3.63E-07 1.23E-06 9.05E-07 6.15E-08

6 2.35E-02 1.76E-05 2.47E-06 1.91E-04 2.78E-04 3.50E-07 1.19E-06 8.74E-07 6.07E-08

7 2.55E-02 1.79E-05 2.52E-06 1.96E-04 2.87E-04 3.50E-07 1.18E-06 8.75E-07 6.20E-08

8 2.59E-02 1.71E-05 2.41E-06 1.88E-04 2.79E-04 3.27E-07 1.10E-06 8.20E-07 5.94E-08

9 2.58E-02 1.60E-05 2.27E-06 1.78E-04 2.66E-04 3.02E-07 1.01E-06 7.57E-07 5.62E-08

10 2.54E-02 1.49E-05 2.12E-06 1.67E-04 2.54E-04 2.76E-07 9.16E-07 6.93E-07 5.27E-08

11 2.50E-02 1.39E-05 1.98E-06 1.57E-04 2.42E-04 2.51E-07 8.29E-07 6.33E-07 4.94E-08

12 2.45E-02 1.29E-05 1.86E-06 1.48E-04 2.32E-04 2.29E-07 7.50E-07 5.79E-07 4.64E-08

13 2.42E-02 1.21E-05 1.75E-06 1.40E-04 2.25E-04 2.10E-07 6.81E-07 5.32E-07 4.38E-08

14 2.34E-02 1.12E-05 1.62E-06 1.30E-04 2.16E-04 1.90E-07 6.09E-07 4.82E-07 4.08E-08

15 2.27E-02 1.04E-05 1.52E-06 1.22E-04 2.09E-04 1.72E-07 5.45E-07 4.37E-07 3.81E-08

16 2.21E-02 9.69E-06 1.42E-06 1.15E-04 2.03E-04 1.56E-07 4.88E-07 3.98E-07 3.59E-08

17 2.16E-02 9.11E-06 1.35E-06 1.09E-04 2.00E-04 1.43E-07 4.40E-07 3.66E-07 3.40E-08

18 2.11E-02 8.57E-06 1.27E-06 1.04E-04 1.98E-04 1.31E-07 3.96E-07 3.36E-07 3.22E-08

19 2.06E-02 8.08E-06 1.21E-06 9.86E-05 1.98E-04 1.20E-07 3.55E-07 3.09E-07 3.07E-08

20 2.00E-02 7.60E-06 1.15E-06 9.37E-05 1.97E-04 1.09E-07 3.17E-07 2.83E-07 2.91E-08

21 1.94E-02 7.17E-06 1.09E-06 8.93E-05 1.97E-04 1.00E-07 2.83E-07 2.60E-07 2.76E-08

22 1.90E-02 6.81E-06 1.04E-06 8.56E-05 1.99E-04 9.19E-08 2.54E-07 2.40E-07 2.64E-08

23 1.87E-02 6.57E-06 9.94E-07 8.34E-05 2.04E-04 8.55E-08 2.31E-07 2.24E-07 2.56E-08

24 1.81E-02 6.23E-06 9.12E-07 7.96E-05 2.08E-04 7.79E-08 2.06E-07 2.05E-07 2.44E-08

25 1.75E-02 5.91E-06 8.37E-07 7.63E-05 2.17E-04 7.08E-08 1.83E-07 1.88E-07 2.32E-08

26 1.75E-02 5.82E-06 7.95E-07 7.57E-05 2.40E-04 6.64E-08 1.68E-07 1.77E-07 2.28E-08

27 1.10E-02 3.62E-06 4.78E-07 4.73E-05 1.67E-04 3.94E-08 9.86E-08 1.06E-07 1.41E-08

28 5.41E-04 1.78E-07 2.32E-08 2.33E-06 8.68E-06 1.90E-09 4.73E-09 5.10E-09 6.93E-10
C-23

Surry Unmitigated STSBO

Plume 

Segment Xe Cs Ba I Te Ru Mo Ce La

1 3.86E-03 6.41E-06 1.13E-06 5.54E-05 8.81E-05 1.60E-07 6.91E-07 2.84E-07 1.18E-08

2 8.94E-03 1.38E-05 2.41E-06 1.24E-04 1.85E-04 3.30E-07 1.41E-06 5.85E-07 2.51E-08

3 1.74E-02 2.53E-05 4.36E-06 2.35E-04 3.28E-04 5.74E-07 2.42E-06 1.02E-06 4.52E-08

4 2.26E-02 3.11E-05 5.28E-06 3.00E-04 3.88E-04 6.68E-07 2.79E-06 1.19E-06 5.43E-08

5 2.77E-02 3.66E-05 6.12E-06 3.64E-04 4.39E-04 7.39E-07 3.05E-06 1.32E-06 6.21E-08

6 2.96E-02 3.78E-05 6.20E-06 3.89E-04 4.33E-04 7.14E-07 2.91E-06 1.28E-06 6.21E-08

7 3.10E-02 3.83E-05 6.20E-06 4.05E-04 4.21E-04 6.79E-07 2.73E-06 1.21E-06 6.13E-08

8 2.98E-02 3.50E-05 5.70E-06 3.78E-04 3.77E-04 5.92E-07 2.35E-06 1.06E-06 5.56E-08

9 2.93E-02 3.30E-05 5.41E-06 3.64E-04 3.47E-04 5.31E-07 2.08E-06 9.54E-07 5.19E-08

10 2.86E-02 3.12E-05 5.14E-06 3.51E-04 3.20E-04 4.76E-07 1.83E-06 8.58E-07 4.85E-08

11 2.69E-02 2.85E-05 4.72E-06 3.27E-04 2.84E-04 4.12E-07 1.55E-06 7.44E-07 4.38E-08

12 2.70E-02 2.79E-05 4.64E-06 3.27E-04 2.71E-04 3.81E-07 1.40E-06 6.90E-07 4.23E-08

13 2.70E-02 2.74E-05 4.52E-06 3.27E-04 2.58E-04 3.52E-07 1.27E-06 6.38E-07 4.09E-08

14 2.44E-02 2.45E-05 4.01E-06 2.97E-04 2.23E-04 2.95E-07 1.04E-06 5.36E-07 3.59E-08

15 2.44E-02 2.43E-05 3.95E-06 2.99E-04 2.14E-04 2.74E-07 9.45E-07 5.00E-07 3.50E-08

16 2.44E-02 2.42E-05 3.90E-06 3.02E-04 2.06E-04 2.54E-07 8.57E-07 4.65E-07 3.41E-08

17 2.28E-02 2.26E-05 3.62E-06 2.86E-04 1.86E-04 2.21E-07 7.32E-07 4.07E-07 3.13E-08

18 2.13E-02 2.07E-05 3.29E-06 2.64E-04 1.69E-04 1.93E-07 6.25E-07 3.56E-07 2.87E-08

19 2.14E-02 1.99E-05 3.12E-06 2.55E-04 1.65E-04 1.82E-07 5.78E-07 3.37E-07 2.84E-08

20 2.08E-02 1.84E-05 2.81E-06 2.37E-04 1.56E-04 1.65E-07 5.15E-07 3.07E-07 2.72E-08

21 2.07E-02 1.74E-05 2.60E-06 2.25E-04 1.52E-04 1.53E-07 4.70E-07 2.87E-07 2.67E-08

22 1.87E-02 1.50E-05 2.24E-06 1.95E-04 1.33E-04 1.29E-07 3.90E-07 2.43E-07 2.38E-08

23 8.79E-03 6.81E-06 1.06E-06 8.87E-05 6.16E-05 5.77E-08 1.73E-07 1.09E-07 1.11E-08

24 6.22E-04 4.76E-07 7.62E-08 6.21E-06 4.33E-06 4.01E-09 1.20E-08 7.59E-09 7.84E-10
C-24

Surry ISLOCA

Plume 

Segment Xe Cs Ba I Te Ru Mo Ce La

1 1.69E-01 4.06E-05 3.31E-07 2.78E-04 2.62E-04 2.94E-07 5.76E-06 3.45E-11 3.18E-11

2 1.69E-01 4.06E-05 3.31E-07 2.78E-04 2.62E-04 2.94E-07 5.76E-06 3.45E-11 3.18E-11

3 3.74E-02 1.90E-03 7.82E-06 1.38E-02 1.17E-02 1.88E-05 2.58E-04 2.12E-09 1.97E-09

4 6.30E-02 3.39E-03 1.38E-05 2.50E-02 2.12E-02 3.49E-05 4.59E-04 3.94E-09 3.67E-09

5 1.84E-03 9.89E-05 3.27E-07 6.88E-04 5.85E-04 7.40E-07 1.42E-05 8.74E-11 8.20E-11

6 1.67E-01 9.01E-03 3.57E-05 6.62E-02 5.65E-02 9.23E-05 1.22E-03 1.04E-08 9.74E-09

7 1.47E-02 8.96E-04 2.50E-06 6.76E-03 5.66E-03 1.07E-05 1.19E-04 1.15E-09 1.06E-09

8 9.90E-02 2.71E-05 1.22E-07 2.20E-04 2.02E-04 4.19E-07 3.30E-06 4.77E-11 4.49E-11

9 9.90E-02 2.71E-05 1.22E-07 2.20E-04 2.02E-04 4.19E-07 3.30E-06 4.77E-11 4.49E-11

10 3.66E-02 1.88E-03 9.81E-06 1.55E-02 1.40E-02 2.90E-05 2.23E-04 3.34E-09 3.23E-09

11 2.36E-02 8.73E-04 6.97E-06 8.57E-03 5.73E-03 2.66E-05 7.64E-05 4.00E-09 3.78E-09

12 6.07E-03 2.36E-04 1.58E-06 2.13E-03 1.59E-03 5.53E-06 2.43E-05 7.65E-10 7.22E-10

13 8.81E-03 3.27E-04 2.62E-06 3.21E-03 2.14E-03 1.01E-05 2.86E-05 1.51E-09 1.42E-09

14 1.27E-02 3.65E-04 3.51E-06 3.78E-03 2.43E-03 1.45E-05 2.81E-05 2.15E-09 2.03E-09

15 2.21E-02 3.33E-06 3.81E-08 3.55E-05 2.30E-05 1.35E-07 2.38E-07 2.11E-11 1.99E-11

16 2.21E-02 3.33E-06 3.81E-08 3.55E-05 2.30E-05 1.35E-07 2.38E-07 2.11E-11 1.99E-11

17 3.61E-03 1.05E-04 2.45E-06 1.19E-03 6.08E-04 1.11E-06 5.30E-06 3.37E-10 3.19E-10

18 1.09E-03 3.24E-05 5.70E-07 3.58E-04 2.14E-04 7.16E-07 1.91E-06 1.57E-10 1.51E-10

19 1.59E-03 4.62E-05 1.01E-06 5.22E-04 2.64E-04 4.84E-07 2.31E-06 1.47E-10 1.39E-10

20 8.99E-04 2.72E-05 1.12E-06 2.49E-04 1.63E-04 2.26E-07 2.06E-06 3.31E-11 3.18E-11

21 4.73E-03 4.89E-07 1.06E-08 5.62E-06 2.16E-06 2.40E-09 2.06E-08 4.53E-13 4.29E-13

22 4.73E-03 4.89E-07 1.06E-08 5.62E-06 2.16E-06 2.40E-09 2.06E-08 4.53E-13 4.29E-13

23 4.37E-04 3.08E-05 6.50E-07 3.07E-04 1.69E-04 2.27E-07 2.26E-06 8.01E-11 7.47E-11

24 2.90E-04 1.65E-05 3.66E-07 1.67E-04 8.59E-05 1.20E-07 1.16E-06 3.72E-11 3.47E-11

25 2.85E-04 2.04E-05 4.43E-07 2.02E-04 1.12E-04 1.54E-07 1.52E-06 5.24E-11 4.88E-11

26 3.46E-03 1.77E-06 3.53E-08 1.71E-05 1.09E-05 2.35E-08 1.44E-07 8.27E-12 7.63E-12

27 3.46E-03 1.77E-06 3.53E-08 1.71E-05 1.09E-05 2.35E-08 1.44E-07 8.27E-12 7.63E-12

28 2.90E-03 5.23E-07 1.14E-08 5.29E-06 4.57E-06 2.77E-08 4.48E-08 6.73E-12 6.33E-12

29 2.90E-03 5.23E-07 1.14E-08 5.29E-06 4.57E-06 2.77E-08 4.48E-08 6.73E-12 6.33E-12

30 8.94E-07 7.45E-08 9.77E-09 8.72E-07 7.60E-07 7.71E-10 2.91E-09 3.71E-08 1.60E-09

31 2.09E-06 7.89E-08 4.17E-09 8.23E-07 7.30E-07 3.00E-09 5.91E-09 1.06E-08 4.58E-10

32 4.25E-07 3.00E-08 3.70E-09 3.56E-07 3.13E-07 3.82E-10 1.34E-09 1.38E-08 5.95E-10

33 3.73E-08 1.25E-08 2.61E-08 3.42E-08 3.02E-08 2.92E-10 1.20E-09 1.49E-08 6.30E-10

34 8.35E-06 3.37E-07 2.33E-08 4.01E-06 4.02E-06 4.75E-09 1.44E-08 6.59E-08 2.92E-09

35 5.24E-06 9.34E-07 4.54E-08 1.26E-05 1.02E-05 2.62E-09 1.10E-08 1.21E-07 7.04E-09

36 7.76E-08 2.01E-08 4.12E-08 5.86E-08 5.57E-08 4.56E-10 1.92E-09 2.31E-08 1.02E-09

37 8.87E-08 1.69E-08 3.45E-08 5.16E-08 5.16E-08 3.75E-10 1.59E-09 1.90E-08 8.71E-10

38 3.99E-06 1.23E-06 4.35E-08 1.70E-05 1.02E-05 5.72E-10 4.81E-09 1.22E-07 7.92E-09

39 9.39E-07 5.34E-07 9.40E-09 7.47E-06 2.62E-06 8.00E-11 1.16E-09 2.53E-08 1.84E-09

40 3.58E-07 2.00E-07 3.63E-09 2.80E-06 9.87E-07 4.37E-11 3.49E-10 9.88E-09 7.14E-10

41 3.80E-07 2.20E-07 3.55E-09 3.10E-06 8.90E-07 4.37E-11 3.49E-10 1.04E-08 7.33E-10

42 9.27E-08 1.24E-08 2.47E-08 4.36E-08 4.04E-08 2.64E-10 1.13E-09 1.35E-08 6.41E-10

43 2.47E-06 1.26E-06 2.33E-08 1.76E-05 5.90E-06 2.28E-10 2.41E-09 6.36E-08 4.57E-09

44 1.39E-06 8.23E-07 1.15E-08 1.16E-05 2.19E-06 1.31E-10 9.31E-10 3.72E-08 2.54E-09

45 4.47E-07 2.60E-07 3.59E-09 3.65E-06 6.80E-07 4.00E-11 3.49E-10 1.15E-08 7.93E-10

46 3.47E-06 4.51E-06 2.63E-08 6.35E-05 6.90E-06 1.99E-10 3.79E-09 5.78E-08 5.02E-09

47 1.86E-06 7.33E-06 1.40E-08 1.03E-04 4.64E-06 7.55E-11 2.40E-09 2.60E-08 2.93E-09
C-25

Surry ISLOCA

Plume 

Segment Xe Cs Ba I Te Ru Mo Ce La

48 1.46E-06 1.06E-05 8.87E-09 1.49E-04 4.47E-06 6.55E-11 1.51E-09 1.70E-08 2.13E-09

49 8.05E-07 5.46E-06 5.14E-09 7.72E-05 2.39E-06 2.18E-11 8.73E-10 9.83E-09 1.22E-09

50 1.58E-06 9.65E-06 9.89E-09 1.36E-04 4.27E-06 5.26E-11 1.68E-09 1.81E-08 2.18E-09

51 4.49E-06 1.58E-05 3.21E-08 2.23E-04 7.14E-06 1.38E-10 2.68E-09 4.08E-08 3.39E-09

52 1.42E-06 5.42E-06 9.50E-09 7.66E-05 2.40E-06 4.73E-11 8.15E-10 1.25E-08 1.07E-09

53 3.41E-06 2.10E-05 1.54E-08 2.97E-04 8.60E-06 8.75E-11 2.32E-09 2.37E-08 2.73E-09

54 1.62E-06 2.08E-05 7.31E-09 2.94E-04 1.02E-05 2.64E-11 1.48E-09 9.50E-09 1.65E-09

55 1.74E-06 1.73E-05 9.86E-09 2.44E-04 2.12E-05 2.91E-11 1.51E-09 8.71E-09 1.76E-09

56 1.48E-06 1.69E-05 7.56E-09 2.38E-04 1.55E-05 1.84E-11 1.28E-09 7.19E-09 1.43E-09

57 1.91E-06 1.46E-05 1.53E-08 2.06E-04 3.25E-05 2.18E-11 1.86E-09 8.41E-09 1.82E-09

58 2.08E-06 1.43E-05 1.61E-08 2.03E-04 3.50E-05 2.17E-11 1.83E-09 8.76E-09 1.90E-09

59 3.90E-06 1.14E-05 4.19E-08 1.62E-04 1.05E-04 3.64E-11 3.96E-09 1.48E-08 3.48E-09

60 2.34E-06 6.81E-06 2.39E-08 9.61E-05 5.96E-05 2.07E-11 2.26E-09 8.55E-09 2.01E-09

61 3.84E-06 2.35E-06 6.81E-08 3.29E-05 1.65E-04 2.18E-11 4.89E-09 1.24E-08 3.20E-09

62 2.23E-06 1.70E-06 3.51E-08 2.38E-05 8.65E-05 1.64E-11 2.60E-09 7.05E-09 1.79E-09

63 5.29E-06 7.39E-07 1.45E-07 9.87E-06 3.54E-04 5.09E-11 8.85E-09 1.48E-08 4.12E-09

64 1.60E-06 2.27E-07 4.38E-08 3.02E-06 1.07E-04 1.09E-11 2.79E-09 4.46E-09 1.25E-09

65 1.73E-06 3.55E-07 4.23E-08 4.84E-06 1.02E-04 1.08E-11 2.64E-09 4.78E-09 1.31E-09

66 7.61E-06 8.07E-07 1.97E-07 1.01E-05 7.22E-04 3.64E-11 1.86E-08 1.84E-08 5.55E-09

67 2.27E-06 2.41E-07 5.90E-08 3.02E-06 2.15E-04 7.28E-12 5.47E-09 5.49E-09 1.66E-09

68 2.29E-06 2.46E-07 7.59E-08 3.04E-06 2.40E-04 1.46E-11 6.26E-09 5.34E-09 1.65E-09

69 6.29E-06 6.79E-07 2.28E-06 7.93E-06 7.39E-04 2.18E-11 2.33E-08 1.31E-08 4.25E-09

70 1.91E-06 2.07E-07 6.78E-07 2.41E-06 2.25E-04 1.09E-11 7.16E-09 3.99E-09 1.30E-09

71 3.40E-06 3.77E-07 2.55E-06 4.26E-06 2.86E-04 1.09E-11 1.48E-08 6.70E-09 2.24E-09

72 1.06E-05 1.09E-06 8.55E-06 1.24E-05 4.93E-04 3.64E-11 4.28E-08 1.86E-08 6.51E-09

73 2.33E-06 2.44E-07 1.90E-06 2.73E-06 1.41E-04 9.09E-12 9.78E-09 4.22E-09 1.45E-09

74 3.29E-06 3.39E-07 2.66E-06 3.83E-06 1.54E-04 1.46E-11 1.32E-08 5.77E-09 2.02E-09

75 5.54E-06 4.65E-07 2.61E-06 5.58E-06 1.88E-04 1.82E-11 1.42E-08 8.30E-09 2.98E-09

76 5.90E-07 5.81E-08 4.53E-07 6.57E-07 2.80E-05 1.99E-12 2.28E-09 9.98E-10 3.48E-10

77 8.43E-06 5.09E-07 6.02E-07 6.76E-06 1.37E-04 1.46E-11 7.33E-09 1.00E-08 3.68E-09

78 2.00E-06 1.33E-07 3.32E-07 1.71E-06 4.54E-05 3.64E-12 2.71E-09 2.55E-09 9.28E-10

79 9.52E-06 6.24E-07 6.17E-07 8.24E-06 1.26E-04 1.82E-11 9.72E-09 1.08E-08 4.09E-09

80 1.12E-05 8.62E-07 6.30E-07 1.13E-05 1.43E-04 2.18E-11 1.41E-08 1.19E-08 4.68E-09

81 2.97E-06 2.06E-07 1.84E-07 2.72E-06 3.90E-05 9.09E-12 3.32E-09 3.18E-09 1.23E-09

82 3.60E-06 2.74E-07 2.02E-07 3.62E-06 4.59E-05 3.64E-12 4.42E-09 3.81E-09 1.50E-09

83 9.49E-06 1.10E-06 5.06E-07 1.47E-05 1.19E-04 1.82E-11 1.48E-08 9.45E-09 3.88E-09

84 1.08E-05 2.89E-06 6.09E-07 3.97E-05 1.39E-04 1.46E-11 2.06E-08 1.00E-08 4.28E-09

85 3.49E-06 9.30E-07 1.97E-07 1.28E-05 4.50E-05 0.00E+00 6.64E-09 3.24E-09 1.39E-09

86 9.67E-06 6.08E-06 5.08E-07 8.46E-05 1.38E-04 1.09E-11 2.20E-08 7.74E-09 3.43E-09

87 1.38E-05 1.78E-05 5.18E-07 2.49E-04 1.97E-04 7.28E-12 3.85E-08 9.01E-09 4.16E-09

88 4.51E-06 5.79E-06 1.69E-07 8.11E-05 6.44E-05 1.09E-11 1.26E-08 2.95E-09 1.36E-09

89 1.12E-05 1.43E-05 3.66E-07 2.00E-04 1.69E-04 7.28E-12 3.56E-08 6.76E-09 3.20E-09

90 8.43E-06 1.74E-06 2.92E-07 2.29E-05 1.43E-04 1.46E-11 2.86E-08 4.49E-09 2.22E-09

91 1.27E-05 2.41E-06 4.58E-07 3.15E-05 2.09E-04 1.09E-11 4.58E-08 6.39E-09 3.23E-09

92 1.14E-05 2.90E-06 3.97E-07 3.83E-05 1.90E-04 0.00E+00 4.71E-08 5.45E-09 2.84E-09

93 1.23E-05 4.49E-06 4.21E-07 6.02E-05 2.17E-04 7.28E-12 6.01E-08 5.66E-09 3.05E-09
C-26

Surry ISLOCA

Plume 

Segment Xe Cs Ba I Te Ru Mo Ce La

94 1.09E-05 5.55E-06 4.05E-07 7.52E-05 1.76E-04 1.46E-11 6.30E-08 4.81E-09 2.68E-09

95 1.16E-05 7.72E-06 4.75E-07 1.05E-04 1.43E-04 3.64E-12 8.27E-08 4.96E-09 2.85E-09

96 1.04E-05 7.01E-06 4.54E-07 9.55E-05 6.92E-05 7.28E-12 9.58E-08 4.26E-09 2.52E-09

97 1.12E-05 5.40E-06 5.25E-07 7.21E-05 2.49E-05 0.00E+00 1.45E-07 4.48E-09 2.72E-09

98 1.21E-05 4.81E-06 5.16E-07 6.34E-05 1.31E-05 3.64E-11 9.60E-05 4.34E-09 2.74E-09

99 3.57E-06 1.41E-06 1.57E-07 1.87E-05 4.95E-06 7.28E-12 1.72E-05 1.30E-09 8.06E-10

100 4.02E-06 1.59E-06 1.72E-07 2.10E-05 4.36E-06 3.64E-12 3.17E-05 1.44E-09 9.10E-10

101 9.43E-06 4.23E-06 3.34E-07 5.63E-05 1.08E-05 4.37E-11 1.02E-04 3.11E-09 2.03E-09

102 1.29E-06 4.87E-07 5.19E-08 6.43E-06 1.39E-06 2.10E-12 7.90E-06 4.29E-10 2.70E-10

103 6.71E-06 3.05E-06 1.26E-07 4.09E-05 8.04E-06 5.09E-11 7.34E-06 1.95E-09 1.37E-09

104 1.90E-06 8.34E-07 3.97E-08 1.12E-05 2.07E-06 7.28E-12 9.25E-06 5.33E-10 3.65E-10

105 7.76E-06 2.48E-06 1.35E-07 3.25E-05 9.91E-06 7.28E-11 6.71E-06 2.16E-09 1.55E-09

106 8.48E-06 1.70E-06 1.43E-07 2.11E-05 1.16E-05 5.82E-11 5.76E-06 2.26E-09 1.67E-09

107 7.34E-06 1.64E-06 1.28E-07 2.05E-05 1.09E-05 5.82E-11 4.12E-06 1.95E-09 1.47E-09
C-27

Surry Mitigated TISGTR

Plume 

Segment Xe Cs Ba I Te Ru Mo Ce La

1 2.59E-06 1.08E-06 9.51E-09 9.49E-07 9.10E-07 4.72E-09 2.76E-07 9.80E-14 9.80E-14

2 0.055065 0.00187 1.69E-05 0.00218 0.00138 7.61E-06 4.68E-04 1.91E-10 1.90E-10

3 1.89E-05 4.81E-06 2.49E-08 4.65E-06 4.80E-06 1.09E-07 1.28E-06 1.85E-12 1.85E-12

4 0.002275 2.77E-04 2.55E-06 2.81E-04 2.88E-04 9.07E-06 7.27E-05 2.37E-10 2.37E-10

5 2.62E-05 7.31E-06 5.41E-08 7.33E-06 7.52E-06 1.97E-07 1.93E-06 4.79E-12 4.79E-12

6 0.013924 7.47E-04 2.18E-05 0.00103 0.00105 6.80E-05 1.96E-04 1.81E-09 1.81E-09

7 3.05E-05 6.87E-06 8.66E-08 7.48E-06 7.63E-06 2.82E-07 1.81E-06 7.39E-12 7.44E-12

8 0.011313 8.78E-04 4.59E-05 0.00132 0.00149 1.29E-04 2.25E-04 6.95E-08 4.28E-09

9 3.40E-05 5.25E-06 8.35E-07 6.06E-06 6.27E-06 3.27E-07 1.38E-06 1.53E-07 2.77E-09

10 0.001501 1.26E-04 5.64E-05 8.95E-05 9.78E-05 1.47E-05 3.38E-05 9.75E-06 2.06E-07

11 3.17E-05 8.52E-07 4.11E-07 9.50E-07 9.84E-07 5.43E-08 2.24E-07 8.03E-08 1.64E-09

12 4.35E-04 5.64E-06 5.95E-07 5.57E-05 2.36E-06 3.33E-08 4.74E-07 6.44E-08 4.02E-09

13 2.18E-04 3.60E-06 3.96E-08 4.98E-05 8.47E-07 1.02E-09 1.96E-08 2.16E-09 7.28E-10

14 1.29E-04 1.75E-06 6.61E-09 2.47E-05 3.21E-07 1.31E-10 2.21E-09 3.47E-10 1.66E-10

15 1.83E-06 1.58E-08 4.37E-11 2.23E-07 2.79E-09 0 0 3.64E-12 1.39E-12

16 1.85E-06 1.30E-08 1.75E-10 1.81E-07 7.45E-09 0 9.71E-08 4.55E-12 4.41E-12

17 6.48E-06 1.12E-07 6.32E-09 1.58E-06 2.22E-07 0 4.21E-06 1.47E-10 1.46E-10

18 8.20E-07 1.03E-07 1.36E-08 1.46E-06 2.74E-07 0 8.92E-06 2.76E-10 2.81E-10

19 5.22E-08 1.03E-07 1.59E-08 1.46E-06 1.98E-07 0 1.01E-05 3.06E-10 3.09E-10

20 7.45E-09 9.31E-08 1.59E-08 1.32E-06 1.29E-07 0 9.80E-06 2.97E-10 3.00E-10

21 0 8.66E-08 1.58E-08 1.23E-06 8.66E-08 0 9.43E-06 2.88E-10 2.92E-10

22 0 7.92E-08 1.55E-08 1.12E-06 6.01E-08 0 9.02E-06 2.79E-10 2.83E-10

23 0 7.08E-08 1.53E-08 1.00E-06 4.05E-08 0 8.70E-06 2.73E-10 2.76E-10

24 0 6.29E-08 1.49E-08 8.83E-07 2.93E-08 0 8.27E-06 2.62E-10 2.66E-10

25 0 5.40E-08 1.43E-08 7.66E-07 2.10E-08 0 7.75E-06 2.48E-10 2.51E-10

26 0 4.80E-08 1.37E-08 6.77E-07 1.54E-08 0 7.30E-06 2.35E-10 2.38E-10

27 0 4.38E-08 1.32E-08 6.17E-07 1.21E-08 0 6.94E-06 2.22E-10 2.25E-10

28 0 3.91E-08 1.31E-08 5.54E-07 8.85E-09 0 6.78E-06 2.16E-10 2.19E-10

29 0 3.73E-08 1.28E-08 5.23E-07 8.38E-09 0 6.56E-06 2.08E-10 2.10E-10

30 0 3.40E-08 1.26E-08 4.84E-07 7.45E-09 0 6.38E-06 1.98E-10 2.02E-10

31 0 3.03E-08 1.25E-08 4.27E-07 6.52E-09 0 6.25E-06 1.95E-10 1.96E-10

32 0 2.56E-08 1.19E-08 3.66E-07 4.66E-09 0 5.92E-06 1.81E-10 1.84E-10

33 0 2.33E-08 1.13E-08 3.24E-07 3.26E-09 0 5.54E-06 1.67E-10 1.69E-10

34 7.45E-09 2.00E-08 1.13E-08 2.83E-07 3.26E-09 0 5.52E-06 1.63E-10 1.66E-10

35 3.73E-08 1.72E-08 1.13E-08 2.46E-07 1.86E-09 0 5.48E-06 1.60E-10 1.62E-10

36 1.42E-07 1.49E-08 1.13E-08 2.14E-07 1.40E-09 0 5.41E-06 1.56E-10 1.58E-10

37 1.71E-07 1.35E-08 1.13E-08 1.88E-07 1.86E-09 0 5.35E-06 1.52E-10 1.54E-10

38 8.20E-08 1.16E-08 1.12E-08 1.64E-07 9.31E-10 0 5.25E-06 1.47E-10 1.50E-10

39 3.73E-08 1.02E-08 1.14E-08 1.44E-07 9.31E-10 0 5.28E-06 1.47E-10 1.49E-10

40 7.45E-09 8.85E-09 1.05E-08 1.24E-07 9.31E-10 0 4.82E-06 1.33E-10 1.34E-10

41 7.45E-09 7.45E-09 1.07E-08 1.09E-07 9.31E-10 0 4.89E-06 1.32E-10 1.34E-10

42 0 6.98E-09 1.17E-08 9.59E-08 4.66E-10 0 5.26E-06 1.41E-10 1.43E-10

43 0 6.05E-09 1.19E-08 8.52E-08 9.31E-10 0 5.31E-06 1.42E-10 1.43E-10

44 0 5.12E-09 1.20E-08 7.45E-08 0 0 5.30E-06 1.38E-10 1.41E-10

45 0 4.66E-09 1.17E-08 6.47E-08 9.31E-10 0 5.13E-06 1.33E-10 1.35E-10

46 0 4.19E-09 1.11E-08 5.68E-08 4.66E-10 0 4.84E-06 1.25E-10 1.26E-10

47 0 3.26E-09 1.10E-08 4.94E-08 0 0 4.74E-06 1.21E-10 1.22E-10
C-28

Surry Mitigated TISGTR

Plume 

Segment Xe Cs Ba I Te Ru Mo Ce La

48 0 2.79E-09 1.04E-08 4.28E-08 0 0 4.43E-06 1.11E-10 1.13E-10

49 0 1.40E-09 4.38E-09 1.68E-08 0 0 1.85E-06 4.64E-11 4.69E-11
C-29

Surry Unmitigated TISGTR

Plume 

Segment Xe Cs Ba I Te Ru Mo Ce La

1 2.59E-06 1.08E-06 9.51E-09 9.49E-07 9.10E-07 4.72E-09 2.76E-07 9.80E-14 9.80E-14

2 0.055065 0.00187 1.69E-05 0.002182 0.001379 7.61E-06 4.68E-04 1.91E-10 1.90E-10

3 1.89E-05 4.81E-06 2.49E-08 4.65E-06 4.80E-06 1.09E-07 1.28E-06 1.85E-12 1.85E-12

4 0.0022745 2.77E-04 2.55E-06 2.81E-04 2.88E-04 9.07E-06 7.27E-05 2.37E-10 2.37E-10

5 2.62E-05 7.31E-06 5.41E-08 7.33E-06 7.52E-06 1.97E-07 1.93E-06 4.79E-12 4.79E-12

6 0.013924 7.47E-04 2.18E-05 0.001034 0.001054 6.80E-05 1.96E-04 1.81E-09 1.81E-09

7 3.05E-05 6.87E-06 8.66E-08 7.48E-06 7.63E-06 2.82E-07 1.81E-06 7.39E-12 7.44E-12

8 0.011313 8.78E-04 4.59E-05 0.001322 0.001489 1.29E-04 2.25E-04 6.95E-08 4.28E-09

9 3.40E-05 5.25E-06 8.35E-07 6.06E-06 6.27E-06 3.27E-07 1.38E-06 1.53E-07 2.77E-09

10 0.0024109 1.36E-04 6.16E-05 1.07E-04 1.10E-04 1.53E-05 3.62E-05 1.07E-05 2.30E-07

11 3.79E-05 3.39E-06 1.63E-06 3.74E-06 3.87E-06 2.17E-07 8.92E-07 3.21E-07 6.45E-09

12 0.0032437 2.29E-05 1.02E-05 5.99E-05 3.03E-05 1.24E-06 5.31E-06 1.93E-06 5.04E-08

13 3.79E-05 1.72E-06 8.32E-07 2.00E-06 2.00E-06 1.10E-07 4.51E-07 1.63E-07 3.38E-09

14 0.0034168 1.21E-05 4.71E-06 4.71E-05 2.04E-05 6.96E-07 2.52E-06 1.04E-06 2.50E-08

15 3.83E-05 8.91E-07 4.25E-07 1.18E-06 1.08E-06 5.69E-08 2.30E-07 8.46E-08 1.78E-09

16 0.0035156 9.70E-06 2.59E-06 6.63E-05 1.50E-05 4.16E-07 1.44E-06 6.22E-07 1.56E-08

17 3.87E-05 5.65E-07 2.56E-07 1.13E-06 7.06E-07 3.48E-08 1.39E-07 5.17E-08 1.11E-09

18 0.0035655 1.17E-05 1.73E-06 1.18E-04 1.35E-05 2.82E-07 9.46E-07 4.21E-07 1.11E-08

19 3.91E-05 4.54E-07 1.74E-07 1.76E-06 5.49E-07 2.41E-08 9.42E-08 3.58E-08 7.89E-10

20 0.0036008 1.74E-05 1.40E-06 2.12E-04 1.55E-05 2.06E-07 8.74E-05 3.08E-07 8.68E-09

21 3.94E-05 4.56E-07 1.30E-07 3.02E-06 5.05E-07 1.80E-08 1.33E-06 2.68E-08 6.09E-10

22 0.0036017 6.49E-06 1.52E-06 6.58E-05 1.81E-05 1.63E-07 0 2.46E-07 7.91E-09

23 3.97E-05 3.95E-07 1.06E-07 2.90E-06 5.28E-07 1.44E-08 0 2.15E-08 5.11E-10

24 0.003629 5.13E-06 1.29E-06 5.19E-05 1.97E-05 1.32E-07 0 1.98E-07 6.77E-09

25 3.99E-05 3.18E-07 8.60E-08 2.42E-06 5.49E-07 1.13E-08 0 1.69E-08 4.22E-10

26 0.0036535 4.56E-06 1.05E-06 4.91E-05 2.30E-05 9.99E-08 0 1.51E-07 5.56E-09

27 4.02E-05 2.38E-07 6.51E-08 1.89E-06 5.63E-07 8.25E-09 0 1.23E-08 3.25E-10

28 0.0036759 3.98E-06 8.17E-07 4.55E-05 2.68E-05 7.08E-08 0 1.08E-07 4.40E-09

29 4.04E-05 1.72E-07 4.75E-08 1.42E-06 5.90E-07 5.75E-09 0 8.62E-09 2.42E-10

30 0.0036984 3.46E-06 6.34E-07 4.15E-05 2.88E-05 5.00E-08 0 7.63E-08 3.54E-09

31 4.06E-05 1.27E-07 3.54E-08 1.10E-06 6.34E-07 4.06E-09 0 6.09E-09 1.84E-10

32 0.0037173 2.96E-06 5.20E-07 3.66E-05 2.78E-05 3.61E-08 0 5.56E-08 2.94E-09

33 4.08E-05 9.77E-08 2.77E-08 8.82E-07 6.72E-07 2.96E-09 0 4.45E-09 1.47E-10

34 0.0037368 2.10E-06 4.41E-07 2.59E-05 2.41E-05 2.71E-08 0 4.20E-08 2.52E-09

35 4.09E-05 7.75E-08 2.26E-08 7.27E-07 6.80E-07 2.23E-09 0 3.36E-09 1.22E-10

36 0.0037562 1.30E-06 4.02E-07 1.55E-05 2.15E-05 2.09E-08 0 3.27E-08 2.22E-09

37 4.10E-05 6.28E-08 1.93E-08 6.11E-07 6.74E-07 1.72E-09 0 2.61E-09 1.05E-10

38 0.0037724 9.50E-07 3.74E-07 1.13E-05 1.90E-05 1.64E-08 0 2.59E-08 1.97E-09

39 4.11E-05 5.15E-08 1.70E-08 5.16E-07 6.56E-07 1.35E-09 0 2.06E-09 9.24E-11

40 0.003791 7.33E-07 3.50E-07 8.66E-06 1.66E-05 1.32E-08 0 2.09E-08 1.78E-09

41 4.12E-05 4.26E-08 1.54E-08 4.37E-07 6.29E-07 1.08E-09 0 1.65E-09 8.26E-11

42 0.0038074 7.72E-07 3.26E-07 9.60E-06 1.44E-05 1.06E-08 0 1.71E-08 1.63E-09

43 4.13E-05 3.62E-08 1.42E-08 3.82E-07 5.92E-07 8.73E-10 0 1.35E-09 7.50E-11

44 0.0038235 8.27E-07 3.04E-07 1.07E-05 1.27E-05 8.67E-09 0 1.41E-08 1.50E-09

45 4.14E-05 3.19E-08 1.32E-08 3.48E-07 5.51E-07 7.13E-10 0 1.11E-09 6.89E-11

46 0.0038394 8.64E-07 2.85E-07 1.14E-05 1.13E-05 7.17E-09 0 1.18E-08 1.38E-09

47 4.14E-05 2.89E-08 1.24E-08 3.26E-07 5.09E-07 5.88E-10 0 9.20E-10 6.38E-11
C-30

Surry Unmitigated TISGTR

Plume 

Segment Xe Cs Ba I Te Ru Mo Ce La

48 0.0038542 9.01E-07 2.60E-07 1.21E-05 9.93E-06 5.97E-09 0 9.97E-09 1.29E-09

49 4.14E-05 2.68E-08 1.17E-08 3.13E-07 4.67E-07 4.90E-10 0 7.72E-10 5.97E-11

50 0.0038672 9.73E-07 2.41E-07 1.32E-05 8.72E-06 5.04E-09 0 8.50E-09 1.21E-09

51 0.0027853 1.70E-06 7.38E-07 2.05E-05 2.83E-05 2.71E-08 0 4.32E-08 3.75E-09

52 4.14E-05 2.54E-08 1.10E-08 3.05E-07 4.25E-07 4.10E-10 0 6.53E-10 5.61E-11

53 0.0038613 1.06E-06 2.32E-07 1.45E-05 7.72E-06 4.21E-09 0 7.25E-09 1.13E-09

54 0.0079969 4.70E-06 2.04E-06 5.80E-05 7.49E-05 6.69E-08 0 1.08E-07 1.03E-08

55 4.11E-05 2.42E-08 1.05E-08 2.98E-07 3.85E-07 3.44E-10 0 5.54E-10 5.27E-11

56 0.0038298 1.21E-06 2.37E-07 1.68E-05 6.85E-06 3.54E-09 0 6.20E-09 1.06E-09

57 0.014864 8.52E-06 3.71E-06 1.07E-04 1.27E-04 1.06E-07 0 1.73E-07 1.82E-08

58 4.04E-05 2.32E-08 1.01E-08 2.92E-07 3.46E-07 2.89E-10 0 4.70E-10 4.95E-11

59 0.0037684 1.42E-06 2.26E-07 1.98E-05 6.07E-06 2.97E-09 0 5.29E-09 1.00E-09

60 0.019254 1.09E-05 4.74E-06 1.40E-04 1.51E-04 1.18E-07 0 1.95E-07 2.27E-08

61 3.95E-05 2.23E-08 9.71E-09 2.86E-07 3.09E-07 2.42E-10 0 3.99E-10 4.66E-11

62 0.0036753 1.40E-06 2.17E-07 1.95E-05 5.40E-06 2.47E-09 0 4.56E-09 9.52E-10

63 0.021255 1.19E-05 5.18E-06 1.56E-04 1.53E-04 1.13E-07 0 1.89E-07 2.44E-08

64 3.80E-05 2.13E-08 9.25E-09 2.78E-07 2.73E-07 2.01E-10 0 3.36E-10 4.34E-11

65 0.0035692 1.41E-06 2.08E-07 1.97E-05 4.86E-06 2.10E-09 0 3.94E-09 9.06E-10

66 0.024223 1.35E-05 5.88E-06 1.79E-04 1.61E-04 1.12E-07 0 1.90E-07 2.71E-08

67 3.75E-05 2.09E-08 9.11E-09 2.77E-07 2.49E-07 1.73E-10 0 2.93E-10 4.20E-11

68 0.0034562 1.38E-06 1.99E-07 1.94E-05 4.46E-06 1.76E-09 0 3.43E-09 8.66E-10

69 0.023745 1.32E-05 5.76E-06 1.77E-04 1.48E-04 9.61E-08 0 1.66E-07 2.62E-08

70 3.58E-05 2.00E-08 8.69E-09 2.68E-07 2.22E-07 1.45E-10 0 2.49E-10 3.96E-11

71 0.0033394 1.34E-06 1.92E-07 1.88E-05 4.10E-06 1.54E-09 0 3.02E-09 8.35E-10

72 0.023485 1.32E-05 5.72E-06 1.78E-04 1.38E-04 8.37E-08 0 1.47E-07 2.58E-08

73 3.45E-05 1.94E-08 8.41E-09 2.62E-07 2.02E-07 1.23E-10 0 2.16E-10 3.79E-11

74 0.0032239 1.32E-06 1.86E-07 1.86E-05 3.80E-06 1.27E-09 0 2.67E-09 8.07E-10

75 0.0222 1.26E-05 5.45E-06 1.71E-04 1.24E-04 7.02E-08 0 1.26E-07 2.43E-08

76 3.33E-05 1.88E-08 8.16E-09 2.57E-07 1.85E-07 1.04E-10 0 1.88E-10 3.64E-11

77 0.0031118 1.25E-06 1.80E-07 1.76E-05 3.54E-06 1.09E-09 0 2.37E-09 7.81E-10

78 0.022319 1.26E-05 5.53E-06 1.73E-04 1.19E-04 6.29E-08 0 1.16E-07 2.46E-08

79 3.20E-05 1.80E-08 7.94E-09 2.48E-07 1.70E-07 8.97E-11 0 1.65E-10 3.52E-11

80 0.0030011 1.16E-06 1.74E-07 1.63E-05 3.33E-06 9.60E-10 0 2.12E-09 7.58E-10

81 0.021661 1.19E-05 5.44E-06 1.64E-04 1.12E-04 5.46E-08 0 1.03E-07 2.41E-08

82 3.08E-05 1.68E-08 7.75E-09 2.33E-07 1.58E-07 7.70E-11 0 1.46E-10 3.42E-11

83 0.0028905 1.11E-06 1.68E-07 1.57E-05 3.18E-06 8.44E-10 0 1.91E-09 7.35E-10

84 0.021648 1.16E-05 5.52E-06 1.61E-04 1.09E-04 4.88E-08 0 9.50E-08 2.43E-08

85 2.96E-05 1.59E-08 7.56E-09 2.20E-07 1.49E-07 6.65E-11 0 1.30E-10 3.33E-11

86 0.0027884 1.05E-06 1.64E-07 1.49E-05 3.06E-06 6.69E-10 0 1.74E-09 7.19E-10

87 0.020122 1.07E-05 5.24E-06 1.49E-04 1.00E-04 4.10E-08 0 8.24E-08 2.31E-08

88 2.86E-05 1.52E-08 7.44E-09 2.12E-07 1.42E-07 5.80E-11 0 1.17E-10 3.28E-11

89 0.0026983 8.68E-07 1.62E-07 1.23E-05 3.03E-06 6.26E-10 0 1.62E-09 7.16E-10

90 0.020926 1.13E-05 5.58E-06 1.59E-04 1.05E-04 3.89E-08 0 8.09E-08 2.46E-08

91 2.75E-05 1.49E-08 7.34E-09 2.09E-07 1.38E-07 5.09E-11 0 1.06E-10 3.24E-11

92 0.0025999 8.47E-07 1.58E-07 1.20E-05 3.11E-06 5.82E-10 0 1.51E-09 7.08E-10

93 0.018819 1.05E-05 5.14E-06 1.48E-04 9.61E-05 3.20E-08 0 6.88E-08 2.27E-08
C-31

Surry Unmitigated TISGTR

Plume 

Segment Xe Cs Ba I Te Ru Mo Ce La

94 2.64E-05 1.48E-08 7.21E-09 2.08E-07 1.35E-07 4.46E-11 0 9.64E-11 3.19E-11

95 0.002504 8.43E-07 1.54E-07 1.20E-05 3.28E-06 4.80E-10 0 1.40E-09 6.99E-10

96 0.018722 1.09E-05 5.22E-06 1.53E-04 9.93E-05 2.91E-08 0 6.51E-08 2.32E-08

97 2.53E-05 1.47E-08 7.07E-09 2.08E-07 1.35E-07 3.90E-11 0 8.77E-11 3.14E-11

98 0.0024121 7.11E-07 1.49E-07 1.01E-05 3.52E-06 4.51E-10 0 1.31E-09 6.95E-10

99 0.017989 1.10E-05 5.10E-06 1.55E-04 1.01E-04 2.55E-08 0 5.97E-08 2.29E-08

100 2.43E-05 1.48E-08 6.88E-09 2.10E-07 1.37E-07 3.42E-11 0 8.03E-11 3.10E-11

101 0.0023166 6.38E-07 1.17E-07 9.05E-06 3.94E-06 3.64E-10 0 1.23E-09 6.87E-10

102 0.017843 1.15E-05 4.96E-06 1.63E-04 1.09E-04 2.31E-08 0 5.66E-08 2.34E-08

103 2.33E-05 1.50E-08 6.47E-09 2.13E-07 1.43E-07 3.00E-11 0 7.37E-11 3.06E-11

104 0.0022242 6.10E-07 1.11E-07 8.66E-06 4.47E-06 3.35E-10 0 1.15E-09 6.80E-10

105 0.016027 1.09E-05 4.35E-06 1.55E-04 1.10E-04 1.90E-08 0 4.89E-08 2.16E-08

106 2.23E-05 1.53E-08 6.05E-09 2.17E-07 1.53E-07 2.63E-11 0 6.80E-11 3.02E-11

107 0.0021344 6.16E-07 1.05E-07 8.75E-06 5.07E-06 2.91E-10 0 1.09E-09 6.72E-10

108 0.015927 1.15E-05 4.24E-06 1.63E-04 1.24E-04 1.73E-08 0 4.70E-08 2.22E-08

109 2.14E-05 1.55E-08 5.69E-09 2.20E-07 1.67E-07 2.32E-11 0 6.29E-11 2.98E-11

110 8.98E-04 2.63E-07 4.40E-08 3.73E-06 2.41E-06 1.31E-10 0 4.51E-10 2.89E-10

111 0.0015566 1.16E-06 4.11E-07 1.65E-05 1.30E-05 1.61E-09 0 4.51E-09 2.20E-09

112 6.97E-07 5.20E-10 1.84E-10 7.39E-09 5.86E-09 6.82E-13 0 1.88E-12 9.88E-13

113 5.17E-04 3.86E-07 1.36E-07 5.49E-06 4.38E-06 5.33E-10 0 1.50E-09 7.34E-10

114 6.96E-07 5.20E-10 1.83E-10 7.39E-09 5.89E-09 6.82E-13 0 1.99E-12 9.88E-13

115 6.84E-05 2.00E-08 3.33E-09 2.86E-07 1.91E-07 0 0 3.27E-11 2.23E-11
C-32

Table C-5 EARLY Parameters Used in the Surry LTSBO, Unmitigated STSBO, TISGTR Mitigated and Unmitigated Scenarios

Variable Description LTSBO Unmitigated 

STSBO ISLOCA TISGTR Mitigated

TISGTR 

Unmitigated

ACNAME Latent Cancer 

Effect

Cancer Type 

1 LEUKEMIA LEUKEMIA LEUKEMIA LEUKEMIA LEUKEMIA

Cancer Type 

2 BONE BONE BONE BONE BONE

Cancer Type 

3 BREAST BREAST BREAST BREAST BREAST

Cancer Type 

4 LUNG LUNG LUNG LUNG LUNG

Cancer Type 

5 THYROID THYROID THYROID THYROID THYROID

Cancer Type 

6 LIVER LIVER LIVER LIVER LIVER

Cancer Type 

7 COLON COLON COLON COLON COLON

Cancer Type 

8 RESIDUAL RESIDUAL RESIDUAL RESIDUAL RESIDUAL

ACSUSC

Population 

Susceptible to 

Cancer

1.0 for all 

cancers

1.0 for all 

cancers

1.0 for all 

cancers

1.0 for all 

cancers

1.0 for all 

cancers

ACTHRE

Linear DoseResponse 

Threshold

0 0 0 0 0

BRRATE

Breathing Rate 

(for all activity 

types)

0.000266 0.000266 0.000266 0.000266 0.000266

CFRISK

Lifetime Cancer 

Fatality Risk 

Factors

Cancer Type 

1 0.0111 0.0111 0.0111 0.0111 0.0111

Cancer Type 

2 0.00019 0.00019 0.00019 0.00019 0.00019

Cancer Type 

3 0.00506 0.00506 0.00506 0.00506 0.00506

Cancer Type 

4 0.0198 0.0198 0.0198 0.0198 0.0198

Cancer Type 

5 0.000648 0.000648 0.000648 0.000648 0.000648

Cancer Type 

6 0.003 0.003 0.003 0.003 0.003

Cancer Type 

7 0.0208 0.0208 0.0208 0.0208 0.0208

Cancer Type 

8 0.0493 0.0493 0.0493 0.0493 0.0493
C-33

Variable Description LTSBO Unmitigated 

STSBO ISLOCA TISGTR Mitigated

TISGTR 

Unmitigated

CIRISK

Lifetime Cancer 

Injury Risk 

Factors

Cancer Type 

1 0.0113 0.0113 0.0113 0.0113 0.0113

Cancer Type 

2 0.000271 2.71E-04 2.71E-04 2.71E-04 2.71E-04

Cancer Type 

3 0.0101 0.0101 0.0101 0.0101 0.0101

Cancer Type 

4 0.0208 0.0208 0.0208 0.0208 0.0208

Cancer Type 

5 0.00648 0.00648 0.00648 0.00648 0.00648

Cancer Type 

6 0.00316 0.00316 0.00316 0.00316 0.00316

Cancer Type 

7 0.0378 0.0378 0.0378 0.0378 0.0378

Cancer Type 

8 0.169 0.169 0.169 0.169 0.169

CRIORG

Critical Organ 

for EARLY 

Phase

L-ICRP60ED L-ICRP60ED L-ICRP60ED L-ICRP60ED L-ICRP60ED

CSFACT

Cloudshine 

Shielding 

Factors

Evacuation 

Shielding 

Factor for All 

but Cohort 4

1 1 1 1 1

Normal 

Activity 

Shielding 

Factor for All 

but Cohort 4

0.68 0.68 0.68 0.68 0.68

Sheltering 

Shielding 

Factor for All 

but Cohort 4

0.6 0.6 0.6 0.6 0.6

Evacuation 

Shielding 

Factor for 

Cohort 4

1 1 1 1 1

Normal 

Activity 

Shielding 

Factor for 

Cohort 4

0.31 0.31 0.31 0.31 0.31

Sheltering 

Shielding 

Factor for 

Cohort 4

0.31 0.31 0.31 0.31 0.31
C-34

Variable Description LTSBO Unmitigated 

STSBO ISLOCA TISGTR Mitigated

TISGTR 

Unmitigated

DCF_FILE

Name of Dose 

Conversion 

Factor File

FGR13GyEq

uivDCF.INP

FGR13GyEquiv

DCF.INP

FGR13GyEq

uivDCF.INP

FGR13GyEquiv

DCF.INP

FGR13GyEquiv

DCF.INP

DDREFA

DoseDependent 

Reduction 

Factor

Cancer Type 

1 2 2 2 2 2

Cancer Type 

2 2 2 2 2 2

Cancer Type 

3 1 1 1 1 1

Cancer Type 

4 2 2 2 2 2

Cancer Type 

5 2 2 2 2 2

Cancer Type 

6 2 2 2 2 2

Cancer Type 

7 2 2 2 2 2

Cancer Type 

8 2 2 2 2 2

DDTHRE

Threshold for 

Applying DoseDependent 

Reduction 

Factor

0.2 0.2 0.2 0.2 0.2

DLTSHL

Delay from 

Alarm Time to 

Shelter

Cohort 1 9900 9900 9900 9900 9900

Cohort 2 9900 9900 9900 9900 9900

Cohort 3 900 900 900 900 900

Cohort 4 9900 9900 9900 9900 9900

Cohort 5 9900 9900 9900 9900 9900

Cohort 6 N/A N/A N/A N/A N/A

DLTEVA

Delay from 

Beginning of 

Shelter to 

Evacuation

Cohort 1 3600 3600 3600 3600 3600

Cohort 2 3600 3600 3600 3600 3600

Cohort 3 3600 3600 6300 3600 3600

Cohort 4 39600 39600 39600 39600 39600

Cohort 5 39600 39600 39600 39600 39600

Cohort 6 N/A N/A N/A N/A N/A

DOSEFA

Cancer DoseResponse 

Linear Factors

1 for all 

organs 1 for all organs 1 for all 

organs 1 for all organs 1 for all organs
C-35

Variable Description LTSBO Unmitigated 

STSBO ISLOCA TISGTR Mitigated

TISGTR 

Unmitigated

DOSEFB

Cancer DoseResponse 

Quadratic 

Factors

0 for all 

organs 0 for all organs 0 for all organs 0 for all organs 0 for all organs

DOSHOT

Hot-Spot 

Relocation 

Dose Threshold

0.05 0.05 0.05 0.05 0.05

DOSMOD Dose-Response 

Model Flag AT AT AT AT AT

DOSNRM

Normal 

Relocation 

Dose Threshold

0.01 0.01 0.01 0.01 0.01

DURBEG

Duration of 

Beginning of 

Evacuation 

Phase

Cohort 1 900 900 900 900 900

Cohort 2 900 900 900 900 900

Cohort 3 900 900 900 900 900

Cohort 4 3600 3600 3600 3600 3600

Cohort 5 3600 3600 3600 3600 3600

Cohort 6 N/A N/A N/A N/A N/A

DURMID

Duration of 

Middle of 

Evacuation 

Phase

Cohort 1 35100 35100 35100 35100 35100

Cohort 2 35100 35100 35100 35100 35100

Cohort 3 1800 1800 1800 1800 1800

Cohort 4 3600 3600 3600 3600 3600

Cohort 5 3600 3600 3600 3600 3600

Cohort 6 N/A N/A N/A N/A N/A

EANAM1

Text Describing 

the EARLY 

Assumptions

SOARCA 

Surry Longterm SBO 

Calculation

SOARCA Surry 

Short-term SBO 

Calculation

SOARCA 

Surry Shortterm SBO 

Calculation

SOARCA Surry 

Short-term SBO 

TI-SGTR 

Calculation

SOARCA Surry 

Short-term SBO 

Calculation

EANAM2

Text Describing 

the Emergency 

Response

Cohort 1 Group 1 Group 1 Group 1 Group 1 Group 1

Cohort 2 Group 2 Group 2 Group 2 Group 2 Group 2

Cohort 3 Group 3 Group 3 Group 3 Group 3 Group 3

Cohort 4 Group 4 Group 4 Group 4 Group 4 Group 4

Cohort 5 Group 5 Group 5 Group 5 Group 5 Group 5

Cohort 6 Group 6 Group 6 Group 6 Group 6 Group 6

EFFACA LD50 for Early 

Fatality Types

A-RED 

MARR 5.6 5.6 5.6 5.6 5.6

A-LUNGS 23.5 23.5 23.5 23.5 23.5
C-36

Variable Description LTSBO Unmitigated 

STSBO ISLOCA TISGTR Mitigated

TISGTR 

Unmitigated

EFFACA ASTOMACH 12.1 12.1 12.1 12.1 12.1

EFFACB

Shape Factor 

for Early 

Fatality Types

A-RED 

MARR 6.1 6.1 6.1 6.1 6.1

A-LUNGS 9.6 9.6 9.6 9.6 9.6

ASTOMACH 9.3 9.3 9.3 9.3 9.3

EFFACY Efficacy of the 

KI Ingestion 0.7 0.7 0.7 0.7 0.7

EFFTHR Threshold Dose 

to Target Organ

A-RED 

MARR 2.32 2.32 2.32 2.32 2.32

A-LUNGS 13.6 13.6 13.6 13.6 13.6

ASTOMACH 6.5 6.5 6.5 6.5 6.5

EIFACA D50 For Early 

Injuries

PRODROM

AL VOMIT 2 2 2 2 2

DIARRHEA 3 3 3 3 3

PNEUMONI

TIS 16.6 16.6 16.6 16.6 16.6

SKIN 

ERYTHRMA 6 6 6 6 6

TRANSEPID

ERMAL 20 20 20 20 20

THYROIDIT

IS 240 240 240 240 240

HYPOTHYR

OIDISM 60 60 60 60 60

EIFACB

Shape Factor 

for Early 

Injuries

PRODROM

AL VOMIT 3 3 3 3 3

DIARRHEA 2.5 2.5 2.5 2.5 2.5

PNEUMONI

TIS 7.3 7.3 7.3 7.3 7.3

SKIN 

ERYTHRMA 5 5 5 5 5

TRANSEPID

ERMAL 5 5 5 5 5

THYROIDIT

IS 2 2 2 2 2

HYPOTHYR

OIDISM 1.3 1.3 1.3 1.3 1.3
C-37

Variable Description LTSBO Unmitigated 

STSBO ISLOCA TISGTR Mitigated

TISGTR 

Unmitigated

EINAME

Early Injury 

Effect Names 

and 

Corresponding 

Organ

PRODROM

AL VOMIT

ASTOMACH A-STOMACH ASTOMACH A-STOMACH A-STOMACH

DIARRHEA ASTOMACH A-STOMACH ASTOMACH A-STOMACH A-STOMACH

PNEUMONI

TIS A-LUNGS A-LUNGS A-LUNGS A-LUNGS A-LUNGS

SKIN 

ERYTHRMA A-SKIN A-SKIN A-SKIN A-SKIN A-SKIN

TRANSEPID

ERMAL A-SKIN A-SKIN A-SKIN A-SKIN A-SKIN

THYROIDIT

IS A-THYROID A-THYROID A-THYROID A-THYROID A-THYROID

HYPOTHYR

OIDISM A-THYROID A-THYROID A-THYROID A-THYROID A-THYROID

EISUSC

Susceptible 

Population 

Fraction

1. for all 

health effects

1. for all health 

effects

1. for all 

health effects

1. for all health 

effects

1. for all health 

effects

EITHRE Early Injury 

Dose Threshold

PRODROM

AL VOMIT 0.5 0.5 0.5 0.5 0.5

DIARRHEA 1 1 1 1 1

PNEUMONI

TIS 9.2 9.2 9.2 9.2 9.2

SKIN 

ERYTHRMA 3 3 3 3 3

TRANSEPID

ERMAL 10 10 10 10 10

THYROIDIT

IS 40 40 40 40 40

HYPOTHYR

OIDISM 2 2 2 2 2

ENDAT2

Control flag 

indicating only 

ATMOS and 

EARLY are to 

be run

.FALSE. .FALSE. .FALSE. .FALSE. .FALSE.

ENDEMP

Time Duration 

for the 

Emergency 

Phase

604800 604800 604800 604800 604800

ESPEED Evaluation 

Speed

Initial 

Evacuation 

Phase, 

Cohort 1

2.235 2.235 2.235 2.235 2.235
C-38

Variable Description LTSBO Unmitigated 

STSBO ISLOCA TISGTR Mitigated

TISGTR 

Unmitigated

ESPEED

Middle 

Evacuation 

Phase, 

Cohort 1

0.447 0.447 0.447 0.447 0.447

Late 

Evacuation 

Phase, 

Cohort 1

8.941 8.941 8.941 8.941 8.941

Initial 

Evacuation 

Phase, 

Cohort 2

2.235 2.235 2.235 2.235 2.235

Middle 

Evacuation 

Phase, 

Cohort 2

0.447 0.447 0.447 0.447 0.447

Late 

Evacuation 

Phase, 

Cohort 2

8.941 8.941 8.941 8.941 8.941

Initial 

Evacuation 

Phase, 

Cohort 3

4.47 4.47 4.47 4.47 4.47

Middle 

Evacuation 

Phase, 

Cohort 3

4.47 4.47 4.47 4.47 4.47

Late 

Evacuation 

Phase, 

Cohort 3

8.941 8.941 8.941 8.941 8.941

Initial 

Evacuation 

Phase, 

Cohort 4

0.447 0.447 0.447 0.447 0.447

Middle 

Evacuation 

Phase, 

Cohort 4

4.47 4.47 4.47 4.47 4.47

Late 

Evacuation 

Phase, 

Cohort 4

8.941 8.941 8.941 8.941 8.941

Initial 

Evacuation 

Phase, 

Cohort 5

0.447 0.447 0.447 0.447 0.447

Middle 

Evacuation 

Phase, 

Cohort 5

4.47 4.47 4.47 4.47 4.47
C-39

Variable Description LTSBO Unmitigated 

STSBO ISLOCA TISGTR Mitigated

TISGTR 

Unmitigated

ESPEED

Late 

Evacuation 

Phase, 

Cohort 5

8.941 8.941 8.941 8.941 8.941

Cohort 6 N/A N/A N/A N/A N/A

ESPGRD

Speed 

Multiplier to 

Account for 

Grid-Level 

Variations in 

Road Network

Table C-6 

Table C-7

Table C-6 

Table C-7

Table C-6 

Table C-7

Table C-6 

Table C-7

Table C-6 

Table C-7

ESPMUL

Speed 

Multiplier 

Employed 

During 

Precipitation

0.7 0.7 0.7 0.7 0.7

EVATYP Evacuation 

Type NETWORK NETWORK NETWORK NETWORK NETWORK

GSHFAC

Groundshine 

Shielding 

Factors

Evacuation 

Shielding 

Factor for All 

but Cohort 4

0.5 0.5 0.5 0.5 0.5

Normal 

Activity 

Shielding 

Factor for All 

but Cohort 4

0.26 0.26 0.26 0.26 0.26

Sheltering 

Shielding 

Factor for All 

but Cohort 4

0.2 0.2 0.2 0.2 0.2

Evacuation 

Shielding 

Factor for 

Cohort 4

0.5 0.5 0.5 0.5 0.5

Normal 

Activity 

Shielding 

Factor for 

Cohort 4

0.05 0.05 0.05 0.05 0.05

Sheltering 

Shielding 

Factor for 

Cohort 4

0.05 0.05 0.05 0.05 0.05

IDIREC

Direction in 

Network 

Evacuation 

Model

Table C-8 Table C-8 Table C-8 Table C-8 Table C-8
C-40

Variable Description LTSBO Unmitigated 

STSBO ISLOCA TISGTR Mitigated

TISGTR 

Unmitigated

IPLUME

Plume Model 

Dispersion 

Code

3 3 3 3 3

KIMODL Model Flag for 

KI Ingestion KI KI KI KI KI

LASMOV

Last Ring in 

Movement 

Zone

17 17 17 17 17

NUMACA

Number of 

Latent Cancer 

Health Effects

8 8 8 8 8

NUMEFA

Number of 

Early Fatality 

Effects

3 3 3 3 3

NUMEIN

Number of 

Early Injury 

Effects

7 7 7 7 7

NUMEVA

Outer Boundary 

of 

Evacuation/Shel

ter Region

Cohort 1 - 15

Cohorts 2 to 

5 - 12

Cohort 1 - 15

Cohorts 2 to 5 -

12

Cohort 1 - 15

Cohorts 2 to 

5 - 12

Cohort 1 - 15

Cohorts 2 to 5 -

12

Cohort 1 - 15

Cohorts 2 to 5 -

12

NUMFIN

Number of Fine 

Grid 

Subdivisions

7 7 7 7 7

ORGFLG

Doses to be 

Calculated for 

Specified Organ

All TRUE for 

FGR-13

All TRUE for 

DOSFAC2 

except for ALower LI and 

L-Liver, 

which are 

FALSE

All TRUE for 

FGR-13

All TRUE for 

DOSFAC2 

except for ALower LI and 

L-Liver, which 

are FALSE

All TRUE for 

FGR-13

All TRUE for 

DOSFAC2 

except for ALower LI and 

L-Liver, 

which are 

FALSE

All TRUE for 

FGR-13

All TRUE for 

DOSFAC2 

except for ALower LI and 

L-Liver, which 

are FALSE

All TRUE for 

FGR-13

All TRUE for 

DOSFAC2 

except for ALower LI and LLiver, which are 

FALSE

OVRRID

Wind Rose 

Probability 

Override

.FALSE. .FALSE. .FALSE. .FALSE. .FALSE.

POPFLG

Population 

Distribution 

Flag

FILE FILE FILE FILE FILE

POPFRAC

Population 

Fraction 

Ingesting KI

Cohort 2 -

1.0

Cohort 1, 3-6 

- 0.0

Cohort 2 - 1.0

Cohort 1, 3-6 -

0.0

Cohort 2 -

1.0

Cohort 1, 3-6 

- 0.0

Cohort 2 - 1.0

Cohort 1, 3-6 -

0.0

Cohort 2 - 1.0

Cohort 1, 3-6 -

0.0

PROTIN 

[E]

Inhalation 

Protection 

Factor –

evacuation for 

all but Cohort 4

0.98 0.98 0.98 0.98 0.98

PROTIN 

[N]

Inhalation 

Protection 

Factor - normal 

activity for all 

but Cohort 4

0.46 0.46 0.46 0.46 0.46
C-41

Variable Description LTSBO Unmitigated 

STSBO ISLOCA TISGTR Mitigated

TISGTR 

Unmitigated

PROTIN 

[S]

Inhalation 

Protection 

Factor –

sheltering for 

all but Cohort 4

0.33 0.33 0.33 0.33 0.33

PROTIN 

[E]

Inhalation 

Protection 

Factor –

evacuation for 

Cohort 4

0.98 0.98 0.98 0.98 0.98

PROTIN 

[N]

Inhalation 

Protection 

Factor - normal 

activity for 

Cohort 4

0.33 0.33 0.33 0.33 0.33

PROTIN 

[S]

Inhalation 

Protection 

Factor –

sheltering for 

Cohort 4

0.33 0.33 0.33 0.33 0.33

REFPNT

Reference Time 

Point 

(ARRIVAL or 

SCRAM)

ALARM ALARM ALARM ALARM ALARM

RESCON

Emergency 

phase 

resuspension 

coefficient

0.0001 0.0001 0.0001 0.0001 0.0001

RESHAF

Resuspension 

Concentration 

Half-Life

182000 182000 182000 182000 182000

RISCAT

Risk by 

WeatherCategory Flag.FALSE. .FALSE. .FALSE. .FALSE. .FALSE.

RISTHR

Risk Threshold 

for Fatality 

Radius

0 0 0 0 0

SKPFAC 

[E]

Skin Protection 

Factors –

evacuation for 

all but Cohort 4

0.98 0.98 0.98 0.98 0.98

SKPFAC 

[N]

Skin Protection 

Factors -

normal activity 

for all but 

Cohort 4

0.46 0.46 0.46 0.46 0.46

SKPFAC 

[S]

Skin Protection 

Factors –

sheltering for 

all but Cohort 4

0.33 0.33 0.33 0.33 0.33
C-42

Variable Description LTSBO Unmitigated 

STSBO ISLOCA TISGTR Mitigated

TISGTR 

Unmitigated

SKPFAC 

[E]

Skin Protection 

Factors –

evacuation for 

Cohort 4

0.98 0.98 0.98 0.98 0.98

SKPFAC 

[N]

Skin Protection 

Factors -

normal activity 

Cohort 4

0.33 0.33 0.33 0.33 0.33

SKPFAC 

[S]

Skin Protection 

Factors –

sheltering for 

Cohort 4

0.33 0.33 0.33 0.33 0.33

TIMHOT

Hot Spot 

Relocation 

Time

86400 86400 86400 86400 86400

TIMNRM

Normal 

Relocation 

Time

129600 129600 129600 129600 129600

TRAVEL

POINT

Evacuee 

Movement 

Option

CENTER

POINT

CENTER

POINT

CENTER

POINT

CENTER

POINT

CENTER

POINT

WTFRAC

Weighting 

Fraction 

Applicable to 

this Scenario

Cohort 1 0.2 0.2 0.2 0.2 0.2

Cohort 2 0.535 0.535 0.535 0.535 0.535

Cohort 3 0.193 0.193 0.193 0.193 0.193

Cohort 4 0.007 0.007 0.007 0.007 0.007

Cohort 5 0.06 0.06 0.06 0.06 0.06

Cohort 6 0.005 0.005 0.005 0.005 0.005

WTNAME

Type of 

Weighting for 

Cohorts

PEOPLE PEOPLE PEOPLE PEOPLE PEOPLE
C-43

Table C-6 Grid-Level Evacuation Speed Multipliers Used in the Surry LTSBO, 

Unmitigated STSBO, TI-SGTR Mitigated and Unmitigated Scenarios for 

Cohorts 1-2

Compass Sector

Radial 

Ring 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16

1 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

2 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

4 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

5 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

6 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

7 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

8 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

9 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

10 1 1 1 1 1 1 3 3 3 3 3 3 3 3 3 3

11 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

12 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

13 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

14 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

15 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

16 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

17 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1
C-44

Compass Sector

Radial 

Ring 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32

1 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

2 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

4 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

5 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

6 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

7 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

8 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

9 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

10 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

11 1 1 1 1 1 1 3 3 3 3 3 3 3 3 3 3

12 1 1 1 1 1 1 1 3 3 3 3 3 3 3 3 3

13 1 1 1 1 1 1 1 3 3 3 3 3 3 3 3 3

14 1 1 1 1 1 1 1 3 3 3 3 3 3 3 3 3

15 1 1 1 1 1 1 1 1 3 3 3 3 3 3 3 3

16 1 1 1 1 1 1 1 1 1 3 3 3 3 3 3 3

17 1 1 1 1 1 1 3 3 3 3 3 3 3 3 3 3
C-45

Compass Sector

Radial 

Ring 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48

1 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

2 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

4 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

5 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

6 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

7 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

8 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

9 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

10 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

11 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

12 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

13 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

14 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

15 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

16 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

17 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3
C-46

Compass Sector

Radial 

Ring 49 50 51 52 53 54 55 56 57 58 59 60 61 62 63 64

1 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

2 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

4 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

5 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

6 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

7 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

8 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3 3

9 3 3 3 3 3 3 3 1 1 1 1 3 3 3 3 3

10 3 3 3 3 3 3 3 1 1 1 1 1 1 1 1 1

11 3 3 3 3 3 3 3 1 1 1 1 1 1 1 1 1

12 3 3 3 3 3 3 3 1 1 1 1 1 1 1 1 1

13 3 3 3 3 3 1 1 1 1 1 1 1 1 1 1 1

14 3 3 3 3 3 1 1 1 1 1 1 1 1 1 1 1

15 3 3 3 3 3 1 1 1 1 1 1 1 1 1 1 1

16 3 3 3 3 1 1 1 1 1 1 1 1 1 1 1 1

17 3 3 3 3 1 1 1 1 1 1 1 1 1 1 1 1
C-47

Table C-7 Grid-Level Evacuation Speed Multipliers Used in the Surry LTSBO, 

Unmitigated STSBO, TI-SGTR Mitigated and Unmitigated Scenarios for 

Cohorts 3-5

Compass Sector

Radial 

Ring 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16

1 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

3 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

4 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

5 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

6 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

7 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

8 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

9 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

10 1 1 1 1 1 1 2 2 2 2 2 2 2 2 2 2

11 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

12 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

13 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

14 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

15 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

16 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

17 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1
C-48

Compass Sector

Radial 

Ring 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32

1 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

3 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

4 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

5 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

6 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

7 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

8 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

9 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

10 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

11 1 1 1 1 1 1 2 2 2 2 2 2 2 2 2 2

12 1 1 1 1 1 1 1 2 2 2 2 2 2 2 2 2

13 1 1 1 1 1 1 1 2 2 2 2 2 2 2 2 2

14 1 1 1 1 1 1 1 2 2 2 2 2 2 2 2 2

15 1 1 1 1 1 1 1 1 2 2 2 2 2 2 2 2

16 1 1 1 1 1 1 1 1 1 2 2 2 2 2 2 2

17 1 1 1 1 1 1 2 2 2 2 2 2 2 2 2 2
C-49

Compass Sector

Radial 

Ring 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48

1 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

3 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

4 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

5 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

6 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

7 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

8 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

9 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

10 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

11 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

12 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

13 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

14 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

15 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

16 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

17 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2
C-50

Compass Sector

Radial 

Ring 49 50 51 52 53 54 55 56 57 58 59 60 61 62 63 64

1 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

3 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

4 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

5 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

6 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

7 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

8 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2

9 2 2 2 2 2 2 2 1 1 1 1 2 2 2 2 2

10 2 2 2 2 2 2 2 1 1 1 1 1 1 1 1 1

11 2 2 2 2 2 2 2 1 1 1 1 1 1 1 1 1

12 2 2 2 2 2 2 2 1 1 1 1 1 1 1 1 1

13 2 2 2 2 2 1 1 1 1 1 1 1 1 1 1 1

14 2 2 2 2 2 1 1 1 1 1 1 1 1 1 1 1

15 2 2 2 2 2 1 1 1 1 1 1 1 1 1 1 1

16 2 2 2 2 1 1 1 1 1 1 1 1 1 1 1 1

17 2 2 2 2 1 1 1 1 1 1 1 1 1 1 1 1
C-51

Table C-8 Evacuation Direction Parameters Used in the Surry LTSBO, Unmitigated 

STSBO, TI-SGTR Mitigated and Unmitigated Scenarios

Compass Sector

Radial 

Ring 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16

1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

2 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

3 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

4 1 1 1 2 2 2 2 2 2 2 2 2 2 2 2 2

5 1 1 1 2 2 2 2 2 2 2 2 2 2 2 2 2

6 1 1 1 2 2 2 2 2 2 2 2 2 2 2 2 2

7 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

8 1 1 1 1 1 2 1 1 1 1 1 1 1 1 1 1

9 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

10 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

11 1 1 1 1 1 1 1 1 1 1 1 1 1 1 4 4

12 1 1 1 1 4 4 4 4 2 2 2 2 2 1 1 1

13 1 4 1 1 1 1 1 1 1 1 1 1 1 1 1 1

14 4 4 2 2 2 2 2 1 4 4 4 4 1 2 2 2

15 1 2 2 2 1 1 4 4 4 4 4 4 4 1 1 1

16 2 2 1 4 4 4 4 1 1 1 1 1 1 1 1 1

17 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1
C-52

Compass Sector

Radial 

Ring 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32

1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

2 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

3 1 1 1 1 1 4 4 4 4 2 2 2 1 1 1 1

4 2 2 2 2 2 2 2 2 2 2 2 2 2 1 1 1

5 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 1

6 2 2 2 2 2 2 2 2 2 2 2 2 2 1 1 1

7 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

8 1 1 1 1 1 1 1 1 2 2 2 1 1 2 2 2

9 1 1 1 1 1 1 1 1 1 1 2 2 2 1 1 1

10 1 1 1 1 1 1 1 1 1 1 1 2 2 2 2 1

11 4 4 4 4 4 4 1 1 1 1 1 1 1 2 2 2

12 1 1 1 1 4 4 1 1 1 1 1 1 1 1 1 1

13 1 1 1 1 1 4 1 1 1 1 2 2 1 1 1 1

14 2 2 2 2 1 4 4 1 1 2 2 2 1 1 1 1

15 1 2 2 2 2 1 1 4 1 2 2 2 1 4 4 4

16 1 1 2 2 1 1 4 2 1 1 1 1 1 1 1 1

17 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1
C-53

Compass Sector

Radial 

Ring 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48

1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

2 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

3 4 4 4 4 4 4 4 4 4 1 1 1 1 1 1 1

4 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

5 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

6 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

7 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

8 1 1 4 4 4 4 1 1 1 1 1 1 1 1 1 1

9 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

10 1 1 1 2 2 2 1 1 1 4 4 4 4 4 4 4

11 2 2 2 2 2 1 1 4 4 2 2 1 1 1 1 1

12 1 1 1 1 1 1 1 4 4 2 2 1 1 1 1 4

13 1 1 1 1 1 1 2 2 1 2 1 4 4 2 2 1

14 1 4 4 4 4 2 2 2 1 2 1 2 2 2 2 2

15 1 1 4 1 1 1 1 1 1 1 1 1 1 1 1 4

16 1 1 1 2 2 2 2 2 2 2 2 2 1 1 1 1

17 4 4 4 4 4 4 2 2 2 1 1 1 2 2 2 1
C-54

Compass Sector

Radial 

Ring 49 50 51 52 53 54 55 56 57 58 59 60 61 62 63 64

1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

2 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

3 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

4 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

5 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

6 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

7 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

8 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

9 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

10 1 2 1 1 1 1 1 1 1 1 1 1 1 1 1 1

11 1 4 4 4 4 1 1 1 1 1 1 1 1 1 1 1

12 4 4 4 4 1 1 1 1 1 1 1 1 1 1 1 1

13 4 4 4 1 1 2 1 1 2 2 1 1 1 1 1 1

14 1 1 4 4 2 2 2 2 2 2 2 2 2 1 1 4

15 4 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

16 1 1 1 1 1 1 2 2 2 2 2 1 4 4 4 2

17 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1
C-55

Table C-9 CHRONC Input Parameters Used in the Surry LTSBO, Unmitigated 

STSBO, TISGTR Mitigated and Unmitigated Scenarios

Variable Description LTSBO Unmitigated 

STSBO ISLOCA TISGTR Mitigated

TISGTR 

Unmitigated

CHNAME CHRONC Problem 

Identification

Surry with no 

Food-Chain 

Modeling

Surry with no 

Food-Chain 

Modeling

Surry with no 

Food-Chain 

Modeling

Surry with no 

Food-Chain 

Modeling

Surry with no 

Food-Chain 

Modeling

CDFRM Farmland Decontamination 

Cost

Level 1 1330 1330 1330 1330 1330

Level 2 2960 2960 2960 2960 2960

CDNFRM Non farmland 

Decontamination Cost

Level 1 7110 7110 7110 7110 7110

Level 2 19000 19000 19000 19000 19000

CRTOCR Critical Organ for CHRONC 

Phase L-ICRP60ED L-ICRP60ED L-ICRP60ED L-ICRP60ED L-ICRP60ED

DPRATE Property Depreciation Rate 0.2 0.2 0.2 0.2 0.2

DLBCST Hourly Labor Cost for 

Decontamination Worker 84000 84000 84000 84000 84000

DPFRCT Farm Production Dairy 

Fraction Surry site file Surry site file Surry site file Surry site file Surry site file

DSCRLT Long-Term Phase Dose 

Criterion 0.04 0.04 0.04 0.04 0.04

DSCRTI Intermediate-Phase Dose 

Criterion 100000 100000 100000 100000 100000

DSRATE Societal Discount Rate for 

Property 0.12 0.12 0.12 0.12 0.12

DSRFCT Decontamination Factors

Level 1 3 3 3 3 3

Level 2 15 15 15 15 15

DUR_INTPHAS Duration of the Intermediate 

Phase 0 0 0 0 0

EVACST Emergency Phase Cost of 

Evacuation/Relocation 172 172 172 172 172

EXPTIM Maximum Exposure Time 1580000000 1580000000 1580000000 1580000000 1580000000

FDPATH COMIDA2 vs. MACCS Food 

Model Switch OFF OFF OFF OFF OFF

FRACLD Fraction of Area that is Land Surry site file Surry site file Surry site file Surry site file Surry site file

FRCFRM Fraction of Area Used for 

Farming Surry site file Surry site file Surry site file Surry site file Surry site file

FRFDL Fraction of Decontamination 

Cost for Labor

Level 1 0.3 0.3 0.3 0.3 0.3

Level 2 0.35 0.35 0.35 0.35 0.35

FRFIM Farm Wealth Improvements 

Fraction 0.25 0.25 0.25 0.25 0.25

FRMPRD Average Annual Farm 

Production Surry site file Surry site file Surry site file Surry site file Surry site file

FRNFIM Nonfarm Wealth 

Improvements Fraction 0.8 0.8 0.8 0.8 0.8

FRNFDL Nonfarm Labor Cost Fraction

Level 1 0.7 0.7 0.7 0.7 0.7

Level 2 0.5 0.5 0.5 0.5 0.5
C-56

Variable Description LTSBO Unmitigated 

STSBO ISLOCA TISGTR Mitigated

TISGTR 

Unmitigated

GWCOEF Long-Term Groundshine 

Coefficients

Term 1 0.5 0.5 0.5 0.5 0.5

Term 2 0.5 0.5 0.5 0.5 0.5

KSWTCH Diagnostic Output Option 

Switch 0 0 0 0 0

LBRRATE Long-Term Breathing Rate 0.000266 0.000266 0.000266 0.000266 0.000266

LGSHFAC Long-Term Groundshine 

Protection Factor 0.26 0.26 0.26 0.26 0.26

LPROTIN Long-Term Inhalation 

Protection Factor 0.46 0.46 0.46 0.46 0.46

LVLDEC Number of Decontamination 

Levels 2 2 2 2 2

NGWTRM

Number of Terms in 

Groundshine Weathering 

Equation

2 2 2 2 2

NRWTRM

Number of Terms in 

Resuspension Weathering 

Equation

3 3 3 3 3

POPCST Per Capita Cost of Long-Term 

Relocation 12000 12000 12000 12000 12000

RELCST Relocation Cost per PersonDay 172 172 172 172 172

RWCOEF Long-Term Resuspension 

Factor Coefficients

Term 1 0.00001 0.00001 0.00001 0.00001 0.00001

Term 2 0.0000001 0.0000001 0.0000001 0.0000001 0.0000001

Term 3 0.000000001 0.000000001 0.000000001 0.000000001 0.000000001

TFWKF Fraction Farmland Worker 

Time in Contaminated Zone

Level 1 0.1 0.1 0.1 0.1 0.1

Level 2 0.33 0.33 0.33 0.33 0.33

TFWKNF

Fraction Non farmland 

Worker Time in 

Contaminated Zone

Level 1 0.33 0.33 0.33 0.33 0.33

Level 2 0.33 0.33 0.33 0.33 0.33

TGWHLF Groundshine Weathering 

Half-Lives

Term 1 16000000 16000000 16000000 16000000 16000000

Term 2 2800000000 2800000000 2800000000 2800000000 2800000000

TIMDEC Decontamination Times

Level 1 5184000 5184000 5184000 5184000 5184000

Level 2 10368000 10368000 10368000 10368000 10368000

TMPACT Time Action Period Ends 158000000 158000000 158000000 158000000 158000000

TRWHLF Resuspension Weathering 

Half-Lives

Term 1 16000000 16000000 16000000 16000000 16000000

Term 2 160000000 160000000 160000000 160000000 160000000

Term 3 1600000000 1600000000 1600000000 1600000000 1600000000
C-57

Variable Description LTSBO Unmitigated 

STSBO ISLOCA TISGTR Mitigated

TISGTR 

Unmitigated

VALWF Value of Farm Wealth 6900 6900 6900 6900 6900

VALWNF Value of Nonfarm Wealth 220000 220000 220000 220000 220000

D. APPENDIX D

E. ISLOCA MODELING DETAILS

D-1

APPENDIX D: ISLOCA MODELING DETAILS

1 INTRODUCTION

This appendix expands on the subjects only mentioned or described briefly regarding the 

Interfacing System Loss of Coolant Accident (ISLOCA) in the main body of the report.

2 SAFEGUARD BUILDING MODELING DETAIL

Table D-1 delineates the flow paths defined in the Surry MELCOR model to represent doors, 

penetrations, etc., interconnecting the Safeguards Area, Containment Spray Pump Area, and 

Main Steam Valve House (MSVH). These flow paths are shown in Figure D-1.

Table D-1 Description of Flow Paths in MELCOR Representation of Surry Safeguards 

Area, Containment Spray Pump Area, and Main Steam Valve House

Flow 

Path 

I.D.

From 

Control 

Volume

To 

Control 

Volume

Size Description of Modeled Door, 

Penetration, etc.

600 850 855 4” radius half circle 

through 1’ thick wall

Wall drain at base of Outside 

Recirculation Spray Pump 2A 

cubicle

601 850 855

1’ diameter hole through

1’ thick wall with 10” 

diameter pipe passing 

through it

Pipe penetration in Outside 

Recirculation Spray Pump 2A 

cubicle wall, centered 3’-8” off 

floor

602 851 855 4” radius half circle 

through 1’ thick wall

Wall drain at base of Low Head 

Safety Injection Pump 1A cubicle

603 851 855

1’ diameter hole through

1’ thick wall with 10” 

diameter pipe passing 

through it

Pipe penetration in Low Head 

Safety Injection Pump 1A cubicle 

wall, centered 3’-8” off floor

604 852 856 4” radius half circle 

through 1’ thick wall

Wall drain at base of Low Head 

Safety Injection Pump 1B cubicle

605 852 856

1’ diameter hole through

1’ thick wall with 10” 

diameter pipe passing 

through it

Pipe penetration in Low Head 

Safety Injection Pump 1B cubicle 

wall, centered 3’-8” off floor

606 853 856 4” radius half circle 

through 1’ thick wall

Wall drain at base of Outside 

Recirculation Spray Pump 2B 

cubicle

607 853 856

1’ diameter hole through

1’ thick wall with 10” 

diameter pipe passing 

through it

Pipe penetration in Outside 

Recirculation Spray Pump 2B 

cubicle wall, centered 3’-8” off 

floor

608 854 855 6’-6” tall, 6’ wide Intra-volume connection, sill at 

floor
D-2

Flow 

Path 

I.D.

From 

Control 

Volume

To 

Control 

Volume

Size Description of Modeled Door, 

Penetration, etc.

609 854 856 6’-6” tall, 6’ wide Intra-volume connection, sill at 

floor

610 857 851 2' wide, 1.5' tall, through

1' thick wall

Electrical opening in Low Head 

Safety Injection Pump 1A cubicle 

wall, sill 13’ off floor

611 860 866 3'-5" wide, 7' tall, through

1' thick wall

Unlatched personnel door at grade 

for entry to Safeguards Area

612 860 819 3" wide, 60'-4" long

Shaker space between Safeguards 

Area roof and Containment 

(covered with flashing), opens at 

37.5 psf differential pressure

613 860 819 10 m2

Safeguards Area gross roof failure, 

opens at 75 psf differential

pressure

614 850 860 2’-6” x 2’-6”

Open manway to Outside 

Recirculation Spray Pump 2A 

cubicle, at grade

615 851 860 2’-6” x 2’-6”

Open manway to Low Pressure 

Injection Pump 1A cubicle, at 

grade

616 857 860 2’-6” x 4’ Open manway to 19'-6" level, at 

grade

617 852 860 2’-6” x 2’-6”

Open manway to Low Pressure 

Injection Pump 1B cubicle, at 

grade

618 853 860 2’-6” x 2’-6”

Open manway to Low Pressure 

Injection Pump 2B cubicle, at 

grade

619 854 857 2’-6” x 4’

Open manway in 19’-6” elevation

floor to 12' level outside of pump 

cubicles

620 854 857

5 total openings each 2'-6" 

square, assume 3/16" steel 

plate ==> each plate 

weighs 47.9 lb and 

requires 0.0532 psi 

differential pressure to 

move

Openings in 19'-6" elevation floor 

covered with unsecured steel plate, 

plates assumed to be wholly 

displaced from the openings by 

0.0532 psi differential pressure

621 854 865 2'-6" x 4' Opening at 12' elevation to Valve 

Pit

622 862 819 3'-4" x 7'-2", through 1’ 

thick wall

Personnel door at grade for entry 

to Containment Spray Pump Area 

and MSVH
D-3

Flow 

Path 

I.D.

From 

Control 

Volume

To 

Control 

Volume

Size Description of Modeled Door, 

Penetration, etc.

623 861 862 30” x 30”

Open manway to lower level of 

Containment Spray Pump Area 

building

624 856 861

2'-6" tall, 2' wide, through

1’ thick wall, estimated to 

be 50% full, sill 12’ off 

floor

Electrical opening between 

Safeguards Area and Containment 

Spray Pump Area

625 856 861

21" tall, 5' long, sill 4’-9” 

off floor, estimated to be 

50% full of piping

Pipe opening between Safeguards 

Area and Containment Spray 

Pump Area

626 862 819 2’ x 2’

4,000 cfm exhaust fan in the roof 

of the Containment Spray Pump 

Area

627 862 864

4'-4 1/2" wide, 8' high, 

through 1’ thick wall, 

opens to MSVH, 1/16” 

wide x 24’-10” tall shaker 

space between buildings 

included as fixed open 

fraction

Closed door between upper levels 

of Containment Spray Pump Area 

and MSVH assumed always 

closed

628 861 863

Doorway is 4' wide x 7' 

high, opening is 6'-6" wide 

x 11’-6” high, thresholds 

at floor

Open doorway & taller opening 

between lower levels of 

Containment Spray Pump Area 

and MSVH

629 862 819

2 holes of 1' diameter

through 2’ thick wall, 

estimated to be 12’ off 

floor

Abandoned pipe penetrations in 

the wall of the upper level of the 

Containment Spray Pump Area 

leading to the environment

630 864 819

8' wide x 4'-6" high but 

area reduced to 1.195 m2

to give zero flow (pre 

transient) from upper floor 

to lower floor of MSVH 

through chained-open 

manway, sill 2’ off floor

Fixed louvers in the wall of the 

upper level of MSVH

631 864 819 4’ x 4’ through 2’ thick 

wall, sill 10’ off floor

13,500 cfm exhaust ventilation fan 

in the wall of the upper level of 

MSVH

632 863 864 3' x 3'-7", assume 25 lbf is 

required to open the cover

Covered manway to lower level of 

MSVH
D-4

Flow 

Path 

I.D.

From 

Control 

Volume

To 

Control 

Volume

Size Description of Modeled Door, 

Penetration, etc.

633 863 810

Assume 4” std piping, 

103’ long, minor loss 

coefficient for elbows, 

branches, etc., of 7.9

4" drain from MSVH pit to 

Auxiliary Building, inlet at 11’-0” 

elevation, attaches to complex 

drain network terminating at 2’ 

elevation

634 863 810

7'-7" wide, 18" tall, 

through 2’ thick wall, sill 

at 8’-3” elevation

Pipe tunnel between MSVH pit 

and Auxiliary Building, filled with 

sprayed penetration sealant, 

assumed to dislodge when water 

fills the pit

636 853 861

1/32" equivalent width,** 

14'-6" tall, extends up 

from floor, assume angle 

iron overlaps concrete by 

3"

Shaker space between Safeguards 

Area and Containment Spray 

Pump Area, 3" shaker space with 

angle iron on each side

642 863 814

1/16" wide x 16' tall, 

extending up from floor, 

assume angle iron overlaps 

concrete by 3"

Shaker space between MSVH and 

Aux Bldg, 3" shaker space with 

angle iron on each side

643 862 819 3" wide x 22'-9" long

Shaker space between 

Containment Spray Pump Area 

roof and Containment (covered 

with flashing), opens at 37.5 psf 

differential pressure

645 857 855

2’-6” tall, 2’ wide, through

1’ thick wall, estimated to 

be 50% blocked

Penetration high in the wall

646 857 856

2’-6” tall, 2’ wide, through

1’ thick wall, estimated to 

be 50% blocked

Penetration high in the wall

647 862 819 10 m2

Containment Spray Pump Area 

gross roof failure, opens at 75 psf 

differential pressure

648 863 864 2’-6” x 2’-6”

Other covered (but chained open) 

manway to lower level (11'-6" 

elevation) of MSVH, the cover to 

this manway was chained open the 

day of NRC/SNL site visit 

(1/18/2011). Modeling assumes 

manway is uncovered
D-5

Flow 

Path 

I.D.

From 

Control 

Volume

To 

Control 

Volume

Size Description of Modeled Door, 

Penetration, etc.

649 860 862

5'-7" tall x 2'-10" wide 

(with half circle top), 

through 1’ thick wall, 

estimated to be 80% 

blocked

Ventilation ducting penetration 

between top levels of Safeguards 

Area and Containment Spray 

Pump Area

812 240 855

132’ of 6” Sch. 160 pipe, 

17’ of 10” Sch. 160 pipe, 

26’ of 10” Sch. 40 pipe, 

stainless steel, 

uninsulated, 2.57” 

diameter cavitating 

venturi, piping connects to 

top of cold leg, break 

centerline 1’-9” off 

Safeguards Area floor, 

7.7” elevation drop from 

top of cold leg to break 

centerline

LHSI piping, Loop A Cold Leg 

backwards to Safeguards Area, 

failed check valves, ISLOCA pipe 

break, active pool scrubbing, DFs 

to manage aerosol turbulent 

deposition and impaction

** For flow path I.D. 636 the width was determined based on review of site construction drawing details.
D-6

3 LOW HEAD SAFETY INJECTION PIPE MODELING DETAIL

The low head safety injection (LHSI) piping that would be subjected to reactor coolant system 

(RCS) pressure, should the two serial check valves in any one of the three cold leg injection lines 

fail, extends backwards from the check valves into the Safeguards Area and then through the 

Containment Spray Pump Area and Main Steam Valve House well into the Auxiliary Building. 

Two MELCOR representations of LHSI piping were utilized in the ISLOCA analysis – a simple 

representation where the fluid volume and metal mass of the piping was unaccounted for and a 

detailed representation where these parameters were accounted for. In the simple representation, 

frictional losses, form losses, and critical flow areas were represented in a single flow path. No 

control volumes or heat structures were included in the simple representation. In the detailed 

representation, flow losses and critical flow areas were portioned among several flow paths, and 

several control volumes and heat structures were included to account for the fluid volume and 

metal mass of the piping. Figure D-1 illustrates the detailed MELCOR model of the LHSI 

piping and roughly identifies the attributes of the physical piping. The specific piping modeled 

is the safety injection piping serving Cold Leg #2.

From the check valves to the isolation motor operated valve (MOV) 1890C, the 6” and 10” 

piping is Schedule 160 with 0.718” and 1.125” wall thicknesses, respectively, and a pressure 

rating higher than RCS operating pressure. From MOV 1890C back, none of the piping is rated 

strong enough to withstand RCS operating pressure. Between MOV 1890C and Flow Element 

(FE) 1945 for LHSI Train A and FE 1946 for LHSI Train B, the piping (10”) is Schedule 40 

having a 0.365” wall thickness. The 8” piping in the Safeguards buildings and all of the way 

back to isolation points within the Auxiliary Building is Schedule 40 having a wall thickness of 

0.322”. Between FEs 1945 and 1946 and the LHSI pump discharge check valves, the piping 

(10”) is Schedule 10 with a 0.165” wall thickness. It is these relatively short sections of thinwalled 10” piping that are judged as most susceptible to rupture given the dual check valve 

failure of the postulated ISLOCA. This piping ranges in centerline elevation from 13’-9” to 15’-

8” (1’-9” to 3’-8” off the floor). Significantly more of this piping exists outside the pump 

cubicles than inside them and slightly more of the piping is at the 13’-9” elevation than at the 

15’8” elevation. Figure D-2 shows the location of the different LHSI pipe sections in the 

Safeguards Area.

The ISLOCA break is assumed to occur centered 1’-9” above the floor with an area equivalent to 

the orifice area of FE 1945 or FE 1946 (each of which are of diameter 7.1469”). Only one of the 

thin-walled 10” piping sections is assumed to rupture (not both). The break is assumed to 

happen outside of a LHSI pump cubicle.
D-7

Figure D-1 Low Head Safety Injection Piping MELCOR Nodalization
D-8

Figure D-2 Low Head Safety Injection Pipe Sections in Safeguard Building
D-9

Noteworthy specifics of this piping and the modeling of it follow.

Size, length, schedule, etc. of the different sections:

All of the piping is either 6” or 10”. Most of the piping is 6”. All of the piping inside 

Containment is 6” while all of the piping in the Safeguards Area is 10”. All of the 6” pipe is 

Schedule 160. The schedule of the 10” piping varies as identified above. All of the pipe is 

stainless steel. Only the 3’ of pipe closest to the cold leg is insulated. This short length of 

insulation was disregarded in the MELCOR modeling because it was of negligible length and 

would not affect the deposition. Potentially, the piping in the Safeguards Area (and hence the 

break location) would be submerged for some time given the postulated ISLOCA.

Number of elbows, fittings, etc. in the different sections:

The following elbows, fittings, etc. exist in the different sections:

• 6” piping from the cold leg to the venturi - 1 entrance, 13 elbows, 2 swing-type check valves

• 6” pipe from venturi to 6”-to-10” transition - 3 elbows, 1 flow-branched tee, 1 expansion

• 10” piping from transition to MOV 1890C - 3 elbows, 1 gate valve (1890C)

• 10” piping from MOV 1890C to FE1946 - 3 elbows, 1 flow-branched tee, 1 gate valve 

(MOV 1864B)

Flow elements:

Flow elements (FE 1945/1946) between the LHSI pumps and MOV 1890C are characterized by 

sharp-edged orifices of diameter 7.1469”. The pipe rupture is assumed to be larger than the 

orifice such that the orifice is the more limiting flow area.

Isolation valve MOV 1890C:

Main LHSI isolation valve MOV 1890C, beyond which only Schedule 160 piping exists 

(reaching all the way to the cold leg), is not of submersible design. Consequently, this valve is 

susceptible to failure in an ISLOCA. This is because water issuing from the ISLOCA pipe break 

will flood the valve’s motor before the valve can be closed against the RCS blowdown 

proceeding through it. The assumption has been made in the ISLOCA analysis that this valve 

fails to close, i.e., remains in its normal open position. 

Venturi:

Each LHSI piping run to an RCS cold leg has a flow venturi incorporated. The venturi is a 

cavitating venturi with a gradual contraction and a gradual expansion. Its minimum flow 

diameter is 2.57”. The smallest flow area in an LHSI piping run is the minimum flow area in the 

venturi. 

Check valves:

Each LHSI piping run to an RCS cold leg has two check valves in series. The check valves are 

swing type. It is the failure of both of these valves in one of the LHSI piping runs that results in 

the ruptured pipe outside of containment in the ISLOCA scenario. The first check valve (in the 

normal flow direction) is 12’-9” from the cold leg. The second check valve is 2’-3” from the 

cold leg.
D-10

Safety injection piping approach to the cold leg:

The safety injection piping to Cold Leg #2 drops 4’-6” from its high point to where it adjoins the 

top of the cold leg. Between the two check valves in the 6” piping, 2” high head safety injection 

(HHSI) piping adjoins such that LHSI and HHSI share the last 3’-7” length of 6” piping leading 

to the cold leg. This commonality of piping between LHSI and HHSI proves important in the 

ISLOCA scenario as the HHSI delivered to the broken injection piping can’t reach the cold leg. 

Instead, it is driven out the pipe break by the strength of the blowdown. This phenomena is 

modeled in the MELCOR calculation by simply delivering the portion of HHSI destined to Cold 

Leg #2 directly to the Safeguards Area (to Control Volume 855).

Noteworthy in relation to this commonality of piping between LHSI and HHSI is that 

accumulator injection is through separate dedicated 12” piping to each cold leg. All accumulator 

injection, therefore, reaches the cold legs. 

Pool scrubbing at potentially underwater break location:

Optional MELCOR pool scrubbing logic (SPARC) was enabled in the flow path representing the 

LHSI piping (Flow Path 812) in the simple representation. This logic removes radionuclide 

aerosols and vapors from a gas as it flows through a pool of water. The flow area of Flow 

Elements 1945/1946 (7.1469” diameter) was specified. A single vent hole with horizontal 

orientation was called out.

The LHSI piping at Surry is not insulated. As the reactor core overheated in an ISLOCA, flow 

backwards through this piping would heat it. Fission product aerosols deposited in the piping 

would heat it further raising its temperature above that of the vapor flowing through it. The 

piping would lose heat from its outer surface via convection and radiation. Both of these heat 

transfer modes are enabled in the MELCOR modeling. For the LHSI piping internal to 

containment, convection heat transfer is to an isolated control volume filled with atmospheric air 

at 20 ̊C and radiation is to an encompassing environment at 20°C with an emissivity of 0.4 

defined for the piping. For the piping internal to the Safeguards Area, heat transfer is to an 

isolated control volume containing a water pool and air. Convection heat transfer is to the pool 

and or the air dependent upon whether the piping is submerged, partially submerged, or above 

the level of the pool. The level and temperature of the pool are time dependent and set equal to 

the level and temperature of the pool in the Safeguards Area in the initial overall ISLOCA 

calculation. Radiation is to an encompassing environment at the temperature of the pool with an 

emissivity of 0.4 defined for the piping. Note that radiation heat loss from the piping would be 

disabled by MELCOR given submergence of the piping in the pool.
D-11

4 SAFETY-RELATED FILTERED VENTILTION SYSTEM MODELING DETAIL

The ventilation system enlists both an exhaust train and a supply train, depicted in Figure D-3. 

There are two operational configurations for the exhaust train, normal and safety-related 

operations. Common to each operational state, the exhaust ventilation has seven inlets from the 

Safeguards Area and one inlet from the Containment Spray Pump Area with a combined rated 

volumetric flow rate of 11,000 cfm. A length of 205 ft is estimated for ducting connecting the 

Safeguards Area and the Containment Spray Pump Area (referred to jointly in this Section as 

Safeguards) with fans and filters located in the Auxiliary Building. 

During normal operations, the 11,000 cfm is drawn by dual 6,000 cfm exhaust fans. Once an 

injection signal has occurred, the exhaust from Safeguards is instead provided by dual 

36,000 cfm fans drawing though HEPA and charcoal filters. In addition to the 11,000 cfm from 

Safeguards, 61,000 cfm is drawn through Central Flow ducting serving the Auxiliary Building 

and other potentially contaminated areas. This 61,000 cfm is simply drawn from the 

environment in the MELCOR model. The safety-related fans and filters are housed in the 

Auxiliary Building. Two parallel filter banks are employed. The filter banks are unique to the 

safety-related filtration fans, unlike the normal exhaust fans. The safety-related fan ducting 

exhausts to the environment through a common stack located on the roof of the Auxiliary 

Building. Figure D-4 presents the safety-related filtration fan curve. 
D-12

Figure D-3 Safeguards Building Ventilation MELCOR Nodalization
D-13

Figure D-4 Auxiliary Ventilation System Safety-Related Fan Curve

The common ducting nodalization is divided into two control volumes, CV879 and CV880. 

CV879 represents the section of the ductwork where each exhaust line, flow paths FL680, 

FL686, and FL688, connect to the ducting. CV880 represents the estimated remaining physical 

volume of the ducting connecting Safeguards to the filters and fans in the Auxiliary Building. 

The ducting is represented by a 1⁄2-cm thick, cylindrical, stainless-steel heat structure and a flow 

area of 4.54 ft2. Original calculations provided by the utility demonstrated a 6 in. H2O gauge

pressure drop due to an orifice in the ducting. FL689 captures this pressure drop as well as the 

remaining frictional losses and flow characteristics throughout the ducting. 

The exhaust ventilation ducts in CV850-CV853 are sized to withdraw 21% of the total flow, 

11,000 cfm, from each of the 4 pump cubicles during normal operations. Similarly FL684, 

FL685, FL686, and FL688 flow areas are defined to withdraw 2%, 3%, 1%, and 11%, 

respectively (see Table D-2).

0

5

10

15

20

25

0 10,000 20,000 30,000 40,000 50,000

Pressure (inH O gauge) 2

Volumetric Flowrate (cfm)

Safety-Related Fan Curve

36,000 cfm fan
D-14

Table D-2 Ventilation Flow Path Flow Rates

Flow Path Vent Area

(ft2) Fraction of Total Flow Rate

FL680 1.36 21%

FL681 1.36 21%

FL682 1.36 21%

FL683 1.36 21%

FL684 0.13 2%

FL685 0.18 3%

FL686 0.11 1%

FL688 0.68 11%

Whether a safety injection signal has occurred will determine which fan/filtration system is 

employed during the calculation. During the steady state analysis, prior to an injection signal, 

the normal exhaust system is functional and the system exhausts to the environment through 

FL670 and FL671. A constant fan head was applied for the normal exhaust fans to produce the 

necessary steady state condition, given the 11,000 cfm reportedly drawn from Safeguards. After 

injection initiates, the normal-operation exhaust fans are isolated and the safety-related filtration 

and fans are utilized. Flow from the common ducting is passed to CV881, which acts as a 

homogeneous mixing volume between the Central Flow ventilation system and Safeguards 

ventilation prior to the atmosphere constituents passing through the filtration system. 

The filtration system is comprised of pre-filters, particulate filters (HEPA filters) and charcoal 

filters. Filter cells are constructed, from inlet to outlet, with one pre-filter, one HEPA filter, and 

two parallel charcoal filters. The filter cells are assembled into filter banks. Each bank is 

10 filter cells wide and 3 filter cells high. There are two filter banks in parallel corresponding 

with the two parallel fans. Each bank was designed for a flow rate of 36,000 cfm.

Due to limited information, the pre-filter is not represented in the model. The HEPA filter 

dimensions are 24” x 24” x 12”. Manufactured clean filter resistance, per Surry technical 

specification, are not to exceed 1 in. H2O gauge for the HEPA filters and 2 in. H2O gauge for the 

charcoal filters at 300 ft/min. A hydraulic diameter of 0.00105 m was specified for the HEPA 

filter to produce the reported upper bound resistance given laminar flow. The charcoal filter is 

modeled with identical laminar flow loss characteristics, but with twice the depth to produce the 

reported 2 in. H2O gauge clear filter resistance, twice the clean filter resistance as the HEPA 

filter. Decontamination factors of 200 and 100 are defined for the aerosol captured by HEPA 

filters and inorganic and elemental iodine vapor species capture by charcoal filters, respectively. 

As aerosols are captured, the flow resistance through the HEPA filters increases. The pressure 

drop due to mass loading is modeled by adjusting the laminar flow coefficient. A 2nd -order 

polynomial least squares fit of the data presented in “The Effects of Media Area on the Dust 

Holding Capacity of Deep Pleat HEPA Filters” [1] was applied to capture the increase in flow 

resistance (see equation below). FL672 and FL673 incorporate the delta pressure contributions 
D-15

from the fans, the particulate filters (clean loss), and the charcoal filters as well as from the 

aerosol loading on the particulate filters.

SLAM ≈ 0.0019 W2 + 0.1943W

where:

SLAM is the laminar flow coefficient and

W is the aerosol mass captured by a HEPA filter within the filter bank (kg) 

Per vendor description, due to the standard use of neoprene in the construction of nuclear 

installation HEPA filter gaskets, the specified maximum continuous service temperature is 

identified as 250 °F. Temperatures in the MELCOR calculation upstream of the filtration 

system, in CV881, were observed relative to this maximum service temperature, but the failure 

criterion is not included in the modeling. Note that any heating of the filters by the decay heat of 

captured fission products is not accounted for in the MELCOR model. Effectively, the 

assumption has been made that as long as the fans are running decay heat in the filters would be 

carried away by the flow through them. 

Pressure switches located just upstream of the safety related fans of the Auxiliary Ventilation 

System will trip the fans should pressure drop below -21 in. H2O gauge at the switches. This 

design feature is included in the MELCOR fan control logic.

The supply air to Safeguards, depicted in Figure D-3, is drawn from the environment. Fans 

originally installed in the supply ducting are no longer energized to ensure a sub-atmospheric 

pressure in Safeguards - negative pressure being maintained by exhaust ventilation operation. 

The supply ducting provides five inlets to the Safeguards Area and four inlets to the Containment 

Spray Pump Area.
D-16

5 BUILDING BOUNDARY FAILURE CRITERIA DETAIL

An ISLOCA at Surry and consequential RCS blowdown into the Safeguards Area would 

pressurize the Safeguards buildings potentially failing building boundaries. Further insults to 

building boundaries could potentially result from subsequent hydrogen burns. A scrutinizing of 

the construct of the Safeguard buildings identified weak points as described here.

Opening of the Safeguards Area personnel door:

The lone personnel door for accessing the Safeguards Area is centrally located in the top floor of 

the building at ground level. The door opens outward and is equipped with a closer. The door is 

36” x 80” and opens to a modest push by hand. It has no latch. Flow Path 611 shown in Figure 

D-5 represents this door in the MELCOR model. Given a meaningful elevation of Safeguards 

Area pressure (i.e., 1 in. H2O gauge or 5.19 psfg), this flow path is opened to the environment. 

The flow path is reclosed upon loss of the elevated pressure. (1” of water would put a force of 

over 100 lb on this door.)

Tearing of the flashing covering Safeguards Area and/or Containment Spray Pump Area roof 

shaker spaces:

Per the Surry UFSAR, certain buildings at Surry are designed to withstand a tornado with winds

up to 360 mph, which equates to a pressure differential of 2.30 psi (the dynamic pressure of a 

360 mph wind). The Safeguards buildings are not such buildings, i.e., the Safeguards buildings 

are not designated as designed to withstand a tornado. The Safeguards buildings are however, 

reinforced concrete buildings and the UFSAR points out that no structural damage is known to 

have resulted to a reinforced concrete building in a tornado. Buildings that are not designed to 

withstand a tornado, also per the UFSAR, are designed to withstand wind loads (on their walls) 

based on their elevation. Roofs are designed for uplift load using 1.25 times the wind load. For 

the Safeguards Area, the design load is 30 psf (0.208 psi) for the walls and the roof uplift load is 

1.25 x 30 psf or 37.5 psf (0.260 psi).

The construction of the Safeguards buildings is reinforced concrete with steel roofing. The walls 

and floors of the buildings are largely below grade. Given the stout construct of the walls, the 

assumption has been made in the ISLOCA analysis that the weak points in the pressure 

boundaries of the buildings are their roofs. This assumption is exclusive of the latch-less 

personnel door to the Safeguards Area described above.

The Safeguards buildings do not have integral back walls. The buildings back up to 

Containment and the Containment cylinder serves as a back wall to the buildings. The steel 

roofing on the buildings extends to within a few inches of the Containment wall leaving a gap 

(i.e., shaker space) that is closed with flashing. This roof flashing has been judged as the weak 

point in the Safeguards building roofs and hence the weakest point in the pressure boundaries of 

the buildings notwithstanding the Safeguards Area personnel door. The flashing has been 

specified in the MELCOR model to tear given a pressure differential across it of 37.5 psf (i.e., 

the design uplift load for the Safeguards Area roof). The flashing is represented with Flow Paths 

612 and 643 for the Safeguards Area and the Containment Spray Pump Area, respectively. Flow 

Path 612 reflects a gap 3” wide by 60’-4” long while Flow Path 643 reflects a gap 3” wide by 

22’-9” long.
D-17

Gross failure of the Safeguards Area and/or Containment Spray Pump Area roof as a 

consequence of a hydrogen burn:

The tearing of roof flashing described above would significantly vent the Safeguards buildings 

given an ISLOCA and consequential RCS blowdown into them. The venting, however, might 

not be sufficient to curtail further damage to the boundaries of the buildings from the blowdown 

or from subsequent hydrogen burns occurring within the buildings.

To address the potential for further damage to the Safeguards buildings due to overpressure, 

additional flow paths have been included in the MELCOR model between the buildings and the 

environment. These flow paths manage a gross roof failure. The flow paths open given a 

pressure greater than twice the design uplift load for the roofs (2 x 37.5 psf or 0.520 psi). The 

factor of two here reflects a reasonable assumption of safety margin in the building’s design and 

the observation that the roofs consist of heavy-gauge corrugated steel topped by waterproof 

fabric, foam board, and concrete pavers. The particular flow paths are 613 and 647 for the 

Safeguards Area and the Containment Spray Pump Area, respectively. The flow paths have an 

area of 10 m2 (each). 
D-18

Figure D-5 Safeguards Building MELCOR Nodalization
D-19

6 IMPACTION MODELING AND TURBULENT DEPOSITION

The new MELCOR models for turbulent deposition in straight pipes and deposition in elbows 

were validated against the results of the LWR Aerosol Containment Experiments (LACE) 

Project. The LACE project, organized by Electric Power Research Institute, performed largescale experiments to investigate aerosol behavior under simulated LWR accident conditions to 

provide a database for testing containment-aerosol and related thermal-hydraulic computer 

codes. The tests studied aerosol behavior under postulated severe accidents conditions not 

adequately addressed by previous test programs. The studied conditions included total 

containment bypass (i.e., ISLOCA). Individual LACE tests that studied ISLOCA conditions 

were CB-1, CB-2, CB-3, LA1, LA3A, LA3B, and LA3C. LACE reports for these tests are the 

following:

• “Aerosol Behavior Under LWR Containment Bypass Conditions—Results of Tests CB-1, 

CB-2 and CB-3,” LACE TR-001, November 1986

• “Aerosol Behavior in LWR Containment Bypass Piping—Results of LACE Test LA3,” 

LACE TR-011, July 1987

• “Summary of Posttest Aerosol Code Comparisons for LWR Aerosol Containment 

Experiment (LACE) LA1,” LACE TR-022, ORNL/M-365, October 1987

• “Summary of Posttest Aerosol Code-Comparisons Results for LWR Aerosol 

Containment Experiment (LACE) LA3,” LACE TR-024, ORNL/M-492, June 1988

A summary of the LACE project is given in “The LWR Aerosol Containment Experiments 

(LACE) Project, Summary Report” [3]. 

The new MELCOR models were applied in a separate-effects calculation that used boundary 

conditions from the integral full-plant calculation to estimate aerosol retention in the LHSI 

piping. The estimated aerosol retention for each fission product class was in turn used to specify 

a decontamination factor (DF) for each MELCOR radionuclide class as input to the integral fullplant calculation which represented the LHSI piping as a single junction. 
D-20

Assessment Against LACE Experiments

The LACE tests experimentally examined the transport and retention of aerosols typical of 

LWRs through pipes with high speed flow and in containment volumes during rapid 

depressurization. In particular, the LA1 and LA3 tests examined deposition in pipe flow under 

conditions of containment bypass to provide a database for validation of aerosol computer 

simulation. Accident scenarios represented by the LACE tests include potentially high 

consequence accidents for which containment may be bypassed altogether, the containment 

function is impaired early in the accident, or a large fission product release occurs 

simultaneously with containment failure. 

Specific objectives of these tests were to provide validation data that would expose important 

dependencies in modeling deposition. In particular the following test conditions were examined:

• Effect of gas velocity through the pipe

• Effect of aerosol composition 

• Effect of aerosol size distribution

Overall test conditions and results are summarized in Table D-4 below. It is important to note 

that the range of gas velocities ranges from 23 m/s to 100 m/s, which comprises a range of 

Reynolds numbers between 30,000 to 300,000. The Wood models [22] implemented into 

MELCOR were validated against the data of Liu and Agarwal [12], which was performed at 

Reynolds numbers of 10,000 – 50,000. The VICTORIA models are based on Friedlander & 

Johnstone’s data [7] with Reynolds numbers of 2800 – 44,000 as well as some of Sehmel’s [17]

experiments with Reynolds numbers of 4200 - 61,000. The conditions of the LA1 and LA3A 

tests were beyond the range of Reynolds numbers in the database used to develop these 

deposition models though it has been observed that the dependence on Reynolds number may be 

small. However, some accident conditions may extend to even higher flow velocities, 

approaching sonic velocities, well beyond the database for these models. As discussed later in 

this section, higher velocities can lead to other effects, such as resuspension and/or entrainment 

of deposited material. There does not appear to be any data that can be used to validate flow 

approaching sonic velocities. 
D-21

Table D-3 Aerosol and Thermal-Hydraulic Conditions at Inlet to Test Pipe

Test Aerosol

CsOH 

Mass 

Fraction

Carrier 

Gas

Gas 

Velocity 

(m/s)

Temp. 

(

o

C)

Aerosol 

Source 

Rate 

(g/s)

Aerosol 

Size 

AMMD 

(μm)

Mass 

Retention 

Fraction

LA1 CsOH/

MnO 0.42 Airsteam 96 247 1.1 1.6 > 0.98

LA3

A

CsOH/

MnO 0.18 N2-steam 75 298 0.6 1.4 > 0.7

LA3B CsOH/

MnO 0.12 N2-steam 24 303 0.9 2.4 > 0.4

LA3C CsOH/

MnO 0.38 N2-steam 23 300 0.9 1.9 > 0.7

Figure D-6 and Figure D-7 are photographs from LACE TR-011 that show pipe deposits 

observed in test LA3C.

Figure D-6 Test LA3C, Downstream end of pipe 6 showing large deposit
D-22

Figure D-7 Test LA3C, downstream end of pipe section 5 showing large bend deposit
D-23

MELCOR Turbulent Deposition Models

MELCOR has long had models for predicting aerosol deposition from gravitational settling, 

diffusion to surfaces, thermophoresis, diffusion to surfaces, and diffusiophoresis [8]. Recently, 

however, several models for turbulent deposition in pipes have been implemented into MELCOR 

1.8.6. In particular, Wood’s models for turbulent deposition in pipes with smooth and rough 

surfaces, as well as the models used by the VICTORIA code were added to MELCOR’s 

deposition modeling capabilities. These models are based on a few fundamental assumptions. 

For example, it is assumed that the concentration of aerosol in the gas is small enough that the 

effect of particle-particle interactions is small. In addition, it is assumed that the influence of the 

aerosol particles on the flow stream is negligible. Not only does this mean that the micro effects 

on turbulent eddies are negligible, but the macro effects from deposition on surfaces with the 

subsequent modification of surface roughness and reduction in flow area is not modeled. 

Particle deposition is modeled in terms of a deposition velocity Vd, defined as the ratio of the 

time-averaged particle flux to the surface to the time-averaged airborne particle concentration in 

the duct. This is then implemented into MELCOR in calculating the rate of deposition on a 

surface: 

V C

dt

dM

A d

C = 1

 

(1)

Where:

Vd - deposition velocity

C - particle mass concentration

MC - Mass deposition rate

A - Surface area of deposition surface

It is common to correlate the deposition velocity with the particle relaxation time, τ. This is the 

characteristic time for a particle velocity to respond to a change in air velocity. For spherical 

particles of diameter dp and density ρp in the Stokes flow regime, it is calculated as:

g

m DpCslip

μ

ρ

τ

18

2

=

 

(2)

Where:

Cslip - slip correction factor (-)

Non-dimensional forms for the deposition velocity and relaxation time are used in the MELCOR 

models : 
D-24

*

*

u

V V d

d = 

(3)

( )

g

g u

μ

τρ

τ

2 *

* = (4)

Where: * u - friction velocity

For flow through smooth cylindrical channels, the friction velocity is found from: 

2

* f u = V

 

(5)

1 4 Re

0.0791 f = 

(6)

g

DH gV

μ

ρ Re =

 

(7)

Where:

DH -hydraulic diameter of conduit

Three models, Wood’s model for turbulent deposition in rough pipes, Wood’s model for 

turbulent deposition in smooth pipes, and Sehmel’s correlation for perfect particle sinks (i.e., the 

VICTORIA model) for turbulent deposition in smooth pipes were implemented in MELCOR and 

assessed against LACE data in this study. In addition, we have plotted the non-dimensional 

deposition velocity as a function of the non-dimensional relaxation time in Figure D-8 and

Figure D-9 along with non-dimensionalized data from numerous studies on turbulent deposition. 

There is significant scatter in the data depicted in Figure D-8 and Figure D-9. Friedlander & 

Johnstone [7] provided some of the earliest experimental data from particle deposition in 

turbulent flows. Their investigations qualitatively demonstrated that deposition velocity 

increased with flow velocity and particle diameter. The data of Liu & Agarwal [12] shows much 

less scatter with a definite trend for increased deposition at higher relaxation times. These 

experiments are frequently referenced because of their credibility due the reproducibility of the 

data and the quality of their methods in obtaining the data. This data set is typically used to 

benchmark correlations for turbulent deposition. The data from Sehmel [17] is the most 

extensive in range of relaxation times, Reynolds number, and the effect of pipe size. However, 

this data shows a significant degree of scatter possibly due to less rigorous experimental 

techniques.
D-25

Figure D-8 Dimensionless deposition velocity plotted as a function of dimensionless 

relaxation time for MELCOR turbulent deposition models
D-26

Figure D-9 Sehmel's correlations compared with experimental data
D-27

Note that Wood’s models are based on the data from Liu and Agarwal whereas the Sehmel’s 

model (VICTORIA) are based on data from Sehmel [17] and Friedlander and Johnstone [7]. For 

τ

* > 1.0, the Wood model for smooth pipes closely reproduces the data from Liu and Agarwal. 

For τ

* < 1.0, the curve depends on the particular test since the model now depends on the Sc 

number and not just the non-dimensional relaxation time. Also, note that the Wood model for 

rough surfaces collapses to the same curves as predicted for the Wood model for smooth surfaces 

when the roughness is zero. When the pipes are rough, the Wood model predicts larger 

deposition velocities for small particles, dependent on the flow conditions of the test being 

modeled. It has been noted that even micro scale roughness can significantly enhance particle 

deposition, particularly for small particles, for which Brownian diffusion is an important 

mechanism for transporting the particle to the wall. These models are discussed in more detail in 

Section 0 of this appendix.

In the section describing the details of the Sehmel model incorporated into MELCOR, it was 

pointed out that Sehmel obtained two curve fits to correlate data from his tests as well as data 

from Friedlander and Johnstone. The first curve in Figure D-9 shows the correlation that was 

implemented into MELCOR based on the VICTORIA code. Note that this equation was fit to a 

limited data set corresponding to those experiments alone where surfaces were treated, often with 

petroleum jelly or viscous oils, to minimize particle bounce. The second curve, also published 

by Sehmel, represents a curve fit to a more complete set of experimental data for real surfaces 

with non-perfect particle sinks. The Sehmel – perfect sink (VICTORIA) model therefore 

predicts higher deposition velocities than was observed for many of the tests on turbulent 

deposition.

MELCOR Bend Impaction Models

In addition, three models for simulation of inertial deposition in bends were added to MELCOR, 

the Pui model (VICTORIA), the INL bend model (Merrill), and McFarland’s bend model. The 

Pui model and McFarland’s model are purely empirical fits to either experimental data (Pui) or 

Lagrangian simulations [13]. The INL model is more theoretical in nature and is based on the 

terminal velocity of a particle due to the centrifugal force acting on a particle traversing a bend. 

Pui’s experiments were for 90o bends in 0.503-0.851 cm diameter tubes and his correlation does 

not correlate any dependency on bend angle or radius of curvature whereas the other models do. 

The dependence of these three bend models on the particle Stokes number is shown in Figure D10 for the case of a 90o bend. Note that McFarland’s model depends on the radius of curvature 

of the bend. For the LA3 tests, this radius of curvature is 2.86, and the penetration of particles 

for bends with this radius of curvature is shown in this plot. Note that there is good agreement 

between the INL model and that of McFarland’s. The Pui model appears to predict more

deposition in the bend in comparison to the other models.

Validation calculations for the LA1 and LA3 tests were made for both the Pui and INL model 

and are reported in the following sections. Deposition from McFarland’s model is similar 

enough to the INL model to give insignificant differences in the MELCOR simulations. 

Sensitivity calculations were also conducted to assess the importance of nodalization of control 

volumes as well as the number of discrete aerosol particle size bins.
D-28

Figure D-10 Bend penetration calculated as function of Stokes # for various bend models
D-29

LACE LA1

6.1.1 Test Conditions and Modeling

The LA1 test was designed to simulate the LWR containment bypass accident (i.e., ISLOCA) 

sequence conditions. CsOH and MnO aerosols were injected into a 0.063-m diameter, 

approximately 30-m-long test pipe. The pipe had six 90o bends. The pipe inlet flow velocity was 

roughly 100 m/s, and the outlet flow velocity was roughly 200 m/s due to the pressure drop 

through the pipe. Aerosols that transported out of the pipe were then allowed to become airborne 

in the 852-m3 CSTP vessel, where their subsequent aerosol behavior was then studied. During 

part of the experiment, (i.e., the aerosol generation period), steam, non-condensable gas, and 

aerosols were vented or "leaked" to a scrubber.

A description of the LA1 test sections is provided in Table D-4 below. For validation of the 

models, a detailed nodalization was performed in which bends were isolated in control volumes 

or combined with adjacent valves, which had small deposition surface areas. However, a 

sensitivity calculation was also performed in which several test sections were lumped into a 

single control volume and deposition was calculated on a single heat structure associated with 

that control volume. Using this nodalization, all surfaces associated with that control volume 

compete equally for aerosol deposition. In reality, those surfaces upstream would remove 

aerosols from the flow that would reduce the source for those surfaces that are downstream. This 

lumped nodalization was used to assess the accuracy of the calculations when multiple heat 

surfaces are lumped into a single volume, as might be performed in a calculation for a 

commercial reactor. Figure D-11 and Figure D-12 show a schematic of the test section where 

pipe section numbers corresponding to MELCOR control volumes are shown in the drawing and 

in the accompanying table.

Source aerosol diameters are nodalized into 10 sections in MELCOR starting with a lower limit 

of 5x10-8 m to an upper bound of 1x10-4 m. The source distribution assumed in MELCOR was a 

lognormal distribution characterized by an aerodynamic mass median diameter (AMMD) and a 

geometric standard deviation (GSD). The LACE data includes both an AMMD and a GSD for an 

assumed lognormal particle size distribution where samples were taken four times during the test 

and analyzed by a cascade impactor with a reported accuracy of 25%. The assumed source 

distributions the LA1 test are shown in Figure D-13 and the AMMD used is the average of the 

four measurements. Notice that 92% of the aerosol source particles have diameters in the three 

size bins between 0.489 μm and 2.24 μm. In addition, a calculation is performed where the 

number of sections is doubled to 20 to assess the sensitivity of the calculations to the section 

nodalization. The 20-section source distribution is also shown in this figure.

Deposition velocities calculated by size bin for various deposition models are also shown in

Figure D-14. These velocities were calculated at representative conditions in pipe section 4. For 

the Wood models depicted, all models clearly display three regimes of deposition. Though the 

Brownian and eddy diffusion regime is not visible for the Victoria model, it is because it predicts 

an earlier onset of the eddy diffusion impaction zone at extremely small aerosol diameters 

beyond the range of the plot. Consequently, deposition velocities predicted by the Victoria 

model for submicron particles are generally greater than predicted by the Wood models. For 

supermicron sized particles, all models predict a relatively constant or slightly decreasing 
D-30

deposition velocity as particle size increases. As would be expected, the Wood model for rough 

pipes shows greater deposition (compared to his model for smooth pipes) for submicron particles 

when the roughness is about 5x10-5 m, but when the roughness is small, the deposition velocity 

for Wood’s model for rough pipes and his model for smooth pipes give approximately the same 

curve. In the MELCOR implementation, the initial surface roughness is specified by the user, 

however, as particles accumulate on the surface, the roughness calculated internally is reduced 

by the mass that accumulates, with a minimum roughness being the particle size. This would 

mean that for significant accumulation on the surface, Wood’s model for rough surfaces would 

give similar results as for his model for smooth pipes. 

Deposition efficiencies are used to characterize deposition in the pipe bends and have been 

plotted in Figure D-14. These velocities were calculated at conditions in pipe section 3. Both 

the Pui and INL bend model are used in these assessment calculations. Both predict a rapidly 

increasing deposition efficiency for supermicron particles where the Pui model predicts larger 

deposition efficiencies than the INL bend model for all particle diameters.
D-31

Table D-4 LA1 Experiment – Test Section Dimension

Section 

No.

LA1 -

Detailed

HS Number

LA1-

Coarse

HS Number

Pipe 

Description

Flow 

Direction

Diameter 

(cm)

Length 

(m)

1a 1011 1011 Straight East 30 1.52

1b 1011 1011 90' Bend --- 30 0.72

1c 1011 1011 Reducer UP 30 - 10 0.28

2 1010 1010 Ball Valve UP 10 0.23

3 1010 1010 Reducer UP 10 - 6.3 0.28

4 1110 1100 Straight UP 6.3 2.26

5 1120 1100 90' Bend UP 6.3 0.36

6V 1120 1200 Ball Valve Horizontal 6.3 0.19

6 1130 1200 Straight Horizontal 6.3 4.2

7 1140 1200 90' Bend Horizontal 6.3 0.36

8 1150 1300 Straight Horizontal 6.3 3.74

9 1160 1400 Straight Horizontal 6.3 4.32

10 1170 1500 90' Bend Horizontal 6.3 0.36

11 1180 1500 Straight Horizontal 6.3 0.62

11V 1190 1500 Ball Valve Horizontal 6.3 0.19

12 1190 1600 90' Bend Vertical 6.3 0.36

13 1200 1600 Straight Vertical 6.3 4.32

14 1210 1700 Straight Vertical 6.3 4.34

15 1220 1800 90' Bend Vertical 6.3 0.35

16 1230 1800 Straight Horizontal 6.3 1.74

17 1240 1900 90' Bend Horizontal 6.3 0.33

18 1240 1900 Ball Valve Horizontal 6.3 0.19

19 1240 1900 Straight Horizontal 6.3 0.71

22 1250 2250 Transition Horizontal 6.3 - 30 1.17

23a 2250 2250 Straight Horizontal 30 1.95

23b 2250 2250 90' Bend UP 30 0.58

Total 35.67
D-32

 

Figure D-11 Detailed LA1 MELCOR Nodalization

MELCOR Model Nodalization:

Control

Volume HS Pipe Sections

CV011 1011 1

CV012 1010 2,3

CV011 1110 4

CV012 1120 5,6V

CV013 1130 6

CV014 1140 7

CV015 1150 8

CV016 1160 9

CV017 1170 10

CV018 1180 11

CV019 1190 11V,12

CV020 1200 13

CV021 1210 14

CV022 1220 15

CV023 1230 16

CV024 1240 17,18,19

CV025 1250 20
D-33

Figure D-12 Coarse LA1 MELCOR Nodalization
D-34

Figure D-13 MELCOR CsOH log-normal source distribution using AMMD = 2.1 

μm and GSD =1.80. (10 particle sizes for base case and 20 particle 

sizes for sensitivity calculation)
D-35

Figure D-14 Calculated deposition velocity for LA1 test using MELCOR models. 

In Figure D-14, the green curve (triangles) shows Wood’s model for rough pipes 

evaluated for a surface roughness of approximately zero.
D-36

Figure D-15 Deposition efficiency calculated for LA1 bends using new 

MELCOR models
D-37

6.1.2 Results

Table D-5 shows that MELCOR predicts a retention factor of 94 -97% which is in good 

agreement with the LA1 test data (>98%). MELCOR also predicts retention in the bends 

that are reasonably accurate, particularly when the bend deposition model is used. 

Table D-5 MELCOR and LA1 Experiment Retention Factors

Experimental

Results

Calculateda

INLd VICe Woodf I-V**g

Retention in 6.3 cm pipeb (%) >98% 94% 97% 94% 95%

Bend Depositionc (%) 37% 29% 35% 42% 37% ** I-V indicates indicate the INL model for the straight sections (I) and the Victoria model for the 

bends (V). a Values reported are for fine nodalization model.

b Values reported are the deposited mass divided by the total source mass less the mass deposited 

upstream of the 6.3 cm pipe (i.e., mass deposited on torches, aerosol mixing vessel (AMV), and 

bypass). c Values reported are the fraction of the deposition in the 6.3 cm pipe that was found in the pipe bends. 

d Uses Wood model for rough pipes and INL bend model. e Uses Sehmel model for perfect particle sinks in straight pipes and Pui bend model. f Uses Wood’s model for smooth pipes and Pui bend model. 

g Uses Wood’s model for rough pipes and Pui bend model.

The following sections graphically present calculation results along the test piping. 

Where available, data is presented alongside the calculated results. In addition, results 

are presented from the code comparison study that was performed as part of the LACE 

project.

Several sensitivity calculations were also performed. Because the deposition velocity is

dependent on particle size, a sensitivity calculation was performed where the particle size 

distribution was modeled with 20 nodes, compared to the default value of 10. A 

calculation was also performed in which several volumes (i.e., bends and straight pipe 

sections) were lumped together for a coarser nodalization. This is important to know 

because it may influence the nodalization choices in building a nodalization for a 

commercial reactor where volume size constraints can affect calculation performance. 

For this calculation, if more than one straight pipe section is lumped together, the 

deposition plotted for each surface is estimated by partitioning the total deposition for the 

combined heat structure by the surface area of each pipe section. Similarly, if two bends 

are lumped together and associated with a single heat structure with combined surface 

area, the total bend deposition calculated is split between the two bends when generating 

the plots. 

Finally, a calculation was performed to investigate the importance of the sticking factor. 

Uncertainty may be associated with this sticking factor so it is important to know how 

sensitive the calculation is to the calculated value. A calculation was performed for 
D-38

which the sticking factor was assumed to be 1.0 and compared to the results generated 

from the calculated sticking factor.

6.1.2.1 Thermal Hydraulic Results

Pressures and velocities along the pipe train are plotted below (i.e., Figure D-16 and

Figure D-17). Note that the calculated velocities agree well with the values reported 

(100 m/s at inlet and 200 m/s at outlet) in the code comparison report.

Figure D-16 Pressure profile along test pipe
D-39

Figure D-17 Velocity profile along test pipe

6.1.2.2 Deposition Profiles

Deposition profiles along the pipe train are plotted below in Figure D-18 through Figure 

D-23. 

Figure D-18 Deposition profile for CsOH along LA1 test section

(Calculated sticking factor)

0

50

100

150

200

250

0 5 10 15 20 25 30 35

Velocity (m/s)

Location along pipe (m)
D-40

Figure D-19 LA1 post test calculation of CsOH aerosol deposition for codes in code 

comparison report including bend deposition models (LACE TR-022)

Figure D-20 Deposition profile for MnO along LA1 test section

(Calculated sticking factor)
D-41

Figure D-21 LA1 post test calculation of MnO aerosol deposition for codes in code 

comparison report including bend deposition models (LACE TR-022)

Figure D-22 shows the integrated deposition at a few points along the piping and 

therefore shows the cumulative deposition for multiple sections. 

Figure D-22 Deposition profile for CsOH along LA1 test section

(Sticking factor = 1)
D-42

Figure D-23 Deposition profile for CsOH along LA1 test section - colored

(Colored regions indicate lumped nodalization)

6.1.2.3 Observations

1. The MELCOR calculations over predict deposition early in the test train even 

though overall retention for the test section is well predicted. This is true for all 

models tested. In contrast, the test data shows that the linear deposition density in 

the straight pipe sections is nearly uniform along most of the test train, though it 

does taper off at the end. To explain this behavior, it may be possible that the 

high flow velocities (100 – 200 m/s) lead to resuspension and deposition further 

down the pipe or flow of deposited mass along the pipe. MELCOR does not 

calculate resuspension of aerosol particles or entrainment of deposited material. 

This is consistent with recommendations by the VICTORIA peer review 

committee [24].

2. When compared with the code comparison calculation (TR-022), the MELCOR 

models estimate the total pipe retention as well as any of the other codes 

participating in that activity. MELCOR tends to overestimate the deposition, 

particularly in upstream test sections where many of the other codes 

underestimate total deposition. 

3. All MELCOR models greatly over predict deposition in test section 4 (< 2.5 m). 

This pipe section happens to be a vertical pipe section and it is possible that the 

orientation of the pipe may be an important role. However, note also that sections 

13 and 14 are also vertical pipe sections. For these pipe sections, MELCOR 

predicts very little deposition though they occur much further downstream after 

airborne aerosol source has been greatly depleted by deposition upstream.
D-43

4. Deposition in bends is important, since about 37% of the deposited mass is found 

in the bends. Both the Pui and the INL model do a reasonable job of predicting 

the bend deposition.

5. The MELCOR calculation of MnO deposition along the piping is similar in 

characteristics to that of CsOH deposition. About 94% of the MnO aerosol was 

deposited in the test section while the MELCOR models predict between 91% 

(Wood’s smooth/Pui) and 98% (Sehmel/Pui).

6. Calculating the sticking factor with Merrill’s model does not significantly affect 

results for these calculations. The calculated sticking factor varies with time. As 

mass is accumulated the sticking factor approaches 1.0 very quickly.

7. Lumping the bends with straight pipe sections leads to a modest reduction in the 

overall retention in the test section and increases the spread between the various 

models. The approximation appears to be less significant for the Wood (rough) 

models. Overall, this approximation was quite reasonable for all models.

LACE LA3

6.1.3 Test Conditions and Modeling

The intent of LA3 experiments was to characterize the aerosol transport and deposition in 

a pipe with the two parameters: gas velocity, and ratio of insoluble to soluble aerosol. 

LA3 was designed to explore the intermediate and low gas velocities (i.e., ~20 to 100 m/s 

at the test section inlet). The ratio of the insoluble to soluble (hygroscopic) aerosol 

determines the absorption of water and therefore the liquid content and stickiness of the 

aerosol. Both LA3A and LA3B were intended to have a higher ratio of 8:1, in 

comparison to LA3C of roughly 2:1. Test conditions are summarized in Table D-3. 

A description of the LA3 test sections is provided in Table D-6. Note that for the purpose 

of these calculations, several test sections were combined into a single control volume 

and deposition was calculated on a single heat structure associated with that control 

volume. However, both a detailed and a coarser lumped calculation (Figure D-24 and 

Figure D-25) were performed to assess the impact of nodalization. 

Source aerosol diameters are nodalized into 10 sections in MELCOR starting with a 

lower limit of 1x10-7 m to an upper bound of 5x10-5 m. The source distribution is 

assumed to be a lognormal distribution with average and geometric standard deviation 

reported for the experiment and the total source is assumed to be released at a constant 

average rate. The assumed source distribution for each LA3 test is shown in Figure D26. Note that the AMMD and GSD reported for each test are based on the average of 

four samples taken during the test and analyzed by a cascade impactor with a reported
D-44

accuracy of 25%. The measured source distribution at each of the four sample times is 

presented in Figure D-27 to Figure D-29. This distribution does not vary much for LA3B 

or LA3C but the variation for LA3A is large enough to affect results. Because of this, a 

more detailed calculation was performed using these four source distributions and the 

time dependent source generation rate, which is shown in Figure D-30.

Deposition velocities calculated by size bin for various deposition models are shown in

Figure D-31 through Figure D-33. These velocities were calculated at conditions in pipe 

section 3. All models are consistent in their trends and three regions of deposition are

seen though there are differences in magnitudes. The VICTORIA models show an earlier 

onset of the eddy diffusion-impaction zone and consequently, deposition velocities for 

submicron particles are greater than predicted by the Wood models. For supermicron 

sized particles, all models predict a relatively constant or slightly decreasing deposition 

velocity as particle size increases. As would be expected, the Wood model for rough 

pipes shows greater deposition (compared to his model for smooth pipes) for submicron 

particles when the roughness is about 5x10-5 m, but when the roughness is smaller, the 

deposition velocity for Wood’s model for rough pipes and his model for smooth pipes 

give approximately the same curve. In the MELCOR implementation, the initial surface 

roughness is specified by the user, however, as particles accumulate on the surface, the 

roughness calculated internally is reduced by the mass that accumulates, with a minimum 

roughness being the particle size. This would mean that for significant accumulation on 

the surface, Wood’s model for rough surfaces would give similar results as for his model 

for smooth pipes. 

Figure D-34 shows the deposition efficiency calculated for LA3 bends using the 

MELCOR Pui, INL and McFarland bend penetration models. The INL and McFarland 

models are in good agreement for all LA3 experiments. The MELCOR Pui model 

consistently is lower than the INL and McFarland models for all LA3 experiments.
D-45

Table D-6 LA3 Experiment – Test Section Dimension

Section 

No.

LA3 -

Detailed

HS Number

LA3 -

Coarse

HS Number

Pipe

Description

Flow 

Direction

Diameter 

(cm)

Length 

(m)

1a 1110 1110 Straight East 30 1.52

1b 1110 1110 90' Bend --- 30 0.72

1c 1110 1110 Reducer UP 30 - 10 0.28

2 1120 1120 Ball Valve UP 10 0.23

3 1120 1120 Reducer UP 10 - 6.3 0.28

4 1130 1120 Straight UP 6.3 2.26

5 1140 1120 90' Bend --- 6.3 0.38

6 1140 1120 Ball Valve West 6.3 0.19

7 1150 1130 Straight West 6.3 4.2

8 1160 1130 90' Bend --- 6.3 0.38

9 1170 1130 Straight South 6.3 2.58

10 1180 1130 Straight South 6.3 4.32

11 1190 1130 90' Bend --- 6.3 0.38

12 1200 1140 Straight East 6.3 4.32

13 1210 1140 Straight East 6.3 3.17

14 1220 1140 90' Bend --- 6.3 0.38

15 1230 1140 Straight North 6.3 1.84

16 1230 1140 Ball Valve North 6.3 0.19

17 1240 1150 90' Bend --- 6.3 0.38

18 1240 1150 Straight Down 6.3 2.15

19 1240 1150 90' Bend --- 6.3 0.38

20 1240 1150 Ball Valve West 6.3 0.19

21 1240 1150 Straight West 6.3 1.09

22 1250 Transition West 6.3 - 30 1.17

23a Straight West 30 1.95

23b 90' Bend UP 30 0.58

Total 35.51
D-46

Figure D-24 Detailed LA3 MELCOR Nodalization
D-47

Figure D-25 Coarse LA3 MELCOR Nodalization
D-48

Figure D-26 Source distribution for MELCOR simulation of LA3 tests using average 

AMMD and GSD
D-49

Figure D-27 Source distribution for LA3A tests at four times of measurement

Figure D-28 Source distribution for LA3B tests at four times of measurement
D-50

Figure D-29 Source distribution for LA3C tests at four times of measurement
D-51

Figure D-30 Source rate reported for each LA3 test
D-52

Figure D-31 Calculated deposition velocity for LA3A test using various models
D-53

Figure D-32 Calculated deposition velocity for LA3B test using various models
D-54

Figure D-33 Calculated deposition velocity for LA3C test using various models
D-55

Figure D-34 Deposition efficiency calculated for LA3 bends using MELCOR Pui, INL and 

McFarland bend penetration model
D-56

6.1.4 Results

Overall retention factors measured for the tests are compared with the MELCOR calculated 

values in Table D-7. The retention factors reported are the fraction of the aerosol mass entering 

pipe section 1 that is deposited in the 6.3 cm pipe sections and components. Also shown is the 

fraction of the CsOH aerosol mass deposited in the 6.3 cm pipe that is found in the bend sections. 

It can be helpful to look at detailed deposition patterns as well as overall deposition factors. For 

example, the overall bend deposition calculated with the INL model is closest to the 

experimental value for LA3B. However, when each pipe section is examined closely, it is found 

that this model over predicts deposition in the first bend and under predicts in subsequent bends. 

Table D-7 Summary of Deposition Calculated for LA3 tests

Experimental Data Calculatedb

Test

Retention in 

6.3 cm pipea

(%)

Bend 

Depositionc

(%)

Retention in 6.3 cm pipea

(%)

Bend Depositionc

(%)

INLd VICe Woodf IVg** INLd VICe Woodf IVg**

CsOH

LA3A 63 46 67 78 57 79 7 10 43 14

LA3B 43 76 50 70 62 63 35 27 65 58

LA3C 47 35 61 76 68 55 95 38 94 92

MnO

LA3A 68 44 94 78 90 95 20 25 42 31 ** I-V indicates indicate the INL model for the straight sections (I) and the Victoria model for the bends 

(V). a Values reported are the deposited mass, including mass deposited in bends, divided by the total source 

mass less the mass deposited upstream of the 6.3 cm pipe (mass deposited on torches, Aerosol Mixing 

Vessel (AMV), and bypass).

b Values reported are for fine nodalization model. 

c Values reported is the fraction of the deposition in the 6.3 cm pipe that was found in the pipe bends.

d Uses Wood model for rough pipes in straight pipes and INL bend model. e Uses Sehmel model for perfect particle sinks in straight pipes and Pui bend model.

f Uses Wood’s model for smooth pipes and Pui bend model. 

g Uses Wood’s model for rough pipes and Pui bend model.
D-57

LA3A

Thermal Hydraulic Results

Pressures and velocities along the pipe train are plotted in Figure D-35 and Figure D-36. 

Figure D-35 Pressure profile along test pipe

Figure D-36 Velocity profile along test pipe

0

20000

40000

60000

80000

100000

120000

140000

160000

0 5 10 15 20 25 30 35

Pressure (Pa)

Position along 6.3 cm diameter pipe (m)
D-58

Deposition Profiles

Deposition profiles along the pipe train are plotted in Figure D-37 through Figure D-44.

Figure D-37 Deposition profile for CsOH along LA3A test section

(Calculated sticking factor)
D-59

Figure D-38 LA3A post test calculation of CsOH aerosol deposition for codes including 

bend deposition models (LACE TR-024)

Figure D-39 Deposition profile for CsOH aerosols along LA3A test section

(Sticking factor = 1)
D-60

Figure D-40 Deposition profile for CsOH + MnO aerosols along LA3A test section

 (Calculated sticking factor)

Figure D-41 LA3A post test calculation of CsOH + MnO aerosol deposition for codes in 

code comparison report including bend deposition models (LACE TR-024)
D-61

Figure D-42 Deposition profile along LA3A test section

 (20 aerosol size bins)

Figure D-43 Deposition profile for CsOH along LA3A test section

 (Colored regions indicate lumped nodalization)
D-62

Figure D-44 Deposition profile for CsOH along LA3A test section using the time 

dependent source rate and the four aerosol source distributions measured 

during the test
D-63

Observations

1. Overall, the MELCOR models do an excellent job of predicting deposition in the test 

train. The results for all models are generally consistent, though the Sehmel model tends 

to predict more deposition upstream than the other models with little or no deposition

near the end of the test train.

2. All MELCOR calculations slightly over predict deposition upstream in the test train, 

though not to the extent observed for LA1. As noted for LA1, this may be an indication 

of resuspension or flow of deposited material along the wall.

3. When compared with the code comparison calculation (TR-022), the MELCOR models 

estimate the total pipe retention better than most of the codes participating in that activity. 

MELCOR tends to overestimate the deposition, particularly in upstream test sections 

where many of the other codes underestimate total deposition. 

4. All MELCOR models over predict deposition in test Section 4 (< 2.5 m). This pipe 

section happens to be a vertical pipe section and it is possible that the orientation of the 

pipe may have an important role not accounted for in these models. 

5. Deposition in bends is important, since about 46% of the deposited mass is found in the 

bends. The Wood (smooth)/Pui model does a reasonable job of predicting the bend 

deposition (54%) while other models predict closer to 30%.

6. The MELCOR calculation of MnO deposition along the piping is similar in characteristic 

to that of CsOH deposition. About 94% of the MnO aerosol was deposited in the test 

section while the MELCOR models predict between 91% (Wood’s smooth/Pui) and 98% 

(Sehmel/Pui).

7. Calculating the sticking factor with Merrill’s model does not significantly affect results 

for these calculations.

8. Lumping the bends with straight pipe sections reduces the overall retention in the test 

section. 

9. Using the four measured aerosol source distributions and the time dependent source rate 

provided for this test shows a small (5-10%) but noticeable improvement in the 

calculated deposition profile. This demonstrates the sensitivity of the calculation to the 

assumed source profile.
D-64

LA3B

Thermal Hydraulic Results

Pressures and velocities along the pipe train are plotted in Figure D-45 and Figure D-46.

Figure D-45 Pressure profile along test pipe

Figure D-46 Velocity profile along test pipe
D-65

Deposition Profiles

Deposition profiles along the pipe train are plotted below in Figure D-47 through D-52.

Figure D-47 Deposition profile for CsOH along LA3B test section

Figure D-48 LA3B post test calculation of CsOH aerosol deposition for codes including 

bend deposition models
D-66

(LACE TR-024)

Figure D-49 Deposition profile for CsOH + MnO aerosols along LA3B test section

 (Sticking factor = 1)

Figure D-50 LA3B post test calculation of CsOH + MnO aerosol deposition for codes 

including bend deposition models (LACE TR-024)
D-67

Figure D-51 Deposition profile along LA3B test section

 (20 aerosol size bins)

Figure D-52 Deposition profile along LA3B test section - colored

 (Colored regions indicate lumped nodalization)
D-68

Observations

1. The deposition calculated by the VICTORIA model for the straight pipe early in the test 

train is larger than was measured in the experiment and exceeds the deposition predicted 

by the Wood model for rough pipes. Investigation showed that most of this deposition 

was from supermicron sized particles. The VICTORIA model uses a separate correlation 

by Sehmel for supermicron sized particles. In addition, for the fine nodalization, the 

VICTORIA model predicts more deposition in the straight section of HS 1120, which is 

the smaller diameter pipe.

2. For the coarse nodalization (see Figure D-25), deposition in the reducer region (i.e., pipe 

section 3) is over predicted by all models. It was reasoned that this could be a result of 

the nodalization for HS 1120. Deposition in the reducer for the fine nodalization cases is

well calculated.

3. Deposition in the bends is under predicted by the INL model for the coarse model but is 

over predicted for the fine nodalization.

4. For this test, the Wood model for smooth pipes gives the best results. The INL model,

which is the Wood model for rough pipes, gives better results for the straight pipe 

sections 2, 4, and 6. 
D-69

LA3C

Thermal Hydraulic Results

Pressures and velocities along the pipe train are plotted below (i.e., Figure D-53 and Figure D54).

Figure D-53 Pressure profile along test pipe

Figure D-54 Velocity profile along test pipe

0

20000

40000

60000

80000

100000

120000

140000

160000

0 5 10 15 20 25 30 35

Pressure (Pa)

Position along 6.3 cm diameter pipe (m)
D-70

Deposition Profiles

Deposition profiles for CsOH along the pipe train are plotted in Figure D-55 and Figure D-56.

Figure D-55 Deposition profile of CsOH along LA3C test section
D-71

Figure D-56 LA3C post test calculation of CsOH aerosol deposition for codes including 

bend deposition models (LACE TR-024)

Observations

1. Figure D-55 shows the VICTORIA model was the same as measured in the experiment 

for the first 3 meters (Sehmel/Pui), and exceeds the deposition predicted by the INL 

model. Investigation showed that most of this deposition was from supermicron sized 

particles. The VICTORIA model uses a separate correlation by Sehmel for supermicron 

sized particles. However, for the fine nodalization (see Figure D-24), the VICTORIA 

model does the best job of predicting deposition in the straight section of HS 1120 which 

is the smaller diameter pipe. This is contrary to the observations for the previous two 

tests.

2. For this test, deposition in the reducer (i.e., pipe section 3 in Figure D-25) is well 

predicted by the VICTORIA model for the coarse nodalization. However, for the finer 

nodalization (see Figure D-24), which would be expected to give better results, 

deposition in the reducer is roughly five times greater than is predicted.

3. Deposition in the bends is under predicted by the INL model for the coarse model but is 

over predicted for the fine nodalization.

4. For this test, the VICTORIA model gives the best results.
D-72

Turbulent Deposition Models

6.1.5 Wood’s Model for Turbulent Deposition on Surfaces

Wood developed a semi-empirical model for predicting turbulent deposition in pipes. His model 

characterizes deposition over three deposition regimes, which are characteristic of particle size: 

1. Turbulent particle diffusion for very small particles where Brownian motion is important 

to transport particles across the viscous sub layer.

2. Eddy Diffusion – Impaction regime for larger particles dominated by eddy diffusion 

where particles are accelerated to the wall due to turbulent eddies in the core and buffer 

layer and coast across the viscous sub layer.

3. Inertia-Moderated Regime – Very large particles, which are subject to reduced 

acceleration by the turbulent core and little or no acceleration to small eddies in the 

buffer near the wall.

In the turbulent particle diffusion regime, Brownian diffusion is important and deposition occurs 

by a combination of Brownian and eddy diffusion. Davies [26] proposed the following equation 

for the deposition velocity in this regime:

ut,s = Sc−2/3v�

14.5�

1

6

1n� (1+ φ)2

1− φ+ φ2�+

1

√3

atan�

2φ−1

√3 �+ π√3 � (8)

where:

 ut,s – Turbulent deposition velocity for submicron particles (m/s)

 v� – Friction velocity (m/s), defined by the following expression:

 v� = U�f

2 (9)

 f – Fanning friction factor (dimensionless)

 φ – SC1⁄3/2.9 

Wood found that for particles that are order of the mean free path or greater, this equation could 

be approximated by:

Vd

∗ = ut,s

v� = 3√3

29π

Sc−2/3 (10)

 

where the deposition velocity is non-dimensionalized by the friction velocity. 
D-73

In terms of the dimensionless relaxation time, τ*, this can be written:

 Vd

∗ = 3√3

29πτ∗

1⁄3 Sc−2/3 τ∗

1⁄3 (11)

As particle size increases, impaction increases and a second term is added to this equation:

 Vd

∗ = 3√3

29πτ∗

1⁄3 Sc−2/3 τ∗

1⁄3 + Kτ∗

2 (12)

The coefficient, K, is derived by solving a diffusion equation written in the form of a turbulent 

version of Fick’s law, i.e.,

N = �Dp + ε� dc

dy (13)

where:

N – particle flux (#/m2-s)

Dp – particle diffusion coefficient (m2/s)

ε – particle turbulent eddy-diffusivity (m2/s)

c – particle concentration (#/m3)

y – distance from surface (m)

For smooth pipes, Wood [22] proposed the following approximation for the deposition velocity:

 Vd

∗ = 3√3

29πτ∗

1⁄3 Sc−2/3 τ∗

1⁄3 + 0.00045τ∗

2 (14)

For large particles, τ* >10, particle inertia becomes important in the inertia-moderated regime 

and the deposition velocity becomes constant, though dependent on the Reynolds number 

through the friction factor:

 

Vd

∗ = �f

2

10 ≤ τ∗ < 270

Vd

∗ = 2.6

√τ∗

�1 − 50

τ∗

� τ∗ ≥ 270

(15)
D-74

For rough pipes, this equation is a little more complicated but was formulated for MELCOR by 

Merrill:

 

Vd

∗ = 1

(Is+IB) for τ∗ ≤ 10

Vd

∗ = 0.69 �2

π erfc � 1

√2 �1 + b+

s+

�� for τ∗ > 20

(16)

Where IS and IB result from integration of the non-dimensional diffusion equation over the buffer 

layer and sub-layer respectively, b+ is the non-dimensional roughness, and s+ is the 

non-dimensional perpendicular stopping distance. This model was originally implemented into 

MELCOR 1.8.0 by Merrill [14] for a branch version of the code for safety analysis of the 

International Thermonuclear Experimental Reactor (ITER).
D-75

6.1.6 VICTORIA Deposition Model

The VICTORIA model also predicts three regimes for turbulent deposition as was observed for 

the Wood models. Similar to the Wood model, deposition in the turbulent particle diffusion 

regime, follows that of Davies [26] (i.e., Equation 8). Though the approximation in Equation 9 

is not used for the VICTORIA model, this does not lead to significant differences in results. 

This term is then added to a term derived by Sehmel [17] for the particle impaction regime:

 ut,s = 1.47 ∗ 10−16 � ρa

1000�

1.01

(

2 ∗104ra

DH

)2.1Re3.02v� (17)

To obtain the following equation for the non-dimensional deposition velocity:

Vd

∗ =

Sc−2

3

14.5�

1

6

1n� (1+ φ)2

1− φ+ φ2�+

1

√3

atan�

2φ−1

√3 �+ π√3 �

+

1.47 ∗ 10−16 � ρa

1000�

1.01

(

2 ∗104ra

DH

)2.1Re3.02

(18)

Note that the correlation reported in Equation 17 was based on a least squares curve fit to a 

restricted data set based on experiments for which surfaces were treated, often with petroleum 

jelly, to simulate a perfect particle sink, by eliminating or drastically reducing particle bounce. 

Sehmel recommended use of another correlation, fit over a more general data set, for untreated 

surfaces:

 ut,s = 1.0 ∗ 10−16 � ρa

1000�

1.83

(

2 ∗104ra

DH

)2.99Re3.08v� (19)

This equation was not used by the VICTORIA code and not implemented into MELCOR, but is 

reported here for completeness.

It should also be pointed out that the VICTORIA user manual indicates Sehmel’s equation is 

used for supermicron particles and Davies model is used for submicron particles. However, this

is an issue with VICTORIA users manual. Doing so leads to a discontinuity in the deposition 

velocity. Examination of the source code indicates that the sum of these two terms is actually 

used.

A maximum is placed on the non-dimensional deposition velocity so that it does not exceed a 

value of 0.1. This leads to the constant deposition velocity characteristic of the inertiamoderated regime. This is also undocumented in the VICTORIA manual.
D-76

Deposition in Pipe Bends

6.1.7 INL Bend Model

For the INL Bend Model [14], to calculate the inertial deposition of aerosols in pipe bends, we 

start with the centrifugal force acting on the particle as the fluid turns a pipe bend. This force is 

given by:

FC = π

6 �ρp − ρf�dp

3 uf

2

rb

≈ mp

uf

2

rb

(20)

Where: 

dp – particle diameter (m)

uf – fluid velocity (m/s)

rb – bend radius of pipe (m)

ρp – particle density (kg/m3)

ρf – fluid density (kg/m3)

mp – particle mass (kg)

Θb – bend turning angle (radians)

S – the particle radial drift (m)

B – the particle mobility

The terminal velocity in the radial direction that a particle will obtain because of this force is 

given by:

up = BFc (21)

where "B" is the particle mobility defined as:

B = 1

(3πμgdp) (22)

where μg, is the carrier gas viscosity. 

The time that it will take for a particle to travel around a bend is given by:

tb = rbΘb

uf

(23)

where Θb is the pipe turning angle in radians. 
D-77

Consequently, the radial distance a particle will drift in this turn is the product of bend travel 

time and the particle radial velocity, which becomes:

S = ΘbBmpuf (24)

Equation 24 is used to express the centrifugal force. By assuming a well mixed particle 

concentration in the pipe (co), the fraction of particles that will collide with the wall in the bend 

is approximately the radial drift distance divided by the pipe diameter (i.e. s/D). The particle 

flux (#/m2-s) for inertial deposition based on this collided fraction, when averaged over the pipe 

surface area, can be expressed as:

Γi = s

D

coufAc

As

(25)

where:

D − pipe diameter (m)

AC − pipe cross-sectional area (m2)

As − pipe surface area (m2)

The deposition velocity associated with this particle flux is as follows:

Vi = 

ri

co

(26)
D-78

6.1.8 Pui Bend Model

The model used in VICTORIA for deposition in 90° pipe bends under turbulent conditions (i.e., 

Re ≥ 2300) is based on the experimental and theoretical work of Pui et al., [25]. Their 

experiments covered a range of Reynolds numbers from 102 to 104

. They found that an 

exponential relationship between Stokes number and deposition efficiency correlated well with 

their data. This relationship is:

ηb = 1 − 10−0.963St (27)

where:

ηb – deposition efficiency due to flow irregularity (dimensionless)

And the particle Stokes number is given by:

St = Ccρpdp

2Uave

9μDh

(28)

Deposition efficiency is defined as the fraction of aerosol particles of a specific size that deposit. 

More specifically for Equation 27, the deposition efficiency represents the fraction of aerosol 

particles that deposit near the pipe bend because of inertial effects induced by curvature of the 

fluid streamlines. Deposition efficiency is converted to deposition velocity in VICTORIA by the 

following definition:

ub = ηb

U

L

VB

A (29)

where: 

ub − deposition velocity for flow through a bend 

VB − volume of bulk gas subregion (m3), as defined in chapter 3

A − surface area for aerosol deposition (m2)
D-79

6.1.9 McFarland Bend Model

McFarland’s model is purely empirical and is based on fitting an equation to data obtained from 

physical experiments and Lagrangian simulations [13]: 





 





+ + +

+ = − b St c St d Sta St

b 2 2 1

4.61 1 0.01exp θ θ θθ

η (30)

Where: Ɵ = the angle of the bend;

a = −0.9526 − 0.0568δ (31)

2 1 0.07 0.0171

0.297 0.0174

δ δ

δ

− +

− − b = (32)

δ δ

1.895 2.0

c = −0.306 + − (33)

2

2

1 0.129 0.0136

0.131 0.0132 0.000383

δ δ

δ δ

− +

− + d = (34)

h

Rbend 2 δ = (35)

and where:

Rbend – radius of the bend in the flow path
D-80

6.1.10 Sticking Factor

Particles that strike a surface may either stick to the surface or bounce and are possibly 

re-entrained if adhesive forces are not sufficient to overcome the incident momentum of the 

particle. An optional factor can be used to calculate the probability of sticking to the surface. 

This factor is calculated [14] by considering both capillary forces and Van der Waals forces:

Fa = 4πrpσ + 3

2

πrpγfr (36)

Where: 

σ – surface tension of possible film on surface (J/m2)

γ – surface energy per unit area for Van der Waals interaction (J/m2

)

fr – surface energy reduction factor due to surface roughness

The minimum required momentum necessary for a particle to overcome adhesion is then given 

by the following:

mpuc = ∫ FSdt = ∫ Fadt + ∫ Fddt ≈

t

0

t

0

t

0 Fata + Fdtd (37)

Where:

uc – critical escape velocity (m/s)

mp – particle mass (kg)

ta – surface adhesion time (s)

td – film residence time (s)

Fd – drag force = 3

2

πrpμfu

Several assumptions are made regarding the surface adhesion time and the film residence time. 

For example, it is assumed that the distance over which these forces act are on the order of the 

magnitude of the particle radius and that this distance divided by the critical escape velocity 

gives the film residence time. It is also assumed that the viscous drag force acts over 

approximately twice the film thickness. With these assumptions, the critical velocity can be 

calculated:

uc = 3πrpμfδf

2mp

+ ��

3πrpμfδf

2mp

�

2

+

πrp

mp

�rσ + 3

2

γfr� (38)

Finally, it is assumed that this critical velocity is the vector sum of the perpendicular velocity 

component and the parallel stream velocity component and equations for standard turbulent 

velocity profiles are then used to determine the fractional area of the pipe for which particle 

velocities would be insufficient to overcome the adhesion forces. The perpendicular velocity 

component is approximated by u = 0.69�v where v� is the friction velocity defined in 

Equation 5.
D-81

New MELCOR RN Package Input Records

RNTURB Record – Deposition Modeling Record

Optional

(1) IMODEL - Deposition Modeling flag for gravitational, thermophoresis, and 

diffusiophoresis components

= 0, Gravitational, thermophoresis, and diffusiophoresis velocities 

are calculated once at the beginning of the calculation.

= 2, Gravitational, thermophoresis, and diffusiophoresis velocities are 

recalculated at each time step.

 (type = integer, default=0, units = none) 

(2) ITURB - Deposition Modeling flag for turbulent component

= 0, MELCOR 1.8.6 deposition modeling

= 1, VICTORIA modeling of deposition in straight pipe sections. If

negative, the sticking factor is assumed to be 1.0

= 2, INL modeling of deposition in straight pipe sections. This is 

essentially Wood’s model for rough pipes and approaches Wood’s 

model for smooth pipes when the roughness is small. If negative, the 

sticking factor is assumed to be 1.0

= 3, INL model for submicron particles, VICTORIA for larger 

particles. If negative, the sticking factor is assumed to be 1.0

= 4, VICTORIA model for submicron particles, INL for larger 

particles If negative, the sticking factor is assumed to be 1.0

= 5, Wood’s model for smooth pipes.

(type = integer, default=0, units = none)

(3) ITRANS - Deposition Modeling flag for impact deposition in transitions

= 0, MELCOR 1.8.6 deposition modeling

= 1, VICTORIA modeling of deposition in bends and other 

transitions. If negative, the sticking factor is assumed to be 1.0

= 2, INL modeling of deposition in bends and other transitions. If 

negative, the sticking factor is assumed to be 1.0

(type = integer, default=0, units = none)
D-82

RNMGnnn Record – RN Turbulent Deposition Record

nnn is a sequence number

Optional

(1) VOL_ID - The volume in which to apply the bend and/or turbulent deposition 

model

(type = integer, default=none, units = none) 

(2) CHARL - characteristic dimension (i.e., pipe diameter)

(type = real, default=none, units = m)

(3) No_Bnd - Number of bends associated with the volume

(type = integer, default=none, units = none)

(7) ANGLE - Turning angle of the bends

(type = real, default=none, units = radians)

(5) RAD_BND - Radius of curvature for bend

(type = real, default=none, units = m)

(6) ROUGH - Surface roughness for the turbulent deposition model (not used in 

VICTORIA model)

(type = real, default=none, units = none)
D-83

RNTRSnnn Record – Transition Deposition Record

nnn is a sequence number

Optional

(1) VOL_ID - The volume in which to apply the bend and/or turbulent deposition 

model

(type = integer, default=none, units = none) 

(2) CHARL - characteristic dimension (i.e., pipe diameter)

(type = real, default=none, units = m)

(3) DODI - Diameter at exit divided by diameter at entrance (<=1)

(type = real, default=none, units = none)

(4) NCONTR - Number of contraction transition

(type = integer, default=none, units = none)

(5) NVENTUR - Number of venturi transitions

 (type = integer, default=none, units = none) 

(6) FACONT - Multiplier on deposition velocity for contraction

(type = integer, default=none, units = none)

(7) FAVENT - Mulitiplier on deposition velocity for venturi transition

(type = integer, default=none, units = none)

RNSTnnn Record – Sticking Factor Options

nnn is a sequence number

Optional

(1) iCOMP - RN Component

(type = integer, default=none, units = none)

(2) SIGMAW - Surface Tension

(type = real, default=none, units = N/m)

(3) ETAF - Surface Viscosity (mu/rho)

(type = real, default=none, units =kg/m-sec)
D-84

6.1.11 New Control Function

RN1-DEPHS-x-s-II.HS /c/ Total radionuclide mass of class x deposited on side s of 

heat structure HS from deposition model II. The deposition 

models that are tracked are as follows:

II = 1, Diffusion deposition

II = 2, Thermophoresis

II = 3, Gravitational settling

II = 4, Turbulent deposition in straight sections

II = 5, Deposition in pipe bends 

II = 6, Deposition in Venturi transitions

II = 7, Deposition in Contraction transitions

(units = kg)
D-85

Summary and Conclusions for LACE Experiments

The LACE tests provide an experimental database for validation of turbulent deposition 

modeling over a wide range of conditions. Entrance flow velocities in these tests ranged from 

23 m/s (LA3C) to 96 m/s (LA1) while the exit flow velocity for LA1 was as high as 200 m/s. 

The Wood’s model for both rough and smooth pipes as well as the VICTORIA model were 

tested. In addition, the Pui model for bends as well as Merrill’s model were examined. 

Deposition profiles along the test section were used to compare the relative importance of these 

deposition mechanisms. The results were also compared against the code comparison study that 

was part of the LACE experimental program (LACE TR-022 and LACE TR-024).

Furthermore, several sensitivity analyses were performed. A calculation utilizing a coarse 

nodalization, where straight pipe sections were lumped with bends by sharing the same control 

volume and heat structure, were tested against experimental results and compared with results 

from the detailed nodalization. In addition, the dependency of results on the number of section 

bins used in the calculation was examined. Finally, the dependency of the results on the assumed 

sticking factor was considered.

In general, the overall deposition in the experiment is captured by the all models. LA1, which 

has the greatest flow velocities, shows approximately 97 % retained in the pipes compared to 

98% in the experiment. For the LA3 tests, the trends in total retention are captured by the 

MELCOR models, though they all slightly over predict deposition. Also, the fraction that is 

deposited in bends does not appear to be consistent between tests. The VICTORIA models 

appear to predict larger depositions than the other modeling approaches. This is consistent with 

the velocity profiles plotted for these models and shown in Figure D-31through Figure D-33.

Some of this difference in the predicted deposition profiles may be an indication of resuspension 

or flow of deposited material along the wall. This could explain why this effect is more

pronounced for LA1 where velocities are greater where it would be expected that resuspension or 

flow of deposited material would be greater for the higher Reynolds numbers. 
D-86

7 REFERENCES

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on the Dust Holding Capacity of Deep Pleat HEPA Filters.” J Dyment AWE plc 

Aldermaston RG7 4PR United Kingdom D Loughborough AEAT Harwell Oxford OX11 

ORA United Kingdom Proceedings of the 24th DOE/NRC Nuclear Air Cleaning and 

Treatment Conference: Held in Portland, Oregon, July 15–18, 1996 (NUREG/CP-0153, 

CONF-960715), Figure 1, 21.0 sq.m 72pl HV curve from 600g-840g of captured dust

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Deposition in Bends with turbulent Flow”, Environmental Science and Technology, 31

(1997) 3371-3377. 
D-87

14. B.J. Merrill and D.L. Hagrman, “MELCOR Aerosol Transport Module Modification for 

NSSR-1”, INEL-96/0081, ITER/US96i/TE/SA-03, 1996.

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16. P.G. Papavergos and A.B. Hedley, “Particle deposition behaviour from turbulent flows”, 

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22. N.B. Wood, “A Simple Method for the Calculation of Turbulent Deposition to Smooth and 

Rough Surfaces”, J. Aerosol Sci. Vol 12. No 3 (1981), 275-290.

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London, A-289, pages 235-46, 1966.




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