MELCOR Best Practices as 

Applied in the 

State-of-the-Art Reactor 

Consequence Analyses 

(SOARCA) Project

Office of Nuclear Regulatory Research

NUREG/CR-7008
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MELCOR Best Practices as 

Applied in the 

State-of-the-Art Reactor 

Consequence Analyses 

(SOARCA) Project

Manuscript Completed: June 2014

Date Published: August 2014

Prepared by:

Kyle Ross*, Jesse Phillips*, Randall O. Gauntt*, 

Kenneth C. Wagner**

*Sandia National Laboratories

Albuquerque, New Mexico 87185

Operated for the U.S. Department of Energy

**dycoda, LLC

Prepared for: 

U.S. Nuclear Regulatory Commission 

Office of Nuclear Regulatory Research 

Washington, DC 20555-0001

NRC Project Manager:

Jonathan Barr

NRC Job Code N6306

Office of Nuclear Regulatory Research

NUREG/CR-7008
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 modeling approach used in the State-of-the-Art Reactor Consequence Analyses (SOARCA) 

project to characterize the release of radionuclides to the environment accompanying a 

postulated severe (core damage) accident is based on plant-specific applications of the 

MELCOR computer code. MELCOR is a state-of-the-art computational model developed by 

Sandia National Laboratories for the U.S. Nuclear Regulatory Commission. Due to large 

uncertainties in many aspects of severe accident behavior, MELCOR provides the code user a 

wide spectrum of options for modeling uncertain physical phenomena and characterizing plant 

response to beyond design basis accident conditions. The selection and exercise of the 

available modeling capabilities are an important aspect of the overall modeling approach. 

This document describes the specific manner in which MELCOR modeling capabilities were 

used to represent important, and in some cases uncertain, aspects of severe accident behavior

in the SOARCA project. This description includes choices made among alternate modeling 

options offered through code input, changes to selected input parameters from those offered as 

‘default’ values, and in some cases, user-generated ‘models’ to represent features of plant 

response to a severe accident that are not directly available in MELCOR. Collectively, these 

features represent the “SOARCA best practice” guidance for using MELCOR to calculate severe 

accident behavior in operating nuclear power plants.

PAPERWORK REDUCTION ACT STATEMENT

The NUREG does not contain information collection requirements and, therefore, is not subject 

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

PUBLIC PROTECTION NOTIFICATION

The NRC may not conduct or sponsor, and a person is not required to respond to, a request for 

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

currently valid OMB control number.

v

TABLE OF CONTENTS

Section Page

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

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

LIST OF FIGURES ....................................................................................................................vi

LIST OF TABLES..................................................................................................................... viii

EXECUTIVE SUMMARY ...........................................................................................................ix

ACKNOWLEDGEMENTS ..........................................................................................................xi

ABBREVIATIONS AND ACRONYMS ...................................................................................... xiii

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

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

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

1.3 Independent Peer Review................................................................................ 1-2

2. TECHNICAL APPROACH ........................................................................................... 2-1

2.1 Analytical Models............................................................................................. 2-1

2.2 Important Differences in Approach from Prior Work ......................................... 2-2

3. MELCOR BEST PRACTICES FOR SEVERE ACCIDENT ANALYSIS......................... 3-1

3.1 Best Modeling Practices................................................................................... 3-2

3.1.1 Generic Light Water Reactor Best Practices......................................... 3-2

3.1.2 Pressurized Water Reactor Best Practices......................................... 3-42

3.1.3 Boiling Water Reactor Best Practices ................................................. 3-54

3.2 MELCOR Code Enhancements for the SOARCA Project............................... 3-56

3.2.1 Updated MELCOR Defaults to Reflect Current Best-estimate 

Modeling Practices ............................................................................. 3-56

3.2.2 Add a Simplified Thermo-mechanical Fuel Collapse Model ................ 3-56

3.2.3 Fission Product Vapor Scrubbing with Aerosol Scrubbing .................. 3-57

4. MELCOR CODE DEVELOPMENT AND VALIDATION BASIS .................................... 4-1

4.1 Selection of Validation Test Cases................................................................... 4-1

4.2 Discussion of MELCOR Validation Tests ......................................................... 4-3

4.3 Comparisons of Code Versions........................................................................ 4-6

4.3.1 Airborne Physics .................................................................................. 4-6

4.3.2 Oxidation.............................................................................................. 4-8

4.3.3 Hydrogen Stratification in Containment ................................................ 4-9

4.3.4 Combustion Modeling......................................................................... 4-11

4.3.5 Containment Pressure Response to Sprays ....................................... 4-12

4.3.6 Fission Product Release..................................................................... 4-13

4.3.7 Molten Core-Concrete Interaction....................................................... 4-14

5. INSIGHTS ................................................................................................................... 5-1

6. REFERENCES............................................................................................................ 6-1

Appendix A Other MELCOR Modeling Best Practices .............................................A-1

Appendix B Updated Default Parameters ................................................................B-1

Appendix C MELCOR Validation Test Suite ............................................................C-1

Appendix D MELCOR Code Version Progression Overview....................................D-1
vi

LIST OF FIGURES

Figure Page

Figure 2-1 Computational methods used to derive source terms in NUREG-0772................2-3

Figure 2-2 Typical spatial nodalization of RCS in the NUREG-0772 methodology ................2-4

Figure 3-1 MELCOR depiction of fuel rod degradation..........................................................3-4

Figure 3-2 Release fractions for different release models – release temperature =

 2000 K.............................................................................................3-14

Figure 3-3 Release fractions at constant temperature for ORNL-Booth versus 

 CORSOR-M3....................................................................................3-15

Figure 3-4 Fractional release rate (%/min) – the time derivative of release fraction ............3-15

Figure 3-5 Vapor pressure of selected species ...................................................................3-16

Figure 3-6 Schematic of the Phebus Test Facility showing test fuel bundle,

heated lines, steam generator tube and simulated containment.........................3-17

Figure 3-7 FPT-1 nuclear and chemical heating history ......................................................3-18

Figure 3-8 FPT-1 maximum bundle temperature history .....................................................3-18

Figure 3-9 Emission gamma tomography of the end-state condition of test FPT-1..............3-19

Figure 3-10 Comparison of ORNL-Booth versus CORSOR-M for Xe release (Class 1)........3-19

Figure 3-11 Comparison of ORNL-Booth versus CORSOR-M for Cs release (Class 2)........3-20

Figure 3-12 Comparison of ORNL-Booth versus CORSOR-M for Ba release (Class 3)........3-20

Figure 3-13 Comparison of ORNL-Booth versus CORSOR-M for I release (Class 4) ...........3-21

Figure 3-14 Comparison of ORNL-Booth versus CORSOR-M for Te release (Class 5) ........3-21

Figure 3-15 Comparison of ORNL-Booth versus CORSOR-M for Ru release (Class 6)........3-22

Figure 3-16 Comparison of ORNL-Booth versus CORSOR-M for Mo release (Class 7) .......3-22

Figure 3-17 Comparison of ORNL-Booth versus CORSOR-M for Ce release (Class 8)........3-23

Figure 3-18 Comparison of ORNL-Booth versus CORSOR-M for La release (Class 9) ........3-23

Figure 3-19 Comparison of ORNL-Booth versus CORSOR-M for UO2 release (Class 10)....3-24

Figure 3-20 Comparison of ORNL-Booth versus CORSOR-M for Cd release (Class 11)......3-24

Figure 3-21 Comparison of ORNL-Booth versus CORSOR-M for Sn release (Class 12) ......3-25

Figure 3-22 Schematic of VERCORS Test Facility for measuring fission

product release from small fuel samples............................................................3-26

Figure 3-23 Comparison of Cs release for modified ORNL-Booth with

CORSOR-M for VI-2 run under steam oxidizing conditions..............................3-27

Figure 3-24 Comparison of Cs release for modified ORNL-Booth with

CORSOR-M for VI-3 performed under steam oxidizing conditions.....................3-28

Figure 3-25 Comparison of Cs release for modified ORNL-Booth with

CORSOR-M for VI-5 performed under steam reducing conditions .....................3-28

Figure 3-26 Comparison of Cs release for modified ORNL-Booth with

CORSOR-M for VERCORS-2............................................................................3-30

Figure 3-27 Comparison of Cs release for modified ORNL-Booth with

CORSOR-M for VERCORS-4............................................................................3-31

Figure 3-28 Comparison of Xe release for modified ORNL-Booth with

CORSOR-M for VERCORS-4............................................................................3-31

Figure 3-29 Comparison of Iodine release for modified ORNL-Booth with

CORSOR-M for VERCORS-4............................................................................3-32

Figure 3-30 Comparison of Te release for modified ORNL-Booth with

CORSOR-M for VERCORS-4............................................................................3-32

Figure 3-31 Comparison of Ba release for modified ORNL-Booth with

CORSOR-M for VERCORS-4............................................................................3-33
vii

Figure 3-32 Comparison of Mo release for modified ORNL-Booth with

CORSOR-M for VERCORS-4............................................................................3-33

Figure 3-33 MELCOR-predicted fission product deposition in FPT-1 circuit

using default CORSOR-M release modeling......................................................3-35

Figure 3-34 MELCOR-predicted fission product deposition in FPT-1 circuit

using default ORNL-Booth release modeling .....................................................3-35

Figure 3-35 Normalized aerosol depletion rate of airborne aerosol in FPT-1

containment.......................................................................................................3-36

Figure 3-36 Predicted and measured aerodynamic mass mean aerosol

diameter in FPT-1 containment..........................................................................3-37

Figure 3-37 MELCOR lower head nodalization .....................................................................3-40

Figure 3-38 Heat transfer from an overlying water pool to an ex-vessel debris bed. .............3-42

Figure 3-39 In-vessel, full-loop, and hot leg natural circulation flow patterns

in a PWR severe accident..................................................................................3-46

Figure 3-40 MELCOR hot leg and steam generator nodalization including the

special natural circulation flow paths..................................................................3-47

Figure 3-41 Westinghouse PWR reactor vessel internals .....................................................3-49

Figure 3-42 MELCOR Westinghouse lower vessel nodalization............................................3-50

Figure 3-43 Discrete regions of the drywell floor to represent debris spreading ....................3-55

Figure 4-1 CSTF Airborne Mass Test AB5............................................................................4-7

Figure 4-2 Depletion of SnO2 in DEMONA-B3 experiment....................................................4-7

Figure 4-3 PHEBUS-B9+ hydrogen generation.....................................................................4-8

Figure 4-4 FPT-1 hydrogen generation .................................................................................4-9

Figure 4-5 Helium stratification calculated for NUPEC M-8-1 for MELCOR 2.x ...................4-10

Figure 4-6 Helium stratification calculated for NUPEC M-8-1 for three MELCOR 

 code versions......................................................................................4-9

Figure 4-7 MELCOR 1.8.6 & 2.1 assessments of CSE A9 ..................................................4-13

Figure 4-8 MELCOR 1.8.3 assessments of CSE A9 ...........................................................4-13

Figure 4-9 MELCOR 1.8.6 & 2.x assessments of ablation depth in SURC-1 Test...............4-14
viii

LIST OF TABLES

Table Page

Table 3-1 Best-estimate time to fuel rod collapse versus cladding oxide temperature .........3-3

Table 3-2 Summary of data from molten debris-coolant interactions experiments. ..............3-6

Table 3-3 CORSOR-Booth, ORNL-Booth, and Modified ORNL-Booth parameters............3-14

Table 3-4 Test conditions for selected ORNL VI tests and VERCORS tests......................3-26

Table 3-5 Release fraction from ORNL VI-2. .....................................................................3-29

Table 3-6 Release fraction from ORNL VI-3. .....................................................................3-29

Table 3-7 Release fraction from ORNL VI-5. .....................................................................3-29

Table 4-1 MELCOR validation tests.....................................................................................4-3

Table 4-2 Historical review of MELCOR assessment studies...............................................4-5

Table 4-3 Hydrogen burn completeness from experiment and MELCOR...........................4-12

Table 4-4 Hydrogen burn times from experiment and MELCOR. .......................................4-12

Table 4-5 Pressure ratio calculated with recent MELCOR code versions

compared to test results.....................................................................................4-12
ix

EXECUTIVE SUMMARY

The purpose of this report is to describe best practices for the MELCOR severe accident 

analysis code, as implemented in the State-of-the-Art Reactor Consequence Analyses 

(SOARCA) project for the Peach Bottom Atomic Power Station and Surry Power Station. By 

applying modern analysis tools and techniques in the SOARCA project, a body of knowledge 

regarding the realistic outcomes of severe reactor accidents has been developed through the 

integrated modeling of accident progression and offsite consequences. The SOARCA project 

provides analyses that use state-of-the-art source term and consequence modeling together 

with consideration of current operational practices and procedures. Collectively, this information 

represents the “best practice” modeling approach for MELCOR accident progression and source 

term analyses.

The objective of this report is to describe the best practices, which are comprised of the model 

improvements, modeling approach, and parameter selections, that support the best-estimate 

analyses in NUREG-1935 [1], NUREG/CR-7110 Volume 1 [62], and NUREG/CR-7110 Volume 

2 [63]. This best practices document provides a compilation of the parameters and inputs used 

in the SOARCA documents. It also includes additional detail on the approach to developing 

some of the input values and methodologies applied in the SOARCA studies. This document 

explains the significance of key modeling improvements and presents a review of the validation 

and phenomena modeling to support the application of MELCOR. This report is intended to 

provide guidance and insights for developers and users of MELCOR.

An accumulation of various efforts, including an expert review panel conducted in August 2006 

and peer review committee meetings during 2009-2011 as well as 25 years of research into 

severe accident phenomenology and security related mitigation improvements, has been 

utilized in codifying both the depiction of the facilities investigated and the current knowledge of 

severe accident phenomena. The entire body of code input reflects the informed judgment on 

how a MELCOR model should be configured to generate a realistic estimate of plant response 

to a severe accident. This information as well as its implementation represents the best practice 

modeling approach used in the SOARCA project for the severe accident progression and 

radionuclide source term calculations.

The best practice modeling features, described herein, fall into two broad categories, input 

controls impacted by the user and modifications and enhancements made to MELCOR in 

support of the SOARCA project.

1. General user input categorizes all input options exercised by the user to perform the 

following functions:

a. Specification of parameters which are not equivalent to the default value, 

typically performed for sensitivity coefficients;

b. Model selection from the available models within MELCOR; and

c. Development of model logic to characterize systems and procedures whereby an 

internal MELCOR model does not exist or is replaced.

2. MELCOR modifications categorize advanced modeling features incorporated into the 

MELCOR source code to accommodate necessary enhancements for the SOARCA 

project.
x

In addition to the best practices detailed within this document, the continued improvement in 

computing resources permits improved fidelity in the physical representation of nuclear power 

reactors and their associated systems. The spatial nodalization of the reactor pressure vessel, 

primary coolant system, and containment in the MELCOR models developed for pressurized 

water reactors (PWRs) and boiling water reactors (BWRs) recognizes much more geometric 

detail. For example, the entire PWR reactor coolant system (RCS) was represented by four

spatial regions in the NUREG-0772, “Technical Bases for Estimating Fission Product Behavior 

During LWR Accidents,” calculations. The MELCOR model of the same 3-loop Westinghouse 

PWR used in the SOARCA analysis uses 25 control volumes for the core region alone; over 100 

control volumes are used to represent the entire RCS. This dramatic increase in detail provides 

much greater resolution of the driving forces governing fission product transport and deposition

and establishes a best practice with regard to model detail.

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

publications including system improvements, improvements in training, emergency procedures, 

offsite emergency response, security-related improvements, and plant changes such as power 

uprates and higher core burnup. Given the advancement in severe accident analyses afforded 

by conducting the SOARCA project, an opportunity to investigate accident phenomena at a 

much greater level of detail was available than analyses performed in the past, which has 

promoted recognition of various insights which benefit the probabilistic risk assessment (PRA)

community. As these insights can be contributed to the implementation of best practices in the 

SOARCA project, they are presented as well; a select few are presented below. A more 

detailed listing is provided in the body of this document.

1. In the Surry station blackout (SBO) scenarios, the most likely first RCS failure occurs at 

the hot leg nozzle prior to significant in-vessel fuel damage. This leads to vessel 

depressurization, accumulator discharge, fuel cooling, and an interruption to the core 

heat-up. A new release pathway for radionuclides is established at the failed hot leg. 

The response of a thermally-induced steam generator tube rupture (TISGTR) is also 

impacted by hot leg failure. The hot leg failure substantially decreases TISGTR flow due 

to the RCS depressurization and the introduction of the larger failure location as the 

primary fission product pathway from the vessel. 

2. With regard to the interfacing systems loss of coolant accident (ISLOCA) modeling and 

the magnitude of predicted radionuclide releases to the environment, a key insight is the 

large amount of deposition of aerosolized radionuclides in the low head safety injection 

(LHSI) piping by means of turbulent deposition. Sustained high velocities in the LHSI 

piping during core degradation drive the importance of this phenomenon. A thorough 

representation of the LHSI piping is necessary to address turbulent deposition of fission 

product aerosols in the piping and revaporization of deposits.

3. The improvements to fuel degradation modeling and 2-dimensional core modeling show 

a delayed heat-up followed by accelerated oxidation. The accelerated oxidation phase 

ends following molten Zircaloy breakout. Without molten Zircaloy breakout, the 

subsequent heat-up is primarily controlled by decay heat. The best practice modeling of 

Zircaloy-oxide collapse creates a debris bed similar to Three Mile Island Unit 2 (TMI-2). 

The debris bed slows oxidation by creating blockages and inhibiting natural circulation. 

The debris bed gradually grows axially and radially, which eventually leads to core plate 

failure.
xi

ACKNOWLEDGMENTS

Contributions to this best practices document were received from NRC and SNL project 

managers and technical experts dedicated to the production of a valuable reference resource 

for the user community. Information received from the Expert Review Panel and Peer Review 

Committee provided insights and information that have influenced the best practices 

documentation. The NRC Project Manager, Jonathan Barr, provided the leadership to ensure 

this project met the objectives of the program. Numerous NRC staff provided technical insights 

supporting key elements of the document. SNL technical staff worked with these experts to 

develop the criteria and document the approach that was used in the SOARCA project and 

described in this report.

xiii

ABBREVIATIONS AND ACRONYMS

ABCOVE

ACRR

Aerosol Behavior Code Validation and Evaluation

Annular Core Research Reactor

AHMED

ANL

Aerosol and Heat Transfer Measurement Device

Argonne National Laboratory

BWR Boiling Water Reactor

CCFL Counter-Current Flow Limitation

CFD Computational Fluid Dynamics

COR

CSARP

CSE

CVTR

CWTI

Core (or Core Package)

Cooperative Severe Accident Research Program

Containment Systems Experiment

Carolinas-Virginia Tube Reactor

Corium-Water Thermal Interactions

DCH

DEMONA

Direct Containment Heating

Demonstration of Nuclear Aerosol Behavior

DOE U.S. Department of Energy

EMUG

EPRI

FLECHT

FP

GE

GRS

HDR

HHSI

HI

HPME

IBRAE

IET

European MELCOR User Group

Electric Power Research Institute

Full-Length Emergency Cooling Heat Transfer

Fission Product

General Electric

Gesellschaft fur Anlagen- und Reaktorsicherheit mbH

Heissdampfreaktor

High Head Safety Injection

Horizontal Induction

High Pressure Melt Ejection

Nuclear Safety Institute of the Russian Federation

Integral Effects Test

IPEEE

ISLOCA

ISP

JAERI

Individual Plant Examination for External Events

Interfacing Systems Loss of Coolant Accident 

International Standard Problem

Japan Atomic Energy Research Institute

LACE

LHF

LHSI

LPRM

LOCA

LOFT

LTSBO

LWR

MAAP

Light Water Reactor Aerosol Containment Experiment

Lower Head Failure

Low Head Safety Injection

Local Power Range Monitor

Loss of Coolant Accident

Loss-of-Fluid Test

Long-Term Station Blackout

Light Water Reactor

Modular Accident Analysis Program

MACCS

MACE

MAEROS

MCAP

MCCI

MEI

MSLB

MELCOR Accident Consequence Code System

Melt Attack and Coolability Experiments

Multi-Component Aerosol Module for CONTAIN

MELCOR Code Assessment Program

Molten Core Concrete Interaction

Moscow Power Engineering Institute

Main Steam-Line Break

NRC U.S. Nuclear Regulatory Commission

NTS Nevada Test Site
xiv

NUPEC Nuclear Power Engineering Corporation

OECD

ORNL

PBF

PNL

Organization for Economic Co-operation and Development

Oak Ridge National Laboratory

Power Burst Facility

Pacific Northwest National Laboratory

PORV

PRA

PRC

PWR

Power-Operated Relief Valve

Probabilistic Risk Assessment

Peer Review Committee

Pressurized Water Reactor

RAMCAP Risk Analysis and Management for Critical Asset Protection

RAS

RCS

RHR

RN

Russian Academy of Science

Reactor Coolant System

Residual Heat Removal system

Radionuclide Package

RPV

RTF

RWST

SBO

SEASET

Reactor Pressure Vessel

Radioiodine Test Facility

Refueling Water Storage Tank

Station Blackout

Systems Effects And Separate Effects Test

SFD

SFP

SG

SNAP

Severe Fuel Damage

Sandia Fuel Project

Steam Generator

Symbolic Nuclear Analysis Package

SNL Sandia National Laboratories

SOARCA State-of-the-Art Reactor Consequence Analyses

SPARC

SQA

SST

STCP

SURC

SV

TISGTR

Suppression Pool Aerosol Removal Code

Software Qualification Assurance 

Siting Source Term

Source Term Code Package

Sustained Urania-Concrete

Safety Valve

Thermally-Induced Steam Generator Tube Rupture

TIP

TMI-2

VANAM

VI

Traversing In-core Probe

Three Mile Island Unit 2

Experiments on the Aerosol Behavior within a Multi-Compartment 

Containment (German translation)

Vertical Induction
1-1

1. INTRODUCTION

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) and others over the last several decades. As a result 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 reexamine

significantly less realistic aspects of previous reactor accident analysis efforts was a major 

motivating factor in the genesis of the State-of-the-Art Reactor Consequence Analyses 

(SOARCA) project. By applying modern analysis tools and techniques, the SOARCA project 

seeks to provide a body of knowledge that will support an informed public understanding of the 

likely outcomes of selected severe nuclear reactor accidents.

The overall objective of the SOARCA project was to develop a body of knowledge regarding the 

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

project has utilized state-of-the-art computational analysis tools, which incorporate knowledge 

gained from the past 25 years of research. These tools require a large amount of input data by 

the code user to describe the physical configuration of the plant and to describe user 

preferences among alternate modeling options for uncertain severe accident phenomena. 

Furthermore, the code also accepts certain types of user input that adds new modeling features

to a MELCOR calculation. For example, user defined functions, which operate on specific timedependent variables exposed to the user, and which can be used to specify boundary 

conditions or to control the operation of various objects defined in the user model; the results of 

these logic and arithmetic operations can be included to define the response of specific plant 

components or systems to changes in thermodynamic or other environmental conditions that 

might occur during a particular calculation. The entire body of code input reflects the informed 

judgment of the code user on how a MELCOR model should be configured to generate a 

realistic estimate of plant response to a severe accident. Collectively, this information 

represents the “best practice” modeling approach for using MELCOR in performing severe 

accident progression and radionuclide source term calculations.

This report documents the best practice approach for performing MELCOR calculations for the 

SOARCA project. Section 2 describes the overall technical approach used in the current 

analysis and compares it to the technical approach used in the analyses documented in 

NUREG/CR-2239, “Technical Guidance for Siting Criteria Development” [45]. Section 3

describes the specific modeling practices used to develop and exercise the MELCOR models of 

pressurized water reactor (PWR) and boiling water reactor (BWR) plants examined in this study. 

Section 4 presents a discussion on MELCOR validation. Finally, Section 5 presents an 

accounting of several insights realized from analyses utilizing the best practices.

1.1 Background

MELCOR [2] is a fully integrated, engineering-level computer code that models the progression 

of severe accidents in light-water reactor nuclear power plants. MELCOR is developed at 

Sandia National Laboratories (SNL) for the NRC as a second-generation plant risk assessment 

tool and the successor to the Source Term Code Package. A broad spectrum of severe 

accident phenomena in both boiling and pressurized water reactors is treated in MELCOR in a 

unified framework. These include thermal-hydraulic response in the reactor coolant system 

(RCS), reactor cavity, containment, and confinement buildings; core heat-up, degradation, and 

relocation; core-concrete attack; hydrogen production, transport, and combustion; fission 

product release and transport behavior.
1-2

Modifications were made to the code throughout the SOARCA studies leading to enhanced user 

options. To take advantage of the modifications discussed throughout the report, MELCOR 

version 1.8.6 subversion 3870 or later will be required. Many of the code modifications and the 

inclusion of new models were available in versions prior to subversion 3870; however, 

subversion 3870 was employed for the last SOARCA simulations. Continual code development 

is ongoing; therefore, confirmation of default user input and models should be performed by 

analysts hoping to employ the practices discussed herein.

1.2 Objective

The objective of this report is to describe the best practices, which are comprised of the model 

improvements, modeling approach, and parameter selections, that support the analyses in 

NUREG-1935 [1], NUREG/CR-7110 Volume 1 [62], and NUREG/CR-7110 Volume 2 [63]. This 

document also explains the significance of key modeling improvements and presents a review 

of the validation and phenomena modeling to support the application of MELCOR.

1.3 Independent Peer Review

Two independent review groups were established for the SOARCA project. The first group was 

an expert panel convened in a public meeting forum August 21-24, 2006. The purpose of the 

meeting was to determine whether the MELCOR and MELCOR Accident Consequence Code 

System (MACCS) modeling approaches were consistent with the project objectives to use stateof-the-art modeling with emphasis on realism in phenomenological and system treatments. The 

review was conducted by two panels drawn from the nuclear industry, the U.S. Department of 

Energy (DOE) national laboratory complex, and Gesellschaft fur Anlagen- und Reaktorsicherheit 

mbH (GRS), with recognized expertise in the MELCOR and MACCS numerical simulation tools.

Observations and recommendations specific to each model were provided. The experts found 

the MELCOR code and the proposed approach for application of the codes (i.e., best practices) 

to be generally appropriate for performing realistic predictions of accident progression and

source term for the SOARCA project. A summary of the expert review panel’s findings may be 

found in “State-of-the Art Reactor Consequence Analyses (SOARCA) Project MELCOR 

Modeling Practices Review” [4].

The second review was conducted by a formal peer review committee (PRC) of internationally 

renowned experts established in 2009-2011 to evaluate and suggest improvements to the 

SOARCA project. The PRC evaluated the project approach, assumptions, results, and 

conclusions. To support the peer review, a series of PRC meetings were held where NRC and 

SNL staff presented technical details regarding the MELCOR accident progression modeling

and MACCS consequence modeling. The PRC was provided draft documents for review prior 

to the meetings. Technical discussions were conducted on all major topics such as the accident 

sequence selection and screening process, seismic analysis, emergency response, offsite 

consequences, and code validation. 

The PRC reviewed the modeling approach, parameter development, parameter selection, and 

consequence results. The PRC provided insights and identified gaps and issues for 

consideration throughout the project. The final letter of each PRC member is provided in 

Appendix B to NUREG-1935 [1]. This MELCOR best practices document was not included in 

the PRC review because it was completed after the publication of the SOARCA project primary 

documentation: NUREG-1935 and NUREG/CR-7110 Volumes 1 and 2.
2-1

2. TECHNICAL APPROACH

The technical approach used to calculate radionuclide release to the environment (i.e., the 

“source term”) accompanying credible, but very low frequency accident scenarios relies on the 

application of the MELCOR computer code [2], which was specifically designed to calculate 

reactor and containment system response to postulated severe accidents. The technical 

approach for evaluating the consequences arising from the release of radioactive material to the 

environment involves a separate computer code (i.e., MACCS), which is described in a separate 

report [3]. Key differences in the approach used in the SOARCA project from those used in past 

nuclear reactor radiological consequence calculations are described in Section 2.2.

2.1 Analytical Models

The technical approach that was adopted to define the quantitative characteristics of 

radiological release to the environment was to calculate temporal changes in reactor and 

containment conditions using MELCOR Version 1.8.6. MELCOR is a large, integrated computer 

code developed at SNL under the direction of the NRC and the joint sponsorship of international 

members of the Cooperative Severe Accident Research Program (CSARP). The code is 

“integrated” in the sense that it combines analytical models for a wide-spectrum of physical 

processes (previously evaluated as separate disciplines) into a single, numerically-coupled 

simulation. Among the technical disciplines addressed by MELCOR are: 

• thermal-hydraulic response of the RCS and containment to the postulated accident 

scenario, 

• fuel (core) heat-up and physical degradation due to melting and loss of mechanical 

strength, 

• fission product release from fuel, and

• transport of fission products (in vapor or aerosol form) away from the core, through the 

RCS and containment, to the environment.

Critical to quantitative evaluation of these disciplines are mathematical models for complex 

physical processes, such as: changes in the physical state (morphology) of core materials, 

generation and combustion of hydrogen as a byproduct of the oxidation of metallic components 

in the core, the erosion of concrete in regions of the containment under the reactor pressure 

vessel due to chemical decomposition by molten core debris, and mechanical failure of major 

structural barriers to fission product release (such as the reactor pressure vessel and 

containment). The MELCOR Reference Manual [64] provides a detailed accounting and 

description of these models. The information contained in this document is designed as a 

companion to the MELCOR code manuals, and provides a brief description of the way in which 

MELCOR models were used to represent aspects of nuclear power plant behavior during a 

severe accident that are (a) difficult to predict with high confidence due to uncertainties in their 

governing phenomena and (b) whose outcome is important to calculated results. Major 

uncertainties in MELCOR models for accident progression were addressed in two steps. First, 

the accident progression analysts developed a list of uncertain phenomena that can have a 

significant effect on the progression of the accident. Second, alternate ways of addressing each 

phenomenon were considered and a ‘best-estimate’ approach was developed by applying 

appropriate and available modeling tools in MELCOR. Calculations performed using the bestestimate approach are referred to in this document as the ‘base case’ analysis, and the manner 
2-2

in which MELCOR models and input parameters were configured to represent uncertain and 

important events and processes is described in Section 3.

2.2 Important Differences in Approach from Prior Work

Radiological source terms to the environment used in NUREG/CR-2239, “Technical Guidance 

for Siting Criteria Development,” were developed from early research on severe accident 

behavior that followed the publication of WASH-1400 in 1975 and the accident at Three Mile 

Island Unit 2 (TMI-2) in 1979. This research was initiated to develop predictive methods for 

calculating fission product release and transport during a wide spectrum of postulated accident 

sequences involving substantial damage to fuel in the reactor core. Results of this work were 

documented in NUREG-0772 [5], which was published in 1981. 

The analysis methods used to generate the ‘Siting Source Terms’ (SSTs) used in 

NUREG/CR-2239 involved the manual integration of calculations from several computer codes, 

as illustrated in Figure 2-1. Each code examined a particular portion of the overall analysis, 

such as RCS thermal-hydraulic response, core heat-up and ‘meltdown’ (MARCH), fission 

product transport and deposition in the RCS (TRAP-MELT) and fission product retention in the 

containment before release to the environment (CORRAL, NAUA, etc.)(1)

. The central element 

of this calculation scheme was the TRAP computer code, which used estimates of fluid 

velocities and RCS surface temperatures calculated by MARCH, combined with formulas for 

vapor and aerosol deposition rates, to characterize the extent to which fission products ‘plate 

out’ on RCS surfaces before being carried to the containment. Similar information regarding 

flow rates and temperatures in the containment were used to estimate deposition on 

containment surfaces(2)

. 

Several major simplifications were involved in the NUREG-0772 calculations that have been 

eliminated in the current approach. Among the most important of these are the following:

• The source term analysis tools illustrated in Figure 2-1 represent a linear progression of 

calculations, in which results of one calculation become ‘input’ to a subsequent 

calculation. As a result, physical dependencies between processes modeled in different 

codes can only be represented in one direction. Feedback mechanisms are not directly 

treated. For example, decay heating of surfaces in the RCS or containment due to 

deposited radionuclides was not addressed in the MARCH calculation of RCS 

thermal-hydraulic response. As a result, long-term revaporization of fission products 

from surfaces was not addressed as a late release mechanism. 

In contrast, models for the processes governing severe accident progression and 

radionuclide release/transport are arranged into an integrated set of computational 

modules in MELCOR, which are solved in a single computational framework. 

Interdependencies among diverse phenomena are captured directly in the numerical 

solution.

• In the early NUREG-0772 approach, the release of fission products from fuel was not 

integrally linked to the calculation of time-dependent changes in fuel temperatures by the 

MARCH code. Results of the MARCH calculations were used to inform the analysts 

 1 A description of these codes can be found in references cited in NUREG-0772. 2 The code system used to perform the calculations documented in NUREG-0772 developed later into the 

‘Source Term Code Package’ (STCP), which was the predecessor of MELCOR as the NRC’s principal tool for 

severe accident analysis.
2-3

about the time at which fuel failure would be expected to begin for a particular accident 

sequence, but fission product release rates were not explicitly (numerically) coupled to 

the calculated fuel temperature history. Rather, an average (constant) release rate was 

defined as input to the TRAP-MELT calculations based on limited data from early 

experimental measurements. Distinct release rates were defined for iodine, cesium, and 

(collectively) all other particulate matter.

In contrast to this simplified approach, MELCOR calculates time-dependent release 

rates for fission products from the core, based on a validated correlation of fission 

product release rate and the temperature history calculated at 50 distinct regions of fuel 

assemblies (five radial rings and ten axial levels).(3)

Figure 2-1 Computational methods used to derive source terms in NUREG-0772

• For the NUREG-0772 analyses, a linear series of control volumes (each with internal 

surfaces) was used to represent fission product transport and deposition within the 

primary coolant system. Fission product retention in the containment was calculated by 

applying a variety of computational models to the entire containment free volume, which 

account for various aerosol and vapor deposition and attenuation mechanisms. An 

illustration of a typical nodalization scheme used for the TRAP-MELT calculation of RCS 

retention is shown in Figure 2-2. 

 3 Previous MARCH results were limited to fewer nodes and a one-dimensional thermal-hydraulic 

representation of the core because of storage and computer runtime limitations. It led to a more coherent melt, 

which was contrary to the end-state of the TMI-2 accident.

ORIGEN

Core Inventory of 

Radioactive Material

TRAP-MELT

Primary Coolant System 

Transport & Retention

MARCH

Thermal Response of 

Reactor & Containment

Fission Product Release 

Rate from Fuel

CORRAL, NAUA, 

HAARM-3, QUICK

Transport of Vapor & Aerosols 

in the Containment Atmosphere

Release of Radioactive 

Aerosol & Vapor to the 

Environment

(Source Term)
2-4

In contrast, the spatial nodalization of the reactor pressure vessel, primary coolant 

system, and containment in the MELCOR models developed for the SOARCA project

recognizes much more geometric detail. For example, the entire PWR RCS was 

represented by four spatial regions in the NUREG-0772 calculations. The MELCOR 

model of the same 3-loop Westinghouse PWR used in the current Surry analysis uses 

25 control volumes for the core region alone; over 100 control volumes are used to 

represent the entire RCS. This dramatic increase in detail provides much greater 

resolution of the driving forces governing fission product transport and deposition. In 

particular, local fluid velocities and temperatures, structural surface temperatures and 

associated temperature gradients are all calculated in greater detail than was available

or practical at the time of the analyses supporting NUREG/CR-2239.

Figure 2-2 Typical spatial nodalization of RCS in the NUREG-0772 methodology

CORE

UPPER

PLENUM LOWER PLENUM

STEAM 

GENERATOR

CONTAINMENT

Release Pathway for 

PWR Cold Leg LOCA

DOWNCOMER

CORE

UPPER

PLENUM

LOWER ANNULUS STEAM DRYERS

UPPER VESSEL

HEAD

UPPER

ANNULUS

CONTAINMENT

Release Pathway for 

BWR ATWS
3-1

3. MELCOR BEST PRACTICES FOR SEVERE ACCIDENT 

ANALYSIS

This section describes the MELCOR best practices used in the SOARCA scenarios selected for 

the Peach Bottom and Surry Nuclear Power Stations. The detailed approach to modeling all 

phenomena governing fuel damage, fission product release, and other aspects of plant 

response to postulated severe accident is not provided here. Rather, a small subset, focused 

on the best practices of the large body of important phenomena, is addressed. The basic 

selection criteria applied to determine which phenomena or modeling practices were included in 

the discussion were the following:

• was the phenomenon or event an important contributor to the progression of the 

accident, 

• was uncertainty perceived in the system response to the phenomenon or event, and 

• was the phenomenon or event addressed in the calculation by user generated models or 

input that differed from the default input parameters.

Although many of the practices described in this section are expressed in terms of their 

implementation into MELCOR models specifically developed for the SOARCA project, they 

involve phenomenological considerations and uncertainties that have broader applicability.

The best practice modeling features described in this chapter fall into two primary categories, 

user input and hard-coded modifications to the MELCOR source code. 

1. User input is a broad category encompassing severe accident modeling features that are 

controlled or influenced by the user through input file modifications. Modeling practices 

within this category can take several forms depending on the specific type of user input 

used to influence or define a MELCOR model. The best practice user input guidance is

documented in Section 3.1.

a. Sensitivity coefficients are an inherent form of user input, native to MELCOR, 

which permit a large variety of modifications which establish the methodology for 

performing the simulation. These parameters span a wide spectrum from 

coefficients of heat transfer correlations to numerical convergence criteria. In 

general, the default values for the sensitivity coefficients represent best-estimate 

settings and are not typically changed.

b. Model selection is an additional form of user input. The user input identifies the 

desired model among the available MELCOR options. An example of this form 

of user influenced modeling practice is the correlation used to calculate the 

release rate of fission products from over-heated fuel. MELCOR offers several 

types of correlations, and the user must select one. 

c. A third form of user influence on a MELCOR calculation is an analytical 

expression or implementation logic developed by the user (typically via control 

functions) to represent an aspect of severe accident progression that is not 

directly represented as a ‘model’ within MELCOR. An example of this type of 

modeling practice is a user-specified correlation or condition that determines if, 

and when, a particular component would fail to function. Failure of a safety/relief 
3-2

valve to reclose due to repeated cycles, or enhanced leakage through an 

over-heated reactor coolant pump seal would fall into this category. 

2. A second broad category of changes to MELCOR modeling practices involve changes to 

the MELCOR source code (Fortran) to accommodate advanced modeling needs for the 

SOARCA project. In some cases, these changes codified modeling practices that had 

been implemented through user input in many past NRC applications of MELCOR. 

Incorporating the user generated input directly into the code simplified user input and 

enhanced the fidelity of the modeling practice among MELCOR models for various 

nuclear power plants. Enhancements to the MELCOR source code that were 

implemented as part of the SOARCA project are described in Section 3.2. 

The best practices documented within this report typify general user input methodologies for 

which the reader must appropriately incorporate for their specific application. The 

implementation of these modeling practices in a plant-specific MELCOR model depends on the 

details of that model and is not described in this report. Important modeling details, including 

possible deviations from best practice, are provided in the documentation associated with the 

plant-specific MELCOR models developed for use in the SOARCA project. 

3.1 Best Modeling Practices

Several of the modeling practices that are defined or controlled by user input are applied in 

calculations of severe accident progression for a PWR or BWR. Others are specific to one 

design or the other. Generic modeling practices (i.e., those applicable to both designs) are 

described in Section 3.1.1. PWR- and BWR-specific modeling practices are then described in 

Sections 3.1.2 and 3.1.3, respectively. The recommended best modeling practices described in 

Sections 3.1.1 through 3.1.3 were presented and reviewed by an independent expert panel. In 

addition to the best modeling practices described in Sections 3.1.1 through 3.1.3, other best 

modeling practices that were less important or had less uncertainty in their implementation are 

simply outlined in Appendix A. 

3.1.1 Generic Light Water Reactor Best Practices

Modeling practices discussed in this section are applied generically to both MELCOR models of 

the SOARCA project Surry Power Station and Peach Bottom Atomic Power Station. The 

specific manner in which they are implemented can vary slightly between the two, but the 

physical processes that are represented in these models are consistent between the two 

designs.

3.1.1.1 Fuel Degradation and Relocation Treatment

As a reactor core overheats in a severe accident, fuel cladding ruptures would occur at relatively 

low temperatures releasing fission product gases from the fuel-cladding gaps. As temperatures

continue to rise, fuel cladding would oxidize and the fuel rods would form outer oxide shells. 

The oxide shells would have a high melting temperature relative to that of unoxidized cladding 

(Zircaloy) and, as evidenced in the Phebus tests, would maintain fuel geometry as Zircaloy,

interior to the shell, melts and drains away. This configuration is illustrated in Figure 3-1. In 

maintaining fuel geometry, the oxide shells would be susceptible to thermal-mechanical 

weakening over time. Modeling was added to the MELCOR code during the course of the 

SOARCA project that acknowledges this thermal-mechanical weakening as a function of time 

and temperature. Prior to the modeling addition, an oxide shell could maintain fuel geometry at 

very high temperatures for a long period of time after interior Zircaloy drains away (ending 
3-3

oxidation and the associated heat generation). An oxide shell would not fail until its temperature 

reaches the eutectic temperature for the UO2/ZrO2 system (2500 K); and with oxidation heat 

generation gone, this could take a long time. The new modeling eliminates this threshold 

temperature failure requirement. The modeling is functionally similar to what is accomplished in 

the Modular Accident Analysis Program (MAAP) [59]. In the new modeling, as cladding oxide 

temperature increases from the melting temperature of Zircaloy (2099 K by default in 

MELCOR), a thermal lifetime function accumulates fractional damage towards an inferred 

thermal-mechanical failure. The cumulative fractional damage is incrementally increased each 

time step by the fractional damage incurred during the timestep considering the current cladding 

temperature and lifetime associated with that temperature, i.e.:





 



 + = +

( ) 1 TtF T

dt Di Di Equation 1

where,

Di+1 is the fractional damage accrued through the current timestep

Di is the fractional damage accrued up to the current timestep

dt is the duration of the current timestep

TtF(T) is the time at the current temperature (T) required to fail oxidized cladding

Fractional damage is accrued in this way locally by axial level and radial ring throughout the 

core. The best practice dependence of time-to-failure as a function of temperature, enforced

through user input, is presented in Table 3-1(4)

.

Table 3-1 Best-estimate time to fuel rod collapse versus cladding oxide temperature

Cladding Temperature Time to Failure

2090 K Infinite

2100 K 10 hrs

2500 K 1 hr

2600 K 5 min

Times to failure intermediate to entries in Table 3-1 are linearly interpolated. Infinite lifetime is 

assumed at cladding oxide temperatures below the melting point of Zircaloy. The relatively 

short time associated with 2500 K and the even shorter time associated with 2600 K reflect the 

melting tendencies of irradiated fuel inferred from the Phebus experiments. Damage function 

accumulation does not begin until unoxidized cladding thickness drops below 10% of nominal 

values.

 4 Sandia National Laboratories conducted an expert review of severe accident progression modeling for the 

SOARCA project in Albuquerque, NM on August 21-22, 2006 [4]. The expert review was conducted by five panelists 

with demonstrated expertise in the analysis of severe accidents at commercial nuclear power plants. The time-tofailure dependencies in Table 3-1 were presented to the panelists as one of the several uncertain modeling practices 

presented. The panelists provided written comments and suggestions, which were incorporated into the subsequent 

analyses.
3-4

Figure 3-1 MELCOR depiction of fuel rod degradation

3.1.1.2 Lower Plenum Debris/Coolant Heat Transfer

Direct interaction between over-heated (possibly molten) core debris and a pool of water can 

occur at two key junctures in the chronology of a severe accident in a light water reactor (LWR). 

The first major juncture is when core debris relocates from a position above the lower core 

support structures; the second juncture is when the reactor vessel lower head fails and core 

debris relocates onto the floor of the containment. The thermo-mechanical interactions between 

core debris and water during these periods of debris relocation can be either benign or 

extremely energetic, depending on several factors:

• the thermodynamic state of the debris (temperature and morphology), 

• composition of the debris (unoxidized metals, ferric oxides, ceramics, etc.),

• debris relocation geometry (coherent pour, massive slump, cascade of particulate, etc.),

• depth and temperature of water pool (saturated or subcooled), and

• initial pressure of the confining vessel.

Proper accounting for the interaction between fuel and coolant at these two junctures can have

a significant effect on the prediction of severe accident chronology, challenges to containment 

integrity and the resulting source terms. For example, relocation of core debris into the lower 

head without significant interaction with residual water below the lower core support structure 

can result in rapid heating and early failure of lower head structures. Conversely, significant 

interaction between core debris and water can significantly delay the time of vessel breach, 

produce large quantities of steam (leading to enhanced oxidation of metallic components) and 

potentially result in a coolable debris bed in the lower head.

Ex-vessel interactions between core debris and water on the containment floor can be equally 

important to severe accident progression. The possibility of avoiding fission product release 

from fuel debris during corium-concrete interactions by quenching core debris upon release 

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)
3-5

from the reactor vessel is a significant enough reason for properly characterizing ex-vessel 

debris-coolant interactions.

Most studies of debris-coolant interactions have viewed the phenomenon as a precursor to 

steam explosion. However, most experiments involving molten debris-coolant interactions do 

not result in a steam explosion. These less energetic events are of equal (perhaps greater) 

value to the analysis of severe accident progression as they provide valuable information on 

debris quenching and long-term coolability. Published literature describing these studies was 

reviewed to determine the depth of water required to sufficiently fragment and cool molten core 

debris. This depth of water could then be applied to typical in-vessel or ex-vessel situations in 

which the coolability of relocating debris needs to be evaluated in MELCOR. Data from five 

different test series and a total of 29 different experiments were examined. Key measurements 

from these tests are listed in Table 3-2.

Analysis of hydrodynamic breakup behavior in these tests resulted in estimates of the vertical 

distance a molten jet must travel in a pool of water to fully quench the molten debris. This 

distance (referred to as the molten jet breakup length) was estimated to be between 20 and 50 

jet diameters for melts without unoxidized metals, and between 10 and 20 diameters for melts 

with unoxidized melts. Test results indicate that molten jet breakup occurs at both the leading 

edge and along the trailing column. Steam production at the leading edge leads to jet breakup. 

Steam moving through the pool alongside the molten jet also contributes to jet breakup. When 

unoxidized metals are present, steam oxidizes metals at the jet surface releasing additional 

energy to the steam/water mixture and enhancing breakup along the trailing jet column. 

If these figures are applied to full-scale reactor conditions, and one can postulate a 

representative diameter of the jet of molten core debris that would emerge from the reactor 

vessel lower head after failure, the minimum depth of water required to quench the debris can 

be estimated. For example, if one assumes the characteristic diameter of the molten jet is 

roughly the diameter of a single ‘unit cell’ of a reactor fuel assembly (i.e., approx. 10 cm), 

fragmentation and quenching of the molten material would be achieved in 2 to 5 m of water (for 

oxidic melts) and 1 to 2 m of water for metallic melts. This distance is well within the range of 

water depth in the lower plenum of a typical BWR at the time lower core support plate failure 

would first occur, initiating large scale in-vessel debris relocation. Therefore, a best-estimate 

characterization of debris behavior operates under the assumption of efficient debris cooling

provided a sufficiently deep pool of water remains in the lower plenum.

Default MELCOR code user input parameters for the ‘falling debris quench’ model were, 

therefore, changed to effect efficient heat transfer. In particular, the debris hydraulic diameter 

was defined to correspond to the average end-state conditions observed in the FARO tests and 

the average ‘fall velocity’ was set to a value that caused the temperature of falling debris to 

decrease by an amount that ensured debris temperatures in the lower head were below the film 

boiling limit. In addition, the one-dimensional counter-current flow limitation (CCFL) was 

removed, through user input, from the overlying debris heat transfer model to represent water 

penetration into the debris bed, perhaps through 2- or 3-dimensional circulation flow patterns. 

This modeling approach resulted in debris cooling if there was a pool of water in the lower 

plenum and delayed heat-up of the vessel lower head until the overlying water had evaporated.
3-6

Table 3-2 Summary of data from molten debris-coolant interactions experiments.
3-7

Table 3-2 (continued)
3-8

3.1.1.3 Safety Relief Valve Cycling and Failure

Safety and relief valves are installed at BWR and PWR installations to provide over-pressure 

protection for the primary loop. The MELCOR user defined models of the valves provide a 

detailed accounting of these valves, including the actuation conditions, support system 

dependencies, and failure characteristics. The various types of safety and relief valves can be 

manually actuated if requisite support systems are available (e.g., electric power, control air, 

etc.). Pressurizer power-operate relief valve (PORV) and safety valves (SVs) as well as steam 

generator PORVs and SVs are included in PWR designs and in the general MELCOR models.

Similarly safety relief valves (SRVs) are included in the BWR models. Failure, in this context 

means failure-to-reclose after successfully opening to relieve pressure.

Two modes of failure are represented in the MELCOR user defined models independently. The 

first failure mode represents stochastic failure of the valves to reclose when pressure reduces 

below the closure set point. Mechanisms for failure-to-reclose are identical to those 

incorporated in the random event captured in the probabilistic risk assessment (PRA) for most 

nuclear power plants. The second failure mode accounts for high temperature operation 

induced seizure in the open state.

Uncertain parameters are modeled as best-estimate values; therefore, stochastic failure is 

represented by the expected value, the mean. Mean failure-to-open and failure-to-reclose 

probabilities for the specific type(s) of valve(s) installed at a particular plant were obtained from 

plant data (if available) or from generic failure data documented in NUREG/CR-6928, “IndustryAverage Performance for Components and Initiating Events at U.S. Commercial Nuclear Power 

Plants”. In all cases examined thus far, the probability for failure-to-reclose was much greater 

than failure-to-open; therefore, failure-to-open is currently ignored in the MELCOR simulations. 

Stochastic failure of the safety relief valve is determined from the failure-to-reclose rate, 3.7x10-3

per demand as presented in NUREG/CR-6928. (Other alternate sources for valve failure

characteristics are available. The PWR simulations used data supplied by the utility.) The 

expected failure is calculated as the inverse of the failure rate, which corresponds to the 270th

cycle or 63rd percentile of the cumulative distribution function.

(5)

 The cumulative probability of 

failure is calculated with each valve cycle. When the maximum tolerable cumulative failure limit 

has been exceeded, the cycling valve remains in the failed open position.

Engineering judgment based on material properties of internal components was employed to 

determine an adequate failure criterion based on temperature. Periodic cycling of a safety or 

relief valve with very high internal gas temperatures will cause the valve body and internal 

components to slowly increase in temperature. At some temperature, thermal expansion or 

yielding of internal components will prevent the valve stem from moving and re-seating. No 

data or models are available that clearly identify the temperature at which seizure would occur. 

However, stainless steel loses its strength at temperatures above approximately 1000 F (811 

K). Temperatures in excess of 900 K reduce the ultimate tensile strength and yield strength by 

30 percent of the design temperature strengths. A reasonable maximum temperature limit of 

900 K was chosen as an approximation for internal component failure [62]. A representative 

heat structure was implemented in the BWR simulations to calculate internal component 

 5 The inverse of the failure rate or the harmonic mean was selected as the best-estimate value for the failure 

rate. The harmonic mean is characteristic of a best-estimate average for rate functions. Due to its importance on the 

BWR accident progression, SOARCA also performed sensitivity studies that examined the impact of alternate failure 

assumptions.
3-9

temperature with regard to the maximum temperature limit criterion. Additional temperature 

induced failure mechanisms were also considered and are discussed in [62].

3.1.1.4 Fission Product Release

Insights have been developed over the past decade as a result of experimental programs and

have been used to update the recommended MELCOR specifications for modeling the release 

of fission products(6) from reactor fuel under severe accident conditions. The new models have 

been incorporated as new defaults in the MELCOR code. Separate specifications are provided 

for use in spent fuel pool release conditions owing to differences in the reduction/oxidation 

potential in air oxidizing conditions. Some review of the motivation for the new modeling 

approach follows with an assessment of the new model against fission product release 

experiments.

Past versions of MELCOR primarily used the CORSOR-M release model for calculating fission 

product release as described in the MELCOR Reference manuals and in a Battelle report by 

Ramamurthi and Kuhlman titled “Refinement of CORSOR – An Empirical In-Vessel Fission 

Product Release Model [6].” Also described in these references are the CORSOR and the 

Booth diffusion release model, implemented in MELCOR as the CORSOR-Booth optional 

release model. The CORSOR and CORSOR-M models are classified as fractional release rate

models, differing only slightly in mathematical form, which specify the fractional release rate of 

the fission product inventory remaining unreleased up to that time. These are empirical models 

that are based largely on the small-scale horizontal induction (HI) and vertical induction (VI)

experiments performed at Oak Ridge National Laboratory (ORNL). 

The Booth diffusion model is by comparison a physics-based model, albeit oversimplified, that 

describes the transport of fission products within fuel grains to the grain surface as a diffusion 

process. In the MELCOR code implementation of the Booth diffusion treatment, an additional 

gas phase transport process is imposed in moving fission products from the grain surfaces to 

the atmosphere. Elements such as molybdenum that are modeled in MELCOR as having very 

low vapor pressures are ultimately released at a low rate regardless of the rate of diffusion 

within the grain. Once released from the fuel, fission product class combinations can be 

defined, such as CsI, in order to represent fission product chemistry and speciation. In the 

present code architecture, multiple combination assignments such as CsI and Cs2MoO4 were 

not foreseen and must be approximated. Once assigned to the chemical class on release, 

generally no additional chemistry is allowed, an exception being CsI chemisorption with 

subsequent revaporization of iodine, leaving the permanently chemisorbed Cs attached to a 

deposition surface.

Critical assessments of these models and their performance have been limited, partly owing to 

lack of additional quality data from which to render a judgment. One assessment performed by 

ORNL with MELCOR 1.8.2 evaluated the performance of the MELCOR default models when 

applied to the VI series of tests [7]. The report observed that while total releases could often be 

adequately predicted that the time-release signature was often not very good. 

Recommendations were provided for code modeling improvements, including provision to vary 

release based on the H2/H2O environment. Recently however, additional experimental data are 

 6 The recent modifications to Version 1.8.6 (Version YR) for the SOARCA project implemented the new 

ORNL-Booth fission product release model as the new default model. In Version 1.8.6, the new defaults are invoked 

with the Version “2.0” keyword. All subsequent 2.X code versions will automatically use the new ORNL-Booth model 

as the default fission product release model.
3-10

becoming increasingly available from international testing programs, in particular the French 

VERCORS program and the Phebus integral experiments, and user assessment of current 

MELCOR release models in the prediction of these tests has illuminated some deficiencies that 

are partly remedied in the recommendations of this report. The Phebus experiments in 

particular reveal shortcomings of the empirical CORSOR and CORSOR-M models with respect 

to release rates during the initial fuel heat-up, and those models have been found to significantly 

overestimate early release rates even though total integral releases might compare reasonably 

well. Additionally, the integral Phebus tests provide release data under conditions that are 

significantly less coherent (and more prototypic) in terms of temperature and oxidation/reduction 

conditions than in the small scale tests (HI, VI and VERCORS) where the fuel sample is small, 

temperatures are uniform and oxidation/reduction conditions controlled and constant. The 

Phebus experiments, we contend, provide conditions for release that are more representative of 

conditions expected in the full scale reactor accident case, and are used as the principal 

reference for judging the performance of the MELCOR release models.

MELCOR Release Models

The various release rate models available in the MELCOR code for user selection are briefly 

summarized as follows. 

CORSOR

The original CORSOR model correlates the fractional release rate coefficient in exponential 

form,

exp( ) i i i T Tj k = A B T for ≥ Equation 2

where ki is the fractional release rate per minute, Ai and Bi are empirical coefficients based on 

experimental data, T is the core cell component temperature in Kelvin, and i indicates the 

specific class. Different values for Ai and Bi are specified for three separate temperature 

ranges. The lowest temperature, Tj, for given range j, as well as the corresponding coefficient 

set, Ai and Bi

, are defined in sensitivity coefficient array 7101. If the cell temperature is below 

the lowest temperature range limit specified, no release is calculated.

CORSOR-M

The CORSOR-M model correlates the same release data used for the CORSOR model using 

an Arrhenius form,

k k (- Q /RT ) i o i i exp = , Equation 3

The values of ko,i, Qi, and T are in units of min-1

, kcal/mole, and K, respectively. The value of R

is 1.987 x 10-3 in kcal/mole-K. The values of ko,i and Qi for each class i are implemented in 

sensitivity coefficient array 7102.

CORSOR-Booth

The CORSOR-Booth model considers mass transport limitations to radionuclide releases and 

uses the Booth model for diffusion with empirical diffusion coefficients for cesium releases. 

Release fractions for other classes are calculated relative to that for cesium. The effective 

diffusion coefficient for cesium in the fuel matrix is given by:

D D (- Q/RT ) o = exp Equation 4
3-11

where R is the universal gas constant (cal/mole-K), T is the temperature (K), Q is an activation 

energy (cal/mol), and the pre-exponential factor D0 is a function of the fuel burn-up (m2

/s). The 

cesium release fraction f at time t is calculated from an approximate solution of the diffusion 

equation for fuel grains of spherical geometry [8],

2 6 3 for D t < 1/π

π ′ ′ ′ = - D t D t f Equation 5

2 2

2 exp for D t > 1/ 6 1 π π

π

f = - ( − D′ t ) ′ Equation 6

where

D′ t = 2 Dt a (dimensionless), and

a = equivalent sphere radius for the fuel grain.

The release rate (in mole/s) of Cs during a time interval t to (t + ∆t) from the fuel grain is 

calculated as:

[ ( ) ( ) ]

F t

f D t f D t V t t t

Cs ∆

′∆ − ′∆ = ∑ +∆ ∑ ρ Release rate Equation 7

Where ρ is the molar density of UO2 in the fuel, V is the fuel volume (m3

), F is the fraction of the 

Cs inventory remaining in the fuel grain, and the summations are done over the timesteps up to 

time, (t + ∆t) and t, respectively.

The release rate formulation in the CORSOR-Booth model is also limited by mass transfer 

through the gas phase. The gas phase mass transport release rate from the fuel rod for species 

i, mk  (in mole/s), is calculated using an analogy from heat transfer as:

( 0) ,

, ⋅ − 















= i eq

fuel

fuel i gas

i P

D RT

A Nu D

m Equation 8

where

Dfuel = diameter of fuel pellet (m)

Afuel = fuel rod flow contact area (m2

)

Di,gas = diffusivity of class i in the gas mixture(m2/s)

Nu = Nusselt number

R = Universal gas constant (J/mol-K)

Pi,eq = equilibrium vapor pressure of class i at temperature T. (Pa)

In the mass transfer term the driving potential is the difference in pressure at the surface of the 

grain and the pressure in the free atmosphere, here assumed to be approximately zero.

The effective release rate for Cs given by Equation 7 is a combination of the rates given by 

diffusion and by gas phase mass transport. Therefore, the contribution from diffusion only is 

taken as:
3-12





 



 −

Cs

Cs m

DIFF 

1

Release rate

1 =

Cs

-1

Equation 9

The diffusion release rate (in mole/s) for species i, other than cesium, is given by multiplying the 

cesium diffusion release rate by an appropriate scaling factor Si for each RadioNuclide Package 

(RN) class i: 

i Cs i DIFF = DIFF S Equation 10

The combined mass transport and diffusion release rate mtot,i  (in mole/s) for class i is then

, 1 1

1

− − + =

i i

tot i DIFF m

m   Equation 11

Inspection of Equation 11 together with Equation 8 reveals that the release predicted by the 

MELCOR models can be mass transfer limited by low vapor pressures even if the diffusive 

transport is large.

Limitations of MELCOR Release Models

The fission product release models implemented in previous versions of MELCOR (i.e., before 

Version 1.8.6 (RO) and the code modifications) are quite simplified and are more than 10 years 

dated as indicated in the principal reference for the MELCOR models. The implemented 

models base the release of all radionuclide chemical classes on the release predicted for Cs, 

which in the Booth model is appropriately considered a diffusion process. Scaling factors are 

used to estimate release of other species based on the data fit to experimentally observed Cs 

release in spite of the fact that it is recognized that likely not all fission product classes diffuse at 

the same rate out of the fuel grains, nor are all principal release mechanisms well represented 

as a diffusion process. Consideration of speciation in MELCOR release models is crude and for 

the most part fixed at the time of release to represent the predominating speciation. The vapor 

pressures of the MELCOR release classes are defined to represent the presumed fission 

product speciation.

A better treatment would be to allow the vapor pressure to be adjusted to account for local 

speciation affected by oxidizing or reducing conditions and to then source these species into 

appropriate chemical classes. Such modifications are probably needed for Ba, Mo, UO2 and 

Ru. Provision does exist to consider the extent of cladding oxidation to attempt to simulate 

retention of Te or Ba, but data are needed to use this provision effectively. Separate diffusion 

coefficients for each of the volatile classes would probably be appropriate, and a UO2 oxidation 

model is needed to account for the effect of stoichiometry on diffusion and to predict fuel 

volatilization. UO2 volatilization may be responsible for release of UO3 as well as other 

non-volatile species owing to physical stripping of the fuel matrix containing the fission products. 

A number of more recently evolved release models consider the effect of fuel stoichiometry on 

the diffusion coefficient as well as the oxidizing/reducing potential of the environment

[9][10][11][12]. The VICTORIA code considers a large number of potential fission product 

species in a thermodynamic equilibrium approach; some simplifications to this potentially 

numerically burdensome approach may be needed [13].
3-13

In the more recent models, often, with respect to release behavior, fission products are 

classified into three main groups, volatile (i.e., Xe, Cs, I, Te), semi-volatile (i.e., Ru, Ba, La, Ce), 

and non-volatile (i.e., UO2 and actinides – Ce and La might belong here also). Volatile fission 

products are released based on the Booth diffusion model where the diffusion coefficient 

includes effects of UO2 hyper-stoichiometry. The hyper-stoichiometry in turn is determined by a 

fuel oxidation model. Release of semi-volatile fission products are strongly affected by vapor 

pressure which in turn is affected strongly by speciation determined by the oxidizing/reducing 

conditions resulting from air/steam/hydrogen/Zr-metal in the release location. Non-volatile 

release may be dominated by UO2 volatilization by formation of UO3, producing fuel matrix 

degradation and fuel vaporization. The French Elsa code follows this approach, using models 

similar to those reported by Lewis et al. [9][10].

A more detailed (and flexible) release modeling was needed for the SOARCA project. The 

importance of accounting for speciation and the ensuing effect on specie volatility (vapor 

pressure) is clear. In the best practices approach, as described in the following section, the 

dominant speciation at the time of release is specified and used globally throughout the core 

region. A more elegant model would allow variation of release speciation as conditions in the 

core change locally and temporally with respect to steam and hydrogen concentrations. In the 

case of air exposure, such as in spent fuel pool accidents, different assumptions about 

speciation, especially concerning Ru release, are needed.

Assessment of MELCOR Default Release Models

Whereas the HI-VI ORNL tests provided the original basis for development of the MELCOR 

fission product release models, the Phebus FPT-1 integral experiment is used as the principal 

basis for evaluation of release modeling options. In previous assessment exercises, in 

particular the ISP-46 (International Standard Problem 46 [14]), while the MELCOR default 

CORSOR-M release model was found to predict reasonable total release fractions for many 

fission products, the empirical model was observed by many MELCOR ISP participants to over 

predict the initial release rates. Similar rapid early release is also observed for the CORSOR 

option. The Booth diffusion treatment for release was thought to be a potentially superior 

release model since it has some basis in a physical transport process, however, investigation of 

the MELCOR CORSOR-Booth option using the default Booth release parameters was found to 

produce inferior results, with total release of Cs and other fission products being significantly 

under predicted in test FPT-1. In view of this, review of the literature revealed numerous other, 

more recent parameter-fits to the Booth solution. 

Modifications to MELCOR Booth Release Modeling

A number of these alternative models are reported in an ORNL report that recommends 

updated values for the previously discussed models [56]. Shown in Figure 3-2 are release 

fractions predicted at a constant temperature of 2000 K by the various release models 

discussed in the ORNL report. From this it can be seen that fractional releases predicted by

CORSOR-M produce the largest total release of all of the models. This trend is consistent with 

observations from analyses considering measured releases from FPT-1. Similarly, the 

CORSOR-Booth diffusion model produces the lowest total release of all of the models. This too 

is consistent with MELCOR analyses of FPT-1 using these modeling parameters. Judging that 

a best fit might lie somewhere in between these extremes, the ORNL-Booth parameters were 

subsequently investigated in MELCOR analyses of FPT-1, wherein significantly improved 

release signatures were obtained. The ORNL-Booth parameters were recommended over the 

CORSOR-Booth parameters in the 1995 ORNL report. The ORNL-Booth model is specified by 

the parameters in Table 3-3. Figure 3-3 shows other comparisons between the ORNL-Booth 

and CORSOR-M release behaviors. The fractional release rate (%/min) for the two models 

obtained by differentiating the release fractions in Figure 3-3 are shown in Figure 3-4.
3-14

Figure 3-2 Release fractions for different release models – release temperature =

2000 K

[Note: CORSOR-M produces largest release whereas CORSOR-Booth produces 

the smallest release.]

Table 3-3 CORSOR-Booth, ORNL-Booth, and Modified ORNL-Booth parameters.

CORSOR-Booth ORNL-Booth Modified

ORNL-Booth

Diffusion coeff. Do 2.5x10-7 m2

/sec 1x10-6 m2/sec 1x10-6 m2/sec

Activation Energy Q 3.814x105

joule/mole

3.814x105

joule/mole

3.814x105

joule/mole

Grain radius, a 6 μm 6 μm 6 μm

Class Scale Factors --- --- ---

Class 1 (Xe) 1 1 1

Class 2 (Cs) 1 1 1

Class 3 (Ba) 3.3x10-3 4x10-4 4x10-4

Class 4 (I) 1 0.64 0.64

Class 5 (Te) 1 0.64 0.64

Class 6 (Ru) 1x10-4 4x10-4 0.0025

Class 7 (Mo) 0.001 0.0625 0.2

Class 8 (Ce) 3.34x10-5 4x10-8 4x10-8

Class 9 (La) 1x10-4 4x10-8 4x10-8

Class 10 (U) 1x10-4 3.6x10-7 3.2x10-4

Class 11 (Cd) 0.05 0.25 0.25

Class 12 (Sn) 0.05 0.16 0.16
3-15

Figure 3-3 Release fractions at constant temperature for ORNL-Booth versus 

CORSOR-M

Figure 3-4 Fractional release rate (%/min) – the time derivative of release fraction

[Note: The discontinuity when transitioning between Equation 5 and 6 

results in the observed spike in the ORNL-Booth 2200K curve.]

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1000 2000 3000 4000 5000 6000 7000 8000

time (sec)

Release Fraction

CORSOR-M 2200K

CORSOR-M 2000K

CORSOR-M 1800K

ORNL-BOOTH 2200K

ORNL-BOOTH 2000K

ORNL-BOOTH 1800K

0.01%

0.10%

1.00%

10.00%

0 1000 2000 3000 4000 5000 6000 7000 8000

time (sec)

Fractional Release Rate (%/min)

CORSOR-M 2200K

CORSOR-M 2000K

CORSOR-M 1800K

ORNL-BOOTH 2200K

ORNL-BOOTH 2000K

ORNL-BOOTH 1800K
3-16

While significant improvements in release behavior were obtained for the analysis of the FPT-1 

test with the as-reported ORNL-Booth parameters, an additional modification to the MELCOR 

release model was pursued. Evidence from the Phebus experiments increasingly indicates that 

the dominant chemical form of released Cs is that of cesium-molybdate, Cs2MoO4. This is 

based on deposition patterns in the Phebus experiment where Cs is judged to be in aerosol 

form at 700o

C, explaining deposits in the hot upper plenum of the Phebus test section, and 

deposition patterns in the cooler steam generator tubes. In recognition of this, the vapor 

pressure of both Cs and Mo classes were defined to be that of Cs2MoO4. While having little 

effect on the net release of Cs, this change had a significant effect on the release of Mo. In 

MELCOR, by default the Mo vapor pressure is so exceedingly low that the net release is limited 

by the vapor transport term, as expressed in Equation 8 and Equation 9. Vapor pressures for 

selected fission product species are shown in Figure 3-5. Defining the Mo vapor pressure to be 

that of Cs2MoO4 produced significantly improved the predicted Mo release rate with respect to 

observed FPT-1 releases, as will be seen in the following section.

Figure 3-5 Vapor pressure of selected species

Assessment of Modified ORNL-Booth Model Against Phebus FPT-1

The Phebus program provides probably the best source of prototypic data on fission product 

release from irradiated fuel, benefiting from many lessons learned from earlier similar 

experimental efforts and from advances in testing technology, instrumentation, etc. A schematic 

of the Phebus test facility is shown in Figure 3-6. A previously irradiated fuel bundle of about a 

meter in length is situated in the irradiation cavity in the Phebus test reactor and caused to 

undergo severe damage from nuclear heating and oxidation by injected steam. Fission 

products released from the test bundle flow through a heated section representing the reactor 

1.E+03

1.E+04

1.E+05

1.E+06

1.E+07

0 500 1000 1500 2000 2500 3000 3500 4000

Pressure (Pa)

Temperature (K)

Fission Product Vapor Pressure

CsI

Cs2MoO4

Mo

I2
3-17

coolant system, through a simulated steam generator tube where strong deposition can occur, 

and into a simulated containment where fission product fallout occurs.

Figure 3-6 Schematic of the Phebus Test Facility showing test fuel bundle, heated 

lines, steam generator tube and simulated containment

Shown in Figure 3-7 is the nuclear heating history that was used in test FPT-1 to heat the 

bundle to simulate severe accident decay heating conditions. The chemical heating produced 

by steam-Zr oxidation is also shown in the figure. The temperature response of the test fuel is 

shown in Figure 3-8 where the temperature transient resulting from the additional oxidation 

heating is clearly evident. During this time, fission products are also released where oxidation 

conditions vary from oxidizing to reducing, depending on elevation in the test bundle. Figure 3-9

shows the end state of the test bundle following conclusion of the experiment. 
3-18

Figure 3-7 FPT-1 nuclear and chemical heating history

Figure 3-8 FPT-1 maximum bundle temperature history

 

0

5

10

15

20

25

30

35

40

5000 7500 10000 12500 15000 17500 20000

time [sec]

Power [KW]

fission power

oxidation power

 

0

500

1000

1500

2000

2500

3000

5000 7500 10000 12500 15000 17500 20000

time (sec)

Temperature [K]
3-19

Figure 3-9 Emission gamma tomography of the end-state condition of test FPT-1

The following figures (Figure 3-10 through Figure 3-21) show the results of using the modified 

ORNL-Booth model for fission product release in the FPT-1 analysis. In most cases significant 

improvement is realized in both the early release time signature as well as for total predicted 

release. Where available, Phebus data is presented. The release for the barium class for the 

ORNL-Booth model is low relative to the data, whereas the release predicted using the 

CORSOR-M model is high. Improvement to this observed release proved elusive and it is 

believed that some adjustments to the vapor pressure for Ba to account for some not yet 

understood barium speciation could produce some improvement. Adjustments to both vapor 

pressure and scaling factors were rationalized for Mo release based on Phebus program 

findings, producing good agreement with experiment. The Ru vapor pressure was increased by 

a factor of 10 arbitrarily to account for some greater volatility attributed to formation of oxides 

under moderately oxidizing conditions, and the Booth scaling factor was adjusted to gain 

agreement with experimental observations. The Booth scaling factor for UO2 was increased 

significantly in order to gain agreement with test observations. This also is rationalized as due 

to effects of fuel oxidation and greater volatility of fuel oxides. Ce and La release parameters 

were not adjusted owing to lack of experimental basis; however, one could reason that their 

releases ought to roughly follow UO2 release if fuel matrix stripping follows from fuel 

volatilization. The following section presents comparisons of the modified ORNL-Booth model 

against ORNL VI tests and more recent VERCORS test data.

Figure 3-10 Comparison of ORNL-Booth versus CORSOR-M for Xe release (Class 1)
3-20

Figure 3-11 Comparison of ORNL-Booth versus CORSOR-M for Cs release (Class 2)

Figure 3-12 Comparison of ORNL-Booth versus CORSOR-M for Ba release (Class 3)

FPT-1 Class 2 Release - Cs

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

5000 7500 10000 12500 15000 17500 20000

time (sec)

Release Fraction

ORNL Booth

data

CORSOR-M
3-21

Figure 3-13 Comparison of ORNL-Booth versus CORSOR-M for I release (Class 4)

Figure 3-14 Comparison of ORNL-Booth versus CORSOR-M for Te release (Class 5)

FPT-1 Class 5 Release - Te

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

5000 7500 10000 12500 15000 17500 20000

time (sec)

Release Fraction

ORNL Booth

data

CORSOR-M
3-22

Figure 3-15 Comparison of ORNL-Booth versus CORSOR-M for Ru release (Class 6)

Figure 3-16 Comparison of ORNL-Booth versus CORSOR-M for Mo release (Class 7)

[Note: The Mo vapor pressure was set to correspond to Cs2MoO4.]
3-23

Figure 3-17 Comparison of ORNL-Booth versus CORSOR-M for Ce release (Class 8)

Figure 3-18 Comparison of ORNL-Booth versus CORSOR-M for La release (Class 9)
3-24

Figure 3-19 Comparison of ORNL-Booth versus CORSOR-M for UO2 release (Class 10)

[Note: The UO2 scaling factor was adjusted to match observed releases. La and Ce 

releases are not expected to be greater than UO2 release, but may be less, owing to 

lower volatility.]

Figure 3-20 Comparison of ORNL-Booth versus CORSOR-M for Cd release (Class 11)
3-25

Figure 3-21 Comparison of ORNL-Booth versus CORSOR-M for Sn release (Class 12)

Comparison to ORNL VI Tests and VERCOR Tests [16]

After optimizing the ORNL-Booth fission product release parameters for the FPT-1 experiment, 

it was useful to compare the modified model to the original ORNL test data upon which the 

CORSOR-M model was developed. The following section explores the application of the 

modified ORNL-Booth release modeling to selected ORNL-VI test results and the VERCORS 

test data. The comparisons are made mainly to the Cs release observed in these experiments 

since all other releases are simply scaled to the Cs release in the Booth implementation in 

MELCOR, and these data were readily available. In the case of VERCORS 4, more data on 

release of other fission products were readily available and comparisons to these releases 

included. The MELCOR models were obtained from a Nuclear Safety Institute of the Russian 

Federation (IBRAE) MELCOR validation exercise [17] investigating the MELCOR default 

release models. The experimental data are taken from Reference [17]. These analyses were 

performed using a MELCOR model of these simple experiments. The present analyses make 

use of the modified ORNL coefficients and compare results with the MELCOR default 

CORSOR-M release model. A schematic of the VERCORS testing facility is shown in 

Figure 3-22; the general layout is similar in the ORNL VI tests. The tests examined are 

summarized in Table 3-4. The tests involved both oxidizing and reducing conditions.

FPT-1 Class 12 Release (Sn)

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

0.45

0.5

5000 10000 15000 20000

time (sec)

Release Fraction

ORNL Booth

CORSOR-M
3-26

Figure 3-22 Schematic of VERCORS Test Facility for measuring fission product release 

from small fuel samples

Table 3-4 Test conditions for selected ORNL VI tests and VERCORS tests.

Test Hydrogen Steam Max 

Temperature

ORNL VI-2 0 1.8 liter/min 2300 K

ORNL VI-3 0 1.6 liter/min 2700 K

ORNL VI-5 0.4 liter/min 0 2740 K

VERCORS 2 0.027 gm/min 1.5 gm/min 2150 K

VERCORS 4 0.012 gm/min 1.5 – 0 gm/min 2573 K

In almost all cases, the modified ORNL-Booth model produces improved signatures, as shown 

in Figure 3-23 through Figure 3-25 for the VI tests and in Figure 3-26 through Figure 3-32 for the 

VERCORS tests. 

In test VI-2 run under steam conditions, the peak temperature attained was ~2300 K. Both 

models over predicted the Cs release for this test, with the modified ORNL-Booth treatment 

performing slightly better (Figure 3-23). Test VI-3 was similar to VI-2 except that higher 

temperatures were attained. In this test, both models were closer to the data, and again the 

modified ORNL-Booth model performed somewhat better (Figure 3-24). From these two tests, it 

would seem that the release rate in the 2300 K range is still slightly over predicted for oxidizing 

conditions. Test VI-5 conducted under reducing conditions was well predicted by both models, 

as shown in Figure 3-25. Table 3-5 through Table 3-7 provides total releases predicted by 

CORSOR-M and ORNL-Booth compared with totals reported for the ORNL VI tests 2, 3 and 5.
3-27

Both VERCORS 2 and 5 were run in mixed conditions with both steam and hydrogen. In 

VERCORS 5, the steam flow was reduced to zero (reducing conditions) for the high 

temperature plateau. Test VERCORS 2, like ORNL VI–2 was performed at a lower temperature 

and produced a comparatively lower Cs release (Figure 3-26). The modified ORNL-Booth 

model captured this lower release where the CORSOR-M model did not. Test VERCORS 4 

was performed under completely reducing conditions during the release phase. In this case 

CORSOR-M under predicted the release, whereas the modified ORNL-Booth model captured 

the release behavior reasonably well.

Figure 3-23 Comparison of Cs release for modified ORNL-Booth with CORSOR-M for 

VI-2 run under steam oxidizing conditions

ORNL VI 2

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 1000 2000 3000 4000 5000 6000 7000 8000

time (sec)

Release Fraction

0

500

1000

1500

2000

2500

3000

Temperature (K)

ORNL-Booth

data

CORSOR-M

Temperature
3-28

Figure 3-24 Comparison of Cs release for modified ORNL-Booth with CORSOR-M for 

VI-3 performed under steam oxidizing conditions

Figure 3-25 Comparison of Cs release for modified ORNL-Booth with CORSOR-M for 

VI-5 performed under steam reducing conditions

ORNL VI 3

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 2000 4000 6000 8000 10000

time (sec)

Release Fraction

0

500

1000

1500

2000

2500

3000

Temperature (K)

ORNL-Booth

data

CORSOR-M

Temperature

ORNL VI 5

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 2000 4000 6000 8000 10000

time (sec)

Release Fraction

0

500

1000

1500

2000

2500

3000

Temperature (K)

ORNL-Booth

data

CORSOR-M

Temperature
3-29

Table 3-5 Release fraction from ORNL VI-2.

7

Experiment CORSOR-M ORNL-Booth

Kr - 0.98 0.92

Cs 0.67 0.98 0.92

Ba 0.18 0.003 0.002

Sr - 0.003 0.002

I 0.4 0.98 0.81

Te - 0.97 0.81

Ru - 1 x 10-7 0.006

Mo 0.86 0.06 0.42

Ce - 1 x 10-8 1.1 x 10-7

Eu - 1 x 10-5 1.1 x 10-7

U 0.003 1 x 10-5 0.001

Sb 0.68 0.04 0.93

Table 3-6 Release fraction from ORNL VI-3.

Experiment CORSOR-M ORNL-Booth

Kr 1 1 1

Cs 1 1 1

Ba 0.3 0.04 0.004

Sr 0.03 0.04 0.004

I 0.8 1 1

Te 0.99 1 0.99

Ru 0.05 10-5 0.03

Mo 0.77 0.15 0.88

Ce 0 2 x 10-6 4 x 10-7

Eu 0 0.0005 4 x 10-7

U 0 0.0005 0.003

Sb 0.99 0.2 0.93

Table 3-7 Release fraction from ORNL VI-5.

Experiment CORSOR-M ORNL-Booth

Kr 1 0.97 0.96

Cs 1 0.97 0.96

Ba 0.76 0.04 0.005

Sr 0.34 0.04 0.005

I 0.7 0.97 0.96

Te 0.82 0.95 0.96

Ru 0 10-5 0.03

Mo 0.02 0.11 0.85

Ce 0.02 3 x 10-6 4 x 10-7

Eu 0.57 0.0008 4 x 10-7

U 0 0.0008 0.003

Sb 0.18 0.19 0.89

 7 NUREG/CR-6261 provides a summary of the ORNL VI and HI test series; however, unreported release 

fractions are present for some elements. The missing data are represented with a ‘-‘ in Table 3-5.
3-30

Figure 3-26 Comparison of Cs release for modified ORNL-Booth with CORSOR-M for 

VERCORS-2

VERCORS 2

0

0.1

0.2

0.3

0.4

0.5

0.6

0 2000 4000 6000 8000 10000 12000

time (sec)

Release Fraction

0

500

1000

1500

2000

2500

Temperature (K)

ORNL-Booth

data

CORSOR-M

Temperature
3-31

Figure 3-27 Comparison of Cs release for modified ORNL-Booth with CORSOR-M for 

VERCORS-4

Figure 3-28 Comparison of Xe release for modified ORNL-Booth with CORSOR-M for 

VERCORS-4

VERCORS 4

0

0.1

0.2

0.3

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0.8

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1

0 2000 4000 6000 8000 10000 12000

time (sec)

Release Fraction

0

500

1000

1500

2000

2500

3000

Temperature (K)

ORNL-Booth

data

CORSOR-M

Temperature

VERCORS 4

0

0.1

0.2

0.3

0.4

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0.6

0.7

0.8

0.9

1

0 2000 4000 6000 8000 10000 12000

time (sec)

Release Fraction

0

500

1000

1500

2000

2500

3000

Temperature (K)

ORNL Booth

CORSOR M

Xe data

Temperature
3-32

Figure 3-29 Comparison of Iodine release for modified ORNL-Booth with CORSOR-M 

for VERCORS-4

Figure 3-30 Comparison of Te release for modified ORNL-Booth with CORSOR-M for 

VERCORS-4

VERCORS 4

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 2000 4000 6000 8000 10000 12000

time (sec)

Release Fraction

0

500

1000

1500

2000

2500

3000

Temperature

ORNL Booth

CORSOR M

I data

COR-TFU.102

VERCORS 4

0

0.1

0.2

0.3

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0.7

0.8

0.9

1

0 2000 4000 6000 8000 10000 12000

time (sec)

Release Fraction

0

500

1000

1500

2000

2500

3000

Temperature (K)

ORNL Booth

CORSOR M

Te data

Temperature
3-33

Figure 3-31 Comparison of Ba release for modified ORNL-Booth with CORSOR-M for 

VERCORS-4

Figure 3-32 Comparison of Mo release for modified ORNL-Booth with CORSOR-M for 

VERCORS-4

VERCORS 4

0.001

0.01

0.1

1

8000 8500 9000 9500 10000 10500 11000 11500 12000

time (sec)

Release Fraction

0

500

1000

1500

2000

2500

3000

Temperature (K)

ORNL Booth

CORSOR M

Ba Data

Temperature

VERCORS 4

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 2000 4000 6000 8000 10000 12000

time (sec)

Release Fraction

0

500

1000

1500

2000

2500

3000

Temperature (K)

ORNL Booth

CORSOR M

Mo data

Temperature
3-34

The use of the modified ORNL-Booth model produces significantly improved predictions for both 

the in-pile Phebus FPT-1 test as well as for the original small scale ORNL VI and French 

VERCORS tests upon which the original CORSOR and Booth models were developed. Barium 

behavior however remains somewhat problematic in that the small-scale tests generally predict 

Ba release greater than is ever observed in the in-piles tests. It is believed this is due to the fact 

that in the small-scale tests, the cladding is almost completely oxidized, whereas considerably 

less coherent conditions are encountered in the in-pile integral tests. It is conjectured that the 

Ba speciation in the small-scale tests is more volatile than that produced in the in-pile tests 

where unoxidized Zr is plentiful.

Given the overall improvement realized in application of the modified ORNL-Booth model, user 

selection of the modified ORNL-Booth model is accepted as the best practice.

3.1.1.5 Evaluation of Fission Product Deposition Modeling and Speciation

Deposition in FPT-1 Circuit (RCS Deposition)

The modified ORNL-Booth release models have been shown to produce favorable release 

signatures when examining the Phebus FPT-1 test and produce good comparisons with the 

ORNL VI and French VERCORS tests. The modifications to the vapor pressures for Cs and 

Mo, which produced favorable release behavior in FPT-1, will have an effect on the subsequent 

deposition of these species in the RCS piping. The effect is illustrated in the following two 

figures showing deposition patterns in the Phebus FPT-1 test circuit and model containment.

Figure 3-33 shows the predicted deposition distribution in the FPT-1 experiment when the 

default CORSOR-M release model was used. While the total Cs release compares reasonably 

well with the measured value, and the total Cs transported to the containment is about right, the 

distribution of Cs deposits in the heated test section above the fuel (upper plenum) and in the 

steam generator tube do not compare well with the experiment. Deposits in the steam 

generator are over predicted and deposits in the heated plenum above the fueled region are 

under predicted. In fact, deposits of Cs in the plenum were never greater than 0.1% and were 

predicted to be completely revaporized before the end of the test. Under predicting deposition 

in the hot plenum region is a big factor in the over predicting of the steam generator tube 

deposits. 

Figure 3-34 shows the Cs distribution predicted for FPT-1 when the modified ORNL-Booth 

model is used. The lower vapor pressure of the presumed Cs2MoO4 results in Cs predicted to 

be in aerosol form in the hot upper plenum region and as a result, Cs deposited in the upper 

plenum remains for the duration of the test. This together with a slightly lower total Cs release 

results in half as much predicted to be deposited in the steam generator tubes, considerably 

closer to the observed tube deposition. The amount reaching the containment remains about 

the same, which from a “release to the environment” point of view, one can observe that either 

model retains about the right amount of fission products within the simulated RCS. The 

changes in Cs deposition within the RCS could alter the decay heat distributions throughout the 

RCS, which in turn could affect revolatilization of other more volatile deposited species, such as 

CsI, which is transported in addition to the presumed dominant Cs2MoO4.
3-35

Figure 3-33 MELCOR-predicted fission product deposition in FPT-1 circuit using 

default CORSOR-M release modeling

[Note: Predicted plenum deposits for this case were less than 0.1 percent, not 

visible on this scale, and were subsequently revaporized.]

Figure 3-34 MELCOR-predicted fission product deposition in FPT-1 circuit using 

default ORNL-Booth release modeling

Cs Distribution as Fraction of Bundle Inventory: CORSOR-M (default)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

5000 7500 10000 12500 15000 17500

time [sec]

fraction of inventory

bundle

plenum

hot leg

steam generator

cold leg

containment

Total

- Plenum

- Hot Leg

- Stm Gen

- Containment

- total

FPT-1

Final

Dist

Cs Distribution as Fraction of Bundle Inventory: ORNL-Booth (modified)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

5000 7500 10000 12500 15000 17500 20000

time [sec]

fraction of inventory

bundle

plenum

hot leg

steam generator

cold leg

containment

Total

- Plenum

- Hot Leg

- Stm Gen

- Containment

- total

FPT-1

Final

Dist
3-36

Deposition within the Phebus Containment

For completeness, the deposition behavior calculated for the FPT-1 containment model is 

shown in the following figures. Shown is the total airborne aerosol mass predicted using the 

sources resulting from the ORNL-Booth release modeling. The suspended mass is normalized 

to the peak value in order to make comparison to the measured data; this normalization was

made necessary because of differences between the magnitude of mass predicted to be 

transported to the containment and the measured value. MELCOR predicted only about half of 

the suspended total mass that was measured. The discrepancy is due to not activating the Ag 

release model for the Ag/In/Cd control rods and the rhenium (Re) in the experiment 

thermocouples.

The overall depletion signature prior to the peak airborne value compares quite well. After 

reaching the maximum value however, the MELCOR predicted suspended mass depletes less 

rapidly than was actually observed. This is apparently due to MELCOR under predicting the 

particle size as shown in Figure 3-36, and consequently under predicting the gravitational 

settling component of containment deposition. Certainly the under prediction of the suspended 

mass by a factor of two also resulted in lower aerosol number concentration, perhaps 

significantly so if the mass is missing from the smaller particle size range, and this may in turn 

have resulted in slower particle agglomeration rate and therefore smaller agglomerated particle 

sizes. If so, this could explain the lower aerosol depletion rate by gravitational settling.

Diffusiophoresis is the other dominant form of aerosol deposition in the FPT-1 containment, and 

may also be under predicted; however, test data do not provide resolution in this respect. 

Under prediction of the containment depletion rate errs on the conservative side since more 

fission products remain suspended in this analysis that might be available for release to the 

environment.

Figure 3-35 Normalized aerosol depletion rate of airborne aerosol in FPT-1 containment

[Note: The under prediction of gravitational settling may be the reason for the low 

prediction of the depletion rate.]

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

10000 15000 20000 25000 30000

time [sec]

normalized airborne mass

MELCOR 1.8.5

FPT-1
3-37

Figure 3-36 Predicted and measured aerodynamic mass mean aerosol diameter in 

FPT-1 containment

[Note: Under prediction of the agglomeration rate from too low of an airborne total 

mass may be responsible for the under prediction of the mean particle size.]

New Speciation Modeling(8)

Based on the chemical analysis of fission products deposited throughout the simulated reactor 

coolant circuit and containment of the Phebus facility, significant amounts of molybdenum were 

transported. This was unexpected because molybdenum has a low vapor pressure and was not 

expected to be released in large quantities. However, the molybdenum combined with the 

cesium and formed cesium molybdate. The vapor pressure of cesium molybdate is much 

higher than molybdenum (see Figure 3-5). In light of the Phebus findings, the chemical 

speciation of cesium was updated for the SOARCA project.

Prior to the Phebus experiments, the common cesium speciation practice was to combine all 

available iodine with the required amount of cesium to form Class 16 (cesium iodide) and place 

the remaining cesium in Class 2 (cesium hydroxide). To incorporate the Phebus findings, Class 

17 has been created to accommodate cesium molybdate. The new speciation of cesium still 

combines all available iodine with cesium as cesium iodide but the remaining cesium is 

combined with molybdenum to form cesium molybdate with the exception of a small portion of 

cesium allocated to Class 2 (cesium hydroxide).

In summary, the best practice speciation manually reconfigures (i.e., through user input 

specifications) the radionuclide and decay heat classes that contain cesium, iodine, and/or

molybdenum as follows:

 8 Whereas, the recent modifications to Version 1.8.6 (Version YR) for the SOARCA project implemented the 

new ORNL-Booth fission product release model as the new default model (see Section 3.1.1.4), the associated 

reconfiguration of the radionuclide masses must be done through user input specification as summarized in this 

section.

0

1

2

3

4

5

10000 15000 20000 25000 30000

time [sec]

AMMD, 

μ

m

MELCOR 1.8.5

FPT-1
3-38

• Class 2 – Characteristic released compound is CsOH with the default inventory 

representative of the cesium in the fuel–cladding gap (5% of the initial core inventory of 

cesium)

• Class 7 – Characteristic released compound is Mo with the default inventory reduced by 

the amount allocated to Class 17

• Class 16 – Characteristic released compound is CsI with the default inventory 

representing all iodine and the necessary cesium to combine with the iodine to form CsI

• Class 17 – Characteristic released compound is Cs2MoO4 with the cesium not in the gap 

(included in Class 2) and not combined with iodine in Class 16, combined with the 

necessary molybdenum to form this class

The prescribed user defined definition of the fuel–cladding gap inventory is specified as follows 

[18]:

• Class 1 – 5% of the noble gases

• Class 2 – 100% of this class equating to the amount of cesium needed additional to the 

cesium allocated to the gap as CsI (in Class 16) to raise the total cesium in the gap to 

5% of the total core inventory

• Class 3 – 1% of the barium inventory

• Class 5 – 5% of the tellurium inventory

• Class 16 – 5% of the CsI inventory

3.1.1.6 Release of Structural Aerosols

Experimental measurements of aerosol release and transport during fuel assembly melting 

consistently show significant releases of non-radioactive species such as tin and (in PWR 

assembly experiments) Ag-In-Cd. The former is released primarily from fuel cladding (tin is an 

alloy constituent of Zircaloy); the latter is released from (PWR) control rods. MELCOR includes 

specialized models to vaporize Ag-In-Cd control materials from PWR control rods. The control 

material is subsequently tracked in the RN like other radionuclides. The control rod radionuclide 

model was used in the SOARCA project and is recommended for PWRs. However, MELCOR 

does not include a release model for tin from the Zircaloy. 

The release of tin from the Zircaloy was included using the in-vessel radionuclide release model 

of “non-fuel” materials in the Core (COR) Package, which is enabled through user input. 

However, the in-vessel structural release model is not a mechanistic treatment of diffusion and 

release of volatile constituents from core structures. Rather, the model provides a simple 

framework for extending the CORSOR models for radionuclide release to examine the effects of 

the additional mass associated with aerosols generated during the oxidation and melting of 

cladding and control rod guide tubes. The tin structural release model was tuned to match 

observations from experiments. 

First, the structural aerosol release model requires specification of non-fuel materials 

recognized by the COR Package to an RN class. The structural aerosols are specified from the 
3-39

list of materials tracked in the COR Package. The structural materials available for conversion 

include the following:

• Unoxidized zirconium

• Zirconium oxide

• Unoxidized steel associated with control rods (blades)

• Oxidized steel associated with control rods (blades)

• Control rod poison

The fraction of the total mass of each of these materials that is available as a potential aerosol 

source is defined. In principle, this fraction should represent the mass fraction of volatile 

constituents in the material. This information is specified on input records for card RNCRCLxx.

The release rate associated with core structure materials is estimated using the active RN 

fission product release model (i.e., the user-selected version of CORSOR.) Scalar multipliers 

are applied to CORSOR to represent differences in the release rates associated with migration 

through a UO2 matrix versus Zircaloy, steel, or other core structural materials. The CORSOR 

release rate multipliers are specified via sensitivity coefficient 7100. The temperature used to 

calculate the release rate of a particular material is tied to the corresponding structure 

temperature in the COR Package (i.e., zirconium to fuel cladding, and steel plus control poison 

to non-supporting structure.) 

Modeling the evaporation and release of control poison in BWRs is not a concern because the 

vapor pressure of boron carbide is very high (sublimation occurs at temperatures above 

2500°C). However, release of non-radioactive vapor and aerosols due to chemical reactions 

between B4C and steam is modeled by the B4C oxidation model in the COR Package. The 

mass of vapor species generated by these reactions is transferred to the CVH Package for 

transport; the mass of condensed aerosols (primarily B2O3) is transferred to the RN Package for 

transport. 

A significant limitation of the non-fuel release modeling framework in MELCOR is that the mass 

of materials represented in the COR Package is not decremented to reflect the amount sourced 

to the RN Package. Further, there is no numerical limit to the amount of structural aerosol that 

can be generated for a given core material. Therefore, material fractions and CORSOR release 

rate multipliers must be defined carefully to prevent non-physical results. 

No direct experimental information on release rates of core structural materials is available to 

aid in selecting appropriate multipliers for the non-fuel release option in MELCOR. Realistic 

values for the volatile mass fraction of tin in Zircaloy were used, namely 1.45 wt/%. Based on 

the observations made by R. Gauntt [60], the release rate of tin from unoxidized Zr was 

assumed to be a factor of ten smaller than the rate from ZrO2 (i.e., SC7100 = 0.1 for Zr and 1.0 

for ZrO2). 

The non-fuel release model is tied to the particular model for fission product release selected for 

a calculation. Reasonable values of the integral release were achieved using unmodified 

CORSOR release coefficients (i.e., the same as fission product release from fuel). Depending 

on the particular release expression used, the total quantity of structural tin released ranges

from 18 to 47 percent of the total available as an alloy agent in Zircaloy fuel cladding. It is 

recommended to assign the released structural tin aerosols a separate RN class dedicated for 

structural aerosols.
3-40

3.1.1.7 Vessel Lower Head Failure and Debris Ejection

The base case approach for modeling lower head failure (LHF) of the vessel and debris ejection 

includes some special non-default modeling options in MELCOR. A schematic of MELCOR’s 

lower head heat transfer model is shown in Figure 3-37. The solid debris convects to the lower 

head wall. Solid debris in the lower plenum is assumed to be wettable by lower plenum water, if 

present. Earlier versions of MELCOR included a one-dimensional model for the CCFL on water 

access to the debris. The one-dimensional CCFL greatly restricted the debris heat transfer and 

was highly susceptible to the lower plenum core cell nodalization (see core plate failure 

discussion). The new heat transfer model recognizes the potential for multi-dimensional flow 

patterns in the lower plenum without a one-dimensional CCFL restriction. Hence, the film and 

nucleate boiling debris bed-to-water heat transfer correlations are applicable for debris 

submerged under water.

Any molten debris will convect to the lower head using the molten debris bed heat transfer 

correlations. A separate lower temperature metallic molten pool (MP1) can exist as well as a 

higher melting temperature mixed oxide molten pool (MP2). There is two-dimensional radial 

and azimuthal conduction through the vessel wall. On the outer surface of the vessel, there is 

heat transfer to the flooded cavity using inverted cylindrical nucleate boiling correlations.

The other key modeling options include the method for modeling LHF and assumptions 

regarding the resultant discharge of debris. Penetration failure is not modeled as a mechanism 

for vessel failure. Rather, only gross creep rupture of the lower head is modeled. In the SNL 

LHF tests [19], gross creep rupture of the lower head was measured to be the most likely 

mechanism for vessel failure. In addition, past observations using MELCOR’s penetration 

model suggest that it lacks sufficient spatial resolution to adequately model the 

multi-dimensional heat transfer effects (i.e., it is a relatively simple lumped capacitance model). 

The lower head creep rupture model uses the code’s default settings. A Larson-Miller 

parameter is calculated using a one-dimensional temperature profile through the lower head. A 

cumulative strain is calculated using a lifetime rule and failure occurs with an 18% strain. Upon 

vessel failure, molten and solid debris are assumed to discharge simultaneously. 

Figure 3-37 MELCOR lower head nodalization
3-41

3.1.1.8 Ex-vessel Phenomena - CCI and Hydrogen Combustion

Following vessel failure, molten core-concrete interactions (MCCI) will take place in the reactor 

cavity of the containment. If there is a pool of water in the reactor cavity, then there is 

simultaneous debris heat transfer to the overlying pool of water and into the concrete. If there is 

inadequate heat transfer to cool the debris, the debris will ablate the concrete and release 

combustible gases. However, if the debris is cooled below the concrete ablation temperature, 

then there is no combustible gas production and, therefore, no accompanying pressure loading 

and combustion potential. 

The simplified one-dimensional geometric configuration of the debris underestimates heat fluxes 

observed in the Melt Attack and Coolability Experiments (MACE) program [20] when default 

MELCOR values are used. In particular, the MACE tests showed cracking of the crust with 

water ingression and multi-dimensional effects that greatly enhanced the amount of cooling 

when water was present, which was not considered within the MELCOR MCCI framework. As 

shown in Figure 3-38, the nominal heat flux calculated by MELCOR, represented by the 1*K

curve, fails to agree with the MACE heat flux data, ranging between 1000-5000 kW/m2

. By 

enhancing conductivity by 10 to 100 times the original MELCOR calculated conductivity (K),

improvement in agreement with the MACE data was observed. Modifications to the MELCOR 

code were performed giving users the ability to enhance the default MELCOR heat transfer 

calculations by specifying constant multipliers to the debris conductivity and surface heat flux 

(i.e., a method to reflect cracks and multi-dimensional effects). A 10x surface heat flux multiplier 

and 5x debris conductivity multiplier for oxide and metallic debris were used to enhance the 

ex-vessel debris-to-water heat transfer model and are the recommended best practice. The 

enhanced heat transfer increased the ex-vessel debris cooling while water was present which 

reduces the rate of concrete ablation. The concrete ablation rate was unchanged in cases

where water was not present.

The default MELCOR combustion model was used by enabling the BUR Package, which is 

performed through user input. Special attention was made within the user input to include 

horizontal and vertical propagation of burns and the time delay for the flame front to span the 

width of the control volume. MELCOR does not include models for detonation. Hence, all burns 

are subsonic deflagrations with appropriate models for steam dilution, hydrogen and oxygen 

concentrations, and propagation to adjacent locations. Finally, in cases without an obvious 

ignition source, sensitivity calculations could be performed that delay combustion to simulate the 

occurrence of a spontaneous ignition source in the containment (e.g., debris ejection at vessel 

failure, reenergizing equipment, etc.). Delaying the ignition event would allow hydrogen and 

other combustion products to accumulate leading to more energetic burns. A user can model 

delayed ignition by adjusting the concentrations capable of supporting a deflagration event.

The associated structural damage from a hydrogen burn is simulated through user-specified 

failure paths. In SOARCA, plant walk-downs and separate structural assessments were 

performed for locations where hydrogen burns were considered possible. For the Surry 

calculations, this included the primary containment and the surrounding buildings for the 

interfacing systems loss of coolant accident (ISLOCA) scenario (i.e., the Safeguards Area, 

Containment Spray Pump Area, and Main Steam Valve House). Hydrogen combustion was not 

observed in the Surry analyses. However, the BWR reactor building experienced a combustion

failure. Leakage and failure flow paths are specified that model failure of blowout panels, walls, 

ceilings, doors, and other vulnerable locations. For example, the resulting damage to the 

reactor building in the Peach Bottom SOARCA station blackout calculations due to a hydrogen 

burn was comparable to what was observed in the Fukushima accident.
3-42

Figure 3-38 Heat transfer from an overlying water pool to an ex-vessel debris bed.

3.1.2 Pressurized Water Reactor Best Practices

Modeling practices discussed in this section are typically applicable to PWR MELCOR models

and were applied to the SOARCA Surry model. The provided best practices are generally 

related to unique hardware found in PWRs but not BWRs.

3.1.2.1 Pump Seal Leakage and Blowout

Based on insights from the Surry Individual Plant Examination, which used the Rhodes 

probabilistic model for seal leakage, a simple model has been incorporated into MELCOR

analyses using user input. The key attributes of the model are implemented as follows. Upon a 

loss of seal cooling in a station blackout sequence, the seals will leak at 21 gpm. In an 

unmitigated station blackout scenario, the fluid exiting the loop seal will approach saturation 

conditions at approximately 2 hours and the seal leakage flow will increase to 182 gpm per 

reactor coolant pump. The seal leakage values are based on normal operating conditions. The 

flow rate will change appreciably as a function of pressure, subcooling, and steam quality.

Corium Crust to Water Heat Flux

0

500

1000

1500

2000

2500

3000

3500

4000

4500

5000

0 0.1 0.2 0.3 0.4 0.5

Time (hr)

Heat Flux (kW/m**2)

1*K

2*K

5*K

10*K

20*K

50*K

100*K

Range observed

In MACE Tests

d d

concrete

corium

water
3-43

3.1.2.2 Loop Seal Clearing and Effects on Progression

The NRC has a separate research program examining thermally induced steam generator tube 

rupture (TISGTR) [65]. Detailed computational analyses are being performed to investigate the 

timing and sensitivity of high temperature natural circulation tube failure versus failures of other 

components in the RCS during severe accident natural circulation conditions. The clearing of 

the loop seal was identified as a key event that could increase the likelihood of tube failure 

under certain circumstances. Furthermore, NUREG-1570 (i.e. “Risk Assessment of Severe 

Accident-Induced Steam Generator Tube Rupture”) previously assessed the potential for tube 

failure in high-pressure station blackout conditions.

MELCOR includes basic thermal-hydraulic models for loop seal clearing and the prediction of 

thermal failure of steam generator tubes. However, the thermal gradients and flow behavior is 

extremely complex in the RCS during natural circulation conditions. The base case response 

utilizes the MELCOR models for natural circulation (i.e., discussed further below), loop seal 

clearing, and user specified RCS component thermally induced failure models. Sensitivity 

calculations were also performed that assume steam generator tube failure before hot leg 

nozzle failure (i.e., the first creep rupture location calculated in the Surry station blackout

calculations).

3.1.2.3 RCS Natural Circulation Treatment

Natural circulation is important in severe accident sequences because circulating steam from 

the core to upper reactor internals, the hot leg, and the steam generators (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 piping connected to 

the vessel could progress to a thermal stress (i.e., creep rupture) failure of the primary pressure 

boundary and a subsequent depressurization. As shown in Figure 3-39 [21], three natural 

circulation flow patterns can be expected during a severe accident for a Westinghouse PWR:

(1) in-vessel circulation, (2) countercurrent hot leg flow, and (3) loop natural circulation. For 

high pressure accidents that do not include RCS pipe breaks (e.g., a station blackout), whole 

loop, single-phase natural circulation flow (i.e., the left hand side of Figure 3-39) is not expected 

during the core degradation phase of the accident. Consequently, the prediction of the first two 

natural circulation flow patterns is critical for predicting severe accident progression. The first 

two natural circulation flow patterns have been studied (a) experimentally in the 1/7th-scale 

natural circulation test program by Westinghouse Corporation for the Electric Power Research 

Institute (EPRI) [22], [23], (b) computationally using the FLUENT computational fluid dynamics

(CFD) computer program [24], [25], and (c) analytically using SCDAP/RELAP5 [54]. 

Subsequently, MELCOR was used to model the 1/7th-scale natural circulation tests [26].

More recently, the NRC has continued improving natural circulation modeling as part of its

steam generator tube integrity program [27], [28]. 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 [26], which closely followed the previous work performed by Bayless [21]. 

The natural MELCOR modeling approach in the Surry model has been recently updated to 

incorporate some of the modeling advances used by Fletcher using SCDAP/RELAP5 [27]. The 

key features of the MELCOR natural circulation models, which were adapted from the recent 

SCDAP/RELAP5 work, are the following:
3-44

• 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

• Explicit modeling of all internal vessel structures for heat transfer

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 of oxide layer thickness

o Variable steel emissivities in the core as a function of temperature

The hot leg is divided in half to represent the counter-current natural circulation flow. 

The flow rate is matched to a Froude Number correlation from the FLUENT CFD 

analysis [27] for the Westinghouse SG,

Q = C

D 

[ g (Δρ / ρ ) D

5 

] 

1/2

Equation 12

where,

Q is the volumetric flow rate in the hot leg,

CD = 0.12 (from FLUENT CFD calculations),

g is the acceleration due to gravity,

ρ is the average fluid density,

Δρ is the density difference between the hot and cold fluid streams, and

D is the pipe diameter

• Steam generator tube-to-hot leg flow ratio tuned results from the FLUENT CFD analysis 

[27]

o Tube mass flow rate/hot leg mass flow rate (M-ratio) = 2

• Explicit modeling of all key heat transfer processes in the hot leg and the steam 

generator

o Augmented convective heat transfer in hot leg based on FLUENT turbulence 

evaluations

o Gas-to-structure radiative exchange in the hot leg and steam generator

o Ambient heat loss through the piping and insulation

• Steam generator mixing fractions based on FLUENT CFD analysis [27]
3-45

o Inlet plenum subdivided into 3 regions for hot, mixed, and cold regions

o Flow ratio into hot tubes tuned to a 0.85/0.15 split

o Flow ratio into cold leg piping tuned to a 0.85/0.15 split 

o Flow divided in the SG tubes in a 0.41/0.59 tube split(9)

• Individual modeling of relief valves

o When the valves are lumped, it creates a very large flow that non-physically 

disrupts natural circulation flow patterns and the timing of the valve openings

• Creep rupture modeling

o Hot leg nozzle carbon safe zone region

o Hot leg piping

o Surge line

o Steam generator inlet tubes

The MELCOR hot leg and SG nodalization for Loop A is shown in Figure 3-40. Control volume 

reactor system codes like MELCOR or SCDAP/RELAP5 have limitations in modeling buoyancy 

plumes associated with natural circulation flow. Hence, the MELCOR system model analyses 

are performed by incorporating flow buoyancy or drag adjustments to the hot leg circulation rate 

to achieve the target value for hot leg discharge coefficient. The drag coefficient was formulated 

based on an experimental correlation for flow through horizontal ducts connecting two tanks 

containing fluids of different densities. The special natural circulation flow paths described 

above are shown in red in the figure. The natural circulation control logic identifies single 

potential single-phase natural circulation conditions and activates the special flow paths to 

achieve the conditions described above.

Finally, it should be noted that the recent work of Fletcher [27] for the NRC steam generator 

tube integrity program revealed a sensitivity of tube failures to the hot and cold tube split, the 

tubes receiving the peak plume temperatures, and the highly refined axial nodalization through 

the tube sheet and into the steam generator. These specific aspects of the modeling 

specifically addressed the potential for a TISGTR, which were not incorporated in the MELCOR 

models.(10)

 To evaluate this potential consequence while acknowledging the potential limitations 

in the MELCOR model and/or vulnerabilities or defects in the plant tubes, the SOARCA project

will perform sequences where TISGTRs were specified to occur prior to other RCS natural 

circulation failures in follow-up SOARCA uncertainty analysis studies.

 9 It was not practical to represent the 41%/59% hot/cold split of the SG tube regions in the MELCOR model 

due to the complications of a single model nodalization for all conditions. A 50%/50% tube split was used. 10 Unlike SCDAP/RELAP5, MELCOR cannot be renodalized at the start of the natural circulation phase with 

a more detailed model due to the control volume and flow path being fixed upon initiation of the simulation. 

Consequently, the MELCOR model must calculate the early two-phase thermal-hydraulic transient, the natural 

circulation phase, the post creep rupture blowdown, the accumulator reflood of the degraded core, and the final 

boil-off and core degradation to vessel failure. It was not practical to use a highly detailed steam generator 

nodalization for the scope of a MELCOR source term calculation.
3-46

Figure 3-39 In-vessel, full-loop, and hot leg natural circulation flow patterns in a PWR 

severe accident
3-47

Figure 3-40 MELCOR hot leg and steam generator nodalization including the special 

natural circulation flow paths

3.1.2.4 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. As discussed above, the hot relocated core materials will 

boil away the water in the lower plenum, which will lead to vessel LHF. The MELCOR 

representation of the Westinghouse core plate assembly includes a separate representation of 

the various supporting structures. At the lowest level is the bottom support casting. The bottom 

support casting is part of the integrated core barrel structure. The mass of the core is 

RPV

CV209

CV211

CV212

CV215

CV213

CV217

CV216

CV214

CV225

CV226

FL233FL294

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.
3-48

transmitted from the core plate via support columns that span the gap to the core support 

forging. Within the gap is a flow mixer plate (see Figure 3-41).

The MELCOR lower head support nodalization is shown in the left side of Figure 3-42. The 

weight of the core material mass is transmitted through the columns to the bottom support 

casting. The local thermal-mechanical failure of the lower core plate, the flow mixer plate, and 

the lower support forging are calculated internally by MELCOR using Roark’s engineering stress 

formulas [64]. The failure is based on exceeding the yield stress at the local material 

temperature conditions. After the core plate fails, it is assumed that the debris falls past the 

columns but is temporarily supported by the flow mixer plate. However, since the flow mixer 

plate is relatively thin, the hot debris will quickly fail the plate (i.e., again according to the Roark 

stress formulas). The debris subsequently falls to the lower support forging, which is very thick 

but eventually fails. The sequential failure of the supporting structures is affected by vessel 

water level, which is also exposed to the sequentially relocating debris. Once the lower support 

forging fails, the debris falls onto the lower head. 

Fully molten materials will relocate through the structures until freezing on supporting structures 

or reaching the lower head. Due to the high melting temperature of U-Zr-O eutectic core debris 

material (i.e., assumed to melt at 2800 K), most of the fuel and cladding debris will be frozen 

during the core support structure failures. However, some unoxidized Zr or control material may 

have enough superheat to relocate through the structures onto the lower head.
3-49

Figure 3-41 Westinghouse PWR reactor vessel internals
3-50

Figure 3-42 MELCOR Westinghouse lower vessel nodalization

CV110

Lower 

Plenum

CV710

CV720

CV730

CV740

CV750

CV701

CV712CV722CV732

CV742

CV752

CV714CV724CV734

CV744

CV754

CV716CV726CV736

CV746

CV756

CV718CV728CV738

CV748

CV758

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

FL701Bypass

FL707

Westinghouse 3-Loop PWR

CVH/FL Nodalization

FL705

FL704

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 m

4.5273 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 Cells

Lower Head 

Nodalization

Flow Path Opens 

at Shroud Failure

Flow Path 

Controlled by 

COR Blockage 

Model

Legend
3-51

3.1.2.5 Interfacing Systems Loss of Coolant Accident

The ISLOCA scenario analyzed in the SOARCA project identified several modeling and 

phenomenological aspects that were important to the timing and magnitude of radionuclide 

releases to the environment. While some of the aspects are probably specific to only the Surry 

units, some may be applicable to modeling an ISLOCA at any plant.

The ISLOCA scenario (in a PWR) involves the pressurizing of low head safety injection (LHSI) 

piping outside of containment to normal reactor operating pressure. The pressurizing results 

from the failure of two serial check valves. The piping outside of containment is not rated to 

normal reactor pressure and so is assumed to rupture causing a loss-of-coolant accident. The 

rupture is not initially isolable due to the inability of isolation valve(s) to close against the 

blowdown. The rupture is not isolable later due to the motors on the isolation valves being 

flooded. The refueling water storage tank (RWST) is eventually exhausted with its contents 

having been pumped into auxiliary buildings outside of containment where the pipe rupture 

occurred. A key assumption specific to the scope of this analysis in SOARCA was that no 

makeup water was provided to the RWST. The reactor boils dry and the core melts releasing 

radionuclides through the failed injection piping to the auxiliary buildings. Hydrogen from core 

oxidation propagates through the piping to the auxiliary buildings and deflagrates failing the 

building boundaries. Radionuclides release through the failed buildings to the environment 

bypassing containment. 

Several aspects of the LHSI piping are important to represent in a computer model (e.g., a 

MELCOR model) used to simulate an ISLOCA, including:

• The size of the flow limiting venturi in the piping (and the choked flow through the 

venturi)

• The length, size, and schedule of the piping

• The fittings (elbows, tees, etc.) in the piping

• The extent to which the piping is insulated

• The section(s) of the piping susceptible to overpressure rupture, i.e., the sections of 

piping outside of containment that are not rated to normal reactor operating pressure

• The orientation of the piping where it connects to a cold leg of the reactor (possibly 

important with respect to eventually ending leakage through the piping, for example, by 

operating the residual heat removal system residual heat removal system (RHR) at midloop water level) 

The interconnected buildings into which the RCS would depressurize given an ISLOCA are 

important to include in a simulation. At Surry these buildings are the Safeguards Area, 

Containment Spray Pump Area, and Main Steam Valve House. In the ISLOCA scenario 

postulated for Surry, the RCS depressurizes into the Safeguards Area and fission products are 

released to the environment through all three buildings.

As the buildings into which the RCS would depressurize in an ISLOCA have recognized 

contamination potential, they are exhausted by a filtered safety-related ventilation system given 

a loss of coolant accident (LOCA). On a safety injection signal, the system is automatically 
3-52

started. The system has particulate and iodine filtration and dual parallel redundant fans. The 

fans exhaust to a stack. This ventilation system is important to model in an ISLOCA simulation 

as evidenced in the SOARCA simulation where substantial amounts of radionuclide aerosols 

were captured in the exhaust ventilation filters.

Several phenomena showed to be important in the SOARCA Surry ISLOCA modeling and the 

best practice suggestion is that they all be addressed in any ISLOCA analysis. The phenomena 

and their impact are discussed below.

Turbulent deposition and impaction

The deposition of aerosolized radionuclides in the LHSI piping by means of turbulent deposition 

and inertial impaction showed to be very meaningful in the SOARCA Surry ISLOCA analysis. 

Turbulent deposition results when velocity fluctuations normal to the generalized direction of a 

turbulent flow push particles carried by the flow to the wall. Inertial impaction results at 

geometrical irregularities, an elbow for example, where the inertia of a particle causes it to stray 

from flow streamlines and collide with the wall. Turbulent deposition and impaction modeling 

was added to the MELCOR code to address these phenomena in the SOARCA Surry ISLOCA 

analysis. The modeling was benchmarked against the Light Water Reactor Aerosol 

Containment Experiment (LACE) series of tests [58].

Resuspension

As aerosol deposition in LHSI piping progressed, the deposits formed might be susceptible to 

breaking loose from the pipe wall given the high velocities that would exist in the piping. 

Consideration needs to be given in an ISLOCA analysis to the associated resuspension 

potential here. In the SOARCA analysis the assumption was made that aerosols, once 

deposited, were no longer in a form that could be readily aerosolized, i.e., the deposits might 

break loose and carry into the auxiliary buildings but would not become aerosolized and so 

would not disperse to the environment.

Revaporization

The radionuclide deposits that built up in the LHSI piping in the SOARCA simulation generated 

significant heat from fission product decay. The heat generation was significant enough to raise 

the temperature of the deposits to the point where the more volatile radionuclide classes, CsI for 

example, revaporized. Upon revaporization, the radionuclides once again transported in the 

piping and escaped through the pipe rupture into the auxiliary buildings.

Gamma heating

The thickness of the LHSI piping walls at Surry are such that as little as 33% of the gamma 

radiation emitted by radionuclide deposits in the piping would be absorbed in the steel walls. 

This proved important with respect to the revaporization of deposits in the SOARCA analysis. 

Pool scrubbing and cooling

The pipe rupture in the SOARCA Surry ISLOCA scenario was determined to be submerged 

during core degradation as was a significant length of the LHSI piping. Meaningful scrubbing of 

aerosols results as the carrier gas bubbles up through the submerging pool. The submerged 

piping was kept cool by the pool which suppressed the revaporization of substantial radionuclide 

deposits that formed in the piping. 

Hydrogen deflagration

The hydrogen produced and released to the auxiliary buildings in an ISLOCA would have the 

potential to destroy the buildings if it were to burn suddenly. This was judged to happen in the 

SOARCA simulation in the form of grossly opening the roof of a mostly subterranean building.
3-53

A few actions that the operators would perform by existing emergency procedure would be 

critical to delaying the onset of core damage in an ISLOCA. At Surry the actions would be the 

following:

• Stopping LHSI Pump A (estimated at 6 min and 17 sec)

• Stopping LHSI Pump B (estimated at 15 min and 44 sec)

• Isolating the LHSI pump suctions from the refueling water storage tank (estimated at 16 

min and 18 sec)

• Stopping and or throttling high head safety injection (HHSI) pumps to conserve RWST 

inventory

With these actions scripted in the analysis, MELCOR predicts the onset of core damage at 12 hr 

and 49 min. Without these actions accomplished, MELCOR predicts the onset of core damage 

dramatically earlier. It is imperative to represent operator actions in ISLOCA modeling.

There are several occurrences to watch for in an ISLOCA simulation as identified in the 

SOARCA Surry ISLOCA analysis, including:

• Equipment flooding, e.g., isolation valve, LHSI pump, and HHSI pump flooding

• Hydrogen deflagrations energetic enough to fail building boundaries

• Excessive particulate loading on ventilation filters considering aerosols originating from;

o Core degradation, e.g., radionuclides released from the fuel and tin released

from cladding

o Core-concrete interactions (concrete dust)

o Fires in the auxiliary buildings (smoke and particulate)

• Exhaust ventilation fan automatic shutdown because of excessive pressure drop across 

particulate-loaded filters

• Exhaust ventilation filter over-temperature failure

• Problematic heat generation from fission product decay in radionuclide-loaded exhaust 

ventilation filters

• Excessive localized aerosol deposition in the LHSI piping from turbulent deposition and 

inertial impaction such that the piping actually indicates to be blocked by the deposits

The ultimate operator action anticipated to end an ISLOCA at Surry is entering RHR with reactor 

water level stabilized at the level consistent with mid-loop operation, i.e., aligning and operating 

RHR in recirculation mode. The mid-loop level is attainable because the LHSI piping connects 

to the top of the cold legs. With the RCS subcooled at atmospheric pressure, leakage through 

the broken LHSI piping would stop with level below the top of the cold legs. This condition 
3-54

should be attainable with managed HHSI and successful RHR entry. It is important to realize 

that the availability of this strategy is somewhat unique to Surry given that the RHR systems are 

entirely separate from the LHSI systems in the Surry units. This is not especially typical of US 

PWRs. It is also important to realize that RHR entry at Surry before exhaustion of the RWST 

would be tenuous and that the net positive suction head margin at the inlets to the RHR pumps 

might be only minimally satisfactory. The feasibility of ending an ISLOCA through the use of 

RHR needs to be investigated in detail in any ISLOCA analysis. 

3.1.3 Boiling Water Reactor Best Practices

Modeling practices discussed in this section are typically applicable to boiling water reactor 

MELCOR models and were applied the SOARCA Peach Bottom model. The provided best 

practices are generally related to unique hardware found in BWRs but not PWRs.

3.1.3.1 Debris Spreading on the Drywell Floor of a BWR Mark I Containment

The floor of a Mark I containment is divided into three distinct regions for the purposes of 

modeling molten-core/concrete interactions, as illustrated in Figure 3-43. The first region (which 

receives core debris exiting the reactor vessel) corresponds to the reactor pedestal and sump 

floor areas (CAV 0). Debris that accumulates in the pedestal can flow out into the second 

region (through an open doorway in the pedestal wall), corresponding to a 90° sector of the 

annular portion of the drywell floor (CAV 1). If sufficient debris accumulates in this region, it can 

spread further into the third region, which represents the remaining portion of the drywell floor 

(CAV 2). Debris within each region is assumed to be fully-mixed, but has a distinct temperature 

and composition from neighboring regions.

Two features of debris relocation among the three regions are modeled by user-defined 

controls. The first models debris overflow from one region to another. The second manages 

debris spreading across the effective radius of the regions outside the reactor pedestal (CAV 1 

and 2). User specified control functions monitor debris elevation and temperature within each 

region, both of which must satisfy user-defined threshold values for debris to move from one 

region to its neighbor. More specifically, when debris in a cavity is at or above the liquidus 

temperature of concrete, all material that exceeds a predefined elevation above the floor/debris 

surface in the adjoining cavity is relocated (6 inches for CAV 0 to CAV 1, and 4 inches for CAV 

1 to CAV 2). When debris in a cavity is at or below the solidus temperature of concrete, no flow 

is permitted. Between these two debris temperatures, restricted debris flow is permitted by 

increasing the required elevation difference in debris between the two cavities (more debris 

‘head’ required to flow).

Debris entering CAV 1 and CAV 2 is not immediately permitted to cover the entire surface area 

of the cavity floor. The maximum allowable debris spreading radius is defined as a function of 

time using control function logic. If the cavity debris temperature is at or above the liquidus

temperature, a maximum transit velocity (corresponding to the shortest transit time) is applied to 

determine the remaining distance between the debris front and the containment liner. Traveling 

at the maximum velocity, the transit times to contact the containment liner are specified as

10 minutes for CAV 1 and 30 minutes for CAV 2. When the debris temperature is at or below 

the concrete solidus, the debris front is assumed to be frozen. A linear interpolation is 

performed to determine the debris front velocity at temperatures between these two values(11)

. 

The specified transit times coincide with the consensus that molten core migrates to the liner if 

 11 The debris spreading model compares debris temperatures to the liquidus and solidus temperatures of 

concrete because MELCOR does not currently allow user access to the debris liquidus and solidus temperatures.
3-55

the bulk temperature permits relocation. This assumption was made concerning a dry cavity 

and may not be appropriate for a flooded cavity where energy extraction from the debris leading 

edge and surface become more relevant.

Figure 3-43 Discrete regions of the drywell floor to represent debris spreading

3.1.3.2 Traversing In-core Probe Guide Tube Leakage

The traversing in-core probe (TIP) system performs in-core instrumentation calibration of the 

local power range monitors (LPRMs). Three fissile material probes are normally stored in a 

shield chamber located outside the containment pressure boundary. When operating, each of 

the probes is driven by a steel cable from the shield chamber through a guide tube to an 

indexing unit. The indexing units serve to direct the TIP probe into various exit tubes connected 

to the LPRMs in the core to perform calibrations. Calibration exercises take approximately 1 

hour and are performed once every four months. The conditional probability of a containment 

bypass due to a station blackout (SBO) coinciding with TIP calibration was precluded by the 

truncation limit for event selection.

A sensitivity calculation was performed to evaluate the effect of unisolated TIP guide tubes on 

accident progression and radionuclide source term for the long term station blackout accident 

sequence. The guide tubes were modeled as a single-lumped control volume to model the 

hydrodynamic volume while flow losses were captured by flow paths, a combined length of 

approx. 150 feet, connected at each end. Heat transfer between flowing gas and the guide 

tube(s) was accounted for; and the tube internal wall was modeled to represent an aerosol 

deposition surface [62]. The diameter of a TIP probe is approximately 0.211 inches. The 

internal diameter of the TIP guide tube is 0.280 inches and the diameter of the drive cable is 

0.258 inches. Therefore, the available cross-sectional area for flow through an operating guide 

is small (0.009 in2 per tube if the probe is inserted and 0.06 in2 if the probe is withdrawn into its 

CAV 0

CAV 2

CAV 1

6.706 m 3.086 m

0.914 m
3-56

shield chamber). When the probes are withdrawn, an alternating current actuated ball valve, 

located on the guide tube, isolates the system. The globe valve functions only when the TIP 

probe is fully withdrawn because the drive cable for the in-core probe runs through the valve 

body. If the probe cannot be retracted, operators can explosively close a direct current actuated 

squib shear valve to shear the drive cable and seal the guide tube.

3.2 MELCOR Code Enhancements for the SOARCA Project

At the start of the SOARCA project, three code development activities were identified to 

enhance MELCOR. Each of these activities had previously been included manually in the best 

practices approach for analysis, through (a) changes to default setting, (b) user-specified control 

logic, or (c) the addition of user-specified filter models, respectively. The incorporation of these 

items as defaults or code models simplified their implementation. Section 3.2.1 summarizes the 

new defaults that had been previously specified as best practice settings in Appendix A. The 

fuel collapse model described in Section 3.1.1.1 was added as a new modeling option. 

Previously, the user specification of control functions for the model required thousands of lines 

of input. Finally, a vapor scrubbing model was added, which is described in Section 3.2.3. 

Previously, user-specified vapor filters were used to scrub fission product vapors exiting BWR 

spargers. 

3.2.1 Updated MELCOR Defaults to Reflect Current Best-estimate Modeling 

Practices

In support of SOARCA MELCOR calculations using Version 1.8.6 and as the new standard for 

Version 2.1, some default values were updated based on best-estimate modeling practices. 

The values are summarized in Appendix B. The new defaults reflect long term practices to 

better model severe accident phenomena, improve numerical robustness, or activate newer 

models. In addition, the standard modeling practices include modeling cesium molybdate using 

RN package Class 17. Previously, Class 17 did not have physical parameters representative of 

any specific substance. Consequently, all physical properties were added as defaults to 

facilitate modeling cesium molybdate without additional user input. The updated ORNL-Booth 

fission product release model and updated modeling parameters are described in 

Section 3.1.1.4. Other updates to the default MELCOR settings are summarized in Appendix B. 

A new “default” record permits the user to specify whether the default parameters correspond to 

prior SOARCA or post SOARCA updated parameters by specifying either ”1.8.6” or “2.1”, 

respectively.

3.2.2 Add a Simplified Thermo-mechanical Fuel Collapse Model

As described in Section 3.1.1.1 under the description of the fuel degradation and relocation 

treatment, a simple parametric model was developed to simulate thermal-mechanical collapse 

of fuel rods only supported by highly oxidized Zircaloy shells at high temperatures. Previously, 

the implementation of the control system logic to perform these calculations required several 

thousands of lines of input. The new model was coded into the MELCOR code to eliminate the 

burden of creating a file of user-specified control logic. The model is activated through a simple 

user directive that identifies the appropriate lifetime failure function table. The lifetime failure 

function table specifies the time remaining to collapse versus the local oxidized cladding 

temperature. The logic is only implemented once the unoxidized cladding thickness drops 

below 0.1 mm, by default, or a user specified thickness.
3-57

3.2.3 Fission Product Vapor Scrubbing with Aerosol Scrubbing

In previous calculations, it was observed that the Suppression Pool Aerosol Removal Code 

(SPARC) fission product scrubbing model in MELCOR would not recognize fission product 

vapors except elemental iodine. The SPARC model automatically scrubbed all aerosols but 

elemental iodine vapor, and all fission product vapors except elemental iodine would flow 

through the pool without any retention. However, the temperature of the carrier gas and fission 

products flowing to pool spargers could have a significant vapor pressure and a mixture of 

vapors and aerosols should be considered for retention. This lack of retention was particularly 

significant for cesium iodide vapor, which can have a relatively high vapor pressure when 

discharged into a BWR sparger deep within the wetwell pool. The model applied in SOARCA 

was updated to calculate scrubbing based on the same parameters used for elemental iodine, 

and now accounts for the non-condensable fraction, bubble size, discharge gas temperature, 

pool subcooling, and pool depth.

4-1

4. MELCOR CODE DEVELOPMENT AND VALIDATION BASIS

MELCOR has been under continuous development by the NRC and SNL where software quality 

assurance (SQA) is an integral part of the development process. The MELCOR SQA program 

is adapted from two internationally recognized standards, CMMI and ISO 9001. These 

standards provide elements of traceability, repeatability, visibility, accountability, roles and 

responsibilities, and objective evaluation. The MELCOR SQA program focuses on reducing 

code error, improving documentation of all processes, and continuous integration of procedures 

into daily work processes. An essential part of SQA is proper validation of physical models 

encoded into analytical tools to provide developers the necessary guidance in developing and 

improving algorithms and numerical methods for describing physical processes. Moreover, 

validation results are essential for code users in order to gain confidence in applying the code to 

real-world applications. It is important that such validation exercises be performed objectively 

by developers, who may better understand the nuances of particular models, as well as users, 

who may have a more distant knowledge of the internal models but may have a greater 

knowledge of real-world applications.

The MELCOR code has been validated against numerous separate effects tests, integral tests 

such as Phebus, and actual accident studies such as Three Mile Island Unit 2 (see references 

[1], [31] - [43], and [61]). When MELCOR 1.8.6 was released, validation calculations were 

performed on many of the same validation test cases used for MELCOR 1.8.5 and results were 

compared between the two code versions and found to be agreeable. Furthermore, when 

existing plant decks were converted from MELCOR 1.8.5 to MELCOR 1.8.6 they were 

scrutinized for deviations in results. Some assessments, performed by SNL or other code users 

have been documented in the MELCOR Code Assessment Program (MCAP) meetings and the 

European MELCOR User Group (EMUG). However, even though internal validations were 

performed for MELCOR 1.8.6, an assessment report was not published because of similarities 

in the physics models between MELCOR 1.8.6 and MELCOR 2.1 which was released very 

shortly after MELCOR 1.8.6. Currently, a validation report is in preparation for MELCOR 2.1, 

covering an expanded set of validation test cases which will include results for Version 1.8.6 for 

comparison.

The purpose of this section is not to provide a complete MELCOR validation test report, but 

rather to provide an overview of the validation program. Code validation supports the 

application of MELCOR for state-of-the-art analyses.

4.1 Selection of Validation Test Cases

An objective of the MELCOR development team is to assess new code models against 

available test data, where that data exists. Both separate effects tests as well as integral tests 

are used for code assessment. Separate effects tests are designed to focus on an individual 

physical process, to eliminate the combined effects of multiple physical models which may 

obscure the validation of a particular model. However, it may be impossible to design a single 

test that isolates a single process and separate effects tests often ignore the important coupling 

between processes that are inherent in real world applications. Integral tests are valuable for 

examining the relationship between such coupled processes. Tests should be selected that are 

applicable to the calculation domain of the code and this domain should be clear to code users. 

Often, this requires significant analytical experience in applying the code to real-world problems 

to understand the calculation domain.
4-2

MELCOR has been assessed against numerous severe accidents performed by the NRC, 

EPRI, DOE, as well as many international research programs. Often, international standard 

problems (ISPs) are used as reference validation cases because they are “standard” problems 

that are assessed against other codes which may have alternate modeling capabilities. These 

ISPs are generally well documented, and may provide code-to-code comparisons to compare 

modeling approaches.

An important aspect of validation is that of coverage. Ideally, it is desirable to target each 

physics model available in the code with one or more validation test cases that can expose the 

capabilities of the model in simulating test conditions and responses. However, limited 

resources require some prioritization of effort in determining those processes that are most 

uncertain and contribute most to the sensitivity of results. There is a significant amount of effort 

involved in developing an input model, and understanding the results in light of the uncertainties 

inherent in the experiment design. This effort involves comparison of important measurements 

to calculated results, interpretation of discrepancies, and variation of model parameters and 

nodalization to best describe the particular case. Often analysts can be tempted to manipulate 

input variables to get the ‘best’ results compared to data. However, it is more desirable to focus 

on what can be learned from the analysis in terms of exposing specific modeling adequacies or 

deficiencies. Furthermore, it is desirable to understand the numerical convergence of such 

calculations by examining both spatial and temporal nodalization of the model. 

More than 50 such validation tests have been proposed for the MELCOR 2.1 assessment 

document. Table 4-1 shows a summary of such tests categorized by physics examined by the 

test, i.e., RN transport, core heat-up and degradation, containment, ex-vessel corium, and 

integral tests. Important physics assessed in this study includes, but is not limited to, heatup/heat transfer, oxidation of materials, reflood cooling, core degradation, molten pool modeling, 

fission product release, vessel failure, critical flow, MCCI, direct containment heating (DCH), 

condensation, containment stratification, hydrogen burn, hygroscopic effects, aerosol 

deposition, radionuclide transport, iodine pool chemistry, suppression pool scrubbing, vent 

cleaning, engineering safety features such as sprays (washing of radionuclides and cooling of 

atmosphere) and ice condensers. These validation tests exercise all the MELCOR physics 

packages to at least some degree, with the exception of the Condenser, Fan Cooler, and 

Passive Autocatalytic Recombiner Packages. Furthermore, specific models such as the point 

kinetics model, high temperature gas reactor models, spent fuel pool models, lower head 

penetration models, mechanical failure models, the integral heat exchanger model, flashing 

models, and the counter-current stratified flow model are not assessed in the current set of 

validation tests.
4-3

Table 4-1 MELCOR validation tests. 

Integral Tests/ 

Accidents 

Core RN Transport Containment Ex-Vessel 

Bethsy LOFT-FP2 FALCON 1 & 2 NUPEC M-8-1, M-8-2 OECD-MCCI 

Flecht-Seaset PBF-SFD VANAM-M3 IET 1 - IET7 and IET 9 - IET 11 SURC 

GE Level Swell CORA-13 LACE-LA4 PNL Ice condenser tests IET-DCH 

RAS MEI DF-4, MP1, MP2 LACE-LA1 & LA3 Wisconsin flat plate 

NEPTUN FPT1, FPT3 STORM DEHBI 

TMI-2 LHF/OLHF AHMED CVTR 

VERCORS ABCOVE HDR V44 

ORNL VI CSE-A9 HDR E-11 

Quench 11 DEMONA NTS-Hydrogen Burn 

RTF ISP-41 GE Mark-III Suppression Pool 

VERCORS Marviken Blowdown Tests 

ORNL VI CSTF Ice Condenser test 

Marviken ATT-4 LOFT-FP2 

4.2 Discussion of MELCOR Validation Tests

Assessment analyses have been performed historically as part of the MELCOR code 

development process. Table 4-2 summarizes the status of code validation tests, including 

aerosol tests, for various versions of MELCOR and the plans for future documentation. It is 

desirable to perform an assessment analysis with each new model added to the code. For 

example, aerosol mechanics for nonhygroscopic aerosols are modeled using the 

Multi-Component Aerosol Module for CONTAIN (MAEROS) code where good verification of 

aerosol agglomeration physics and gravitational depletion was demonstrated in early versions of 

MELCOR based on Marviken, Aerosol Behavior Code Validation and Evaluation (ABCOVE),

and the LACE testing. MELCOR Version 1.8.5 introduced extensions to treat hygroscopic 

aerosol effects where good validation against the Experiments on the Aerosol Behavior within a 

Multi-Compartment Containment (VANAM) M3 test (similar to the Demonstration of Nuclear 

Aerosol Behavior (DEMONA) test) as well as the Aerosol and Heat Transfer Measurement 

Device (AHMED) experiments was demonstrated. The Containment Systems Experiment

(CSE) A9 test was used to validate the containment spray scrubbing modeling in MELCOR 

Version 1.8.5 in the CONTAIN-MELCOR parity assessment study. The CONTAIN-MELCOR 

parity study introduced numerous other containment behavior assessments including the

Nuclear Power Engineering Corporation (NUPEC) mixing tests, the Nevada Test Site (NTS) 

hydrogen burn tests, and the Integral Effects Testing (IET) DCH experiments. Fission product 

release from fuel, including mixed oxide and high burnup were assessed against ORNL HI/VI 

tests and against more recent VERCORS experiments and documented in the Phebus 

Synthesis report using MELCOR Version 1.8.5. In Version 1.8.5 fission product release models 

were adjusted using sensitivity coefficient overrides to the Version 1.8.5 models. These were 

formalized as code options and defaults for code Version 1.8.6. MELCOR Version 1.8.6 also 

introduced expanded modeling detail for core melt progression processes, including molten pool 

convection treatments. These extensions provided improved prediction of the TMI-2 accident, 

some of which are still currently under assessment. The Phebus FPT-1 test stands as the most 

comprehensive integral assessment of core damage progression, hydrogen generation, fission 

product release and RCS deposition and containment natural depletion processes. This test 

provides good assessment of key deposition behavior in the reactor RCS and for containment 

depletion. Other code assessments for code Version 2.1 have been performed by IBRAE 

during the code Version 2.1 conversion process as indicated in Table 4-2. Detailed descriptions 
4-4

as well as results and findings can be found in the most recent Volume III of the MELCOR 

manuals [61]; an ongoing effort to update the demonstration problems as well as include 

additional experiments is underway for MELCOR Version 2.1. A comparison of key phenomena 

modeled with various MELCOR code versions is provided in the following section to 

demonstrate good agreement is maintained as the development of the code progresses.
4-5

Table 4-2 Historical review of MELCOR assessment studies. 

Assessment\Code 

Version

1.8.1 (1991)1.8.2 (1992)1.8.3 (1994)1.8.4 (1995)1.8.5 (1996)1.8.6 (2005)2.1 (2011-2012)M 1.8.5 Volume 3CONTAIN ParityPhebus SynthesisIBRAE AssessmentsSeparate12AssessmentsM 2.1 Volume 3

ABCOVE Tests x x x x

ACRR DF-4 x x

ACRR MP-1/MP-2 x x

Ahmed Hydroscopic 

Tests x x x x x

BETHSY (ISP-38) x x x

BWR Mk-III Vent 

Clearing Tests x x x x x

CORA 13 x x x x

CSE Spray Experiments x x x x x

CVTR x x x x x

DEMONA x x x

FALCON Tests x x x

FLECHT-SEASET x x x x x

GE Level Swell Tests x x x x x

HDR E-11 x x x

HDR V44 x x x x

HI/VI FP Tests x x x x

IET DCH Experiments x x x

JAERI x x x

LACE-LA1, LA3 x x

LACE-LA4 x x x x

LOFT-FP2 x x x x x x

Marviken x x x

NEPTUN Experiment x x x

NTS H2 Burn Tests x x x x

NUPEC Mixing Tests x x x x x

OECD MCCI x

PBF-SFD1-4 x x x

Phebus B9+ x x x

Phebus FPT-1 x x x x x x x

PNL Ice Condenser Test x x x x

Quench 6 x x

RAS MEI Tests x x x

RASPLAV Salt Tests x x

RTF Iodine Tests 

(ISP41) x x x

STORM x x x x

SURC MCCI x x x

TMI-2 x x x x x x

VANAM-M3 (ISP37) x x x x x

VERCORS 1-6 & HT/RT 

FP Tests x x x x

 12 “Separate Assessments” means that there is a standalone report (SAND or NUREG report) that 

documents the work.
4-6

4.3 Comparisons of Code Versions

Many of the validation analyses referenced in this report, see Appendix C, were performed with 

earlier versions of the MELCOR code, since the MELCOR 1.8.6 validation report was not 

published. However, given the level of maturity in many of the existing MELCOR physics 

models, essential validation exercises for the most part are not strongly dependent on the code 

version. Even so, small modeling changes and coding errors can impact results. Therefore a 

discussion of code version and the impact on validation is presented here.

SNL is currently updating the validation report for MELCOR Version 2.1. MELCOR Version 2.1 

is largely identical to Version 1.8.6 with respect to model pedigree; the main difference being 

conversion of the source code to FORTRAN 95. Changes made to 2.0 subsequent to its 

release have mainly affected new modeling for high temperature gas reactors. Significant code 

corrections made to 2.x were also made in the 1.8.6 version and made available to the 

SOARCA analysis team. The published MELCOR 2.1 validation report will also present 

validation results using MELCOR 1.8.6 for many of these analyses.

In order to better appreciate the significance of the historical validation analyses, an evolution of 

code development with code versions is required. 

Appendix D provides a list of major code modifications that were made during the development 

cycle. Note that this list only considers those physics models that may be directly related to the 

assessments in Table 4-2 and the SOARCA project. It does not contain many usability features 

and physical models that were not used in the SOARCA project, such as the point kinetics or 

the intermediate heat exchanger models. It also does not catalog model corrections and other 

bugs that were addressed.

Finally, since many of the historical validation cases have already been updated with Version 

2.1, the following sections provide comparisons of select computational results with the historic 

code assessment analysis. However, it is not the intention of this report to reproduce the details 

of the validation report here. Instead, the following discussions focus on some key physical 

models assessed.

4.3.1 Airborne Physics

MAEROS is a multisectional, multicomponent aerosol dynamics code that evaluates the size 

distribution of each type of aerosol mass, or component, as a function of time. MELCOR uses 

the MAEROS code for modeling aerosol agglomeration and deposition processes of 

nonhygroscopic aerosols. The MAEROS models have been in the code since MELCOR 

Version 1.8.0 with only error corrections and extension since. Hygroscopic models were added 

to the code in Version 1.8.4.

Agglomeration of non-hygroscopic aerosols from condensation of water vapor is assessed in 

the ABCOVE [31] and DEMONA experiments. Figure 4-1 shows the non-hygroscopic aerosol 

mass calculated for 1.8.2, 1.8.6, and 2.x, together with data from the AB-5 test. Note for this 

simple one volume calculation, the results have not changed noticeably since the early versions 

of the code. Similarly, the DEMONA test shows depletion of SnO2 due to condensation on the 

non-hygroscopic aerosol. These examples demonstrate the version independence of such 

calculations.
4-7

Figure 4-1 CSTF Airborne Mass Test AB5

Figure 4-2 Depletion of SnO2 in DEMONA-B3 experiment

M 182

Test
4-8

4.3.2 Oxidation 

Metal oxidation is calculated using standard parabolic kinetics, with appropriate rate constant 

expressions for Zircaloy and steel, limited by gaseous diffusion considerations if necessary. For 

the Zircaloy-H2O reaction, the rate constant is evaluated using the Urbanic-Heidrich constants. 

Though these constants and equations have not changed since they were first implemented into 

the code, other changes to the code can lead to changes in clad temperature, surface areas, 

and oxidation thickness histories. Therefore, changes in results are not so much indicative of 

changes to the oxidation models as they are changes in the core heat-up and degradation 

modeling. Figure 4-3 and Figure 4-4 show the hydrogen generation calculated for the PhebusB9+ and FPT-1 assessment cases [61] respectively, using MELCOR Versions 1.8.5, 1.8.6, and 

2.1. Note that only minor differences are observed for these three code versions. There is a 

slight trend in the data showing that MELCOR Version 1.8.5 predicted higher hydrogen 

generation than MELCOR Version 2.1 and MELCOR Version 1.8.6 and that all three versions 

slightly over predict the cumulative hydrogen generation.

Figure 4-3 PHEBUS-B9+ hydrogen generation
4-9

Figure 4-4 FPT-1 hydrogen generation

4.3.3 Hydrogen Stratification in Containment

Because of its lower density than surrounding air, hydrogen would concentrate in higher regions 

of the containment. It is important to be able to capture this stratification to predict local regions 

of flammability. The NUPEC M-8-1 [61] mixing test provides an excellent validation of 

MELCOR’s capabilities for calculating stratification of helium in a large, compartmentalized 

containment. The MELCOR Version 1.8.5 input deck was converted to MELCOR Version 1.8.6 

and then to MELCOR Version 2.1, using the Symbolic Nuclear Analysis Package (SNAP) as 

the converter. Though there are noticeable discrepancies between calculations and test data, it 

is important to observe that all three code versions give identical results. Overall, MELCOR 

does a reasonable job of capturing helium stratification for these tests.
4-10

Figure 4-5 Helium stratification calculated for NUPEC M-8-1 for MELCOR 2.x

Figure 4-6 Helium stratification calculated for NUPEC M-8-1 for three MELCOR code 

versions
4-11

4.3.4 Combustion Modeling

MELCOR uses relatively simple models for burning of premixed gases without modeling the 

actual reaction kinetics or tracking the actual flame front propagation based on the HECTR 1.5 

code; the models were implemented into the code before MELCOR Version 1.8.0. These 

models have a high level of maturity and only minor code corrections have been made to these 

models in recent code versions.

Table 4-3 through Table 4-5 show burn characteristics calculated for the NTS hydrogen burn 

tests. These tests were sponsored by the NRC and performed by the EPRI and were used as 

part of the MELCOR - CONTAIN parity study. No significant changes are observed among 

those tests included in the assessment study.

The MELCOR calculated burn times differ significantly from the experimental values. A likely 

explanation would be that the complexity of the flame propagation is not modeled. For example, 

if the first flame propagates upward and then down along the wall of the spherical pressure 

vessel, such behavior could not be captured in a lumped-parameter code. It should be noted 

however, that similar errors were obtained by analogous calculations performed with the 

CONTAIN code.
4-12

Table 4-3 Hydrogen burn completeness from experiment and MELCOR.

Burn Completeness (%)

Test Experiment M 1.8.5 M 1.8.6 M 2.1

NTSP01 32.0 36 35.67 35.67

NTSP12 58.0 74 72.94 72.94

NTSP15 100.0 100 100.0 100.0

NTSP20 100.0 100 100.0 100.0

Table 4-4 Hydrogen burn times from experiment and MELCOR.

Burn Time (s)

Test Experiment M 1.8.5 M 1.8.6 M 2.1

NTSP01 68.5 2.0 1.9 1.9

NTSP12 27.0 9.0 9.2 9.2

NTSP15 6.0 1.7 1.2 1.2

NTSP20 2.0 6.0 4.0 4.0

Table 4-5 Pressure ratio calculated with recent MELCOR code versions compared to 

test results.

Test ID and Initial H2 & H2O

Concentrations

P(max)/P(initial) 

Test ID H2, v/o H2O, v/o M 1.8.5 M1.8.6 M2.1 Test

Standard Tests

NTSP01 5.3 4.2 1.71 1.70 1.70 1.48

NTSP15 9.9 4.2 4.11 4.08 4.08 3.61

Steam-Laden Tests

NTSP12 6.9 28.3 2.37 2.36 2.36 1.831

NTSP20 12.9 27.8 3.97 3.95 3.95 3.87

4.3.5 Containment Pressure Response to Sprays

A series of experiments were conducted in the CSE vessel to evaluate the performance of 

aqueous sprays as a means of decontaminating containment atmospheres [33]. Measurements 

were obtained which provide a suitable basis for judging the ability of various mathematical 

models to predict spray performance in large nuclear power plant buildings. Assessments have 

been performed with MELCOR Version 1.8.3, 1.8.6, and 2.1 models for the A9 experiment.

The containment pressure response is shown in Figure 4-7 and Figure 4-8 for all modern code 

versions. These calculations indicate that the modeling of heat removal from sprays has not 

significantly changed in these recent code versions.
4-13

Figure 4-7 MELCOR 1.8.6 & 2.1 assessments of CSE A9

Figure 4-8 MELCOR 1.8.3 assessments of CSE A9

4.3.6 Fission Product Release

Fission product release rates are validated by comparison to several experimental series, 

principally the ORNL VI tests, Phebus FPT-1, and VERCORS-2 and 4. The release rates are 

set relative to Cs based on the VI tests. The releases were then adjusted based on FPT-1, 

which has the most complete data for the various fission products. The resulting release 

coefficients, termed “modified ORNL-Booth,” were then compared to the VERCORS tests. 
4-14

VERCORS-2 has data only on Cs release, and VERCORS-4 has some others also. The last 

comparisons to VI and VERCORS were done with MELCOR Version 1.8.5 and 1.8.6; 

comparisons to FPT-1 have been done with 1.8.5, 1.8.6 and 2.1. In general, differences 

observed were due to the switch to the modified ORNL-Booth release coefficients in 1.8.5 rather 

than to any version differences.

4.3.7 Molten Core-Concrete Interaction

The source term during the late phase of a severe accident is dominated by the molten debris –

concrete interactions that occur in the reactor cavity. CORCON-MOD3 was implemented into 

MELCOR Version 1.8.3 and, aside from a few changes in default sensitivity coefficients, has 

largely remained unchanged. The SURC-1 test examines the one-dimensional ablation front 

from overlying core debris. Results of simulations for MELCOR Version 1.8.6 and 2.1 are 

shown in Figure 4-9. These results show that the CORCON models continue to give good 

results in predicting the ablation front for these tests. The MELCOR Version 2.1 assessment 

report will investigate more recent tests such as the Organization for Economic Co-operation 

and Development (OECD) MCCI tests.

Figure 4-9 MELCOR 1.8.6 & 2.x assessments of ablation depth in SURC-1 Test
5-1

5. INSIGHTS

Conducting a state-of-the-art analyses project, such as the SOARCA project, provides an 

opportunity to investigate accident phenomena at a much greater level of detail than analyses 

performed in the past. Throughout the process, advanced knowledge and insights were gained 

that will benefit the PRA community in performing accident sequence analyses and the further 

refinement of branch probabilities. The methodology applied in the analyses established a 

framework for future applications and demonstrated that use of complex severe accident codes 

has advanced into risk informed application capable of evaluating response actions and timing 

to mitigate accident progression. The following insights were obtained during the SOARCA 

project. The general order of the presented insights is the authors’ perception of importance, but 

many of the Surry ISLOCA insights were grouped as a matter of convenience for the reader.

1. In the Surry SBOs, the most likely first RCS failure occurs at the hot leg nozzle prior to 

significant in-vessel fuel damage. This leads to vessel depressurization, accumulator 

discharge, fuel cooling, and an interruption to the core heat-up. A new release pathway 

for radionuclide is established at the failed hot leg. The response of a TISGTR is also 

impacted by hot leg failure. The hot leg failure substantially decreases TISGTR flow due 

to the RCS depressurization and the introduction of the larger failure location as the 

primary fission product pathway from the vessel. 

2. With regard to the ISLOCA modeling and the magnitude of predicted radionuclide 

releases to the environment, a key insight is the large amount of deposition of 

aerosolized radionuclides in the LHSI piping by means of turbulent deposition. Sustained 

high velocities in the LHSI piping during core degradation drive the importance of this

phenomenon. A thorough representation of the LHSI piping is necessary to address 

turbulent deposition of fission product aerosols in the piping and revaporization of 

deposits. 

3. The improvements to fuel degradation modeling and 2-dimensional core modeling show 

a delayed heat-up followed by accelerated oxidation. The accelerated oxidation phase 

ends following molten Zircaloy breakout. Without molten Zircaloy breakout, the 

subsequent heat-up is primarily controlled by decay heat. The best practice modeling of 

Zircaloy-oxide collapse creates a debris bed similar to TMI-2. The debris bed slows

oxidation by creating blockages and inhibiting natural circulation. The debris bed 

gradually grows axially and radially, which eventually leads to core plate failure.

4. Upon core plate failure, the lower plenum debris response is consistent with 

experimental research and leads to delayed vessel failure versus the previously default 

0-dimensional counter-current flow limiting models. The lower plenum debris cools as it 

transfers heat to the water and steel in the lower plenum. Once the water has 

evaporated, the debris bed heats the lower head towards creep rupture failure. Drain 

line, control rod drive penetration, etc. lower head failure modes are considered for the 

Peach Bottom reactor but could only be modeled parametrically and do not lead to 

significant changes in the source term magnitude.

5. Natural circulation processes substantially delay the heat-up of the fuel as heat is 

transmitted from the core into the vessel internals and to the steam generator in PWRs. 

The slower heat-up leads to high hydrogen production and more extensive transport of 

aerosols in the primary system.
5-2

6. The failure of an in-core instrument tube leads to an early radionuclide release which 

does not impact the overall source term. A reviewer cited strong evidence of early noble 

gas releases through instrument tubes for the TMI-2 and Fukushima accidents. The 

approach is fully discussed in the NUREG/CR-7110, Vol. 1 and can be important for 

accident management and early release timing but less important for source term 

magnitude characterization.

7. The inclusion of control rod materials and tin from the Zircaloy increases the aerosol 

density, especially at the start of the core degradation. The control rod and structural 

aerosols enhance in-vessel deposition by increasing agglomeration rates.

8. The incorporation of structural, creep failure models for the core plate and enhanced incore debris bed modeling leads to higher temperature debris relocating to the lower 

plenum. Following core plate failure, the application of the best practices demonstrates

vigorous but not explosive interaction between the relocating debris and the lower 

plenum.

9. Although the modified ORNL-Booth radionuclide release model has a slower release 

rate than the default CORSOR-M model, the new fuel degradation modeling leads to 

higher temperatures and a slower collapse than previous core degradation models. The 

net effect is essentially complete release of volatile fission products from the fuel.

10. The BWR containment failure best practices include simultaneous consideration of 

multiple failure modes. Although each failure mode can have temporal significance, the 

impact on the source term is dominated by the liner melt-through failure.

11. The potential of combustion failure of the PWR containment is significant in one scenario 

which included emergency containment sprays. While delayed combustion can lead to 

containment failure or increased leakage, the formation of combustible quantities can

only occur through sustained spray operation that simultaneously captures

radionuclides. Hence the impact is comparable to cases without delayed combustioninduced containment failure.

12. For Peach Bottom, two modes of SRV failure were considered in the SOARCA project –

failure due to excessive cycling (stochastic failure to close) and failure due to 

overheating (thermal failure to close). The best-estimate number of SRV cycles 

permissible before failing in an open position was taken as 270. The best-estimate 

temperature an SRV valve stem could withstand before seizing with the valve in an open 

position was taken as 900 K. Due to inclusion of these best-estimate failure mode 

values, the failure of the lowest set-point SRV occurs prior to the onset of core damage 

as a result of the excessive cycling criterion.

13. For Surry, two modes of primary SRV and PORV failure are considered in the SOARCA 

project – failure due to excessive cycling (stochastic failure to close) and failure due to 

overheating (thermal failure to close). The 50th percentile number of SRV and PORV 

cycles permissible before failing in an open position was taken as 256 and 247, 

respectively. The best-estimate temperature an SRV or PORV valve stem could 

withstand before seizing with the valve in an open position was taken as 10 cycles 

above 1000 K for both valve types. Due to inclusion of these best-estimate failure mode 

values, the failure of the lowest set-point primary SRV occurs prior to the onset of core 

damage as a result of the excessive cycling criterion for the unmitigated LTSBO with late 
5-3

RCP seal failure, the mitigated and unmitigated STSBO, and the mitigated and 

unmitigated STSBO with a TISGTR.

14. The formation of cesium-molybdate leads to a higher release of molybdenum, lower 

volatility of released cesium, and significantly different deposition and revaporization 

behavior. In the Surry ISLOCA analysis, the cesium release to the environment is 

strongly impacted by revaporization of radionuclide deposits formed in the LHSI piping. 

The majority of the cesium is in the form of cesium-molybdate, which is retained in the 

LHSI piping.

15. The necessity of modeling in detail the interconnected buildings into which the RCS 

would vent given an ISLOCA proved valuable and provided insights regarding the depth 

of the water pool formed in the buildings. The rupture in the Surry scenario was 

determined to be submerged in a pool formed from water leaked into the auxiliary 

buildings through the rupture. A key insight is that pool cooling of submerged LHSI 

piping sections strongly suppresses revaporization of radionuclide deposits in those 

sections and the pool scrubbing of radionuclides at the LHSI pipe rupture results in a 

meaningful reduction in release to the environment. Had the piping sections not been 

determined in the Surry scenario to be submerged, radionuclide releases to the 

environment would have been greater.

16. The degree of revaporization described above is strongly affected by to what extent

gamma radiation emitted by the radionuclide deposits is transmitted through the LHSI 

pipe wall. In certain sections of the piping, including the gamma transmission or not 

meant the difference between not having and having revaporization, respectively.

17. The resuspension of radionuclide deposits formed in the LHSI piping is an important 

phenomenon that was not addressed mechanistically in the ISLOCA analysis. The 

phenomenon is important because of the large deposition in the LHSI piping predicted 

by MELCOR and the sustained high-velocity gas flow in the piping that can conceivably 

tear the deposits from the pipe wall and carry them along the piping and out the rupture. 

Resuspension is not addressed mechanistically because, while MELCOR includes 

modeling to predict deposition in the piping (turbulent deposition and impaction), it does 

not include modeling to predict resuspension. During the ISLOCA analysis, it was 

judged that that the deposits would likely break away from the pipe wall to some extent 

and be carried out the piping to buildings outside of containment, but that the material 

would be in a form difficult to aerosolize and so would not release to the environment. 

This is because the break is under water for the entire simulation time, and the aerosols 

generated by a melting core are very small since they are formed by vapor condensation 

but the aerosols formed by re-suspending the deposited materials by high speed vapor 

flow across the surface of the pipe are so large that the aerosols do not travel far 

downstream before re-depositing. The insight gained regarding the resuspension of 

deposited material in the LHSI piping is that a judgment regarding resuspension needs 

to be made when employing a tool such as MELCOR because the code models 

deposition but not resuspension.

18. Another fundamental insight gained from the ISLOCA analysis is that hydrogen 

deflagrations in the building(s) into which the RCS would vent (given an ISLOCA) were 

important because the resulting breaches in building boundaries were the ultimate paths 

of radionuclide releases to the environment. The best practices include pressurization 

and/or hydrogen deflagration failure logic for the BWR reactor building boundary. The 
5-4

ineffective retention of fission products within the reactor building was confirmed by 

March 2011 events at Fukushima.

19. Aerosol (radionuclide, concrete, soot) deposition in the HEPA filters of the safety-related 

exhaust ventilation system serving the building(s) into which the RCS would vent was 

important to the Surry ISLOCA scenario. Deposition was determined to be high on the 

filters but not so high to suggest that the filters would fail or that the system would shut 

down because of excessive loading on the filters. This is insightful as the progression of 

the scenario (and the releases to the environment) would have been substantially 

different if the system shut down or if the filters failed but the fans continued to run.

20. The necessity of modeling in detail the filtered safety-related exhaust ventilation system 

serving the buildings into which the RCS would vent is another valuable insight gained 

from the ISLOCA analysis. Substantial amounts of fission products were captured in the 

filters of this system in the Surry analysis such that had the system not been 

represented, predicted releases to the environment would have been substantially 

greater.

21. A key insight from the ISLOCA analysis was the importance of operators completing 

certain critical procedural actions and the necessity of representing these actions in an 

ISLOCA simulation, see NUREG/CR-7110 [63]. At Surry, the timing of actions such as 

stopping the LHSI pumps, isolating the LHSI pump suctions from the refueling water 

storage tank, and stopping and/or throttling HHSI pumps to conserve RWST inventory, 

which strongly influence accident progression, were taken directly from operator 

response to the ISLOCA scenario in the control room simulator. With these actions 

accomplished in a timely manner, MELCOR predicts a slow progression to core 

damage; therefore, the operator responses and corresponding timings were of high 

importance in modeling the sequence. Analyses for individual sites should look closely at 

the timing of procedural actions as well as simulated response times, if possible.
6-1

6. REFERENCES

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[11] H. Manenc, P. Mason and M.P. Kissane, “The modeling of fuel volatilization in 

accident conditions,” J. Nuc. Mat. (294) pp 64-68, 2001. 

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(255) pp 59-66, 1998. 

[13] N.E. Bixler, “VICTORIA 2.0: A Mechanistic Model for Radionuclide Behavior in a 

Nuclear Reactor Coolant System Under Severe Accident Conditions,”

NUREG/CR-6131, SAND93-2310, 1998. 

[14] B. Clement and T. Haste, “Comparison report on International Standard Problem 

ISP-46 (Phebus FPT-1),” Note Technique SEMAR 03/021, Draft Final Report, April 

2003. 
6-2

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Cooperative Severe Accident Research Meeting, U.S. Nuclear Regulatory 

Commission, 1996.

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presentation of the program and synthesis of VERCORS 1-6 results,” Nuc. Eng. 

And Des. (208) pp 191-203, 2001.

[17] V. Strizov, “MELCOR 1.8.5 Validation: Modeling of Fission Product Release,” 

NSI-SARR-149-03, Russian Academy of Sciences, Nuclear Safety Institute, Dec.

2002.

[18] Soffer, L., et al., “Accident Source Terms for Light-Water Nuclear Power Plants,” 

NUREG-1465, U.S. Nuclear Regulatory Commission, 1995.

[19] NUREG/CR-5582, “Lower Head Failure Experiments and Analyses,” SAND98-2047, 

Sandia National Laboratories, 1998.

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Investigating 2-D Core-Concrete Interaction and Debris Coolability,” Proc. Proc. Int. 

Conf. Adv. Power Plants, ICAPP’04, Pittsburgh, Pennsylvania, June 2004.

[21] Bayless, P.D., et al., “Severe Accident Natural Circulation Studies at the INEL,” 

NUREG/CR-6285, INEL-94/0016, February 1995.

[22] Stewart, W.A., Pieczynski, A.T., and Srinivar, V., "Natural Circulation Experiments for 

PWR Degraded Core Accidents," EPRI Report NP 6324 D, 1989.

[23] Stewart, W.A., Pieczynski, A.T., and Srinivar, V., "Natural Circulation Experiments for 

PWR High Pressure Accidents," EPRI Report TR 102815, 1993. 

[24] Boyd, C.F. Hardesty, K., “CFD Analysis of 1/7th Scale Steam Generator Inlet 

Plenum Mixing During a PWR Severe Accident,” NUREG-1781, October 2003.

[25] Boyd, C.F., Helton, D. M., Hardesty, K., “CFD Analysis of Full Scale Steam 

Generator Inlet Plenum Mixing During a PWR Severe Accident,” NUREG-1788, U.S. 

Nuclear Regulatory Commission, May 2004.

[26] Wagner, K. C., “MELCOR 1.8.5 Analysis of Natural Circulation Flow in the 

Westinghouse High-Pressure SF6 Experiments,” Draft SAND Report, Sandia 

National Laboratories, June 2001.

[27] Fletcher, C. D., and Beaton, R. M., “SCDAP/RELAP5 Base Case Calculation for the 

Station Blackout Uncertainty Study,” Letter report to D. M. Helton, US NRC, August 

2006. 

[28] Fletcher, C. D., and Beaton, R. M., “Evaluation of Uncertainties in SCDAP/RELAP5 

Station Blackout Simulations,” Letter report to D. M. Helton, US NRC, August 2006. 

[29] Petrie, L.M et al., “Standard Composition Library,” NUREG/CR-0200, Rev. 6, Vol. 3, 

ORNL/NUREG/CSD-2/V3/R6, Oak Ridge National Laboratory, September 1998.

[30] Downes, K.W., Beck, C.K., Cowan, F.P., Fleck, J.A., Kuper, J.B.H., McLaughlin, J., 

Singer, I., Smith, M., “Theoretical Possibilities and Consequences of Major 
6-3

Consequences in Large Nuclear Power Plants,” WASH-740, United States Atomic 

Energy Commission, March 1957. 

[31] Souto, F.J., Haskin, F.E., Kmetyk, L.N., “MELCOR 1.8.2 Assessment: Aerosol 

Experiments ABCOVE AB5, AB6, AB7, and LACE LA2,” SAND94-2166, 1994.

[32] Tautges, T.J., “MELCOR 1.8.2 Assessment: The MP-1 and MP-2 Late Phase Melt 

Progression Experiments,” SAND94-0133, 1994.

[33] Kmetyk, L.N., “MELCOR 1.8.3 Assessment: CSE Containment Spray Experiments,” 

SAND94-2316, 1994.

[34] Tills, J., Notafrancesco, A, Longmire, P., “An Assessment of MELCOR 1.8.6: Design 

Basis Accident Tests of the Carolinas Virginia Tube Reactor (CVTR) Containment 

(Including Selected Separate Effects Tests),” SAND2008-1224, 2008.

[35] Tautges, T., “MELCOR 1.8.2 Assessment: The DFI-4 BWR Damaged Fuel 

Experiment,” SAND93-1377, 1993.

[36] Tautges, T., “MELCOR 1.8.3 Assessment: GE Large Vessel Blowdown and Level 

Swell Experiments,” SAND94-0361, 1994.

[37] Kmetyk, L.N., “MELCOR 1.8.2 Assessment: IET Direct Containment Heating Tests,” 

SAND93–1475, 1993.

[38] Kmetyk, L.N., “MELCOR 1.8.1 Assessment: LACE Aerosol Experiment LA4,” 

SAND91–1532, 1991.

[39] Kmetyk, L.N., “MELCOR 1.8.1 Assessment: LOFT Integral Experiment LP-FP-2,” 

SAND92–1373, 1992.

[40] Kmetyk, L.N., “MELCOR 1.8.1 Assessment: Marviken-V Aerosol Transport Tests 

ATT-2b/ATT-4,” SAND92–2243, 1993.

[41] Gross, R.J., “PNL Ice Condenser Aerosol Experiments,” SAND92–2165, 1993.

[42] Kmetyk, L.N., “MELCOR 1.8.1 Assessment: FLECHT SEASET Natural Circulation 

Experiments,” SAND91-2218, 1991.

[43] Kmetyk, L.N., “MELCOR 1.8.1 Assessment: ACRR Source Term Experiments ST1/ST- 2", SAND91-2833, 1992.

[44] "Reactor Safety Study, An Assessment of Accident Risks in U.S. Commercial 

Nuclear Power Plants," NUREG-75/014, U.S. Nuclear Regulatory Commission, 

October 1975. 

[45] Aldrich, D.C., Sprung, J.L., Alpert, D.J., Diegert, K.V., Ostmeyer, R.M., Ritchie, L.T., 

Strip, D.R., Johnson, J.D., Hansen, K., Robinson, J., “Technical Guidance for Siting 

Criteria Development,” NUREG/CR-2239, U.S. Nuclear Regulatory Commission, 

Washington, DC, 1982.

[46] Corradini, M.L. et al, “An Analysis of Containment Failure by Steam Explosion 

Following a Postulated Core Meltdown in a Light Water Reactor,” Nuclear 

Engineering and Design, 66:287-298, 1981.
6-4

[47] Theofanous, T.G., Yuen, W.W., “The Probability of alpha-mode containment failure,” 

Nuclear Engineering and Design.155:495-473, 1995. 

[48] Turland, B., Fletcher, D.F., et al., “Quantification of the probability of containment 

failure caused by an in-vessel steam explosion for the Sizewell B PWR,” Nuclear 

Engineering and Design 155:445-458, 1995.

[49] Huhtiniemi, I., Magalon, D., Hohmann, H., “Results of recent KROTOS FCI tests: 

alumina versus corium melts,” Nuclear Engineering and Design 189:379-389, 1999.

[50] Huhtiniemi, I., Magallon, D., “Insight into steam explosions with corium melts in 

KROTOS,” Nuclear Engineering and Design, 204:391-400, 2001.

[51] Song, J.H., et al., “Experiments on the interactions of molten ZrO2 with water using 

TROI facility,” Nuclear Engineering and Design, 213:97-110, 2002.

[52] Song, J.H., Park, I.K., Shin, Y.S., Kim, J.H., Hong, S.W., Min, B.T., Kim, H.D., “Fuel 

coolant interaction experiments in TROI using a UO2/ZrO2 mixture,” Nuclear 

Engineering and Design, 222(1):1-15, 2003.

[53] NEA/CSNI, “Technical Opinion Paper on Fuel-Coolant Interaction,” NEA/CSNI/R(99), 

2000.

[54] Nuclear Regulatory Commission, “Severe Accident Risks: An Assessment for Five 

U.S. Nuclear Power Plants,” NUREG-1150, U.S. Nuclear Regulatory Agency, 

Washington, DC, 1990.

[55] Pilch M.M.; Yan, H., and Theofanous T.G., “The probability of containment failure by 

direct containment heating in Zion," Nuclear Engineering and Design, Vol. 

164, Number 1. August 1996.

[56] R. A. Lorenz and M. F. Osborne, “A Summary of ORNL Fission Product Release 

Tests with Recommended Release Rates and Diffusion Coefficients,” 

NUREG/CR-6261, 1995.

[57] Natesan, K. and Soppet, W. K. “Air Oxidation Kinetics for Zircaloy-4,” Argonne 

National Laboratory, NUREG/CR-6846, July 2004.

[58] Wright , A. L. and Arwood, P. C., “Summary of posttest aerosol code-comparison 

results for LWR Aerosol Containment Experiment (LACE) LA3”, LACE TR-024, 

ORNL/M-492, June 1988.

[59] Fauske & Associates, LLC, MAAP (Modular Accident Analysis program), 

http://www.fauske.com/pdf/MAAP.pdf

[60] Gauntt, R. O., “Synthesis of VERCOS and Phebus data in Severe Accident Codes 

and Applications,” SAND2010-1633, Sandia National Laboratories, April 2010.

[61] Gauntt, R. O., Cole, R. K., Erickson, et.al., “MELCOR Computer Code Manuals, Vol. 

3: Demonstration Problems, Version 1.8.5,” NUREG/CR-6119 Rev. 0, 

U.S.SAND2001-0929P, Nuclear Regulatory Commission, Washington, DC, 2001.
6-5

[62] US. NRC. “State-of-the-Art Reactor Consequence Analysis Project, Volume 1: Peach 

Bottom Integrated Analysis,” NUREG/CR-7110, Volume 1, U.S. Nuclear Regulatory

Commission: Washington DC, 2012.

[63] US. NRC. “State-of-the-Art Reactor Consequence Analysis Project, Volume 2: Surry 

Integrated Analysis,” NUREG/CR-7110, Volume 2, U.S. Nuclear Regulatory

Commission: Washington DC, 2012

[64] Gauntt, R. O., Cole, R. K., Erickson, et.al., “MELCOR Computer Code Manuals, Vol. 

2: Reference Manuals, Version 1.8.5,” NUREG/CR-6119 Rev. 2, U.S. SAND2000-

2417/1, U.S. Nuclear Regulatory Commission: Washington DC, 2001.

[65] US. NRC. “Research Activities FY2012-2014,” NUREG-1925, Revision 2, U.S. 

Nuclear Regulatory Commissin: Washington DC, 2013.

A-1

Appendix A Other MELCOR Modeling Best Practices 

Several user input options were enabled or adjusted for the SOARCA project that do not 

warrant detailed descriptions or justifications of their use. This Appendix presents these generic 

adjustments within Table A-1 as a reference for simple input options which readers may wish to 

review with consideration for implementation.

Table A-1 Standard MELCOR modeling practices, modeling parameters, and 

sensitivity coefficients for analysis of severe accidents

Item Record Field Value(s) used in 

SOARCA Description

1. BUR000 IACTV 0 (Active) Burn package activation

2. BUR1xx

(xx = CV)

IGNTR 86 for CVs where 

ignition is to be 

prohibited.

Apply to RCS control volumes to preclude 

combustion.

3. BUR1xx

(xx = CV)

TFRAC 1.0 Time fraction of burn before propagation to 

neighboring CV is allowed. Value of 1.0 

means a flame must travel the radius of the 

control volume before propagating to its 

neighbor.

4. FLnnn00 ZFM,

ZTO

ZFM – ZTM != 0.0

(For vertical 

containment flow 

paths only)

To insure that MELCOR properly estimates 

vertical burn propagation in containment, 

and adjacent buildings, it is necessary to 

define “vertical” flow path “from” and “to” 

elevations with a small dZ. If the “from” 

and “to” elevations are set equal (which 

has been historical practice to ensure 

complete vertical pool drainage), the 

MELCOR burn package uses criteria for 

horizontal burn propagation.

5. FLnnnFF KFLSH 1 Calculate superheated pool flashing for all 

liquid LOCA connections to initially dry 

containment regions. KFLSH activates the 

model. Activate RN1Ikkk as needed for 

impact into specified heat structures.

6. FLnnn02 IBUBF & 

IBUBT

-1

+2

Vapor heat transfer in pools for RCS FLs. 

SPARC scrubbing in pools for spargers, 

quencher, vents, and BWR downcomers.

7. RN2FLTXX00 FPVAPOR Various geometric 

values

MELCOR SPARC pool scrubbing model 

was modified to scrub all gaseous RN 

classes for 

8. COR00001 DRGAP 0.0 Thickness of gas gap between fuel pellets 

and cladding set 0.0 to account for swelling 

of operating fuel.

9. COR00001A ILHTYP

ILHTRN

0

BWR =0, PWR =1

Lower head is a hemisphere

Transition is at RCOR (BWR) 

or RVES (PWR)
A-2

Item Record Field Value(s) used in 

SOARCA Description

10. COR00009 HDBPN

HDBLH

MDHMPO

MDHMPM

TPFAIL

CDISPN

100 W/m2-K

100 W/m2-K

‘MODEL’

‘MODEL’

9999 K

1.0

This record activates the internal molten 

pool to lower head heat transfer models 

and provides reasonable solid debris to 

lower head heat transfer coefficient.

11. COR00012 HDBH2O

VFALL

2000 W/m2-K

0.01 m/s

HTC in-vessel falling debris to pool 

(W/m2-K)

Velocity of falling debris (m/s). ). Perhaps 

not correct for shallow pools and not 

necessary in deep pools since adoption of 

no 1-D CCFL limitation via the 

one-dimensional Lipinski model.

12. CORCR0 IAICON 2 For PWRs only

Activate control rod release model, 2 = 

Model is active, vaporization is allowed 

from both candling material and 

conglomerate.

13. CORZjj01 PORDP 0.4 Porosity of particulate debris

14. CORijj04 DHYPD Core - 0.01 m

LP - 0.002 m

Particulate debris equivalent diameter (LP 

values for DHYPD, HDBH2O, VFALL 

tuned to get appropriate end-of-pour debris 

temperature. 2mm based on FAERO 

fragmented debris size). Perhaps not 

correct for shallow pools.

15. CORZjjNS TNSMAX 1520 K

1700 K

Control blades failure temperature (BWR)

Core top guide failure temperature (BWR)

16. CORijjDX FBYXSS Calculated. For BWRs only. Fraction of lower head 

COR cells normally displaced by control 

rod guide tubes should be ‘excluded’ from 

volume available to particulate debris. 

Volume recovered when tubes (as 

supporting structure) fails.

17. SC-1132(1) TRDFAI 2800 K Fuel rod collapse temperature (addressed 

with CORijjFCL records)

18. SC-1141 (2) GAMBRK 0.20 kg/m-s The maximum molten Zr breakout flow rate 

parameter adjusted to yield 2 mm/s as 

evidenced in CORA experiments

19. SC-1701 (1) 0.01 Open volume fraction for subnode 

blockage criterion. This is the default 

setting.

20. SC-4401(3) XPASMX 15 Maximum number of iterations permitted 

before solution is repeated with a 

decreased (subcycle) timestep.

Table A-2 Standard MELCOR modeling practices, modeling parameters, and sensitivity 

coefficients for analysis of severe accidents (continued)
A-3

Item Record Field Value(s) used in 

SOARCA Description

21. DCHNEMnn00 ELMNAM

ELMMAS

Use ORIGEN 

results for core, if 

available.

Elemental fission product mass at 

shutdown for calculation of decay heat.

22. DCHNEMnnmm DCHEAT Use pre-combined 

methodology for 

Cs, I, and Mo

Elemental fission product decay heat per 

unit mass (based on shutdown RN 

inventory).

• Define specific decay heat for CsI 

(Class 16) as 0.51155 of value for 

Class 2 (Cs) plus 0.48845 of value for 

Class 4 (I).

• Define specific decay heat for 

Cs2MoO4 (Class 17) as 0.7348 of value 

for Class 2 (Cs) plus 0.2652 of value 

for Class 7 (Mo).

If ORIGEN results are not available for the 

core, perform an input deck with BE 

burn-up and cycle history. Redistribute RN 

mass as follows,

• Class 2 initial mass represents the 

NUREG-1465 Cs gap mass not 

already included in Class 16.

• Class 4 initial mass is empty (10-6 kg)

• Class 7 initial mass is remaining Mo 

mass not included in Class 17.

• Class 16 has all I and an appropriate 

amount of Cs mass for CsI 

stoichiometry.

• Class 17 has the remaining Cs not 

included in Classes 2 and 16 plus Mo 

for Cs2MoO4 stoichiometry.

23. DCHCLSnnn0,

DCHCLSnnnm

RDCNAM, 

CLSELM

New RN

definitions for

Classes 1-12, 

16-18

If ORIGEN results are available, 

synthesize ORIGEN data to define a single 

representative element for each class with 

decay heat data that reflects decay heat for 

all elements within the class 

(DCHNEMxxxx input.) Redefine each 

class to include only the representative 

element.

24. DCHDEFCLS0 DEFCLS 13, 14, 15 Specifies that MELCOR DCH default 

classes are to be used.

Table A-3 Standard MELCOR modeling practices, modeling parameters, and sensitivity 

coefficients for analysis of severe accidents (continued)
A-4

Item Record Field Value(s) used in 

SOARCA Description

25. DCHCLNORM CLSNRM ‘No’ when ORIGEN 

results are 

available.

‘Yes’ when 

MELCOR is used 

to estimate initial 

inventories.

New ORIGEN input for elements/classes 

defines the total core decay heat.

Otherwise, let MELCOR normalize the 

elemental decay heats to the rated power.

Do not use RN1DCHNORM. Default 

behavior normalizes Class 10 (Uranium).

26. HSccccc400 &

HSccccc600

CPFPL

CPFAL

See discussion Minimum value of CVH pool fraction such 

that heat transfer is calculated to 

Pool/Atmosphere. For heat structures 

within the RPV, use 0.9. For PWR SG 

Tubes, use 0.1. All other structures 

modeled use default value of 0.5. 

27. HSccccc401

HSccccc601

EMISWL

RMODL

PATHL

0.27

EQUIV-BAND

0.1 m

Mean emissivity of SS type 316.

Equivalent band radiation model.

Nominal optical distance in steam (m).

For SS heat structures within the reactor 

vessel and those being monitored for 

creep-rupture failure.

28. HSDGccccc0 ISRCHS

ISDIST

GASNAM

HS #

1

SS

Heat structure for application of degas 

model.

Degassing model requires 1 mesh.

Name of released gas.

For SS boundary structures modeled with 

the HS package that are coupled to the 

core.

29. HSDGccccc1 RHOSRC

HTRSRC

TEMPL

TEMPU

7930 kg/m3

2.63x105 J/kg

1695 K

1705 K

Gas source density.

Gas source heat of reaction.

Lower temperature for degassing.

Upper temperature for degassing.

For SS boundary structures modeled with 

the HS package that are coupled to the 

core.

Table A-4 Standard MELCOR modeling practices, modeling parameters, and sensitivity 

coefficients for analysis of severe accidents (continued)
A-5

Item Record Field Value(s) used in 

SOARCA Description

30. MPMATxxxx MLT 2800 K

2800 K

Uranium-dioxide

Zirconium-oxide

Because of the interactions between 

materials, liquefaction can occur at 

temperatures significantly below the melt 

point. The interaction between ZrO2 and 

UO2 results in a mixture that is fluid at 

above about 2800 K (compared to the 

melting temperatures of

3113 K and 2990 K, respectively, for the 

pure materials). Similarly, although pure 

B4C melts

at 2620 K, interaction with steel produces a 

mixture that is fluid at above about 1700 K.

31. RN1001 NUMSEC

NUMCMP

NUMCLS

10

2

20 (PWR)

18 (BWR)

Default

Default

For BWR & PWR: 16 = CsI, 17 = Cs2MoO4

Now Class 17 includes default settings for 

Cs2MoO4. 

32. BWR structural 

tin release 

RN/DCH data 

for RN Class 18

For BWR: RN Class 18 = SnO2

(non-radioactive)

Define SnO2 (DCHCLSnnn0)

18 = ‘SnO2’

SnO2decay heats (DCHNEMnn00) 

0 W/kg (no decay heat)

SC(7110) vapor pressures

SnO2: Log10(P(mm Hg)) = 15400/T + 8.15

SC(7111) diffusion coefficients

SnO2: Sigma = 3.617, E/K = 97

SC(7120) elem./compound molecular weights

Sn: MW = 150.7 kg/kg-mole

Table A-5 Standard MELCOR modeling practices, modeling parameters, and sensitivity 

coefficients for analysis of severe accidents (continued)
A-6

Item Record Field Value(s) used in 

SOARCA Description

33. PWR control 

rod RN data for 

RN Classes 18, 

19, and 20

For PWR RN Class 18 = Ag, 19 = In, 20 = 

Cd

Define Ag, In, Cd (DCHCLSnnn0)

18 = ‘Ag-CR’, 19 = ‘In-CR’, 20 = ‘Cd-CR’

Ag, In, Cd decay heats (DCHNEMnn00) 

0 W/kg (no decay heat)

SC(7110) vapor pressures

Ag: Log10(P(mm Hg)) = 1000/T + 1.26x104 + 

7.989

In: Log10(P(mm Hg)) = 400/T + 1.27x105 + 

8.284

Cd: Log10(P(mm Hg)) = 500/T + 5.31x103 + 7.99

SC(7111) diffusion coefficients

Ag: Sigma = 3.48, E/K = 1300

In: Sigma = 3.61, E/K = 2160

Cd: Sigma = 3.46, E/K = 1760

SC(7120) elem./compound molecular weights

Ag: MW = 107.8 kg/kg-mole

In: MW = 114.8 kg/kg-mole

Cd: MW = 112.4 kg/kg-mole

34. RNCA100 ICAON 1 (Active) Chemisorption model is active (default).

35. RN1002 IHYGRO 1 (Active) Hygroscopic model activation. 

(RNACOND set to default, 0 = 

condensation of water onto all aerosols.

36. RNCRCLxx

SC7100

ICRMT/ 

ICLSS/ 

FRAC

(2) Zr

(3) ZrO2

(4) steel

(5)steel ox.

(6) B4C

2 / 18 / 0.0145

3 / 18 / 0.0145

0.1

1.0

0.0

0.0

0.0

For BWRs, apply the non-fuel release 

model. Assign aerosol generated from Zr 

and ZrO2 to RN Class 18 (SnO2). The 

mass will be added as a non-radioactive 

mass to this class. The fraction of material 

mass available for release as an aerosol 

from these materials is 0.0145 (Sn fraction 

in Zirc-2 and -4.) 

Note: must also add input for the release 

rate (SC7103) for RN Class 18. Values 

should be identical to those used (default) 

for Class 12 (fission product Sn).

Multipliers for various structural material 

types

Table A-6 Standard MELCOR modeling practices, modeling parameters, and sensitivity 

coefficients for analysis of severe accidents (continued)
A-7

Item Record Field Value(s) used in 

SOARCA Description

37. RNFPNijjXX NINP

RINP1

RINP2

Use ORIGEN 

results, if available.

NINP = RN Class, RINP1 = mass, RINP2 = 

axial peaking factor. Distributes mass 

based on distribution developed with 

ORIGEN.

If ORIGEN results are unavailable, NINP = 

0, RINP1 = axial peaking factor, RINP2 = 

radial peaking factor. Where, 

ΣiΣj RINP1i * RINP2j = 1.

38. RNGAPijjnn NINP

RINP1

RINP2

1 (Xe) = 0.05

2 (Cs) = 1.00

3 (Ba) = 0.01

5 (Te) = 0.05

16 (CsI) = 0.05

Where, NUREG-1465 recommends the 

following gap quantities,

• Xe = 5%

• Cs = 5%

• Ba = 1%

• Te = 5%

39. RN2FLTXX00 FPVAPOR Various geometric 

values

For all flow paths entering pools via 

quenchers or spargers, specify the flow 

path to scrub all gaseous RN classes.

Table A-7 Standard MELCOR modeling practices, modeling parameters, and sensitivity 

coefficients for analysis of severe accidents (continued)

B–1

Appendix B Updated Default Parameters

In review of the various input options that were modified for SOARCA applications, several user 

input options warranted a permanent adjustment to the default MELCOR value. This Appendix 

presents the new defaults, Table B-1, which reflect long term practices to better model severe 

accident phenomena, improve numerical robustness, or activate newer models.

Table B-1 Summary of Updated Default Parameters

# Description Parameter(s) Field(s)

Value(s) used 

in SOARCA

Current 

Default 

Value(s)

1 COR package candling 

heat transfer coefficient. COR00005

HFRZUO

HFRZZR

HFRZSS

HFRZZX

HFRZSX

HFRZCP

7500 W/m2-K

7500 W/m2-K

2500 W/m2

-K

7500 W/m2

-K

2500 W/m2

-K

2500 W/m2-K

1000 W/m2-K

1000 W/m2-K

1000 W/m2

-K

1000 W/m2

-K

1000 W/m2

-K

1000 W/m2-K

2 COR package radiation 

heat transfer parameters COR00003 FCELRFCELA

0.1

0.1

0.25

0.25

3

COR package min. 

porosity for flow and 

heat transfer

SC1505 (1)

(2)

0.05

0.05

0.001

0.001

4 COR package min. 

CVH volume fraction SC4414 (1) 0.01 0.001

5 COR package 1-dim. 

stress/strain distribution SC1600 (1) 1.0 0.0

6 COR package min yield 

stress temperature SC1603 (2) 1700.0 K 1800.0 K

7

COR package temp. for 

enhanced debris to 

lower head conduction

SC1250 (1) 2800.0 K 3200.0 K

8

CVH/FL direct versus 

iterative solution 

algorithm

SC4415 (1) 1.0 0.5

9 HS temperature 

convergence criterion SC4055 (2) 0.5 5.0x10-4

10

CAV package emissivity 

of oxide, metallic, and 

surrounding materials

CAVnnak

EMISS.OX

EMISS.MET

EMISS.SUR

0.9

0.9

0.9

0.6

0.6

0.6

11

Multipliers for surface 

boiling heat transfer and 

material (oxide/metallic) 

conductivity

CAVnnak

BOILING

COND.OX

COND.MET

10.0 

(multiplier)

5.0

5.0

CORCONMod3

1.0

1.0

12

DCH package default 

classes – new default 

class 17 (Cs2MoO4)

* arrays 

initialized with 

17 classes

* arrays 

initialized with 

16 classes
B–2

# Description Parameter(s) Field(s)

Value(s) used 

in SOARCA

Current 

Default 

Value(s)

13 RN class 17 physical 

properties

SC7120

SC7120

SC7170

SC7170

SC7170

(1,17)

(2,17)

(3,17)

(4,17)

(9,17)

351.75 kg/kgmole

425.75 kg/kgmole

0.67 kg/kgH2O

0.67 kg/kgH2O

4030.0 kg/m3

28.97 kg/kgmole

28.97 kg/kgmole

0.0 kg/kg-H2O

0.0 kg/kg-H2O

1000.0 kg/m3

Table B-2 Summary of Updated Default Parameters (continued)
C–1

Appendix C MELCOR Validation Test Suite

A list of the assessment analyses performed for the MELCOR severe accident analysis code is

presented below. The MELCOR assessment is comprised of analyses performed to simulate 

test series, integral and separate effects, as well as the TMI accident to demonstrate sufficient 

characterization of the relevant phenomena. 

• ABCOVE: AB5 & AB6 [31]

o General Description:

 Simulation of the dry atmosphere conditions of a liquid metal fast breeder 

reactor containment with a sodium fire, i.e., sodium combustion product 

aerosols. AB6 modeled fission product aerosols, sodium iodide, in the 

presence of sodium combustion product aerosol.

o Important Physics: 

 Agglomeration behavior of two aerosol species (hygroscopic and nonhygroscopic), condensation of water vapor

o Results and Findings:

 MELCOR adequately predicts deposition of the sodium combustion, 

NaOX, aerosols.

 MELCOR over predicts aerosol depletion of sodium iodide in AB6, 

possibly due to lack of resuspension modeling.

• ACRR: MP1 & MP2 [32]

o General Description:

 Tests investigated late phase core melt progression and examined 

material interactions and rod degradation for an intact rod / dense Zr-UO2

crust / rubblized debris bed geometry.

o Important Physics:

 Heat transfer in a degraded core geometry, core degradation, material 

interactions

• AHMED: AMMD

o General Description:

 A series of hygroscopic aerosol experiments were conducted at the 

AHMED Test Facility by injecting NaOH in aerosol form into an 

atmosphere with controlled humidity. 

o Important Physics:

 Hygroscopic effects under differing humidity conditions and the impact on 

aerosol masses available for release

• Bethsy-6.9c (ISP-38) 

o General Description:

 The purpose of the Bethsy test was to study the accident transient 

following the loss of the RHR during mid-loop operation with the primary 

circuit open at the pressurizer and steam generator outlet plenum 

manways. The Bethsy facility is a three-loop PWR core and primary 

circuit, with the elevations scaled 1:1 and the volume scaled to 1:100.

o Important Physics:

 Entrainment and retention of water in the pressurizer caused by steam 

flow through the pressurizer manway

 Low pressure pool boiling
C–2

 Level swell in the upper head

 Expulsion of water through the steam generator manway

 Level of pressurization

 Reflooding of the core from the gravity and forced emergency core 

cooling water injection

• CORA-13 (ISP-31) [61]

o General Description:

 CORA-13 permitted analysis of the heat-up and meltdown phases of a 

PWR type fuel element in the CORA test facility. The CORA facility 

consists of a fuel rod bundle with heated and unheated rods under 

controlled thermal-hydraulic boundary conditions with a steam supply to 

provide superheated steam and a quench capability.

o Important Physics:

 Oxidation/hydrogen generation, fragmentation of rods, relocation of core 

materials, formation of blockages, forced convection, conduction, 

radiation, and fluid-structure heat transfer

o Results and Findings:

 MELCOR was unable to predict significant hydrogen production during 

the quench phase. MELCOR does not have any models to simulate 

quench-induced fracturing of the otherwise protective oxide layer on the 

cladding surface. Some experiments in the QUENCH facility suggest that 

such fracturing can result in high transient oxidation rates owing to the 

exposure of fresh metallic Zircaloy following cool-down fracturing of the 

oxide layer.

• CSE-A9 [33]

o General Description:

 Eight experiments have been performed in the CSE containment vessel 

to evaluate the performance of aqueous sprays as a means of 

decontaminating containment atmospheres. 

o Important Physics:

 Cesium and uranium aerosol and iodine vapor washout by sprays. 

aerosol depletion by gravity, thermal-hydraulic response to containment 

sprays

o Results and Findings:

 MELCOR does an adequate job of predicting the thermal/hydraulic 

response to the spray injection with differences explained by the 

fundamental assumptions of fully mixed spray and that the droplets fall 

through a volume atmosphere at rest. This leads to a slight 

overestimation of steam condensation.

• CVTR: Test 3, 4, and 5 [34]

o General Description:

 Design basis simulation of a postulated main steam-line break (MSLB) 

inside a large dry PWR containment was performed. The 

Carolinas-Virginia Tube Reactor (CVTR) facility is a decommissioned 

reactor containment building.

o Important Physics:

 Multi-component gas compression/expansion, thermal/hydraulic response 

to containment sprays, atmosphere cooling by fan cooler, jet-plume gas 
C–3

interaction, buoyancy/stratification, 1-D heat transfer to HS, free 

convection, forced convection

• DEMONA: B3 

o General Description:

 Investigation of the transport and deposition behavior of aerosols in the 

containment was performed. The test was performed in the Battelle 

model containment (total volume 640 m3

) using an open (quasi one-room) 

geometry and condensation aerosols from a plasma torch generator.

o Important Physics:

 Effects of steam condensation on aerosol settling

• DF: DF4 [35]

o General Description:

 The purpose of the Damaged Fuel, or DF, series of experiments was to 

investigate core melt progression. This experiment investigated the 

behavior of BWR-type fuel materials and configurations in a hightemperature oxidizing environment typical of the conditions during a 

LOCA.

o Important Physics:

 Eutectic interaction between the control poison material (B4C) and the 

stainless steel control blade sheath and tubes, and the oxidation of 

Zircaloy in the cladding and canister

• FALCON: 1 & 2 (ISP 34) 

o General Description:

 Heating of a bundle of six fuel specimens and six absorber specimens in 

steam-helium environment containing boric acid was performed. The ISP 

provided information concerning deposition along a controlled thermal 

gradient tube and containment structure.

o Important Physics:

 Physical and chemical behavior of fission products under simulated 

severe accident conditions and multi-component aerosol effects, vaporaerosol interactions, thermophoretic deposition

• FLECHT-SEASET (Natural Circulation) [42]

o General Description:

 The facility design is scaled to a typical Westinghouse PWR on a 1:307 

volume basis, with prototypic full lengths and full heights. The loop piping 

consists of two flow paths representing the unbroken, or intact, three 

loops and the broken loop of a 4-loop PWR. However, for the natural 

circulation tests the broken loop is not connected to a containment tank, 

simulating a break, but is connected to the downcomer extension to 

provide a normal, uninterrupted, flow path from the upper plenum through 

the steam generator, the loop pump seal and the cold leg to the 

downcomer.

o Important Physics:

 Pool boiling in core, natural circulation, steam condensation and reflux

• FPT1 (ISP 46) 

o General Description:
C–4

 The FPT-1 system consisted of an in-pile fuel bundle assembly and upper 

plenum region, an external circuit including a steam generator U-tube and 

connecting lines, and a containment section. The objective of the fuel 

bundle assembly was to assess fuel degradation and fission product 

release from a degraded fuel assembly. In the circuit, the objective was 

to determine fission product transport and deposition in steam generator 

tubes.

o Important Physics:

 Thermal modeling was assessed from thermocouple responses and 

temperature profiles 

 Oxidation (thermocouple responses and measurements of hydrogen 

generation)

 Material relocation (thermocouple and radiography and transmission 

tomography for the end state)

 Fission product release, transport, and deposition (Emission tomography 

of the fuel bundle and steam generator as well as measurements of 

activity along the external line to the containment)

o Results and Findings:

 Fuel and clad temperatures very close

 Hydrogen generation rate close

 Fission product release timing and amount close (Xe, I, Cs)

• Modified CORSOR-Booth model

 Fission product deposition in hot leg also close (new Ag-In-Cd control rod 

poison release model)

• GE Mark III Suppression pool

o General Description:

 Purpose of test was to obtain validation data for Mark III suppression pool 

vents during DBA.

o Important Physics:

 Vent clearing times, pressures in drywell/wetwell, LOCA DBA conditions

• GE Level Swell [36]

o General Description:

 A number of blowdown tests were conducted, some with blowdown 

occurring near the top of the vessel (vapor blowdown) and others with 

blowdown occurring near the bottom of the vessel (liquid, two-phase, 

vapor transient). These experiments were conducted in the “large 

blowdown vessel” (4.5 m3

).

o Important Physics:

 Vessel blowdown, level swell, critical flow

o Results and Findings:

 Level swell is better predicted by a single control volume than from a 

finely subdivided stacked volume

 Blowdown flow and vessel depressurization are strongly dependent on 

the break discharge coefficient used

 Discrepancies in top blowdown calculations due to maximum allowed 

pool bubble fraction, C4407(11)

• SNL/IET: IET-1, 3, 6 [37]

o General Description:
C–5

 Series of experiments performed at the Surtsey test facility at Sandia 

(1:10 linear scale) and in the corium-water thermal interactions (CWTI) 

test facility at ANL (1:40 linear scale), to evaluate the effects of high 

pressure melt ejection (HPME) on DCH.

o Important Physics:

 High pressure melt ejection, DCH, oxidation and hydrogen generation, 

hydrogen combustion

• RTF (ISP41) [61]

o General Description:

 Objective was to develop data on the behavior of iodine in reactor 

containment pools. The experiment consisted of a pool in a stainless 

steel vessel with a radioisotope dose source and aqueous iodine provided 

by adding CsI to the pool. The pool pH was controlled during the 

experiment by adding acid and base chemicals to the pool.

o Important Physics:

 Speciation of iodine in the aqueous and gaseous phases, effect of 

radiation on H2O2 and H2 concentrations, adsorption/desorption of iodine 

on surfaces

• Quench-6 (ISP45) 

o General Description:

 The objective of the Quench-6 test was to assess the capability of severe 

accident codes to simulate delayed reflood situations in which a preoxidized LWR fuel rod bundle is quenched by water inserted from the 

bottom.

o Important Physics:

 Oxidation of metallics, bottom reflood cooling

• JAERI Spray Tests: PHS-1, 6 

o General Description:

 Pressure suppression spray tests were conducted in Japan during the 

late 1970s in a 700 m3 steel vessel (20 m high, 7 m in diameter). PHS-6 

was a single nozzle test where PHS-1 was a 6 nozzle test. Vessel walls 

are hot so that droplets contacting walls are vaporized, degrading spray 

effectiveness.

o Important Physics:

 Containment pressure reduction by sprays

o Results and Findings:

 Comparison of single cell and multi-cell models confirm appropriate 

treatment of spray droplets falling through stacked volumes. 

 Calculation sensitive to assumed spray/vessel contact.

• LACE Turbulent Deposition: LA1 & LA3 

o General Description:

 The LACE LA1 and LA3 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. Specific 

objectives of these tests were to provide validation data that would 

expose important dependencies in modeling deposition. The effects of 

gas velocity, aerosol composition and aerosol size were considered.
C–6

o Important Physics:

 Turbulent deposition of aerosols in pipes and deposition of aerosols in 

pipe bends

• LACE: LA4 [61]

o General Description:

 The purpose of the experiment was to determine the disposition of 

aerosols in the containment building under conditions of high steam 

concentrations. Of particular interest was the difference in aerosol 

disposition between hygroscopic (water-soluble) aerosols such as CsOH 

and nonhygroscopic aerosols such as MnO in a high steam 

concentration.

o Important Physics:

 Hygroscopic effects, deposition of aerosols on surfaces, heat transfer to 

surfaces, steam condensation on surfaces

o Results and Findings:

 Calculates pressures very well but slightly over predicts pool temperature

 Aerosol removal of hygroscopic CsOH well calculated until late in test

 Prior to venting aerosol removal of non-hygroscopic MnO well-predicted, 

however after, MELCOR under predicts aerosol removal 

 Early in the calculation, the dynamic film tracking overestimates film 

drainage from heat structures. The quasi-steady thickness provides a 

better calculation of pool mass.

• LOFT: LP-FP-2 

o General Description:

 Experiment LP-FP-2 models the V-sequence accident, defined as a 

rupture in a low pressure injection system line outside the containment 

with simultaneous failure to isolate the system. The experimental 

subsystems include the reactor vessel, the intact loop, the broken loop, 

the blowdown suppression tank system, and the emergency core cooling 

system.

o Important Physics:

 Heat conduction/convection (temperatures and pressures measured),

hydrogen generation (hydrogen mass measured), fission product release, 

flow blockage in degraded core, break flow, choked flow

• MARVIKEN: ATT-4 [40]

o General Description:

 Test ATT-4 studied fission product transport in the presence of a 

structural aerosol simulant; in addition to the fissium aerosol, a “corium” 

aerosol was produced that was composed of Ag and Mn. Corium vapors 

were mixed with vaporized fissium and steam in the lowest portion of the 

reactor vessel to form aerosols which were transported through the 

simulated large-scale primary piping.

o Important Physics: 

 Thermal hydraulics of a PWR, aerosol and vapor transport and deposition

• MARVIKEN Blowdown Tests: CFT-21 & JIT-11 

o General Description:
C–7

 Large scale tests intended to provide data for analysis of critical flow from 

vessel blowdown were performed at the Marviken facility. The CFT-221 

test was designed for validation of subcooled and two-phase flow through 

a discharge nozzle whereas JIT-11 tested a saturated steam flow. 

o Important Physics:

 Vessel blowdown, critical flow of vapor, subcooled liquid, and two-phase 

flow

o Results and Findings:

 Vapor critical flow well calculated by MELCOR (sonic flux as minimum 

section)

 Subcooled liquid critical flow well calculated by Henry-Fauske critical flow 

model

 Moody model predicts a larger flow rate than observed for two-phase flow

• NEPTUN: 5006 & 5007 

o General Description:

 The NEPTUN experiments were designed to measure the rate of boil-off 

and additionally, the heat-up of fuel rods during two-phase uncovery of 

the core in a severe accident.

o Important Physics: 

 Boil off, fuel rod heat-up, level swell

o Results and Findings:

 Level swell calculated for stacked control volumes highly dependent on 

nodalization (see also General Electric (GE) Level Swell)

• NTS Burn: NTSP01, NTSP12, NTSP15, NTSP20 

o General Description:

 Premixed hydrogen combustion experiments with hydrogen 

concentrations ranging from 5 to 13% (by volume) and steam 

concentrations from 4 to 30% were performed. 

o Important Physics:

 Combustion burn completeness, burn time, vessel pressurization

o Results and Findings:

 MELCOR slightly over predicts the burn completeness for oxygen limited 

tests, likely because it is assumed a homogeneous mixture in a single 

control volume. Consequently, MELCOR predicts slightly higher peak 

pressures and temperatures, for these cases.

 Because MELCOR is a lumped parameter code, it does not predict burn 

times and flame propagation well

• NUPEC: M-8-1, M-8-2 [61]

o General Description:

 The tests explored the response of a 1:4 scale containment to steam 

injection and containment spray actuation (M-8-2) with helium as a 

surrogate for hydrogen gas.

o Important Physics:

 Pressure response, temperature distribution and stratification, and 

hydrogen mixing

o Results and Findings:

 Without sprays, MELCOR calculation predicted higher pressures (10%) 

and temperatures in dome
C–8

 Tracked helium concentrations in the dome and upper compartments well

 In lower compartments, calculation predicted decreased mixing after 15 

minutes where data showed strong mixing throughout test

• PBF SFD: 1-4 

o General Description:

 The Severe Fuel Damage (SFD) tests were performed at the Power Burst

Facility (PBF) to investigate fuel rod and core response as well as the 

release of fission products and hydrogen generation during degraded 

core accidents.

o Important Physics:

 Fission product release

 Oxidation and hydrogen generation

• PHEBUS: B9+ (ISP 28) [61]

o General Description:

 The B9+ test was designed to provide data principally on fuel 

degradation. It consists of a driver reactor core to provide neutronic 

heating to the test bundle, a fluid supply system to inject steam and 

helium into the test bundle, and associated cooling systems for the 

bundle and driver core. 

o Important Physics:

 Heat conduction/convection (temperatures measured), hydrogen 

generation (hydrogen mass measured), and fuel degradation (no direct 

measurement) 

o Results and Findings:

 Excellent agreement with thermocouple data

 Excellent agreement with hydrogen generation

 Improvements with MELCOR Versions 1.8.6 and 2.1

• PNL Ice Condenser: 11-6 & 16-11 [41]

o General Description:

 A series of large-scale experiments conducted at the High Bay Test 

Facility at Pacific Northwest Laboratory (PNL) to investigate the extent to 

which an ice condenser may capture and retain air-borne particles. In 

Experiment 11-6, the low flow rate induced a natural circulation flow 

between the diffuser outlet and the ice condenser. Experiment 16-11 was 

performed with every compartment full of ice and was a high flow test with 

no recirculation.

o Important Physics:

 Aerosol deposition, heat transfer in ice condenser containment, natural 

circulation, ice phase transition

o Results and Findings:

 MELCOR captured well, the thermal response of the experiment

• The HS characteristic length should reflect the physical diameter 

of a typical ice cube in the condenser.

• HS heat transfer coefficient is design specific

• Because the model is not a moving boundary model, it is 

recommended that only 2 nodes are modeled

• Model must be nodalized to capture natural convection
C–9

• TEMPL should be set to 274 K (rather than 273.15 K) to avoid 

numerical problems with CVH when temperatures are near 

freezing

 MELCOR adequately captured the limited data for aerosol retention

• Aerosol retention calculated is extremely sensitive to particle 

density

• STORM: SR-11 (ISP 40) 

o General Description:

 The Simplified Test of Resuspension Mechanisms (STORM) test SR-11, 

was intended for examining aerosol deposition and resuspension in pipes 

and included two distinct phases: (1) the aerosol deposition by 

thermophoresis and eddy impaction, and (2) aerosol resuspension under 

a stepwise increasing gas flow. MELCOR does not have a resuspension 

model and the second phase was not modeled.

o Important Physics:

 Aerosol deposition from thermophoresis and eddy impaction as well as 

resuspension (not modeled in MELCOR)

• SURC: SURC-1 & SURC-2 

o General Description:

 The Sustained Urania-Concrete (SURC) experiments were designed to 

measure and assess releases due to interactions between core materials 

and concrete in containment structures.

o Important Physics:

 Ablation of concrete, release of reactant gases, temperature response

o Results and Findings:

 Reasonable prediction of concrete ablation front

• MCCI: CCI-1 & CCI-2

o General Description:

 The CCI Phase 1 experiments were designed to measure concrete 

ablation with different types of concrete.

o Important Physics:

 Ablation of concrete, release of reactant gases, temperature response

o Results and Findings:

 Reasonable prediction of concrete ablation front

• TMI-2

o General Description:

 Though not an experiment, the TMI-2 accident serves as an excellent 

resource for code validation. 

o Important Physics:

 The accident conditions stress the capabilities of the code for predicting 

core degradation, formation of a debris bed in the upper core, formation 

of a molten pool in the core, relocation of molten corium to the lower 

plenum, the response of the lower head, and reflood and quench of the 

degraded core.

• VANAM: M3 (ISP 37) [61]

o General Description:
C–10

 The objectives of the VANAM-M3 test was to provide data on 

containment-building response to severe accident conditions with 

particular emphasis on characterizing the depletion rate of hygroscopic 

aerosol under varying humidity and thermal-hydraulic conditions.

o Important Physics:

 Multi-compartment geometry, stratified atmosphere, atmosphere mixing 

by forced convection loops, thermal energy balance, structural heat 

transfer, steam condensation effects, and aerosol behavior

o Results and Findings:

 Calculated room temperatures indicated code’s ability to correctly capture 

the forced and natural circulation patterns that occur at different times in 

the test

 Calculated pressure behavior demonstrated the ability of the code to 

calculate the response of a multi-room building to sources and sinks of

steam and air

 Aerosol depletion results demonstrated adequacy of aerosol growth 

modeling, aerosol transport, and depletion from gravitational settling
D-1

Appendix D MELCOR Code Version Progression Overview

The discussion regarding the validation history for MELCOR was focused solely on the 

experimental comparisons between various MELCOR versions as well as side-by-side 

comparison between MELCOR versions to demonstrate evolution in models or consistency 

during the MELCOR development. As development progressed, MELCOR has received 

various updates which were not presented directly alongside with the validation discussion. 

This Appendix presents a summary of the major modification made to MELCOR for the major 

version milestones since MELCOR Version 1.8.2.

MELCOR 1.8.3

• CORCON-MOD3 (including VANESA) was added to MELCOR to replace the separate 

CORCON-MOD2 and VANESA models.

MELCOR 1.8.4

• Previous versions of MELCOR were known to predict too-early collapse of reactor cores. 

A model for retention of molten metals behind oxide shells (particularly, molten Zircaloy 

on fuel rods), with ultimate failure by another mechanism was added to correct that 

behavior.

• A creep rupture model was added for the lower head, together with the capability to 

model external cooling of the lower head in a flooded cavity.

• A “flow blockage” model was added to account for redistribution of flow through a reactor 

core as a result of changed flow resistance when intact geometry is lost and a debris 

bed or pool forms.

• A capability was added to calculate radiative heat transfer between pairs of heat 

structure surfaces.

• Models were added for the behavior of hygroscopic aerosols and for the chemisorption 

of Cs onto the surfaces of heat structures.

• The SPARC 90 pool decontamination model replaced a previous “preliminary” version of 

SPARC (SPARC 87). In addition to other improvements, the new model includes 

removal of iodine vapor.

MELCOR 1.8.5

• A diffusion flame model was added to calculate the combustion of hydrogen flowing 

through flow paths during direct containment heating.

• Previous versions of MELCOR required the use of a single component (called “Other 

Structure”, OS) to represent all support structures, control structures, and miscellaneous 

structures in the core in addition to fuel rods and BWR canisters. This approach had 

serious deficiencies, and none of the structures could be realistically represented. New 

components referred to as “Supporting Structure” (SS) and “Non-supporting Structure” 

(NS) were introduced. Both parametric and mechanistic, load-based, failure models 

were added for SS, which can support other core components. NS is subject to simpler 

failure models, but these have sufficient flexibility to represent BWR control blades, PWR 

control rods, and structures such as filler rods in experiments.
D-2

• Optional models were added for convective heat transfer to water pools from the top and 

bottom surfaces of SS plates, and for radiative heat transfer between the bottom of such 

a plate and the water pool or lower head below it.

• Previous versions of MELCOR did not properly differentiate between debris in the 

channel and debris in the bypass of a BWR. This was resolved by introduction of a 

“Particulate debris in the Bypass” (PB) component. After failure of the fuel canisters in a 

BWR that separate PD from PB, the two debris fields were allowed to mix and 

equilibrate. A debris exclusion model (with flexible user control) was implemented to 

control the relocation of particulate debris (PD and/or PB) based on the presence or 

absence of intact structures that could prevent it (for example, solid debris cannot, as a 

general rule, enter the small spaces between fuel rods.)

• The flow blockage model was much improved as a result of the ability to distinguish 

particulate debris in the channel of a BWR from that in the bypass, and a model was 

added to allow the opening of a flow path on failure of a channel box (canister).

• Improvements were made to the implementation of candling and debris slumping models 

and to those for conductive, radiative, and candling heat transfer.

• Cesium iodide was added as a default class.

• Substantial improvements were made in the model for hygroscopic aerosols.

• A model for the chemical behavior of iodine in water pools was added to MELCOR. It 

includes models for pH, including transport of nitric and hydrochloric acid formed by 

radiolysis off air and plastic in cables, respectively. The effects of different surface 

coatings on containment structures are also modeled.

MELCOR 1.8.6

• Flexibility was added to allow the user to enhance quenching of ejected debris through 

conductivity multipliers.

• New LM-CREEP and PIPE-STR CF types were added to make it far less difficult for 

users to model pipe ruptures.

• Modeling of the lower plenum and head was heavily revised. The curvature of the head 

and its effect on lower plenum volumes can now be consistently modeled (this was not 

possible in previous versions). The head can take the form of a cylinder, hemisphere, or 

hemispherical segment. Heat transfer and failure models were improved. Because the 

new model contains all of the capabilities of the separate BH package, this package was 

eliminated.

• Models were added for formation of stratified molten debris pools, both in the core and in 

the lower plenum. These include circulation-driven convective heat transfer in the case 

of coherent pools.

• A core periphery model was introduced for PWRs to allow proper modeling of the core 

baffle (shroud) and core formers, and the bypass region between the baffle and the core 

support barrel. Such modeling was impossible in previous versions of MELCOR.
D-3

• A model was added for quenching of core structures by reflood of water from below. A 

model was also added to evaluate oxidation of the submerged but unquenched surfaces 

that could be predicted by this model.

• More realistic models were added for behavior of control poison in a PWR. One involves 

oxidation of B4C control poison, the other models release of AgInCd control poison, 

including formation of aerosols.

• The local fluid temperature model (also known as “dT/dz” was improved to reduce 

problems with small stagnant volumes that had forced falsification of geometry in some 

previous input decks.

• Treatment of support structures modeling columns was modified to allow better 

representation of the support in a typical PWR. The package now allows user-defined 

“flavors” of support structures, allowing further flexibility.

• A model was added to calculate breakaway oxidation of Zircaloy in air.

• The default modeling of collapse of BWR canisters was modified; previous code 

versions predicted survival to unreasonably high temperatures.

• The previous approach to specification of the inner and outer areas of BWR canisters 

did not always allow a correct representation; these areas may now be directly input.

• Current best practices for modeling reactor cores involve reducing the melt temperatures 

used for ZrO2, UO2, and B4C from handbook values to account for the fact that they do 

not typically appear as pure materials. Redefinition of all the tables and other MP input is 

tedious and time consuming, particularly if one is interested in the effects of changes in 

these reductions. So-called “interacting materials”, ZRO2-INT, UO2-INT, and B4C-INT 

were added to the MP package. Their properties differ from those of the pure materials 

only in the melt temperatures. Initialization in MELGEN uses the modified melt 

temperature to generate complete and consistent properties tables from those of the 

pure materials; the melt temperature of any of these materials can be modified from its 

default value with a single input record.

• Creep data have been added to the MP package.

• The user can now specify (via Control Functions) the failure criteria for COR 

components and add arbitrary heat sources in and heat transfer paths between them. 

• A new fuel collapse model was added to allow a user to supply a time at temperature 

lifetime failure table to determine rod collapse. Fuel collapse was previously specified by 

a failure temperature.

• The user has more control over the flow resistance calculated from the Ergun equation 

in the flow blockage model to account for phenomena such as fuel swelling.

• New sensitivity coefficients are available to modify the surface emissivities used in the 

radiation model.
D-4

• A new, optional model was added to treat flashing of superheated water entering a 

volume, either through a flow path or from a volume source. It improves the partition of 

mass and enthalpy of the water between the volume pool and atmosphere, and includes 

the formation of water aerosols.

• Previous versions of MELCOR could exhibit unphysical behavior if the volume of either 

hydrodynamic field (pool or atmosphere) in a volume became very small or was absent. 

A new thermodynamic model has been added to better model these situations.

• Heat structure surfaces may be partially covered by a water pool. In previous code 

versions, a rising pool surface acted as a “squeegee”, increasing the film thickness by 

maintaining its total mass rather than subsuming the covered portion of the film. This 

was corrected in MELCOR 1.8.6.

• Mechanistic models are used for draining of thick films from surfaces that are involved in 

a “film tracking network”, but previous versions of MELCOR used a simple maximum film 

thickness to remove water from too-thick films on other surfaces. The mechanistic 

treatment is now used universally, replacing the maximum limit on isolated structure 

surfaces. In effect, an isolated structure surface is treated as if it form a film tracking 

network of length one.

• A new version of the CORSOR Booth release model corrects an obvious error in the 

previous implementation based on published material from Battelle Columbus.

• The algorithm previously used to interpolate aerosol agglomeration and deposition 

kernels could lead to significant errors if a calculation spanned a wide range of 

temperatures and/or pressures. A much-improved interpolation algorithm has been 

implemented.

• The aerosol filter model was extended to allow specification of decontamination factor by 

particle size as well as by class. Each decontamination factor may now be defined by 

either a constant or a control function.

• A model was added to calculate deposition of water from a jet impacting on the surface 

of a heat structure.

• New turbulent deposition models were added along with bend impaction model.



UNITED STATES

NUCLEAR REGULATORY COMMISSION

WASHINGTON, DC 20555-0001

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OFFICIAL BUSINESS
NUREG/CR-7008 MELCOR Best Practices as Applied in the State-of-the-Art Reactor 

Consequence Analyses (SOARCA) Project