Unclassified

NEA/CSNI/R(2009)3

Organisation de Coopération et de Développement Économiques
Organisation for Economic Co-operation and Development

03-Aug-2009________________________________________________________________________________________________________English text only
NUCLEAR ENERGY AGENCYCOMMITTEE ON THE SAFETY OF NUCLEAR INSTALLATIONSNEA/CSNI/R(2009)3
Unclassified
ABILITY OF CURRENT ADVANCED CODES TO PREDICT CORE DEGRADATION, MELTPROGRESSION AND REFLOODING
Benchmark Exercise on an Alternative TMI-2 Accident Scenario

English text only

JT03268437

Document complet disponible sur OLIS dans son format d'origine
Complete document available on OLIS in its original format

NEA/CSNI/R(2009)3

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NEA/CSNI/R(2009)3

Ability of Current Advanced Codes to
Predict Core Degradation, Melt Progression
and Reflooding
Benchmark Exercise on an
Alternative TMI-2 Accident Scenario

Participating Organizations and Authors

ENEA
GRS
IKE
IRSN
IVS
SNL
SNU
UPI

Giacomino Bandini
Henrique Austregesilo
Michael Buck
Olivier Marchand & Florian Fichot
Miro Branak & Peter Matejovic
Larry Humphries
Kune Suh
Sandro Paci

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TABLE OF CONTENTS

1.

ABBREVIATIONS ............................................................................................................................... 8

2.

LIST OF FIGURES ............................................................................................................................... 93.

EXECUTIVE SUMMARY ................................................................................................................. 13
3.1
3.2
3.3
3.4
3.5

Background ................................................................................................................................ 13
Objectives .................................................................................................................................. 13
Definition of the Alternative Scenario and Analysis of Benchmark Results ............................. 13
Main Results .............................................................................................................................. 14
Conclusions and recommendations ............................................................................................ 14

4.

INTRODUCTION ............................................................................................................................... 17

5.

PLANT DESCRIPTION AND ALTERNATIVE TRANSIENT DEFINITION ................................ 19
5.1
5.2
5.3
5.4
5.5

6.

GEOMETRY ....................................................................................................................................... 23
6.1
6.2
6.3
6.4

7.

Free volumes .............................................................................................................................. 23
Heat transfer with secondary side .............................................................................................. 23
Initial core geometry .................................................................................................................. 23
Vent valves................................................................................................................................. 24

CHOICE OF PHYSICAL PARAMETERS TO COMPARE .............................................................. 25
7.1
7.2

8.

Initiating Event........................................................................................................................... 19
End of calculation ...................................................................................................................... 19
Main Assumptions ..................................................................................................................... 19Power ......................................................................................................................................... 20
Boundary Conditions ................................................................................................................. 21

Parameters selected for comparison........................................................................................... 25
Criteria for the quality of agreement .......................................................................................... 26

COMPARISONS BETWEEN PARTICIPANTS RESULTS ............................................................. 29
8.1
8.2

Initial Steady State ..................................................................................................................... 29
Initial thermal-hydraulic phase up to the stop of main pumps ( ~5000s) .................................. 30
8.2.1 Mass flow rate at the break ........................................................................................... 30
8.2.2 Primary mass ................................................................................................................ 30
8.2.3 Heat transfers with secondary side ............................................................................... 31
8.2.4 Water levels in the core ................................................................................................ 31
8.2.5 SG behaviour ................................................................................................................ 31
8.2.6 Core temperatures ......................................................................................................... 31
8.2.7 Pressurizer level............................................................................................................ 31

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8.3

8.4

8.5
9.

Core degradation phase (up to reflooding) ................................................................................ 31
8.3.1 Water levels in the core ................................................................................................ 32
8.3.2 Mass flow rate at the break ........................................................................................... 32
8.3.3 Primary mass ................................................................................................................ 32
8.3.4 Primary pressure ........................................................................................................... 32
8.3.5 Core temperatures ......................................................................................................... 32
8.3.6 Hydrogen production .................................................................................................... 33
8.3.7 Cladding failure and melt relocation ............................................................................ 33
8.3.8 Fuel melting and relocation .......................................................................................... 33
8.3.9 State of the core ............................................................................................................ 34
Reflooding phase ....................................................................................................................... 34
8.4.1 Water levels in the core ................................................................................................ 34
8.4.2 Primary pressure ........................................................................................................... 35
8.4.3 Core temperatures ......................................................................................................... 35
8.4.4 Hydrogen production- Melt oxidation .......................................................................... 35
8.4.5 State of the core ............................................................................................................ 35
Figures for comparison .............................................................................................................. 36

SENSITIVITY STUDIES ................................................................................................................... 49
9.1
9.2

Selection of parameters .............................................................................................................. 49
Results of sensitivity studies ...................................................................................................... 52

10. CONCLUSIONS ................................................................................................................................. 55
11. PROPOSALS FOR FUTURE EXTENSIONS OF THE WORK........................................................ 58
12. REFERENCES .................................................................................................................................... 59
13. APPENDIX: SCHEDULE OF THE PROJECT .................................................................................. 60
14. APPENDIX: PARTICIPATING ORGANIZATIONS AND CODES USED..................................... 61

14.1 ENEA ............................................................................................................................ 61
14.1.1 Brief Description of the ASTEC V1.3 Code .................................................... 61
14.2 GRS ............................................................................................................................... 62
14.2.1 Brief Description of the ATHLET-CD Code ................................................... 62
14.3 IVS ................................................................................................................................ 63
14.3.1 Brief Description of the ASTEC V1.3 Code ..................................................... 63
14.4 NRC - SNL .................................................................................................................... 63
14.4.1 Brief Description of the MELCOR Code .......................................................... 64
14.5 University Pisa .............................................................................................................. 65
14.5.1 Brief Description of the MELCOR V1.8.5 Code ............................................. 65
14.6 IRSN .............................................................................................................................. 65
14.6.1 Brief Description of the ICARE/CATHARE V2 Code ..................................... 65
14.7 Seoul National University (SNU) ................................................................................. 66
14.7.1 Brief Description of the MAAP4 Code 4.03 ..................................................... 66
14.8 IKE ................................................................................................................................ 67
14.8.1 Brief Description of the ATHLET-CD/MEWA Code ...................................... 68
15. APPENDIX : ENEA CALCULATION WITH ASTEC V1.3 ............................................................ 69
15.1 TMI-2 Plant Modelling .............................................................................................................. 69
15.2 Updated Initial Steady-State Conditions .................................................................................... 70
15.3 Standard Physical Parameters of the Code ................................................................................ 71

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15.4 Chronology of Major Events ..................................................................................................... 72
15.5 RESULT ANALYSIS ................................................................................................................ 73
15.6 Major Discrepancies with Other Code Results .......................................................................... 74
15.7 Synthetic Views of the Core at Selected Instants....................................................................... 75
15.7.1 Standard Case ............................................................................................................... 75
15.7.2 Sensitivity Case ............................................................................................................ 83
16. APPENDIX : GRS CALCULATION WITH ATHLET-CD .............................................................. 89
16.1 TMI-2 Plant Modelling .............................................................................................................. 89
16.2 Main modelling options ............................................................................................................. 90
16.3 Initial steady state conditions ..................................................................................................... 91
16.4 Discussion of Results ................................................................................................................. 92
16.5 Synthetic Views of the Core at Selected Instants..................................................................... 101
16.6 Additional calculations ............................................................................................................ 106
16.7 Final remarks ........................................................................................................................... 106
17. APPENDIX : UPI CALCULATION WITH MELCOR 1.8.5........................................................... 109
17.1 TMI-2 Plant Modelling ............................................................................................................ 109
17.2 Updated Initial Steady-State Conditions .................................................................................. 111
17.3 Standard Physical Parameters of the Code .............................................................................. 111
17.4 Chronology of Major Events ................................................................................................... 112
17.5 RESULT ANALYSIS .............................................................................................................. 113
18. APPENDIX : IVS CALCULATION WITH ASTEC V1.3............................................................... 137
18.1 TMI-2 Plant Modelling ............................................................................................................ 137
18.1.1 Reactor vessel: ............................................................................................................ 137
18.1.2 Circulation loops: ....................................................................................................... 138
18.1.3 Pressurizer and surge line: ......................................................................................... 138
18.1.4 Steam generator – secondary side: ............................................................................. 139
18.1.5 Core, lower plenum and downcomer model:.............................................................. 139
18.2 Initial Steady-State Conditions ................................................................................................ 140
18.3 Physical Parameters Used in the Analysis ............................................................................... 141
18.4 The results of the analysis ........................................................................................................ 142
19. APPENDIX : NRC-SNL CALCULATION WITH MELCOR 1.8.6 ................................................ 159
19.1 TMI-2 Plant Modeling ............................................................................................................. 159
19.2 Geometry ................................................................................................................................. 160
19.2.1 Vent valves ................................................................................................................. 162
19.3 Updated Initial Steady-State Conditions .................................................................................. 162
19.4 Standard Physical Parameters of the Code .............................................................................. 163
19.5 Chronology of Major Events ................................................................................................... 164
19.6 RESULT ANALYSIS .............................................................................................................. 165
19.7 Synthetic Views of the Core at Selected Times ....................................................................... 167
20. APPENDIX : IRSN CALCULATION WITH ICARE/CATHARE V2 ............................................ 169
20.1 TMI-2 Plant Modeling ............................................................................................................. 169
20.2 Physical Parameters of the Code .............................................................................................. 172
20.3 Initial Steady State Conditions................................................................................................. 173
20.4 Analysis of code results ........................................................................................................... 174
20.4.1 The thermal hydraulics phase (from t = 0 s to t = 6250 s) .......................................... 174
20.4.2 The core degradation phase (from t = 6250 s to t = 9831 s) ....................................... 175

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20.4.3 The core reflood phase (from t = 9831 s to 15000 s) ................................................. 175
20.5 Comparisons to other code results ........................................................................................... 176
21. APPENDIX : SNU CALCULATION WITH MAAP4 ..................................................................... 185
21.1 TMI-2 Plant Modelling ............................................................................................................ 185
21.2 Initial Steady State Conditions and Scenarios ......................................................................... 187
21.3 Sensitivity Study ...................................................................................................................... 188
21.4 Chronology of Major Events ................................................................................................... 189
21.5 Results Analysis ....................................................................................................................... 189
21.6 Synthetic Views of the Core .................................................................................................... 191
21.6.1 Standard Case ............................................................................................................. 191
21.6.2 Sensitivity Case .......................................................................................................... 195
22. APPENDIX : IKE CALCULATION WITH ATHLET/MEWA ....................................................... 199
22.1 TMI-2 Plant Modelling ............................................................................................... 199

22.2 Main modelling options .............................................................................................. 199
22.3 Discussion of Results .................................................................................................. 199
22.4 Additional calculations .......................................................................................................... 207

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1. ABBREVIATIONS

AFW
CCFL
CL
CSNI
DC
DOE
EC
ECCS
GAMA
GRS
HL
HPI
IKE
IRSN
ISP
IVS
LOCA
LOFT
MATPRONPPNRCPHÉBUSPWR
QUENCHSAM
SARNET
SB-LOCA
SG
SGTR
SNU
TMI-2
TSO
UPI
US

Auxiliary Feed Water
Counter Current Flow Limitation
Cold Leg
Committee on the Safety of Nuclear Installations
Downcomer
Department of Energy
European Community
Emergency Core Coolant System
Working Group on Accident Management and Analysis
Gesselschaft für Anlagen und Reaktorsicherheit mbH (Germany)
Hot Leg
High Pressure Injection
Institut für Kernenergetik und Energiesysteme - Universität Stuttgart
Institut de Radioprotection et de Sûreté Nucléaire (France)
International Standard Problem
Inzinierska Vypoctova Spolocnost (Slovakia)
Loss Of Coolant Accident
Loss-of-Fluid Test (experimental facility at Idaho National Engineering Laboratory, USA)
Materials Properties correlations and computer subcodes
Nuclear Power Plant
U.S. Nuclear Regulatory Commission
Experimental facility at Cadarache, France
Pressurized Water Reactor
Experimental facility at Forschungszentrum Karlsruhe, Germany
Severe Accident Management
Severe Accident Research Network
Small Break-Loss of Coolant Accident
Steam Generator
Steam Generator Tube Rupture
Seoul National University (Korea)
Three Mile Island – Unit 2
Technical Support Organisation
University of Pisa (Italy)
United States of America

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2. LIST OF FIGURES

5-5:

Fig. 1.SG Level regulation ............................................................................................................. 21
Fig. 2 SG Pressure regulation ......................................................................................................... 21

8.5:

Figures for comparison ................................................................................................................... 36
Break flow rate .......................................................................................................................... 36
Primary mass ............................................................................................................................. 36
Swollen water level ................................................................................................................... 37
Collapsed water level ................................................................................................................ 37
SG A pressure............................................................................................................................ 38
SG A level ................................................................................................................................. 38
SG A power exchanged ............................................................................................................. 39
Temperature at the bottom of the core ...................................................................................... 39
Temperature at mid-core ........................................................................................................... 40
Temperature at the top of the core............................................................................................. 40
Pressurizer pressure ................................................................................................................... 41
Pressurizer level ........................................................................................................................ 41
Total mass of molten materials ................................................................................................. 42
Total mass of molten metal ....................................................................................................... 42
Mass of molten pool .................................................................................................................. 43
Total mass of dissolved UO2 .................................................................................................... 43
Cumulated hydrogen production ............................................................................................... 44
Fraction of non oxidized Zr ....................................................................................................... 44
Fraction of non oxidized Zr in molten materials ....................................................................... 45
Instantaneous Hydrogen Production ......................................................................................... 45
Temperature at the bottom of the core ...................................................................................... 46
Temperature at mid-core ........................................................................................................... 46
Temperature at the top of core .................................................................................................. 47
Temperature at entrance of hot leg ............................................................................................ 47

15-1:
15-7:

TMI-2 plant nodalization scheme ................................................................................................... 69
Synthetic Views of the Core at Selected Instants
Standard case
Beginning of oxidation (6431 s) – Global view ........................................................................ 75
Beginning of oxidation (6431 s) ................................................................................................ 76
Time of first melt relocation (7370 s) – global view ................................................................. 77
Time of first melt relocation (7370 s) ....................................................................................... 78
Time of reflooding (9825 s) – Global view ............................................................................... 79
Time of reflooding (9825 s) ...................................................................................................... 80
End of calculation (12000 s) – Global view .............................................................................. 81
End of calculation (12000 s) ..................................................................................................... 82

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Sensitivity case
Beginning of oxidation (6431 s) – Global view ........................................................................ 83
Time of first melt relocation (7428 s) – Global view ................................................................ 84
Time of reflooding (9825 s) – Global view ............................................................................... 85
Time of reflooding (9825 s) ...................................................................................................... 86
End of calculation (10928 s) – Global view .............................................................................. 87
End of calculation (10928 s) ..................................................................................................... 88
16-1: ATHLET-CD nodalization scheme for TMI-2 (primary circuit) ....................................................... 89
16-2: Pressurizer pressure ........................................................................................................................... 96
16-3: Break mass flow rate .......................................................................................................................... 96
16-4: Total primary liquid inventory ........................................................................................................... 96
16-5: Core collapsed liquid level - inner ring 1 ........................................................................................... 97
16-6: Core collapsed liquid level - outer ring 4 ........................................................................................... 97
16-7: Coolant temperatures in upper plenum .............................................................................................. 97
16-8: Fuel cladding temperatures at top of core (elevation 3.0 m) .......................................................... 98
16-9: Fuel cladding temperatures at mid-core (elevation 1.7 m) ............................................................. 98
16-10: Fuel cladding temperatures at bottom of core (elevation 0.6 m) .................................................... 98
16-11: Total hydrogen production ............................................................................................................. 99
16-12: Total mass of molten materials in the core ..................................................................................... 99
16-13: Total mass of molten Zr in the core ................................................................................................ 99
16-14: Total mass of liquid melt in the core ............................................................................................ 100
16-15: Total mass of refrozen materials in the core................................................................................. 100
16-16: Average axial mass distribution at t = 11500 s ............................................................................. 100
16-17: Secondary pressure ....................................................................................................................... 101
16-18: Secondary level ............................................................................................................................ 101
16-19: TMI core at beginning of oxidation excursion (T~1850 K) ......................................................... 102
16-20: TMI core at time of first ceramic melt relocation......................................................................... 103
16-21: TMI core at start of reflooding ..................................................................................................... 104
16-22: TMI core at the end of calculation (30 min of reflooding) ........................................................... 105
17-1: TMI-2 plant nodalization scheme ................................................................................................. 110
17-2: TMI-2 core simulation with MELCOR ........................................................................................ 110
17-3: Steam generator A level ............................................................................................................... 115
17-4: Steam generator A pressure .......................................................................................................... 115
17-5: Pressuriser pressure ...................................................................................................................... 116
17-6: Main pump A void fraction .......................................................................................................... 116
17-7: Steam generator A removed power .............................................................................................. 117
17-8: Total primary water mass ............................................................................................................. 117
17-9: Break mass flowrate ..................................................................................................................... 118
17-10 Loop A mass flowrate .................................................................................................................. 118
17-11 Core swollen water level .............................................................................................................. 119
17-12 Cold leg A temperature ................................................................................................................ 119
17-13: Hot leg A temperature .................................................................................................................. 120
17-14: Fuel rod temperature at the top of the core ................................................................................... 120
17-15: Fuel rod temperature at the middle of the core ............................................................................. 121
17-16 Fuel rod temperature at the bottom of the core ............................................................................ 121
17-17: Total mass of molten metals ......................................................................................................... 122
17-18: Total mass of debris...................................................................................................................... 123
17-19: Fraction of non oxidised Zr .......................................................................................................... 123
17-20: Cumulated mass of Hydrogen ...................................................................................................... 123

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17-21: Hydrogen production comparison ................................................................................................ 124
Synthetic Views of the Core at Selected Instants

Standard case
Cladding temperature maps
17-22: Cladding temperature (t = 6600. s) ............................................................................................... 125
17-23: Cladding temperature (t = 8200. s) ............................................................................................... 125
17-24: Cladding temperature (t = 10000. s) ............................................................................................. 126
17-25: Cladding temperature (t = 12000. s) ............................................................................................. 126
Linear mass of oxidised Zircalloy
17-26: ZrO2 linar mass (t = 6600. s)........................................................................................................ 127
17-27: ZrO2 linar mass (t = 8200. s)........................................................................................................ 127
17-28: ZrO2 linar mass (t = 10000. s)...................................................................................................... 128
17-29: ZrO2 linar mass (t = 12000. s)...................................................................................................... 128
Molten Zircalloy mass
17-30: Molten Zircalloy mass (t = 7900. s) ............................................................................................. 129
17-31: Molten Zircalloy mass (t = 8200. s) ............................................................................................. 129
17-32: Molten Zircalloy mass (t = 12000. s) ........................................................................................... 130

Sensitivity case
Cladding temperature maps
17-33: Cladding temperature (t = 6600. s) ............................................................................................... 131
17-34: Cladding temperature (t = 8200. s) ............................................................................................... 131
17-35: Cladding temperature (t = 10000. s) ............................................................................................. 132
17-36: Cladding temperature (t = 12000. s) ............................................................................................. 132
Linear mass of oxidised Zircalloy
17-37: ZrO2 linar mass (t = 6600. s)........................................................................................................ 133
17-38: ZrO2 linar mass (t = 8200. s)........................................................................................................ 133
17-39: ZrO2 linar mass (t = 10000. s)...................................................................................................... 134
17-40: ZrO2 linar mass (t = 12000. s)...................................................................................................... 134
Molten Zircalloy mass
17-41: Molten Zircalloy mass (t = 7900. s) ............................................................................................. 135
17-42: Molten Zircalloy mass (t = 8200. s) ............................................................................................. 135
17-43: Molten Zircalloy mass (t = 12000. s) ........................................................................................... 136
18-1: Nodalization of reactor vessel ...................................................................................................... 137
18-2: Nodalization of loop with pressurizer .......................................................................................... 138
18-3: Nodalization of SG secondary side .............................................................................................. 139
18-4: Nodalisation of reactor core, lower plenum and downcomer ....................................................... 140
18-5: Pressures ....................................................................................................................................... 144
18-6: Pressures (fine scale) .................................................................................................................... 144
18-7: Cold leg mass-flows ..................................................................................................................... 145
18-8: Break mass-flow ........................................................................................................................... 145
18-9: Primary masses ............................................................................................................................. 146
18-10: Reactor outlet temperatures .......................................................................................................... 146
18-11: SG-s outlet .................................................................................................................................... 147
18-12: SG-s levels .................................................................................................................................... 147
18-13: Pressurizer level............................................................................................................................ 148
18-14: RCP void fraction ......................................................................................................................... 148
18-15: Core collapsed level ..................................................................................................................... 149
18-16: Maximum temperatures ................................................................................................................ 149
18-17: Hydrogen production .................................................................................................................... 150
18-18: Hydrogen production rate ............................................................................................................. 150

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18-19: Masses in the core ........................................................................................................................ 151
18-20: Powers .......................................................................................................................................... 151
18-21: Core temperature field [K] ........................................................................................................... 152
18-22: Debris saturation in the core [-] .................................................................................................... 153
18-23: Magma saturation in the core [-] .................................................................................................. 154
18-24: Gas temperature in the core [K] ................................................................................................... 155
18-25: Water level in the core [3] ............................................................................................................ 156
18-26: Gas porositry in the core [-].......................................................................................................... 157
19-1: TMI-2 Plant Modeling .................................................................................................................. 160
19-7 : Synthetic Views of the Core at Selected Times............................................................................ 167
20-1: Reactor Plant Nodalization ........................................................................................................... 170
20-2: Reactor Vessel Meshing ............................................................................................................... 171
20-3: View of the Reactor Vessel at the beginning of oxidation ........................................................... 177
20-4: View of the Reactor Vessel at time of first melt relocation ......................................................... 178
20- 5: View of the Reactor Vessel at the start of the reflooding ............................................................. 179
20- 6: View of the Reactor Vessel at the end of the calculation ............................................................ 180
20-7: Core Degradation Phase: Relocation Process (t = 9250 s  t = 9750 s) ..................................... 181
20-8: Core Reflood Phase: Core By-Pass Phase (t = 9800 s  t = 10100 s) ........................................ 182
20-9: Core Reflood Phase: Top Core Quenching Phase (t = 10100 s  t = 10600 s) .......................... 183
20-10: Core Reflood Phase: Bottom Core Quenching Phase (t = 10600 s  t = 11000 s) ..................... 184
21-1:
21-6:

Figs. 1 and 2 : Primary System Nodalization. .............................................................................. 186
Synthetic Views of the Core ....................................................................................................... 191
Standard caseAt the time of beginning of oxidation ..................................................................................... 191
At the time of first melt relocation .......................................................................................... 192
At the time of reflooding ......................................................................................................... 193
At the end of calculation ......................................................................................................... 194
Sensitivity caseAt the time of beginning of oxidation ..................................................................................... 195
At the time of first melt relocation .......................................................................................... 196
At the time of reflooding ......................................................................................................... 197
At the end of calculation ......................................................................................................... 198

22-1: Water volume fraction in TMI core during reflooding ................................................................. 201
22-2: Water volume fraction in TMI core during reflooding (continued). ............................................ 202
22-3: Temperature distribution in TMI core during reflooding. ............................................................ 203
22-4: Temperature distribution in TMI core during reflooding (continued). ......................................... 204
22-5: Porosity distribution in TMI core during reflooding. ................................................................... 205
22-6: Development of pressure in the pressurizer. ................................................................................ 205
22-7: Cumulated hydrogen production. ................................................................................................. 206
22-8: Development of total mass of molten materials (i.e. sum of liquid and refrozen melt) ............... 206
22-9: Mass of molten pool (liquid melt). ............................................................................................... 207
22-10: Development of water fraction in during reflooding for the case with 1 mm representative
particle diameter. .......................................................................................................................... 208
22-11: Temperature (left) and porosity distribution (right) at the end of the calculation for the casewith 1 mm representative particle diameter. ................................................................................ 208

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3. EXECUTIVE SUMMARY

3.1 BackgroundThe experimental database on core degradation and melt relocation (and their consequences onhydrogen production, vessel rupture) is limited to small-scale experiments which are only partiallyrepresentative of what could occur in a reactor. As a consequence, there is uncertainty in the capabilityof codes to predict core degradation in postulated severe accident transients of nuclear power plants. The
GAMA has launched an action in order to determine the ability of current advanced codes to predict
core degradation in nuclear reactors. The TMI-2 scenario was selected as the case to analyze since it
concerns the only full scale Pressurized Water Reactor to have experienced core degradation. Data fromthe code calculations were compared to the TMI-2 end-state to determine the codes‟ predictive
capability. The study was completed in 2004, and is documented in ref. 2. One conclusion of the study isthat variability in the codes predictions existed in part because initial conditions of the tMI-2 scenario
were not well defined. It was concluded that code variability could be better evaluated if these
conditions were better defined. Therefore, an additional task was proposed to benchmark the codes. This
phase II study evaluates the variability in the codes‟ results using a postulated core degradation scenarioof the TMI-2 reactor. The scenario was specified with simple initial and boundary conditions so that the
influence of uncertainty of these conditions was minimized and the variability in the codes‟ results ismore readily determined.
3.2 Objectives
This phase II exercise is the first benchmarking of severe accident codes promoted by NEA in almost 20
years. The objective of Phase II of the exercise was to perform a code benchmark on a well-definedplant (similar to TMI-2) and with prescribed boundary conditions. Therefore, the benchmark avoided
additional and unwanted sources of discrepancies between code calculations, so as to focus on thevariability of the codes calculations of core degradation. This report includes all the information
concerning the preparation of the benchmark, the results, and the conclusions of the exercise.
3.3 Definition of the Alternative Scenario and Analysis of Benchmark Results
The first meeting of the group of participants in Phase II was held in November 2005 and the final
results were received in January 2008, which gives a total duration of slightly more than 2 years for thisphase. For the new benchmark exercise, an alternative scenario was proposed. The standard TMI-2 plant
is modelled, with the complete primary circuit (loops A and B) and a simplified secondary circuit. The
initial plant state corresponds to the standard TMI-2 accident sequence. The accident is initiated by asmall break located at 4m along the hot leg A. It is followed by a stop of primary pumps when the massof water contained in the primary system is lower than 85 tons. The HPI operation is delayed until 5000s
after the stop of primary pumps. This scenario leads to a significant degradation and core melting before
the reflooding of the core. The complete sequence was thus divided in three phases. Participants met andhad group discussions between each phase. Relevant variables were selected for comparison, andbenchmark results were assessed. Previous benchmark exercises on TMI-2, and other benchmarks and13

NEA/CSNI/R(2009)3

International Standard Problems (ISP) were taken into account in the assessment too. Sensitivity
analysis of some input and model parameters were performed by participants.
3.4 Main Results
At the end of this exercise, it is possible to conclude that:

A reference severe accident scenario, initiated by a SB-LOCA, based on a TMI-2 plant model,
and with prescribed boundary conditions was defined.

A set of output variables was proposed and accepted by the participants, which allowed theassessment of the quality of the calculation results.

This scenario can be used as a reference calculation to assess the agreement between current
codes or even for an adequate training of a code user.

The main results of the exercise, regarding the ability of the participating codes‟ calculations of severe
accidents are the following:

For the initial transient, up to the primary system pumps trip, the calculated results are in good
agreement. The states of the core and the primary circuit calculated by all codes are very similar
at the time of pump trip. The main discrepancies are in the calculation of the void fraction in thecore, and the calculation of the time of pump stop, but those discrepancies are considered
acceptable.



For the degradation phase, up to the reflooding of the core, the results show a rather good
agreement between all participants for global calculation of results such as total hydrogenproduction and total mass of molten materials. The variability in these results is not only muchbetter than the results obtained in the benchmark exercise performed 20 years ago, but is also
comparable or even better than the variability obtained in recent benchmarks on integral tests(PHEBUS-FPT1 and QUENCH-11). This result shows a significant improvement of the codesin the last two decades.

For the reflooding phase, there is a general agreement on the calculated pressure increase, on the
total hydrogen mass produced during this phase, and on the increased rate of core degradation.
However, although all codes agree, some results may be questionable as they are apparently incontradiction with experimental observations (LOFT and QUENCH for example) and with the
TMI-2 assumed evolution. In addition, there is a lack of agreement on the calculated efficiencyof quenching. It can be concluded that, despite considerable improvement in codes abilities,
more modelling and assessment should be done before codes can be considered reliable enoughto calculate the reflooding phase.

Sensitivity studies performed by participants have shown that variations of some key empiricalmodel (such as the cladding failure criteria) could induce a variation in some calculation results(such as hydrogen production) obtained by a single participant, which is of the same order ofmagnitude as the variation obtained when comparing different codes and/or participants. This
indicates that some physical processes are still poorly known and inadequately modeled.

3.5 Conclusions and recommendations
The surprisingly reduced scatter of the global calculation results is probably due to the fact that code
models have been assessed globally and adjusted on the same basis of integral tests but also to theparticipation of very experienced code users. Such conclusion raises again the issue of the importance ofadequate user training, in particular for utilities and TSOs who submit safety studies to regulatory

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NEA/CSNI/R(2009)3

bodies. Also it raises the issue of having good user manuals and guidelines, which must reduce to aminimum user‟s uncertainty and “guessing” when building inputs.
Overall, the results of this exercise are quite encouraging. First, all the codes succeeded in calculating
the scenario from the beginning to the end, with very little tuning of parameters or optimization of input
decks. This shows the robustness of current codes, which is a great progress compared to the state ofcodes 20 years ago when several codes could not calculate the complete TMI-2 sequence. Moreover, itappears that the codes‟ global results are significantly more consistent than they were 20 years ago.
The code scattering in the calculation of some phenomena revealed some model weaknesses. Three
major phenomena are concerned. The first one is oxidation of molten mixtures and their relocation. The
second one is the prediction of UO2 melting and its interaction with molten corium. The third one is thecore coolability and the behaviour of hot corium when the vessel is entirely reflooded. The weaknessesof codes in the prediction of those phenomena are not surprising because experimental data are scarceand the physical understanding is still incomplete.
The benchmark exercise involved some of the most important system codes which are currently used byutilities, TSO‟s, and regulatory agencies. Because of that, the benchmark has provided substantial andimportant insights. However, due to the limited number of participants and the fact that only one
scenario was calculated, the conclusions of this benchmark exercise must be considered with caution.

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NEA/CSNI/R(2009)3

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NEA/CSNI/R(2009)3

4. INTRODUCTION

The experimental database on core degradation and melt relocation (and their consequences onhydrogen production, vessel rupture) is limited to small-scale experiments which are only partiallyrepresentative of what could occur in a reactor. As a consequence, there is uncertainty in the capabilityof codes to predict core degradation in postulated severe accident transients of nuclear power plants.
A first benchmark exercise was conducted 20 years ago by OECD/NEA and US/DOE, involving
participants from several countries and various codes to calculate the TMI-2 accident transient. The
results were made available in 1988. This was followed by other reactor benchmarks, mainly in theframework of EC programs.
The GAMA has launched an action in order to determine the ability of current advanced codes to predict
core degradation in nuclear reactors. The TMI-2 scenario was selected as the case to analyze since it
concerns the only full scale Pressurized Water Reactor to have experienced core degradation. Data fromthe code calculations were compared to the TMI-2 end-state to determine the codes‟ predictive
capability. The study was completed in 2004, and is documented in ref. 2. One conclusion of the study isthat variability in the codes predictions existed in part because initial conditions of the tMI-2 scenario
were not well defined. It was concluded that code variability could be better evaluated if these
conditions were better defined. Therefore, an additional task was proposed to benchmark the codes. This
phase II study evaluates the variability in the codes‟ results using a postulated core degradation scenarioof the TMI-2 reactor. The scenario was specified with simple initial and boundary conditions so that the
influence of uncertainty of these conditions was minimized and the variability in the codes‟ results ismore readily determined. Considering the interest of this Phase 2, it was decided to open the exercise
also to organizations which were not directly involved with GAMA but belonged to the EC SARNET
network where a similar activity on reactor benchmarks was also planned.
Several participants (from GAMA and SARNET), all of them experts in using severe accident codes,
have been involved in the activity. The codes used are the most widely used system codes throughout
the world, with maybe the only exception of the SCDAP/RELAP code. The strategy of the project was
to first define a schematic scenario or “alternative TMI-2 scenario” that would lead to core uncovery anda large degradation before reflooding the vessel. Then, this initial scenario was slightly refined in orderto obtain the expected state of degradation and to allow a proper comparison of all results. It was also an
opportunity for all participants to review extensively their input decks for the TMI-2 reactor and to
update them.
In this report, complete descriptions of the modelling of the core, the steady state, the initiating event,the boundary conditions and the transient scenario are provided. The physical parameters selected forcomparison, and the criteria used to judge the quality of the comparisons are also listed and discussed.
The results are compared and conclusions about the predictive qualities of the codes are drawn.
Sensitivity studies were also made and their results are discussed. Finally, general conclusions of thisexercise are drawn and proposals to continue this benchmark activity are given. In the Appendix to thisreport, all participants have provided their complete individual results and their views and opinions
about the comparisons.

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NEA/CSNI/R(2009)3

Twenty years ago, one of the conclusions of the first benchmark exercise was the expectation that“significant progress in the severe accident code capabilities could be achieved in the future”. The main
goal of this report is to measure the progress made in twenty years.

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5. PLANT DESCRIPTION AND ALTERNATIVE TRANSIENT DEFINITION

For the accurate prediction of the TMI-2 transient, the proper definition of boundary conditions andplant characteristics is essential. However, some of these data are either unknown or difficult toestimate. In particular, the data for the make-up and let-down flows were not recorded during theaccident. Although these data do not bring any improvement in the understanding of severe accident
processes, they have required important efforts from code users who have tried to estimate them. To
avoid such problems, a new benchmark exercise is proposed, based on an alternative scenario. The
objective is to do the calculations on a well-defined plant (similar to TMI-2, for convenience) and withprescribed boundary conditions, in order to avoid additional and unwanted sources of discrepancies
between code predictions.
The standard TMI-2 plant is modelled, with the complete primary circuit (loops A and B) and asimplified secondary circuit. The details are given below. The geometry of the circuits is provided in thefollowing section. The initial state corresponds to the standard TMI-2 sequence (see table below for thethermal-hydraulic conditions)
5.1 Initiating Event
Loss of main feedwater.
Opening of a small break on hot leg A:size : 0.001 m2
position : 4 m along hot leg A

5.2 End of calculation
The calculation is stopped a few thousand seconds after the HPI operation or as soon as the core iscompletely cooled down.
5.3 Main Assumptions
No PORV failure.



Stop of primary pumps when the primary water mass is less than 85 tons.

Delayed HPI operation: 5000s after the stop of primary pumps. 30 kg/s, per loop (60kg/s total).



No let-down flow.

All other parameters are standard TMI-2 parameters (in particular the behaviour of vent valves
between the upper plenum and the top of the downcomer: details are provided below).

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NEA/CSNI/R(2009)3

5.4 Power








Initial core power = 2700 MW
Core power transient as shown in the table below
Thermal heat losses from primary system to containment atmosphere are not taken into accountTime (s)
0
1
4
10
40
100
400
800
1000
2000
4000
8000
10000
20000

Power (W)2700e6167.94e6
147.96e6
130.14e6
103.14e6
86.13e665.34e652.92e649.95e642.39e634.56e628.35e626.05e621.46e6

Z (m)
0
0.15
0.302
0.607
0.912
1.217
1.522
1.826
2.131
2.436
2.893
3.503
3.81
4

Factor
0
0
0.675
0.857
1.037
1.153
1.202
1.231
1.241
1.209
1.102
0.595
0
0

Ring
1
2
3
4
5
6

Factor
1.2572
1.2127
1.14691.0596
0.951
0.8198

 Core heat flux profile
Axial profile:

Radial profile

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NEA/CSNI/R(2009)3

5.5 Boundary Conditions
Boundary conditions for primary system make-up flow = 3.0 kg/s (total).
Boundary conditions for secondary system are given by the regulation of steam generator pressures andwater levels (see figures 5-1 and 5-2 below).
SG Level Regulation
2,00

Level (m)

1,80
1,60
1,40
1,20
1,00
0,80
0,60
0,40
0,20
0,00
0

100

200

300

400

500

600

700

800

900

1000

700

800

900 1000Time (s)Figure 5-1 : SG Level regulation

SG Pressure Regulation
80,00

Pressure (bar)

75,00
70,00
65,00
60,00
55,00
50,00
0

100

200

300

400

500

600

Time (s)Figure 5-2 : SG Pressure regulation

21

NEA/CSNI/R(2009)322

NEA/CSNI/R(2009)3

6. GEOMETRY

6.1 Free volumesPrimary system volume without the pressurizer = 294.6 m3
Pressurizer volume = 43 m3
RPV volume = 116.5 m3
Secondary side free volume of 1 SG (up to the SG exit) = 34.4 m3

6.2 Heat transfer with secondary side
SG tubes surface (secondary side)/SG = 12302.5 m2
 Number of tubes/SG = 15530
6.3 Initial core geometry
Main characteristics of the core









Number of fuel bundles of type 15x15 = 177
Active core length = 3.66 m
Total core length = 4.00 m (from core bottom: 0.15+3.66+0.19)
Type of fuel lattice = square, pitch = 0.01443 m
Number of fuel rods/assembly = 208External diameter of fuel rod = 0.0109 m
Fuel pellet radius = 0.0047 m
Fuel rod cladding thickness = 0.000673 m

Initial core material inventory








UO2 mass = 93650 kg (over the 3.66 m of core active length)
Zircaloy mass = 23050 kg (H2 total mass if converted from total zircaloy = 1011 kg (mZr/22.8))
AIC mass (Ag+In+Cd) = 2750 kg
Core baffle internal diameter = 3.28 m
Core baffle external diameter = 3.33 m
Core barrel internal diameter = 3.584 m
Core barrel external diameter = 3.683 m

Core meshing

Radial rings = 5 or 6
Axial meshes (over total core length) = 20 , mesh size = 0.20 m (all meshes with equal height)

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NEA/CSNI/R(2009)3

6.4 Vent valves
These valves, between the upper plenum and the top of the downcomer, are designed to avoid a directloss of water by the hot leg. They should be modelled, as they were shown to have an influence on thetransient. In the ICARE/CATHARE code model, these valves are modelled as follows:

If p < 414 Pa, the valves are closed.
If p > 1724 Pa, the valves are fully open, which corresponds to a total section of 0.794 m2.
If 414 < p < 1724, the valves are considered partly open, with a cross section area increasing
linearly with p.

24

NEA/CSNI/R(2009)3

7. CHOICE OF PHYSICAL PARAMETERS TO COMPARE

7.1 Parameters selected for comparison
A selection of parameters to compare was made, according to several criteria. The first criterion is therelevance for safety and/or severe accident management (i.e. primary pressure, hydrogen production,
core exit temperature),. The second criterion is the monitoring of the reactor state (i.e. water inventory,
mass of molten materials). The third and last criterion is the significance for more detailed comparisonsof models (i.e. break flow rate, mass of dissolved UO2). The parameters selected for comparison arelisted in the following table, with the criteria for their selection.
Variables

Criterion of selection

Steam Generator A Pressure

Monitoring of circuits

Steam Generator A Level

Monitoring of circuits

Steam Generator B Pressure

Monitoring of circuits

Steam Generator B Level

Monitoring of circuits

Pressuriser Pressure

Accident management

Pressuriser Level

Monitoring of circuits

Loop A Flow Rate

Monitoring of circuits

Loop B Flow Rate

Monitoring of circuits

Pump A void fraction

Comparison of models

Pump B void fraction

Comparison of models

Hot Leg A Temperature (at the position 12m)

Monitoring of circuits

Hot Leg B Temperature (at the position 12m)

Monitoring of circuits

Cold Leg A Temperature

Monitoring of circuits

Cold Leg B Temperature

Monitoring of circuits

Break Flow Rate

Comparison of models

Total Primary Mass

Monitoring of circuits

Swollen Water Level in the Core (2)Comparison of models

Collapsed Water Level in the Core (2)Comparison of models

Cumulated Hydrogen production

Accident management

Power exchanged with SG A

Monitoring of circuits

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NEA/CSNI/R(2009)3

Variables

Criterion of selection

Power exchanged with SG B

Monitoring of circuits

Instantaneous Hydrogen production

Comparison of models

Total mass of molten materials (i.e. materials which havemelted but may have frozen at lower levations)Monitoring of core state

Total mass of molten metals (same as above)

Comparison of models

Total mass of dissolved UO2 (i.e. removed by chemicalinteraction with molten Zr)

Comparison of models

Fraction of non-oxidized Zr in the molten materials (i.e. ratio
between the mass of molten Zr and the molten mass ZR+ZrO2)

Comparison of models

Total mass of debris (if available in the modelling)

Monitoring of core state

Total mass of molten pool (if any, but it may also beapproximated as the total mass of liquid materials)

Monitoring of core state

Rod temperature at the bottom of the core

Monitoring of core state

Rod temperature in the middle of the core

Monitoring of core state

Rod temperature at the top of the core

Monitoring of core state

Gas Temperature at the entrance of the Hot Leg

Accident management

Total decay power

Monitoring of core state

Water level in downcomer

Monitoring of circuits

Water level in bypass

Monitoring of circuits

Fraction of non-oxidized Zr in all the core (i.e. ratio between
the initial total mass of Zr and the total mass ZR+ZrO2)

Monitoring of core state

7.2 Criteria for the quality of agreement
To be able to claim that a set of results is in good agreement with another one, one has to define
subjective criteria reflecting the level of uncertainty that is acceptable for a specific physical parameter
calculateed by a code. The choice can be made either from a safety analysis point of view, or with amore pragmatic point of view, by comparing with uncertainties obtained in previous benchmarks.
Because there are no regulations or rules which define what can be considered as “acceptable”
uncertainty in safety studies, the first criterion was not used. Rather, what is essential from theregulatory point of view is that license applications based on best-estimate calculations, include athorough and complete evaluation of uncertainties using accepted methods. Because of that, the second
option was selected and the following table provides the uncertainty ranges obtained in the current
benchmark, to be compared with the uncertainties obtained in the previous TMI-2 benchmark, in the
PHEBUS-FPT1 benchmark, which is relevant for late phase degradation and in the QUENCH-11
benchmark, which is relevant for transients involving core dry-out followed by reflooding. For somevariables, no criteria could be found in previous benchmarks to judge the quality of agreement (theyappear with the mention “NA” in the following table). For those variables, the acceptable range of
uncertainty was decided among the participants by “expert judgement”.

26

NEA/CSNI/R(2009)3

Current
benchmarkPrevious
benchmarks

Time of pump stop+/- 3 %

specified

Minimum Steam Generator Pressure+/- 11 %NA

Maximum Steam Generator Level+/- 11 %NA

Primary Pressure before reflooding+/- 14 %> 30% TMI2

Loop Flow Rate before pump stop+/- 40 %NA

Variables

Loop Flow Rate during core reflood> 200% Q11

Pump A void fraction (at pump stop)+/- 10 %NA

Maximum Core Temperature+/- 300K~11 %+/- 350K TMI2+/- 150K FPT1+/- 500K Q11

Maximum Hot Leg Temperature+/- 10 %NA

Break Flow Rate+/- 30 %NA

Total Primary Mass at pump stop+/- 28 %+/- 50% Q11

Swollen Water Level in the Core (before HPI
operation)+/- 10 %NA

Collapsed Water Level in the Core

NA

Time of oxidation runaway+/- 42 %17% TMI2

Cumulated Hydrogen production+/- 18 %+/- 33% TMI2+/- 10% FPT1+/- 20% Q11

Power exchanged with SGNA

Max. Instantaneous Hydrogen production+/- 30 %NA

Total mass of molten materials+/- 35 %+/- 67% FPT1+/- 100% Q11

Total mass of molten metals+/- 16 %NA

Total mass of dissolved UO2+/- 80 %NA

Fraction of non-oxidized Zr in the moltenmaterials+/- 100 %NA

Fraction of non-oxidized Zr (total)+/- 10 %NA

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NEA/CSNI/R(2009)3

8. COMPARISONS BETWEEN PARTICIPANTS RESULTS

The main objective of the benchmark exercise was to compare the calculations of core degradation andreflooding. However, any severe accident sequence starts with an initial, purely thermalhydraulic,transient phase which can last a significant time, depending on the scenario. The impact of that initialtransient on the subsequent degradation is quite important because it determines the time of core
uncovery and heat-up. Therefore the comparisons include that phase too. The sequence was divided into
3 phases: initial thermalhydraulic phase (until main pumps stop), core degradation phase and reflooding
phase. Calculation of the sequence was preceded by the calculation of an “Initial Steady State”.
8.1 Initial Steady State
The steady state is calculated by all codes. Differences in the calculated values are not significant andcan be considered as acceptable. The following table indicates the steady-state values obtained by eachparticipant and also for comparison the “reference values” obtained from TMI-2 measurements.

ENEA

GRS

DIMNP

ASTEC
1.3

ATHLETCDMELCOR MELCOR ICARE/
1.8.5
1.8.6
CATHARE

SNL

IRSN

IVS

SNU

IKE

ASTEC
1.3

MAAP4

ATHLET/
MEWA

TMI-2

Reactor Power

W

2700

2663

2770

2700

2700

2700

2700

2700

Primary Pressure

MPa

15.2

14.9

15.2

15.0

15.02

15.32

15.2

15.2

Temperature Hot
Leg A

K

592.1

592.2

592

592

592.3

592.9

579

592

Temperature Hot
Leg B

K

592.1

592.2

592

592

592.3

592.9

579

592

Temperature
Cold Leg A

K

563.8

564.8

563.6

564

566.3

565.1

559

548561

Temperature
Cold Leg B

K

564.1

564.8

563.6

564

564.7

565.1

559

565

Mass Flow Rate
– Loop A

kg/s

8290

8638

8520

8737

8805

8501

?

8280

Mass Flow Rate
– Loop B

kg/s

8560

8675

8520

8828

8846

8498

?

8560

Pressurizer Level

M

5.78

5.78

5.68

5.77

5.60

5.69

6.78

5.77

Total Primary
Mass

ton

22512022330023085023127323038

221515211170-Pressure SG A

MPa

6.80

6.34

6.41

6.42

6.38

6.247

6.38

7.31

Pressure SG B

MPa

6.80

6.20

6.41

6.27

6.24

6.246

6.38

7.24

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NEA/CSNI/R(2009)3

ENEA

GRS

DIMNP

ASTEC
1.3

ATHLETCDMELCOR MELCOR ICARE/
1.8.5
1.8.6
CATHARE

SNL

IRSN

IVS

SNU

IKE

ASTEC
1.3

MAAP4

ATHLET/
MEWA

TMI-2

Steam
Temperature
SG A

K

574

568

569.5

568

584.9

578.8

508

586

Steam
Temperature
SG B

K

575

568

569.5

567

583.6

578.0

508

586

Collapsed Level
SG A

M

3.7

4.03

?

5.6

5.26

3.81

3.28

-

Collapsed Level
SG B

M

3.5

3.54

?

5.1

5.39

3.65

3.28

-

Liquid Mass
SG A

Kg

19210

14630

?

18700

18425

15303

17506

-

Liquid Mass
SG B

Kg

18410

13140

?

16900

18989

14409

17506

-

Feedwater Flow
SG A

Kg/s

742

743.7

761.6

750

700

746.5

756

723

Feedwater Flow
SG B

Kg/s

759

738.5

761.6

770

742

737.2

756

717

SG Feedwater
Temperature

K

513

503

510.9

513

513

513

508

-

8.2 Initial thermal-hydraulic phase up to the stop of main pumps ( ~5000s)
The first phase is governed mainly by the mass flow rate at the break and the heat transfer to thesecondary circuit.
8.2.1 Mass flow rate at the break
Looking at the curves, we can observe a very good coincidence of the results for most of the codes. Fiveof the curves are almost identical. Discharge from the break can be divided in thre phases: liquid
discharge up to the reaching of saturated conditions at about 200 sec, two-phase discharge until about
4500 or 5000 sec when the main pumps are stopped, and steam discharge up to the start of the
reflooding phase at about 10000 sec. The UPI-MELCOR calculation shows a slightly different
behaviour of the break mass flow rate during the two-phase discharge, which becomes constant aftersome time, instead of slowly decreasing like the other ones. This leads to reach earlier the criterion forthe stop of pumps. The SNU-MAAP4 calculation shows a slightly lower flow rate after the initial rapid
depressurization and the latest time for pump stop.
8.2.2 Primary mass
As a result of the agreement for the previous parameter, there is a very similar evolution of the primary
mass calculated by all participants. This is a surprisingly good result because larger discrepancies were
expected, considering that there was almost no “tuning” of the codes or the input decks, and the slight
discrepancies in the break flow rate for some codes. Five calculations indicate almost the same time ofpump stop, just before 5000s. The UPI-MELCOR calculation predicts it 500s earlier and the SNUMAAP4 calculation predicts it 400s later.

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NEA/CSNI/R(2009)3

8.2.3 Heat transfers with secondary side
Again, most of the curves coincide, indicating that the energy balance of the primary circuit is very
similar for all calculations. However, the ASTEC calculations (ENEA and IVS) show a higher value ofthe heat transfers (approximately 100% higher than the others, at the end of the thermal hydraulics
phase). This is probably due to the overestimation of pump heat loss into the primary coolant. The SNUMAAP4 results show very strong oscillations which do not appear in the other calculations. No
explanation of that behaviour has been provided, but it seems related to oscillations which appear also in
SG A pressure and SG A level until the pumps stop.
8.2.4 Water levels in the coreMore significant differences appear when comparing the water levels. Most codes calculate that the coreremains wet during all the thermal hydraulics phase (swollen level at the top of the core) although the
GRS-ATHLET calculation shows an early and steady decrease of the dry-out level. However, this may
just be a misinterpretation of the parameter because the core remains at saturation temperature for allcodes, indicating that the core is not dry. The comparison of the collapsed level shows more
discrepancies, indicating that the calculated average core void fractions are different for each code. Thismay be due to several differences in the modelling of heat transfers between coolant and assemblies, ofpump behaviour or of fluid stratification in the primary circuit.
8.2.5 SG behaviour
The SG water level and pressure were controlled by regulations in that phase, therefore no significant
differences are observed, except for the water level calculated by SNU-MAAP4 calculations which
increases instead of remaining constant. It may be due to an improper regulation.
8.2.6 Core temperatures
The core remains at saturation temperature everywhere, for all codes, indicating that the core is not dry.
8.2.7 Pressurizer level
Most codes calculate a fast decrease of the pressurizer level down to a minimum value, close to zero.
The UPI-MELCOR results show very strong oscillations which do not appear in the other calculations.
These oscillations are due to the loop-seal configuration of the pressurizer surge line (a key issue topredict how the pressurizer water level was maintained during the TMI-2 accident). The surge line at
TMI-2 drops down from the bottom of the pressurizer, traverses several meters, and then rises to where
it connects to hot-leg piping. This downward and then upward routing of the surge line is captured in the
UPI MELCOR model with the inclusion of a control volume representing the surge line and capturing
the hydrostatic characteristics of the water loop seal. Following the complete emptying of thepressurizer, some water is trapped inside the loop-seal up to the first 4500 s of this benchmark and verysmall oscillations in the primary pressure cause the oscillations in the pressuriser level
Overall, the states of the core and the primary circuit calculated by all codes are very similar at the timeof pumps stop. The main discrepancies are in the void fraction in the core and the time of pump stop
itself but those discrepancies are considered acceptable.
8.3 Core degradation phase (up to reflooding)
When the pumps stop, the mass flow rate at the break is significantly reduced due to the separation of
steam and liquid, and the primary mass decreases more slowly. The heat transfer to the secondary side isstrongly reduced and the evolution of the primary pressure becomes independent of the secondarypressure. This phase is characterized by a progressive dry-out and uncovery of the core, followed by the31

NEA/CSNI/R(2009)3

increase of temperature of the rods and oxidation of the claddings. The details of the degradation
processes are given below.
8.3.1 Water levels in the core
After the pumps stop, water in the loops is drained down into the vessel, which leads to the increase ofwater volume in the vessel. All codes calculate this increase of the collapsed level. Following this rapid
variation of the level, a progressive decrease of the water level is calculated, corresponding to the dryout and core uncovery. The curves calculated by each code are very similar, in particular the slope(which is approximately proportional to the steam production in the core) and the final value of thewater level. Despite differences in the time of beginning of core uncovery, most codes calculate astabilization of the water level around 60-70 cm above the bottom of the core. The SNU-MAAP4
calculation shows a remarkably different behaviour, probably due to the fact that the MAAP4 swollen
core level is of parametric nature. The UPI-MELCOR calculation calculates a complete dry-out of theactive part of the core. In general, all the results agree on the prediction of the core dry-out.
8.3.2 Mass flow rate at the breakAll results show a slow decrease of the mass flow rate at the break, from a value of approximately 10
kg/s to a rather stable value of approximately 4 kg/s before reflooding. SNU-MAAP4 drops to zero andremains to the end of the calculation. However, the shapes of the curves are different and, during thetransient, the calculated mass flow rate can differ by a factor 2 between the minimum estimated value
and the maximum one.
8.3.3 Primary mass
As a result of the previous observation, the decrease of primary mass differs from one result to the other.
Looking at the absolute value of the primary mass calculated just before reflooding, the relative
discrepancy is limited to 9%. However, it is more relevant to compare the relative decrease with respect
to the value at the time of pumps stop (limit of 85 tons in the primary system). In this case, the relative
discrepancy between the calculated values is 28%, which is quite significant.
8.3.4 Primary pressure
There are significant discrepancies in the evolution in the primary pressure calculated by theparticipants. Some results indicate that the primary pressure decreases continuously after pump stopwhereas other codes show a transient increase of the primary pressure (between 8 and 15 bars) during
approximately 1000s, due to the fact that heat transfer to the secondary system is greatly reduced. The
time when this pressure increase starts also vary from one result to the other, due to the differences in
the calculation of the time when pumps stop. The SNL-MELCOR calculation shows a swelling inpressuriser level when the pumps stop, probably due to flashing, although calculated primary pressure
remains constant. Other participants calculate an increase in pressure, but without any swelling.Those
discrepancies were not clearly explained. They could depend on the efficiency of decay heat removal by
the secondary side through residual natural circulation of steam in the primary circuit and condensationin the steam generators.
The average final primary pressure (before reflooding) is predicted at 57 bars, with a variance of 14%,which is acceptable.
8.3.5 Core temperatures
The evolution of core temperatures calculated by participants show several differences. The curves may
be divided into three parts: heat-up before oxidation, oxidation runaway and “late-phase” heat-up
following the first relocation of molten materials.

32

NEA/CSNI/R(2009)3

For the heat-up phase before oxidation, the slope is very similar for all curves, except for the IVSASTEC and the ENEA-ASTEC calculations which show a delayed but steep heat up, in particular at the
top of the core (the IVS-ASTEC middle core temperature is in fact maximum core temperature). The
main discrepancy is the time of beginning of heat-up which varies (in the middle of the core) from 500s
to 1800s after pump stop, with a variance of 40%. Discrepancies are smaller for the evolution oftemperatures at the top of the core. The delay between curves is partly due to the delay of core uncovery(discussed previously). Another cause is probably the differences in the melt relocation models whichlead to a different velocity of propagation of the “hot” oxidation front.
For the oxidation run-away, we can observe a surprising agreement between several results which
predict it between 7500s and 8000s. One calculation predicts it earlier (IVS-ASTEC at 6500s) and one
predicts it later (IRSN-ICARE/CATHARE at 9500s). The increase of temperature due to temperatureescalation is very comparable for most results. The maximum calculated temperature after escalation isapproximately 2300K.
The heat-up calculated after the temperature escalation has approximately the same slope as before theescalation, as shown in almost all curves. The main difference between results in that phase is the
maximum temperature reached, which is closely related to the choice of the melting temperature of
UO2.
8.3.6 Hydrogen production
The calculated rate of hydrogen production is very similar for most results (approximately 0.6 kg/s inaverage) and the main differences are the times of maximum production. This is of course linked to thedelays observed in the core temperature curves.
As a result, the shapes of the cumulated hydrogen mass curves are very similar, in particular the slope atthe oxidation runaway. Once again, it is interesting to notice that 5 out of 7 results calculate theoxidation runaway approximately at the same time (around 8000s). The main differences are observed
after the oxidation runaway where some results show a continuation of oxidation whereas other codes
show a clear reduction of the oxidation after the runaway. The SNU-MAAP4 calculation shows, acontinuation of oxidation after the runaway. There is a rather good agreement on the cumulatedhydrogen production calculated before reflooding, with a lower than 20% variation.
The fraction of non-oxidized Zr, which was another parameter chosen for comparison, showsinconsistencies with the cumulated hydrogen mass. The curves should be almost symmetrical becausehydrogen is mainly produced from Zr oxidation. Those inconsistencies need to be explained. Despite ofthat, the agreement between calculations is very good with an average oxidized fraction of 0.43 with avariation of only 10% between all results.
In contrast, the fraction of oxidized Zr in the melt (which could not be provided by all participants)
shows an extreme variability: from 0.18 to 0.96. This indicates that models for oxidation of melts are
significantly different from one code to the other.
8.3.7 Cladding failure and melt relocation
The cladding failure and first melt relocation are also calculated at approximately the same time for mostcodes. This is a direct consequence of the good agreement on the time of oxidation run-away.
8.3.8 Fuel melting and relocation
The evolution of the total mass of molten materials shows interesting differences. For some participants,
the calculated molten mass increases simultaneously and almost proportionally to the cumulatedhydrogen mass, which means that the energy brought by oxidation is directly converted into latent heat33

NEA/CSNI/R(2009)3

to melt the oxides. For other participants, the molten mass increases smoothly, even a long time after theoxidation runaway, which indicates that the oxidation energy is “stored” in the core materials and does
not cause a rapid melting. One reason for such discrepancy is the choice of melting temperature of theoxides. If this temperature is far from the temperature reached after the end of the oxidation runaway(approximately 2300K), then the melting must occur later than the oxidation. Despite the differences in
the history of melting, there is a rather good agreement on the final mass of molten materials. The
maximum variation obtained is 35%, which is surprisingly low considering that the molten massdepends on many uncertain phenomena: core heat-up, Zr oxidation, melting of oxides, etc. This may
indicate that there is some sort of compensation between all the phenomena leading to core melting andthat the calculated global energy balance of the core is approximately the same for all participants.
The SNU-MAAP4 calculation is a surprising exception: it predicts the maximum hydrogen mass but
also the minimum molten mass. This apparently contradictory result was not explained.
There is an even better agreement on the calculated mass of molten metal with a variation of only 16%.
This could be explained by the fact that the melting temperatures of metals are known with lessuncertainty than those of the oxides.
8.3.9 State of the core
The states of the core calculated by the participants at the beginning of the oxidation runaway appear to
be rather consistent. The temperature distributions are similar, the core is divided into a cold lower partand a hot upper part which is uncovered. The radial variations of temperature are consistent and the mostexternal ring is calculated to be colder than the other ones by all participants. The main difference is the
position of the maximum temperature. Some participants calculate it at the top of the core and others
calculate it slightly below the top (approximately 50cm below).
The states of the core calculated at the beginning of melt relocation, show greater differences, and thelocation of the first damages or melting varies significantly from one calculation to the other. Someparticipants calculate the first melting at the top (3.66m) while others calculate it closer to the middle(between 2.5m and 3m). In addition to that, a difference is also observed in the calculated level of the“cold” zone (covered with water): most participants calculate it around 1m but some participants
calculate it up to 1.5m. Such differences are quite significant and are likely due to the modelling ofthermalhydraulics in the core. Differences in pressure drops or heat transfer coefficients and flow
distribution (axial/radial) are probably the main causes for the observed discrepancies.
The states of the core calculated just before the reflooding, show both a very different distribution ofmaterials and a very different temperature field. Some calculations show a large empty area at the top of
the core, resulting from the collapse or melting of fuel rods. Other calculations show standing rods at the
top of the core with a more limited molten area in the center of the core. Almost all calculations show avery compact region of accumulated materials at approximately 1m elevation. This may be interpreted
as a “crust” or “crucible” made of relocated materials. The elevation of that crust varies significantly
from one calculation to the other, between 0.6m to 2m. Some calculations predict high temperatures inthe “crust” region, while other calculations predict the maximum at the center of the core. The ASTECIVS calculation shows the formation of a molten pool before reflooding while the MELCOR-SNL
calculation shows the formation of a large debris bed.
8.4 Reflooding phase
8.4.1 Water levels in the core
All codes calculate a fast increase of the water level in the core and the core by-pass, with approximately
the same velocity. In some calculations, discontinuities in the increase of the water level are observed

34

NEA/CSNI/R(2009)3(MELCOR-SNL, ICARE/CATHARE-IRSN and ASTEC-ENEA) but they are probably due to thenumerical tracking of the water level. In comparison, the increase of water level in the by-pass issmoother and continuous. At the end of the calculation, most calculations indicate a stable water level inthe core but the ASTEC-ENEA calculation shows a surprising decrease of the water level (this is likely
due to some flow instabilities in the primary loops during reflooding with subsequent redistribution ofvoids in the primary circuit). The stabilized level differs from one calculation to the other (from 3m to4m). This was not explained.
8.4.2 Primary pressure
The average calculated pressure increase is 50 bars, with a variance of 38%. In the real TMI-2 scenario,
the pressure increase after the restart of pumps was also of the order of 50 bars, but it was attributed toboth a strong steam generation and a strong oxidation which produced a lot of non-condensablehydrogen. In the present exercise, participants calculate a small amount of hydrogen generated during
reflood and therefore the increase of pressure results mainly from the strong steam formation.
8.4.3 Core temperatures
Most codes calculate a fast quenching of the core, at all elevation. However, the ATHLET/CD-GRS
calculation predicts the non coolability of the inner ring (central part) at the bottom of the core and both
ASTEC-IVS and ICARE/CATHARE-IRSN predict the non coolability of the middle of the core. The
reasons for such non coolability should be investigated more thoroughly. The issue of coolability is
crucial for safety studies and the existence of such discrepancies between code calculations reveals aweakness of current models dealing with reflooding of a degraded core.
8.4.4 Hydrogen production- Melt oxidation
A few codes calculate some hydrogen production during reflooding but it is quite limited. The largest
amount is calculated by ASTEC-ENEA (approx. 29 kg) while ATHLET/CD-GRS andICARE/CATHARE-IRSN calculate a slow continuation of oxidation for a long time after reflooding,because of the existence of non-quenched areas in the higher temperature areas of the core as explained
above. For the other calculations, oxidation is predicted to stop as soon as the reflooding starts and no
additional hydrogen is produced during reflooding.
For all codes, those calculations are obviously wrong as it was demonstrated by QUENCH program thatthe reflooding of a very hot but still undestroyed core (i.e. a core with a maximum temperature above
2100K but below 2500K) should lead to a large hydrogen production. And the same observation wasmade during the TMI-2 accident where half of the total hydrogen output (160kg out of 300kg) was
supposedly produced during reflooding. Recent studies suggest that hydrogen produced during
reflooding mainly comes from the oxidation of melt containing Zr. Therefore we may assume that there
is an inadequate modelling of this phenomenon in current codes. This may be amplified by an inaccurate
modelling of melt formation and relocation.
8.4.5 State of the core
For almost all the participants, the final calculated state of the core is very similar to the state before
reflooding. There is no significant increase of the degradation or core melting during reflooding. Such
result might not be conservative at all, as it was observed in QUENCH program that the reflooding of avery hot core (i.e. the maximum temperature is above 2000K) could lead to the melting of the claddings
and a significant relocation of materials. However, QUENCH tests are not fully representative of a core
since there are no UO2 pellets. Moreover, the average calculated core temperature in this scenario is
approximately 2500K, i.e. much higher than the temperature obtained in QUENCH tests. Therefore it is
not possible to conclude, solely on the basis of experimental results, that code calculations are wrong.

35

NEA/CSNI/R(2009)3

8.5 Figures for comparison

Break Flow Rate
1,00E+02
ENEA

9,00E+01

IVS
GRS

Mass Flow Rate (kg/s)

8,00E+01

IKE

7,00E+01

UPI

6,00E+01

IRSN

SNL
IRSN-dSNU

5,00E+01
4,00E+01
3,00E+01
2,00E+01
1,00E+01
0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

8000

9000 10000 11000 12000 13000 14000

Time (s)Primary Mass
2,50E+05
2,25E+05
2,00E+05

Mass (kg)

1,75E+05
1,50E+05
1,25E+05

ENEA
IVS

1,00E+05

GRS
IKE

7,50E+04

UPI
SNL

5,00E+04

IRSN
IRSN-d

2,50E+04

SNU

0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

Time (s)3680009000 10000 11000 12000 13000 14000

NEA/CSNI/R(2009)3

Swollen Water Level
5,00E+00

Water Level (m)4,00E+00

3,00E+00

2,00E+00
ENEA
IVS
IKE

1,00E+00

UPI
SNL
IRSN

0,00E+00

IRSN-dSNU-1,00E+00
0

1000

2000

3000

4000

5000

6000

7000

8000

9000 10000 11000 12000 13000 14000

Time (s)Collapsed Water Level
4,00E+00

Water Level (m)3,00E+00

ENEA

2,00E+00

IVS
GRS
IKE
UPI
SNL

1,00E+00

IRSN
IRSN-dSNU
0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

Time (s)378000

9000 10000 11000 12000 13000 14000

NEA/CSNI/R(2009)3

SG A Pressure
8,00E+06
7,00E+06

Pressure (Pa)6,00E+06
5,00E+06
ENEA

4,00E+06

IVS
GRS
IKE

3,00E+06

UPI
SNL

2,00E+06

IRSN
IRSN-d

1,00E+06

SNU

0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

8000

9000 10000 11000 12000 13000 14000

Time (s)SG A Level
2,00E+00
1,80E+00
1,60E+00

Level (m)1,40E+00
1,20E+00
1,00E+00
ENEA
IVS

8,00E-01

GRS
6,00E-01

IKE
UPI

4,00E-01

SNL
IRSN

2,00E-01

IRSN-dSNU

0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

Time (s)388000

9000 10000 11000 12000 13000 14000

NEA/CSNI/R(2009)3

SG A Power Exchanged
2,00E+08
ENEA

1,80E+08

IVS
GRS

1,60E+08

IKE
UPI

1,40E+08

SNL

Power (W)

1,20E+08

IRSN
IRSN-d

1,00E+08

SNU

8,00E+07
6,00E+07
4,00E+07
2,00E+07
0,00E+00-2,00E+07
0

1000

2000

3000

4000

5000

6000

7000

8000

9000 10000 11000 12000 13000 14000

Time (s)Temperature at the Bottom of Core
3,50E+03
ENEA
IVS

3,00E+03

GRS
IKE
UPI

Temperature (K)2,50E+03

SNL
IRSN

2,00E+03

IRSN-dSNU

1,50E+03

1,00E+03

5,00E+02

0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

Time (s)398000

9000 10000 11000 12000 13000 14000

NEA/CSNI/R(2009)3

Temperature at mid-Core
3,50E+03
ENEA
IVS

3,00E+03

GRS
IKE
UPI

Temperature (K)2,50E+03

SNL
IRSN

2,00E+03

IRSN-dSNU

1,50E+03

1,00E+03

5,00E+02

0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

8000

9000 10000 11000 12000 13000 14000

Time (s)Temperature at the Top of Core
3,50E+03
ENEA
IVS

3,00E+03

GRS
IKE
UPI

Temperature (K)2,50E+03

SNL
IRSN

2,00E+03

IRSN-dSNU

1,50E+03

1,00E+03

5,00E+02

0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

Time (s)40

80009000 10000 11000 12000 13000 14000

NEA/CSNI/R(2009)3

Pressurizer Pressure
1,80E+07
ENEA
1,60E+07

IVS
GRS

1,40E+07

IKE
UPI

Pressure (Pa)

1,20E+07

SNL
IRSN
IRSN-d

1,00E+07

SNU
8,00E+06
6,00E+06
4,00E+06
2,00E+06
0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

8000

9000 10000 11000 12000 13000 14000

Time (s)Pressurizer Level
9,00E+00
ENEA

8,00E+00

IVS

Level (m)7,00E+00

GRS

6,00E+00

IKE

5,00E+00

SNL

UPI
IRSN

4,00E+00

IRSN-dSNU

3,00E+00
2,00E+00
1,00E+00
0,00E+00-1,00E+00

0

1000

2000

3000

4000

5000

6000

7000

-2,00E+00-3,00E+00

Time (s)418000

9000 10000 11000 12000 13000 14000

NEA/CSNI/R(2009)3

Total Mass of Molten Materials
7,00E+04
ENEA

6,50E+04

IVS

6,00E+04

GRS

5,50E+04

IKE

5,00E+04

UPI
SNL

Mass (kg)

4,50E+04

IRSN

4,00E+04

IRSN-d

3,50E+04

SNU

3,00E+04
2,50E+04
2,00E+04
1,50E+04
1,00E+04
5,00E+03
0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

8000

9000 10000 11000 12000 13000 14000

Time (s)Total Mass of Molten Metal
2,00E+04
ENEA
IVS

1,75E+04

GRS
IKE

1,50E+04

UPI
SNL

Mass (kg)

1,25E+04

IRSN
IRSN-dSNU

1,00E+04
7,50E+03
5,00E+03
2,50E+03
0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

Time (s)4280009000 10000 11000 12000 13000 14000

NEA/CSNI/R(2009)3

Mass of Molten Pool
3,00E+04
ENEA

2,75E+04

IVS

2,50E+04

GRS
IKE

2,25E+04

UPI

Mass (kg)2,00E+04

SNL
IRSN

1,75E+04

IRSN-d

1,50E+04
1,25E+04
1,00E+04
7,50E+03
5,00E+03
2,50E+03
0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

8000

9000 10000 11000 12000 13000 14000

Time (s)Total Mass of Dissolved UO2
3,25E+04
ENEA

3,00E+04

IVS

2,75E+04

GRS

2,50E+04

IKE

2,25E+04

SNL

UPI
IRSN

Mass (kg)2,00E+04

IRSN-d

1,75E+04

SNU

1,50E+04
1,25E+04
1,00E+04
7,50E+03
5,00E+03
2,50E+03
0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

Time (s)438000

9000 10000 11000 12000 13000 14000

NEA/CSNI/R(2009)3

Cumulated Hydrogen Production
6,00E+02
ENEA
IVS
GRS

5,00E+02

IKE
UPI
SNL

H2 Mass (kg)4,00E+02

IRSN
IRSN-dSNU

3,00E+02

2,00E+02

1,00E+02

0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

8000

9000 10000 11000 12000 13000 14000

Time (s)Fraction of non Oxidized Zr
1,00E+00
9,00E-01
8,00E-01

Mass Fraction

7,00E-01
6,00E-01
5,00E-01

ENEA
IVS

4,00E-01

GRS
IKE

3,00E-01

UPI
SNL

2,00E-01

IRSN
IRSN-d

1,00E-01

SNU

0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

Time (s)448000

9000 10000 11000 12000 13000 14000

NEA/CSNI/R(2009)3

Fraction of non Oxidized Zr in Molten Materials
1,00E+00
9,00E-01
8,00E-01

Mass Fraction

7,00E-01
6,00E-01
5,00E-01
ENEA
4,00E-01

IVS
GRS

3,00E-01

IKE
UPI

2,00E-01

IRSN
IRSN-d

1,00E-01
0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

8000

9000 10000 11000 12000 13000 14000

Time (s)Instantaneous Hydrogen Production
2,00E+00
ENEA

H2 Production Rate (kg/s)

1,80E+00

IVS

1,60E+00

GRS

1,40E+00

UPI

IKE
SNL

1,20E+00

IRSN
IRSN-d

1,00E+00

SNU

8,00E-01
6,00E-01
4,00E-01
2,00E-01
0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

Time (s)4580009000 10000 11000 12000 13000 14000

NEA/CSNI/R(2009)3

Temperature at the Bottom of Core
3,50E+03
ENEA
IVS

3,00E+03

GRS
IKE
UPI

Temperature (K)2,50E+03

SNL
IRSN

2,00E+03

IRSN-dSNU

1,50E+03

1,00E+03

5,00E+02

0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

8000

9000 10000 11000 12000 13000 14000

Time (s)Temperature at mid-Core
3,50E+03
ENEA
IVS

3,00E+03

GRS
IKE
UPI

Temperature (K)2,50E+03

SNL
IRSN

2,00E+03

IRSN-dSNU

1,50E+03

1,00E+03

5,00E+02

0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

Time (s)46

8000

9000 10000 11000 12000 13000 14000

NEA/CSNI/R(2009)3

Temperature at the Top of Core
3,50E+03
ENEA
IVS

3,00E+03

GRS
IKE
UPI

Temperature (K)2,50E+03

SNL
IRSN

2,00E+03

IRSN-dSNU

1,50E+03

1,00E+03

5,00E+02

0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

8000

9000 10000 11000 12000 13000 14000

Time (s)Temperature at entrance of Hot Leg
2,50E+03
ENEA
IVS
GRS

2,00E+03

IKE

Temperature (K)

UPI
SNL
IRSN

1,50E+03

IRSN-dSNU

1,00E+03

5,00E+02

0,00E+00
0

1000

2000

3000

4000

5000

6000

7000

Time (s)478000

9000 10000 11000 12000 13000 14000

NEA/CSNI/R(2009)348

NEA/CSNI/R(2009)3

9. SENSITIVITY STUDIES

Discrepancies between code results may have different origins. One of them is the so-called “usereffect”. This has been observed in several previous benchmarks, in particular in the frame of OECD
International Standard Problems (ISP). User effect may result from an improper checking of the data(plant data are particularly complex) or from a wrong use of the code by an insufficiently trained user. In
order to avoid the user effect, recommended values of the parameters used in the models are usuallyprovided to code users in calculation benchmarks. All the reference calculations of this benchmark weredone with the recommended values of physical parameters. Taking into account the differences whichwere observed in the results of the reference calculation, especially in the core degradation phase, it was
decided to perform a second calculation with the same core degradation parameters for all codes. Thus, aset of parameters known to have a significant impact on degradation calculations were selected andgiven to participants. Those will be called “alternative physical parameters”. They were chosen from theconclusions of ISP46 (PHEBUS-FPT1) and from the latest findings on core degradation phenomena.
Those parameters are listed in the next section.
9.1 Selection of parameters

Zr oxidation correlation : Urbanic-Heidrick correlation was the most widely used byparticipants of ISP46. Although there are better alternatives now (following Volchek and
Schanz work), it is decided to choose Urbanic correlation for this sensitivity calculation.



Cladding failure criteria : these criteria have a tremendous impact on the hydrogenproduction because they control the release and relocation of molten Zr towards lower andcolder elevations. Based on PHEBUS analysis, it was found that relevant criteria for the
failure of cladding are, under steam starved conditions :oT > 2280K and e(ZrO2) < 0.2mm (thickness of zirconia layer)oT > 2450K and e(ZrO2) < 0.3mm

Melting temperature of oxides (UO2 and ZrO2) : this parameter also has a strong impact on
the results. While it is theoretically around 2800K, it was found that the best agreement with
FPT1 was obtained with a greatly reduced relocation temperature of 2500K.



Debris formation criteria : there are very few data about this parameter (and no informationfrom PHEBUS FPT1). It may be assumed that debris is formed during reflooding, because of
the thermal shock. A simple criterion based on temperature is chosen, with a threshold valueof 2000K.



Debris porosity and diameter : once again, there are very few data about these parameters.
Based on the analysis of TMI-2 debris, a diameter of 2 mm and a porosity of 0.4 are chosen.

49

NEA/CSNI/R(2009)3

The values used by each participant are summarized in the table below:
Zr Oxidation correlation
ENEA
ASTEC 1.3

Correlations proposed by Schanz et al. (2001)o Cathcart correlation if T < 1800 K
o Prater correlation if T > 1900 K
Linear interpolation between 1800 – 1900 K

GRS
ATHLET-CD

o Cathcart correlation if T < 1800 K
o Urbanic correlation if T > 1900 K
Linear interpolation between 1800 – 1900 K

DIMNPMELCOR 1.8.5

Urbanic-Heidrick

SNL
MELCOR 1.8.6

Urbanic Heidrick

IVS
ASTEC 1.3

Urbanic-Heidrick

IRSN
ICARE/CATHARE

Correlations proposed by Schanz et al. (2001)o Cathcart correlation if T < 1800 K
o Prater correlation if T > 1900 K
Linear interpolation between 1800 – 1900 K

SNU
MAAP4

Baker-Just model (Tclad > 1875 K)
Cathcart model (Tclad < 1850 K)
Interpolated inbetween

IKE
ATHLET-CD/MEWA
Cladding failure criteria
ENEA
ASTEC 1.3oo

T > 2300 K and e(ZrO2) < 300 μm
T > 2500 K

GRS
ATHLET-CDoo

T > 2300 K and e(ZrO2) < 300 μm
T > 2500 KDIMNPMELCOR 1.8.5

o

T > 2400 K and e(ZrO2) < 100 μm

SNL
MELCOR 1.8.6oo

T > 2500 K and e(ZrO2) < 100 μm
T > 2800 K

IVS
ASTEC 1.3

o
o
o
o
o

T > 2260K and e(ZrO2) < 160 μm
T > 2280K and e(ZrO2) < 200 μm
T > 2340K and e(ZrO2) < 220 μm
T > 2380K and e(ZrO2) < 240 μm
T > 2450K and e(ZrO2) < 300 μm

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IRSN
ICARE/CATHAREoo

T > 2300 K and e(ZrO2) < 300 μm
T > 2500 K

SNU
MAAP4

o
o

T > 2500 K ore(ZrO2) > 400 μm

IKE
ATHLET-CD/MEWAUO2-ZrO2 melting temperature
ENEA
ASTEC 1.3

2830K

GRS
ATHLET-CD

2600K

DIMNP
MELCOR 1.8.5

UO2: 3113K and ZrO2: 2990K

SNL
MELCOR 1.8.6

2800K

IVS
ASTEC 1.3

Phase Diagram ?

IRSN
ICARE/CATHARE
SNU
MAAP4

2500K

IKE
ATHLET-CD/MEWA

Debris formation criteria
DIMNPMELCOR 1.8.5

Conglomerate debris forms when molten corematerial resolidifies

SNL
MELCOR 1.8.6

Debris forms when clad fails

IVS
ASTEC 1.3

code default criteria, debris models and reflooding
model in ASTEC are still under development

IRSN
ICARE/CATHARE
SNU
MAAP4
IKE
ATHLET-CD/MEWA

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Debris diameter
DIMNP
MELCOR 1.8.5

0.5 m is the particular debris equivalent diameterused to calculate total debris surface area during
quenching (it is applied to the zircaly resolified in
the colder zones)

SNL
MELCOR 1.8.6

Core – 10mm
Lower Plenum - 2 mm

IVS
ASTEC 1.3

5 % of 2 mm diameter
15 % of 4 mm diameter
60 % of 6 mm diameter
15 % of 8 mm diameter
5 % of 10 mm diameter

IRSN
ICARE/CATHARE
SNU
MAAP4
IKE
ATHLET-CD/MEWA

Debris porosity
DIMNP
MELCOR 1.8.5

0.1

SNL
MELCOR 1.8.6

0.4

IVS
ASTEC 1.3

0.6

IRSN
ICARE/CATHARE
SNU
MAAP4

0.4

IKE
ATHLET-CD/MEWA

9.2 Results of sensitivity studies
The details of sensitivity studies are given in the individual descriptions of calculations provided in the
Appendix. Because standard core degradation parameters are somewhat different for each code, theparameters varied in the sensitivity studies performed were not the same for all participants. Therefore,52

NEA/CSNI/R(2009)3

comparisons between sensitivity calculation results can not be made. But interesting conclusions can be
drawn from those studies, in particular to answer the following questions:

Are the uncertainties in important core degradation results coming from changes in model
parameters larger or smaller than the uncertainties coming from the use of different codes?

Among the parameters chosen, which ones have greater importance, in terms of their impact on
significant core degradation results?

To summarize the results obtained by the participants, we can draw the following conclusions:

The variation of the melting temperature of oxides has a strong impact on the final mass ofmolten material (which is quite obvious) causing a variation which is of the same order ofmagnitude as the variation obtained with different codes and/or users. This shows that thephysical understanding of fuel rod melting and collapse at is still too limited. Therefore there isno correct mechanistic modelling of the process which relies on user parameters which are very
sensitive.

The hydrogen production is more sensitive to the cladding failure criteria than to the choice ofoxidation kinetics. The variation of cladding failure criteria also causes a variation which iscomparable to the variation obtained with different codes and/or users. Again, this shows thatthe physical understanding of cladding failure and liquid Zr relocation is still too limited.
Therefore there is no correct mechanistic modelling and the whole process relies on user
parameters which are very sensitive.



The impact of the variation of the selected parameters on the reflooding phase and on the final
state of the core seems to be more limited.



The models involving debris formation introduce an additional criterion for fuel rods collapse
which is likely to increase the variability of the results. However, this point was not investigatedenough. This is mainly due to the fact that the definition of “debris” differs from one code to theother.

Due to the limited time frame of this exercise, it was not possible to make of thourough analysis of the
impact of “debris models” to treat the late phase of degradation. Only two calculations used suchmodels: MELCOR-SNL and ASTEC-IVS. Two more calculations were expected (ATHLET/MEWAIKE and an alternative ICARE/CATHARE-IRSN) but they could not be performed. Since those models
are considered as more advanced than the regular models used up to now in current codes, it would bevery interesting, in a further extension of this exercise, to focus on the impact of such “debris models”.
It also appears that some model parameters cannot be varied independently because they depend on thesame physical processes. For example, if the melting temperature of oxides is reduced to 2400K, thecladding failure criteria should also include a temperature condition lower than 2400K. Sensitivity
analyses take into account the combined effect of a consistent set of parameters. This issue should be
investigated more thouroughly in the future.
Taking into account that, from the regulatory point of view, it is essential that uncertainties in bestestimate calculations be well studied and assessed, it is suggested that systematic uncertainty analysis be
included in future extension of this benchmark, using uncertainty analysis techniques. The BEMUSE
program action, which is being carried out in the frame of GAMA, could be a good reference for that.

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10. CONCLUSIONSThis phase II exercise is the first benchmarking of severe accident codes promoted by NEA in almost 20
years. The objective of Phase II of the exercise was to perform a code benchmark on a well-definedplant (similar to TMI-2) and with prescribed boundary conditions. Therefore, the benchmark avoided
additional and unwanted sources of discrepancies between code calculations, so as to focus on thevariability of the codes calculations of core degradation. Although the initial objectives of the Phase 2 ofthis action were apparently clear, the participants realized that the task needed several preliminary steps
and complementary actions:

Make sure that all participants had a correct input deck for the TMI-2 reactor.

Define a reasonable scenario that would include a large core degradation followed by areflooding phase.



Select the parameters for comparison.



Analyse the results of the last TMI-2 exercise and of previous benchmarks (the recent PHEBUSFPT1 and QUENCH-11 benchmarks were chosen) in order to be able to measure the quality of
the comparisons made in the present benchmark and the progress made.



List the physical criteria which are relevant for PSA studies and define the uncertainty ranges
which can be considered acceptable for current safety studies.

The first four actions were completed with success and it is possible to conclude that:

A reference severe accident scenario, initiated by a SB-LOCA, based on a TMI-2 plant model,
and with prescribed boundary conditions was defined.

A set of output variables was proposed and accepted by the participants, which allowed theassessment of the quality of the calculation results.

This scenario can be used as a reference calculation to assess the agreement between current
codes or even for an adequate training of a code user.

The last action was not finished, because there are no regulations or rules which define what can beconsidered as “acceptable” uncertainty in safet studies. Rather, what is essential from the regulatory
point of view is that license applications based on best-estimate calculations, include a thorough andcomplete evaluation of uncertainties using accepted methods.
The main conclusions of the exercise, regarding the ability of the participating codes to calculate severe
accidents are summarized below.
For the initial transient, up to the primary system pumps trip, the calculated results are in good
agreement. The states of the core and the primary circuit calculated by all codes are very similar at thetime of pump trip. The main discrepancies are in the calculation of the void fraction in the core, and thecalculation of the time of pump stop, but those discrepancies are considered acceptable.

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For the degradation phase, up to the reflooding of the core, the results show a rather good agreement
between all participants for global calculation of results such as total hydrogen production and total
mass of molten materials. The variability in these results is not only much better than the results
obtained in the benchmark exercise performed 20 years ago, but is also comparable or even better thanthe variability obtained in recent benchmarks on integral tests (PHEBUS-FPT1 and QUENCH-11). This
result shows a significant improvement of the codes in the last two decades. They now include assessed
models to describe the main physical processes of degradation as well as robust and consistentnumerical schemes which make them reliable to calculate a severe accident sequence with acceptable
results. The surprisingly reduced scatter of the global calculation results is probably due to the fact thatcode models have been assessed globally and adjusted on the same basis of integral tests but also to theparticipation of very experienced code users. Such conclusion raises again the issue of the importance ofadequate user training, in particular for utilities and TSOs who submit safety studies to regulatory
bodies. Also it raises the issue of having good user manuals and guidelines, which must reduce to aminimum user‟s uncertainty and “guessing” when building inputs.
For the reflooding phase, there is a general agreement on the calculated pressure increase, on the totalhydrogen mass produced during this phase, and on the increased rate of core degradation. However,
although all codes agree, some results may be questionable as they are apparently in contradiction withexperimental observations (LOFT and QUENCH for example) and with the TMI-2 assumed evolution.
In addition, there is a lack of agreement on the calculated efficiency of quenching. Some codes calculatea fast and complete quenching while other codes calculate the existence of uncoolable regions in thecore. Compared to the previous TMI-2 benchmark exercise where almost none of the codes were able tocalculate the reflood phase, the abilities and robustness of codes have been considerably improved.
However, it is not enough yet to produce reliable predictions. It can be concluded that, despite
considerable improvement in codes abilities, more modelling and assessment should be done before
codes can be considered reliable enough to calculate the reflooding phase.
Overall, the results of this exercise are quite encouraging. First, all the codes succeeded in calculating
the scenario from the beginning to the end, with very little tuning of parameters or optimization of input
decks. This shows the robustness of current codes, which is a great progress compared to the state ofcodes 20 years ago when several codes could not calculate the complete TMI-2 sequence. Moreover, itappears that the codes‟ global results are significantly more consistent than they were 20 years ago.
They are also equally or more consistent than predictions made on integral tests, usually dedicated to thestudy of specific phenomena. However, this surprising and very positive result must be considered
carefully because of the limited number of participants and the fact that they were all very experienced
code developers or users, which obviously helped in avoiding to get unphysical or unreasonable results.
Despite of that, a few minor physical inconsistencies were observed which indicated that some mistakes
may remain in some input decks or some models or in the post-processing of the results. Finally, an
unexpected difficulty was met: some codes could not provide plots of the evolution for some requested
physical parameters. It is suggested that code developers improve the standard outputs and postprocessing of code results in order to help the users to make a sensible and useful analysis of theircalculations.
Sensitivity studies performed by participants have shown that variations of some key empirical model(such as the cladding failure criteria) could induce a variation in some calculation results (such ashydrogen production) obtained by a single participant, which is of the same order of magnitude as thevariation obtained when comparing different codes and/or participants. This indicates that some physicalprocesses are still poorly known and inadequately modeled. It also appears that some model parameters
cannot be varied independently and that sensitivity analyses should look at the combined effect of aconsistent set of parameters. This issue was not investigated further.

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The code scattering in the calculation of some phenomena revealed some model weaknesses. Three
major phenomena are concerned. The first one is oxidation of molten mixtures and their relocation. The
second one is the prediction of UO2 melting and its interaction with molten corium. The third one is thecore coolability and the behaviour of hot corium when the vessel is entirely reflooded. The weaknessesof codes in the prediction of those phenomena are not surprising because experimental data are scarceand the physical understanding is still incomplete. Experimental programmes are currently under way orplanned to provide a better insight of those phenomena. It should help to improve the physical models
and the overall quality of prediction of codes.
The benchmark exercise involved some of the most important system codes which are currently used byutilities, TSO‟s, and regulatory agencies. Because of that, the benchmark has provided substantial andimportant insights. However, due to the limited number of participants and the fact that only one
scenario was calculated, the conclusions of this benchmark exercise must be considered with caution. It
is believed that some important codes did not participate in the benchmark. In particular
SCDAP/RELAP was absent although it is used in several countries. And also there were many major
nuclear actors absent of this exercise (in particular utilities but also TSOs from several countries).
Therefore, additional participants are strongly encouraged to join the group of participants if suchbenchmark action is continued.

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11. PROPOSALS FOR FUTURE EXTENSIONS OF THE WORKThe conclusions of this benchmark exercise must be considered with caution because of the limited
number of participants and because only one scenario was studied. However, the work performed during
this project constitutes a solid basis to continue such benchmark activity in order to confirm the current
conclusions. The proposed directions for future work are the following:
1. Improvement of the definition of the boundary conditions and plant description to eliminate any
remaining uncertainty and validate this sequence as a reference calculation to be used for further
benchmarks or even in training programmes for code users.
2. Extension of the number of parameters to compare, in particular boundary conditions for thecontainment: hydrogen mass flow rate at the break, mass of non-coolable corium, time of vessel
failure (if it happens), etc…
3. Better definition of the criteria to compare and improvement of the outputs of the codes (somecomparisons were difficult in the present exercise).
4. Different branch scenarios involving various SAM operations or system failures starting fromthe same initial conditions: depressurization, delayed start of HPI, loss of AFW, corium
relocation into the lower plenum and vessel failure.
5. Taking into account that, from the regulatory point of view, it is essential that uncertainties inbest-estimate calculations be well studied and assessed, it is suggested that uncertainty analysis
be investigated in future extension of this benchmark, using well known uncertainty analysis
methods. The BEMUSE program action, which is being carried out in the frame of GAMA,
could be a good reference for that.
6. Extension of the number of participants to involve:More countries
More users (in particular utilities and TSO‟s).

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12. REFERENCES

1. TMI-2 Analysis Exercise Final Report, NEA/CSNI/R(91)8, 1991.
2. Progress made in the Last Fifteen Years through Analyses of the TMI-2 Accident performed in
Members Countries, NEA/CSNI/R(2005)1, 2005.
3. ISP-46 PHEBUS FPT1, NEA/CSNI/R(2004)18, 2004.
4. SARNET Benchmark on QUENCH-11 – Final Report, FZKA 7368 – SARNET CORIUM P008,
2008.

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13. APPENDIX: SCHEDULE OF THE PROJECT

1st meeting in November 2005 in Aix en Provence

2nd meeting in June 2006 in Aix en Provence

3rd meeting in January 2007 in Garching

4th meeting in June 2007 in Paris

5th meeting in November 2007 in Karlsruhe

Draft of final report prepared in February 2008

Final report.

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14. APPENDIX: PARTICIPATING ORGANISATIONS AND CODES USED

A few Members Countries have designated experts able to analyze the progress made in the predictionsof one or several codes. Each expert is either a code developer or a code user, with a long experience.
The operational sequence of events during the TMI-2 accident is well known by all of them. The
modelling of the circuits and reactor core has been rather well assessed by all experts. The main
uncertainties in the actual TMI2 sequence are the boundary flows, which were not recorded and havebeen adjusted by each expert independently. Therefore an alternative scenario with fixed boundary
conditions has been defined for this benchmark exercise.
All the participants to this analysis are shortly introduced below, with a mention of the code(s) they haveused.
14.1 ENEA
ENEA has been involved in several studies of the TMI-2 transient with different codes, since the lastOECD benchmark. An important step was the complete transient calculation with SCDAP/RELAP5,
which showed a significant progress in the capabilities of a severe accident code. More recently, ENEA
particpated to a benchmark exercise in the frame of the COLOSS European Project and calculated thefirst phases of the transient with ASTEC V1 (developed by GRS and IRSN) and ICARE/CATHARE V1(developed by IRSN).
In the present exercise, ENEA has used ASTEC V1.3.
14.1.1 Brief Description of the ASTEC V1.3 CodeThe European ASTEC code is being jointly developed by IRSN and GRS. It is an integral code able toassess the whole severe accident sequence in a nuclear power plant, from the initiating event up to
fission product release and behaviour in the containment.
The code includes several coupled modules that can deal with the different severe accident phenomena:
thermal-hydraulics in the reactor system, core degradation and melt release, fission product release andtransport, ex-vessel corium interaction, aerosols behaviour and iodine chemistry in the containment, etc.
Among them, the CESAR module is used to compute the thermal-hydraulics in the primary andsecondary systems of the reactor. Such module is coupled to the DIVA module that computes coredegradation, melt relocation and behaviour in the lower head up to vessel failure.
The CESAR module allows a detailed representation of all components of primary and secondary
circuits including auxiliary, emergency and control systems. CESAR is a two-phase flow thermalhydraulic code. The gas phase can be a mixture of steam and hydrogen. The solution of the problem isbased on two mass equations, two energy equations, one equation for steam velocity, and a drift flux
correlation for water velocity. The state variables computed by CESAR are: total pressure, void fraction,
steam and water temperature, steam and water velocity, and partial pressure of hydrogen. All hydraulics
components can be discretized by volumes (one mesh) or axial meshed volumes and connected by61

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junctions. The volumes can be homogeneous or with a swollen level. Thermal structures are used tomodel the walls of the components, and compute thermal heat exchange between primary and secondary
systems and heat losses to the environment.
The DIVA module, which is a simplified version of the ICARE2 core degradation code, can model thethermal-hydraulics in the part of the vessel below the top of the core: downcomer, lower plenum and thecore itself including the core bypass. There are simplified thermal-hydraulic models specific for DIVA
based on a simplified 2D gas (steam plus non condensable gases) model with water swollen level. The
simplified model of the lower head of DIVA has one single mesh for fluids, three layers for corium(pool, metal and debris), and a 2D meshing for the vessel.
The DIVA module is activated to compute core heatup and degradation, in coupled mode with CESAR,at the onset of core uncovery. Before DIVA activation, the thermal-hydraulics in the vessel and the core
is computed by CESAR through an automatic vessel model creation based on DIVA input deck.
The convective and radiative heat exchanges between core components and structures are computed by
DIVA. Most important core degradation phenomena are dealt with in DIVA including: (1) core material
oxidation and hydrogen generation, (2) control rod material interaction, melting and relocation, (3)zircaloy clad melting and fuel dissolution, (4) fuel rod clad failure and metallic melt relocation, (5)debris bed and molten pool formation and spreading in the late degradation phase (these last models,
still under development and qualification, were not applied in the present analysis; therefore, fuel rod
melting and relocation in the late phase was calculated by decanting and candling models in rod-likegeometry).
14.2 GRS
GRS is developing the code ATHLET-CD with the aim to simulate severe accidents in the reactor
cooling system. The TMI-2 accident has been selected to assess the code for plant applications.
14.2.1 Brief Description of the ATHLET-CD CodeThe system code ATHLET-CD (Analysis of THermal-hydraulics of LEaks and Transients with Core
Degradation) is designed to describe the reactor coolant system thermal-hydraulic response during
severe accidents, including core damage progression as well as fission product and aerosol behaviour, tocalculate the source term for containment analyses, and to evaluate accident management measures. It isbeing developed by GRS in cooperation with the Institut für Kernenergetik und Energiesysteme (IKE),
University of Stuttgart. ATHLET-CD includes also the aerosol and fission product transport code
SOPHAEROS which is being developed by the French Institut de Radioprotection et de Sûreté
Nucléaire (IRSN).
The ATHLET-CD structure is highly modular in order to include a manifold spectrum of models and tooffer an optimum basis for further development. ATHLET-CD contains the original ATHLET models
for comprehensive simulation of the thermo-fluid-dynamics in the coolant loops and in the core. The
ATHLET code comprises a thermo-fluid-dynamic module, a heat transfer and heat conduction module,
a neutron kinetics module, a general control simulation module, and a general-purpose solver ofdifferential equation systems called FEBE. The thermo-fluid-dynamic module is based on a six-equation
model, with fully separated balance equations for liquid and vapour, complemented by massconservation equations for up to 5 different non-condensable gases and by a boron tracking model.
Alternatively, a five-equation model, with a mixture momentum equation and a full-range drift-flux
formulation for the calculation of the relative velocity between phases is also available. Specific models
for pumps, valves, separators, mixture level tracking, critical flow etc. are also included in ATHLET.

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The rod module ECORE consists of models for fuel rods, absorber rods (AIC and B4C) and for the fuelassemblies including BWR-canisters and -absorbers. The module describes the mechanical rod
behaviour (ballooning), zirconium and boron carbide oxidation (Arrhenius-type rate equations), Zr-UO2
dissolution as well as melting of metallic and ceramic components. The melt relocation (candlingmodel) is simulated by rivulets with constant velocity and cross section, starting from the node of rod
failure. The model allows oxidation, freezing, re-melting, re-freezing and melt accumulation due toblockage formation. The feedback to the thermal-hydraulics considers steam starvation and blockage
formation. Besides the convective heat transfer, energy can also be exchanged by radiation between fuelrods and to surrounding core structures.
The release of fission products is modelled by rate equations or by a diffusion model within the module
FIPREM. The release of aerosols is described by rate equations. The release of control rod materials(Ag, In, Cd) is based on temperature functions taking into account the partial pressure of the material
gases. The transport and retention of aerosols and fission products in the coolant system are simulatedby the code SOPHAEROS.
For the simulation of debris beds a specific model MEWA is under development with its own thermalhydraulic equation system, coupled to the ATHLET-thermo-fluid-dynamics on the outer boundaries ofthe debris bed. The transition of the simulation of the core zones from ECORE to MEWA depends onthe degree of degradation in the zone. The code development comprises also late phase models for coreslumping, melt pool behaviour and vessel failure.
The code system ATHLET/ATHLET-CD is coupled to the containment code system COCOSYS, and itis the main process model within the German nuclear plant analyzer ATLAS. The ATLAS environment
allows not only a graphical visualisation of the calculated results but also an interactive control of dataprocessing
The code validation is based on integral tests and separate effect tests, proposed by the CSNI validation
matrices, and covers thermal-hydraulics, bundle degradation as well as release and transport of fission
products and aerosols. Recent post-test calculations have been performed for the out-of pile bundle
experiments QUENCH-07, QUENCH-08, QUENCH-09, QUENCH-10 and QUENCH-11 as well as forthe in-pile experiments Phébus FPT2 and FPT3. The TMI-2 accident is used to assess the code forreactor applications.
The development and the validation of ATHLET-CD are sponsored by the German Federal Ministry of
Economics and Technology (BMWi).
14.3 IVS
IVS has been involved in TMI-2 study in the frame of bilateral co-operation with IRSN devoted to
ASTEC validation. The first TMI-2 calculations were performed with ASTEC V1.0 code in 2002-2003.
In the present exercise IVS has used ASTEC V1.3 code.
14.4.1 Brief Description of the ASTEC V1.3 CodeThe code is the same as the one used by ENEA. It is described in section 14.1.
14.4 NRC - SNL
The main computational tool for the NRC severe accident analysis studies is MELCOR 1.8.6.

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14.4.1 Brief Description of the MELCOR CodeMELCOR is a fully integrated computer code that is capable of modeling the progression of severe
accidents in light water reactor nuclear power plants. It has been developed for the USNRC by Sandia
Laboratories since 1985. The development of MELCOR fulfills NRC's objective of having an analyticalcapability for predicting the complete evolution of postulated severe accidents in nuclear power plants insome reasonable level of detail.
MELCOR models a broad spectrum of severe accident phenomena in both boiling and pressurized water
reactors is treated in MELCOR in a unified framework. Various code 'packages' models:

thermal-hydraulic response of the primary reactor coolant system, the reactor cavity, thecontainment, and the confinement buildings,

core uncovering (loss of coolant), fuel heatup, cladding oxidation, fuel degradation (loss of rod
geometry), and core material melting and relocation,

heatup of reactor vessel lower head from relocated fuel materials and the thermal andmechanical loading and failure of the vessel lower head, and transfer of core materials to thereactor vessel cavity,

core-concrete attack and ensuing aerosol generation,

in-vessel and ex-vessel hydrogen production, transport, and combustion,fission product release (aerosol and vapor), transport, and deposition



behavior of radioactive aerosols in the reactor containment building, including scrubbing inwater pools, and aerosol mechanics in the containment atmosphere such as particle
agglomeration and gravitational settling, and,

impact of engineered safety features on thermal-hydraulic and radionuclide behavior

Although MELCOR is modular in structure the code is highly integrated and flexible. For example,
there are no models for reactor components such as steam generators and pressurizers. Instead, thesecomponents are constructed from control volumes, flow paths, and heat structures. In addition, there is
one package for modeling the thermalhydraulics that is used in all aspects of the reactor system.
The governing equations for thermal-hydraulic behavior in MELCOR are the equations of conservationof mass, momentum, and energy. A semi-implicit (linearized) formulation of the governing equations isused to permit timesteps greater than the acoustic Courant limit. Also, MELCOR uses a full two-fluid
treatment rather than a drift-flux formulation and the resulting equations are iterated when necessary sothat the result is fully implicit with respect to pressures used in the momentum equation. A significant
feature of this method is that the resulting equations are exactly conservative (to within machineroundoff) with respect to masses and to thermal energy.
The MELCOR COR package calculates the thermal response of the core and lower plenum structures,
including the portion of the lower head directly beneath the core, and models the relocation of corematerials during melting, slumping, and formation of molten pool and debris. Fuel pellets, cladding, grid
spacers, canister walls (for boiling water reactors [BWRs]), core baffles and formers (for pressurizedwater reactors [PWRs]), other structures (e.g., control rods or guide tubes), molten pools, and particulate
debris are modeled separately within individual cells, the basic nodalization unit in the COR package.
Many new modeling enhancements have been added to the MELCOR 1.8.6 COR package to improvethe capabilities of the code to better represent the late-phase behavior of severe accidents. These new64

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models include hemispherical lower head geometry, models for simulating the formation of molten
pools both in the lower plenum and the upper core, crust formation, convection in molten pools,
stratification of molten pools into metallic and oxide layers, and partitioning of radionuclides between
stratified molten pools.
14.5 University Pisa
The University of Pisa was regularly involved in severe accident studies, mainly using SCDAP/RELAP5and MELCOR codes. The experience gained from these studies helped to achieve a calculation of thefirst two phases of the TMI-2 accident. The same nodalisation has been employed, with the obvious
modifications to the boundary conditions, for the simulation of this benchmark.
14.5.1 Brief Description of the MELCOR V1.8.5 CodeThe code is the same as the one used by NRC-SNL. It is described in section 1.7.1. However, UPI has
used the version 1.8.5 which differs from the version 1.8.6, mainly by the modelling of late phasebehaviour.
14.6 IRSN
The ICARE/CATHARE system is developed by IRSN to meet requirements in light water reactor safety
analysis (PWR1, VVER2, EPR3, etc.) and level 2 probabilistic safety analysis (PSA). This system
results from the combination of the ICARE2 mechanistic code for core degradation developed by IRSN
and the CATHARE2 thermohydraulic code developed in collaboration by CEA, IRSN, EDF and
FRAMATOME-ANP.
This system is also an excellent synthesis tool for all phenomenological knowledge on core degradation.
ICARE/CATHARE is designed to :
evaluate the consequences of a severe accident upon the primary system of a PWR (from theinitiating event to vessel failure) ,calculate experimental programmes conducted by IRSN and French or international partners.

14.6.1 Brief Description of the ICARE/CATHARE V2 CodeICARE/CATHARE is composed of a series of modules that each deal with a specific phenomenon :thermohydraulics, thermics, mechanics, chemistry, fission products, movement of materials, debris beds,
core meltdown, etc.
The V1 version (released from 1999) has been used for IRSN level 2 PSAs applied to 900 MWe
reactors.
The first V2 version (V2.1) combined with a 3D thermohydraulic model has been released in 2007 andsimulates all physical phenomena, including vessel failure. The system is designed to facilitate theintegration of new models and perform calculations both for experimental programmes and reactors.

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The validation phase carried out in partnership with about ten foreign organisations includes :

approximately one hundred separate effect tests (chemistry, mechanics, reflooding, etc.),

about thirty integral degradation tests making it possible to validate core & structure degradation
models (PHEBUS, MAESTRO, PBF, QUENCH, etc.) and enabling coupling with the primarysystem (LOFT),reactor sequence calculations.

Most of the thermohydraulic models benefit from the validation of the CATHARE code.
Development is achieved in a high-level scientific environment: fundamental research (theses & post
doctoral studies), collaboration with universities and French & foreign research organisations [(CEA,
IMFT/Toulouse, Kurchatov & IBRAE Institutes (Russia), ENEA (Italy), etc.)], not to mention
participation in several European R&TD Framework Programmes and international projects (OECD
MASCA). International seminars for the Users‟ Club are held on a regular occasion and unite about fifty
users from various different countries.
14.7 Seoul National University (SNU)
14.7.1 Brief Description of the MAAP4 Code 4.03
The Modular Accident Analysis Program (MAAP4) was developed by Fauske & Associates, Inc.
MAAP4.03 is a fast running, integrated severe accident analysis code to simulate a transient andspecifically accounts for system events that occur during the transient, including operator actions.
MAAP4, as a severe accident code, is most widely used by nuclear utilities and vendors because of its
short run time and reduced requirements for code expertise. In addition, MAAP4 can be used by anexisting plant to simulate how a proposed modification would affect plant operations. Also, plantdesigners could use the code to predict the performance of future plants if a certain set of conditions
were imposed on those plants.
MAAP4 is a fully integrated severe accident analysis code and includes models for important thermal
hydraulic and fission product phenomena which may occur during a postulated accident in a pressurizedwater reactor plant.
The underlying models for MAAP4 address the following phenomena:









Reactor coolant system (RCS) thermal hydraulics
Cladding water reaction
Reactor core heatup, melting and relocation
Containment thermal hydraulics
Fuel coolant interactions
Molten core concrete interactions
Hydrogen combustion
Fission product release, transport and deposition.

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MAAP4 has capability to model the following in-vessel recovery and advanced reactor systems:






Modeling of core melting and relocation; and separately tracking UO2, Zircaloy, ZrO2, andcontrol rod materials
Reactor vessel cooling by containment water surrounding the vessel
Reactor vessel lower head creep rupture
Generalized containment model
Generalized user-defined event codes.

MAAP4 solves a set of lumped parameter, nonlinear, first order, coupled, ordinary differential equations
in time. These equations generally express the conservation of mass and energy. Momentum balances in
MAAP are assumed to be quasi-steady. MAAP4 then calculates thermal hydraulic and fission productvariables at each time step.
The modeling of regions, components, and phenomena is fully integrated into the code. For example,
thermal hydraulic and fission product models are generally applicable to such major regions as thereactor coolant system, pressurizer and containment nodes treated in the code. The thermal hydraulicparameters are calculated at each time step, and the fission product parameters are overlaid on thethermal hydraulic models. This feedback permits full interaction between the models. In addition, all of
the models are run in parallel. The integrated reactor coolant system and containment modeling is ideal
for the passive plant application.
The execution of the MAAP4 code requires the development of two input files: a parameter file and asequence file. The parameter file describes the nonvariant, plant specific data including geometry,
performance data, and system setpoints. The sequence file describes the severe accident sequence to beanalyzed including the initiating event, system operability status, and operator actions. MAAP4
automatically initiates system operation, based on system availability described in the sequence file andsetpoints described in the parameter file. Similarly, operator actions can be modeled in the sequence fileby employing the user-defined event codes to describe the timing and implementation of emergency
actions.
The results of the MAAP4 analysis describe the severe accident progression and fission product releases
associated with accident sequences analyzed. MAAP4 evaluates the following details of the severe
accident:Time histories of the important state parameters, such as pressure, temperature, and inventories
in each control volume of the reactor coolant system and containment models

Timing of engineered safety system initiation and termination

Timing and magnitude of severe accident phenomena predicted to occur in an accident sequenceTime history for fission product locations and states, including fission product releases to theenvironment

14.8 IKE
IKE is developing models for the German system code ATHLET-CD, especially for the late phase ofcore degradation and for the coolability of core debris. In the benchmark calculations, IKE was applying
the ATHLET-CD code to model the plant behaviour. Up to the onset of quenching, standard models
were used. During the quenching phase, the MEWA module, which is based on a porous medium

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approach, was activated, in order to describe the behaviour of a particulate debris bed which was
assumed to be formed during flooding.
14.8.1 Brief Description of the ATHLET-CD/MEWA CodeSince the ATHLET-CD code is already described in more detail in the GRS contribution, here only abrief descriptin of the MEWA module is provided.
The MEWA module is being developed and integrated in the German system code ATHLET-CD for thedescription of late phase core melting. It is developed from a combination of the MESOCO and WABE
models. The MEWA model is designed to describe the processes of late phase core degradation withmassive melting, melt relocation, molten pool formation and behaviour up to the relocation of melt tothe lower head, taking into account the effects of possible core refloding. Emphasis is on the potential
accumulation of larger melt masses and on the resulting modes of outflow from the core, which
determine the chances of coolability by water injection and the subsequent processes in the lower head.
The processes in a strongly degraded core are described in a quasi-continuum approach in twodimensional cylindrical geometry. Solid, melt and two-phase fluid (gas/steam) are modelled as separate
phases with thermal non-equilibrium between all phases. Several material components can beconsidered in the solid and melt phase, especially metallic and ceramic components are distinguished.
The gas phase can be composed of steam and hydrogen to take into account oxidation of metallic
components. Melt velocities are calculated according to the Ergun approach for flow in a porous
medium. Formation of a molten pool is detected from the quasi-continuum model based on melt volume
fractions. The thermalhydraulics of the molten pool are not simulated directly. The thermal behaviour isdescribed either by empirical correlations or by a representative model assuming an axial temperature
profile in the center and a boundary layer flow along the crust of the pool.
Simplified momentum conservation equations are also used for the two-phase fluid, assuming dominantfriction between the solid matrix and the phases and neglecting temporal and spatial derivatives of thevelocities. For the particle-fluid drag forces the model of Ergun is adapted which is presently extended
for two separated fluids by the introduction of relative permeabilities and passabilities according toReed. Alternatively, a more detailed formulation of drag forces according to Tung and Dhir can bechosen which takes into account flow pattern for bubbly, slug and annular flow and an explicit
formulation of interfacial drag between steam and water. The explicit formulation of interfacial drag isconsidered as essential when multidimensional flow occurs.
In integrated calculations with ATHLET-CD, the modelling can be switched from the description by therod-oriented ECORE model to the porous media model MEWA, based on user-defined criteria on
porosity, melt fraction or temperature. This can be done progressively for parts (groups of meshes) of
the core. Since the thermalhydraulics of the coolant in the debris bed is described by MEWA, the
ATHLET thermal-hydraulic modelling is switched off in the respective region. The coupling is provided
through the exchange of mass and enthalpy flows over the region boundaries.
The development and the validation of the MEWA late phase module for ATHLET-CD are sponsoredby the German Federal Ministry of Economics and Technology (BMWi).

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15. APPENDIX : ENEA CALCULATION WITH ASTEC V1.3

15.1 TMI-2 Plant Modelling
The TMI-2 plant nodalization scheme employed in this ASTEC V1.3 application is shown in the figure
below. The plant nodalization includes a detailed modelling of the primary coolant system with:






the reactor pressure vessel volumes and structures, including the VENT valve between the cold
collector and the hot collector;
the two primary coolant loops (1 hot leg and 2 cold legs in each loop) with once-through steam
generators and main coolant pumps;
the pressurizer with surge-line, PORV, heaters, spray-line and valve;
the external walls of the primary circuit;main emergency and control systems.

Modelling of the secondary systems is limited to the secondary side of steam generators, the steam lines
with isolation valves, and main feedw-ater and auxiliary feedwater injections.
Breaklocation

OTSG-AOTSG-B

Pressurizer

Steam
Line

Upper head

feedwater

Upper plenum

VesselSteamLine

Check valve

feedwater
Pump

Pump

Make-up
Core

LOOP B

LOOP A

Lower plenum

TMI-2 plant nodalization scheme

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The core is represented by 6 radial rings plus the by-pass, and axially discretized in 20 equal meshes.
The core baffle, the barrel and the thermal shield at the core periphery are also represented. One
representative fuel rod component and one control rod component are modelled in each ring. The control
rod component simulates all the full and part-length control rods, all the guide tubes (including thosecontaining burnable poison rods) and all the instrument tubes.
The plant geometry, the boundary conditions and the accident scenario have been strictly definedaccording to alternative TMI-2 scenario benchmark specifications provided by IRSN, as well asrecommended values of code physical parameters for the sensitivity study.
15.2 Updated Initial Steady-State Conditions
The TMI2 plant state at transient initiation (initial conditions at t = 0) is obtained by a steady-state code
run lasting 500 s and starting from plant thermal-hydraulic parameter values close to the ones specified
for TMI-2. During the steady-state calculation, some regulations in the primary and secondary sides areactivated to facilitate the achievement of stable conditions, as close as possible to the TMI2 plant state at
t = 0 (turbine trip in TMI2 accident).
The regulated parameters are:

The pressurizer pressure by turning on the heaters when the pressure is lower than the nominal
value;

The pressurizer liquid mass, by water injection or draining, in order to obtain the precise liquid
level;

The feedwater mass flowrate at the steam generator downcomer inlet according to the steam
mass flowrate at the dome outlet;

The primary loop mass flowrate, according TMI2 data, varying the pump rotation speed during
the first 100 s of steady-state run.

The TMI2 plant initial conditions calculated by ASTEC V1.3 are compared with TMI2 accident data at
turbine trip in the table below. The primary circuit conditions are very well reproduced by ASTEC.
Since the turbin trip in TMI2 accident was provoked by instabilities and sudden pressure increase in thesecondary circuits, the steady-state conditions calculated by ASTEC for the secondary side cannot
precisely match the TMI2 data.
Parameter

Unit

ASTEC V1.3

TMI-2

Reactor Power

W

2700

2700

Primary Pressure

MPa

15.2

15.2

Temperature Hot Leg A

K

592.1

592

Temperature Hot Leg B

K

592.1

592

Temperature Cold Leg A

K

563.8

548-561

Temperature Cold Leg B

K

564.1

565

Mass Flow Rate – Loop A

kg/s

8290

8280

Mass Flow Rate – Loop B

kg/s

8560

8560

Pressurizer Level

M

5.78

5.77

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Parameter

Unit

ASTEC V1.3

TMI-2

Total Primary Masston

225120-Pressure SG A

MPa

6.80

7.31

Pressure SG B

MPa

6.80

7.24

Steam Temperature SG A

K

574

586

Steam Temperature SG B

K

575

586

Collapsed Level SG A

M

3.7

-

Collapsed Level SG B

M

3.5

-

Liquid Mass SG A

Kg

19210

-

Liquid Mass SG B

Kg

18410

-

Feedwater Flow SG A

Kg/s

742

723

Feedwater Flow SG B

Kg/s

759

717

SG Feedwater Temperature

K

513

-

15.3 Standard Physical Parameters of the CodeMain core degradation physiscal parameters used in the standard calculation with ASTEC are
synthetized in the table below. The physical parameters used in this standard calculation are the ones
used for the interpretation of the first two phases of the TMI2 accident with ASTEC V1.2 code in the
SARNET projectParameter

ASTEC V1.3

Zr Oxidation correlation

Correlations proposed by Schanz (FZK):Cathcart correlation if T < 1800 K
Prater correlation if T > 1900 K
Linear interpolation between 1800 – 1900 K

Cadding failure criteria

T > 2300 K and e(ZrO2) < 300 μm
T > 2500 K

UO2-ZrO2
temperature

melting 2830 K

Debris formation criteria

No debris bed and molten pool formation
modelling (standard and sensitivity case)

Debris diameter-

Debris porosity-

Clad deformation and burst due to overpressure is calculated by the creep model derived from French
EDGAR experiments. The consequent release of fission products from failed pins and their transport
through the primary circuit is not addressed in the present analysis.

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The Urbanic-Heidrick correlation is used by default in ASTEC V1.3 for zircaloy oxidation. Other
correlations (Cathcat-Pawel and Prater-Courtright) can be used or defined by the code user. In thisstandard calculation, the correlation proposed by Schanz (FZK – Germany) has been defined andapplied to compute zircaloy oxidation and hydrogen generation. This correlation is now implemented as
default in the ICARE2 code that will be part of the next ASTEC code release.
Control rod failure and relocation occurs at the melting point of stainless steel of absorber rod cladding
through dissolution of the zircaloy guide tube by molten steel. Oxidation of steel is modelled byparabolic laws, once the melt relocates downwards on the external surface of zircaloy guide tube.
When fuel rod temperature exceeds the melting point of zircaloy, the Kim-Olander convective model isapplied to evaluate the fuel dissolution by molten zircaloy; the simultaneous dissolution of oxide layer isevaluated by the Hofmann diffusive model. The limit in fuel dissolution is based on the liquidus line ofternary U-Zr-O diagram. During this phase, fuel rod clad failure and metallic melt relocation is assumed
to occur if clad temperature exceeds 2300 K and oxide layer thickness is lower than 300 μm, or, in any
case, when clad temperature exceeds 2500 K. Once the clad fails, the molten U-Zr-O mixture is
relocated downwards on the external surface of cladding, at constant velocity (0.6 m/s), until its
solidification in a colder region of the core.
Debris bed and magma models are not applied in the late core degradation phase; therefore, fuel rod
melting and relocation is computed in rod-like geometry by decanting and candling models. Fuel rod
melting is calculated at 2830 K, according to the eutectic temperature of UO2-ZrO2 binary phase
diagram.
15.4 Chronology of Major Events
The chronology of major events calculated by ASTEC is presented in the table below and discussed in
the following section 5.
Time (s)EventStandard caseSensitivity case

Break opening and loss of feedwater0

0

Reactor scram

20

20

Pressurizer is empty

145

145

Primary pump void fraction > 0.6

4740

4740

Primary pump shutdown (M < 85000 kg)

4825

4825

Onset of core uncovery

5580

5580

Onset of core heatup

5950

5950

Beginning of oxidation (DIVA start)

6431

6431

First fuel rod clad burst

7145

7130

First melt relocation (U-Zr-O)

7370

7428

First ceramic melting (UO2-ZrO2)

8820

8550

Onset of reflooding

9825

9825

End of calculation

12000

10928

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15.5 Result Analysis
Despite the hot leg break opening at 0 s, the contemporary feedwater trip with consequent loss of heatremoval by the secondary side results in primary pressure increase. The pressure rise leads to reactor
scram (P > 16.3 MPa) by the protection system at 20 s. After reactor scram the primary pressure starts todecrease. The primary pressure decrease is accelerated when the pressurizer is empty at 145 s, and theprimary pressure approximates the secondary pressure after about 300 s.
Before pump shutdown, a mixture of liquid water and steam circulates in the primary circuit and throughthe core in saturated conditions, while the void fractions in the primary system increase due to the lossof fluid (mostly liquid) from the hot leg break. In this conditions, the core decay power and pump heat
are removed by the steam generators. The primary pump void fraction is greater than 60% after 4740 s.
According to benchmark specifications, the primary pumps are shutdown a little bit later at 4825 s, when
the total primary mass decreases down below 85000 kg.
After pump shutdown, the liquid water starts to settle down and in the primary circuit and inside thevessel. By this time, the break flow rate quickly reduces, changing from mostly liquid to pure steam
flow due to liquid stratification in the primary system with hot leg draining. The core starts to uncover at5580 s, while core heatup starts at 5950 s at the core top, because the core decay power is no more
removed by natural circulation in the primary circuit.
When the maximum core temperature exceeds 773 K at 6431 s, the DIVA module of ASTEC is
activated to compute in-vessel core degradation and melt relocation. By this time onwards, the oxidationof the claddings is calculated by DIVA. First fuel rod clad burst due to overpressure is calculated by
ASTEC at 7145 s in the standard calculation, when the maximum clad temperature reaches 1250 K at
the top of the core centre. In the sensitivity calulation, first clad burst occurs a little bit earlier at 7130 s
due to higher oxidation rate below 1200 K with Urbanic-Heidrick correlation. Oxidation runaway withsudden temperature excursion occurs first at the top of the core at 7360 s in the standard calculation. It isfollowed after few seconds by clad failure (T > 2300 K) and first melt relocation. First clad failure andmelt relocation is calculated later on at 7428 s in the sensitivity calculation due to the lower oxidation
rate of Urbanic-Heidrick correlation above1200 K.
The different oxidation kinetics is also likeky the main reason for the different ratio of UO2 dissolution
in the two ASTEC calculations (565 kg and 1575 kg in standard and sensitivity calculations,
respectively). In fact, higher oxidation rate above the melting point of zircaloy in the standard
calculation results in earlier clad failure and consequent lower duration of molten zircaloy and UO2contact. The slightly difference in clad failure criteria does not seem to contribute significantly to thedifference in the amount of UO2 dissolved.
Fuel rod melting starts much earlier in the sensitivity calculation, at 8550 s, due to the lower melting
point of UO2-ZrO2 considered in this calculation. The total mass of molten UO2 before reflood is30590 kg (including dissolved UO2), much higher than the one of the standard calculation that is 8900
kg. The total mass of molten metals before reflood (17190 kg and 14510 kg in standard and sensitivity
calculations, respectively) is about 15% higher in the standard calculation due to larger axial extent ofdegradation in the more external core rings, as evidenced in synthetic core degradation views in
section 6.
The mass of hydrogen produced before reflood is 331 kg in the standard calculation and 278 kg in thesensitivity calcualtion. The large difference is partly due to different oxidation kinetics and partly causedby different axial extent of oxidized zone at core periphery.

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Core reflood by HPI injection is started at 9825 s, 5000 s after pump trip according to benchmark
specifications. By this time, the collapsed core water level is about 60 cm above the core bottom in both
calculations. The reflood of the core results in sudden primary pressure increase due to steam formationand hydrogen generation. Maximum pressure peak is larger in the standard calculation (10.4 MPa) thanin the sensitivity one (9.2 MPa) due to enhanced hydrogen release during reflood.
After reflood, the core is rapidly and completely quenched in about 8 minutes. The hydrogen mass
produced during reflood is 29 kg in the standard calculation, about 8% of the total H2 mass release (360kg), and only 7 kg in the sensitivity calculation (total mass release = 285 kg). The difference seems
related to material mixture oxidation, as confirmed by the sharp decrease of the fraction of non-oxidized
Zr in the molten material (approximately 5%) observed in the standard calculation during reflood.
Core degradation and increase in the total amount of molten materials during reflood is not verysignificant in both standard and sensitivity calculations (1960 kg and 520 kg, respectively), as indicatedby synthetic core views before and after reflood (see section 6).
Note: The boundary conditions of the steam generators deviate from the specification values after 6700 s
and 10700 s (SG pressure decrease and water level increase) owing to steam condensation phenomena inthe secondary side in isolated conditions (no auxiliary feedwater injection).
15.6 Major Discrepancies with Other Code Results
In general the thermal-hydraulic behavior of the primary system calculated by ASTEC is consistent withother code results. The main discrepancies concern:
The heat transfer to the secondary side before pump stop, which is much larger with ASTEC likely due
to overestimation of total pump heat loss into the primary circuit.
The primary pressure increase after pump stop, which is avoided in ASTEC calculation bysteam circulation in the primary circuit and condensation on steam generators with heat decay
heat removal.

The primary pressure after reflooding that follows the steam generator pressure (heat removal by
the secondary side).



The core uncovery that starts with some delay with respect to other codes (likely much water
settle down inside the vessel after primary pump shutdown).

The rod heatup at core top is larger than other code results, likely due to lower natural
circulation calculated by ASTEC inside the core.

Cumulated hydrogen production and total mass of molten materials before and after reflood is consistent
with other code results.

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15.7 Synthetic Views of the Core at Selected Instants
15.7.1 Standard Case

STANDARD CASE: Beginning of oxidation (6431 s) – Global view75NEA/CSNI/R(2009)3

STANDARD CASE: Beginning of oxidation (6431 s)

Temperature distribution

Void fraction distribution

Geometry and degradation76

NEA/CSNI/R(2009)3

STANDARD CASE: Time of first melt relocation (7370 s) – Global view77

NEA/CSNI/R(2009)3

STANDARD CASE: Time of first melt relocation (7370 s)

Void fraction distribution

Temperature distribution

Geometry and degradation

78

NEA/CSNI/R(2009)3

STANDARD CASE: Time of reflooding (9825 s) – Global view79

NEA/CSNI/R(2009)3

STANDARD CASE: Time of reflooding (9825 s)

Temperature distribution

Void fraction distribution

Geometry and degradation

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NEA/CSNI/R(2009)3

STANDARD CASE: End of calculation (12000 s) – Global view81

NEA/CSNI/R(2009)3

STANDARD CASE: End of calculation (12000 s)

Temperature distribution

Void fraction distribution

Geometry and degradation82

NEA/CSNI/R(2009)3

15.7.2 Sensitivity CaseSENSITIVITY CASE: Beginning of oxidation (6431 s) – Global view

Same Results as Standard CaseSENSITIVITY CASE: Time of first melt relocation (7428 s) – Global view83

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SENSITIVITY CASE: Time of first melt relocation (7428 s)

Temperature distribution

Void fraction distribution

Geometry and degradation

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SENSITIVITY CASE: Time of reflooding (9825 s) – Global view85

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SENSITIVITY CASE: Time of reflooding (9825 s)

Void fraction distribution

Temperature distribution

Geometry and degradation

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SENSITIVITY CASE: End of calculation (10928 s) – Global view87

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SENSITIVITY CASE: End of calculation (10928 s)

Temperature distribution

Void fraction distribution

Geometry and degradation

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16. APPENDIX : GRS CALCULATION WITH ATHLET-CD

16.1 TMI-2 Plant Modelling
The nodalization adopted for the ATHLET-CD benchmark calculations is shown in Fig. 16-1. It consistsof the reactor pressure vessel (RPV), the two coolant loops A and B with the once-through steam
generators, four cold legs with main coolant pumps, four high pressure injection lines connected to thecold legs and one let-down in loop A1, as well as the pressurizer with the surge line connected to the hotleg of loop A, heaters, spray line and pilot operated relief valve (PORV).

PORV

PB-SG-IN

PB-HL

PA-SG-IN

PA-HL

PB-SG

PA-SG

PRESS

PV-UH
HPIA2

HPIB1
PV-UP-T

HPIA1

HPIB2
PV-UP-M1 PV-UP-M2

LEAK

PV-DC-MA
SURGE

PV-CORE1
PV-CORE2

PV-DC-B
PB-CL1

PV-CORE3
PV-CORE4
PV-CORE5
PV-REF

PB-CL2

PA-CL2
PA-CL1

PV-LP
PV-LH
PB-SG-OUT

PA-SG-OUT

LETDOWNA

Fig. 16-1: ATHLET-CD nodalization scheme for TMI-2 (primary circuit)

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NEA/CSNI/R(2009)3

The RPV comprises the downcomer, lower and upper plenum, upper head, the core region and the core
bypass. The vent valves between downcomer and upper plenum are modelled as check valves. The coreis modelled by five concentric rings with 22 axial nodes (20 within the active core region) and withcross flow connections to allow flow deflection due to fuel rod deformation and blockage formation
caused by refreezing of molten material. The three inner core rings include fuel and AIC control rods.
The fourth ring contains only fuel rods. The fifth channel contains no rods and is defined to avoid acomplete core flow blockage in case of strong melt relocation.
The simplified model of the secondary system consists of two components (loop A and B) simulating
the riser, with 16 axial volumes, and steam dome as well as the boundary conditions for feedwater
injection and steam outlet flow to control the secondary side water levels and pressures according to thebenchmark specifications. In total, the nodalization comprises 376 control volumes, 407 flow paths, aswell as 153 heat slabs (not including fuel and control rod components) for the modelling of RPV
structures and pipe walls.
16.2 Main modelling options
The benchmark calculations were performed with the code version ATHLET-CD Mod. 2.1 Cycle A,
released in August 2006. Only the modules ATHLET (thermal-hydraulics) and ECORE (coredegradation) have been applied.
The system thermal-hydraulics was simulated with the six-equation model (fully separated balance
equations for liquid and vapour), complemented by an additional mass conservation equation forhydrogen as a non-condensable gas, except for the pressurizer and for the steam generator secondarysides, where the five-equation approach (one mixture momentum equation) together with the mixture
level tracking model was used. The quench front model, which takes into account both top and bottom
reflooding, has been applied for all rod components.
According to the benchmark specifications, two sets of calculations have been performed, one with input
parameters for core degradation as recommended in the code User‟s Manual (referred as “standard
calculation”) and the other using values deduced from the analyses of the PHÉBUS experiments(referred as “alternative calculation”). As it can be seen in Table 16-1, both sets of input parameters arevery similar, except for a 50 K lower failure criterion for oxidized cladding and a 100 K lower UO2ZrO2 melting temperature in the alternative calculation.
Fuel rod relocation is simulated in rod-like geometry (candling model) assuming a constant relocation
velocity of 6 cm/s for metallic melt and 3 cm/s for ceramic melt, respectively. Radial melt spreading
outside a core ring is not taken into account.
Parameter

Unit

Standard Calculation

Alternative Calculation

Start of fuel dissolution by Zirconium

K

2250

2250

Clad failure temperature (δox < 0.3 mm)

K

2300

2280

Clad failure temperature (δox ≥ 0.3 mm)

K

2500

2450

Start of ceramic fuel and Zirconia melting

K

2600

2500

Correlation for cladding oxidation at low
temperature range (T<1800 K)-

Cathcart

Cathcart

Correlation for cladding oxidation at hightemperature range (T>1900 K)-

Urbanic-Heidrick

Urbanic-Heidrick

Tab. 16-1: ATHLET-CD code parameters relevant to core degradation

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16.3 Initial steady state conditions
The system code ATHLET/ATHLET-CD provides a steady state calculation capability which initializes
the complete system starting from a limited amount of user-supplied initial and boundary conditions. To
check the adequacy of this steady state solution, a zero-transient run (without any external disturbances)lasting 100 s was performed. At the end of this period, the main calculated plant parameters were stable
and thus this time point was chosen for transient initiation (break opening and loss of main feedwater).
Table 16-2 compares the main plant parameters at transient initiation calculated by ATHLET-CD with
TMI-2 accident data at turbine trip. The primary circuit conditions show a good agreement with theplant data. The secondary parameters reflect the estimated plant conditions before turbine trip. Due tothe simplified modelling of the steam generator secondary sides, the degree of steam overheating at thesteam generator outlets is underestimated, but this does not affect the results of the transient
calculations.
Parameter

Unit

GRS
(ATHLET-CD)

TMI-2(Turbine Trip)

Primary System
Reactor Power

MW

2663

2700

Primary pressure

MPa

14.9

15.2

Temperature Hot Leg A

K

592.2

592

Temperature Hot Leg B

K

592.2

592

Temperature Cold Leg A

K

564.8

548-561

Temperature Cold Leg B

K

564.8

565

Mass Flow Rate - Loop A

kg/s

8638

8280

Mass Flow Rate - Loop B

kg/s

8675

8560

Pressurizer level

m

5.78

5.77

Total primary masskg

223300-

Secondary System
Pressure SG A

MPa

6.34

7.31

Pressure SG B

MPa

6.20

7.24

Steam temperature SG A

K

568

586

Steam temperature SG B

K

568

586

Collapsed level SG A

m

4.03

-

Collapsed level SG B

m

3.54

-

Liquid mass SG A

kg

14630

-

Liquid mass SG B

kg

13140

-

Feedwater flow SG A

kg/s

743.7

723

Feedwater flow SG B

kg/s

738.5

717

SG feedwater temperature

K

503

-

Tab. 16-2: Initial conditions for TMI-2 benchmark calculation

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16.4 Discussion of Results
The main results of the ATHLET-CD calculations are depicted in Figs. 16-2 to 16-18. The blue, solid
curves refer to the “standard calculation”, the red, dashed ones to the “alternative calculation”. A
chronology of the main events is presented in Table 16-3. Synthetic views of the core status at selected
time points are shown in Section 16.5. The calculations have been performed up to about 30 min after
start of ECC injection.
Phase 1: Small break LOCA until start of core heat-up (t = 0 s to t = 5100 s)
The postulated accident starts with the opening of a small break at hot leg A, simultaneously with thecomplete loss of main feedwater. Due to the strongly reduced heat transfer to the secondary side,primary pressure increases (Fig. 16-2). The pressure set-point for PORV opening is reached within 7 s.
Shortly afterwards, reactor scram signal due to high pressure is generated. After reactor scram, primarypressure starts to decrease and the relief valves close (t = 15 s). About 200 s after break opening thepressurizer empties, and the hot legs reach saturation conditions.
At about 250 s flashing starts at the suction lines of the main coolant pumps. In comparison to the othercodes, ATHLET-CD predicts a stronger pump performance degradation under two-phase flow
conditions and thus considerably lower loop mass flow rates. The void fraction at pump inlet reaches
60% at ca. t = 4000 s (original benchmark criterion for pump trip).
The primary coolant mass inventory reaches 85 ton at t = 4676 s (Fig. 16-4) and the pumps are switched
off. With the breakdown of forced circulation, liquid coolant settles down very quickly, mainly in thereactor pressure vessel (Fig. 16-5) and in the pump loop seals. Due to liquid stratification, steam withentrained water flows through the break (Fig. 16-3).
The heat transfer to the secondary side is reduced abruptly, resulting in a considerable primary pressure
increase. Two-phase natural circulation is prevented by the liquid accumulation in the loop seals.
Afterwards the liquid level within reactor vessel decreases continuously, leading to core depletion (Figs.
16-5, 16-6) and start of core heat-up at t ~ 5100 s (Fig. 16-7), considerably earlier than predicted byother codes (ASTEC, ICARE/CATHARE). In these calculations, the liquid level within reactor vessel
remains practically constant for a quite long period (ca. 600 to 800 s) after pump trip. Reasons for suchdeviations may be found mainly in the different approaches to describe phase separation and countercurrent flow phenomena in the core region, but also in the modelling of the vent valves.
Phase 2: Core degradation phase (t = 5100 s to t = 9676 s)
After start of core dryout, fuel cladding temperatures at higher core elevations increase rapidly (Fig. 168). First control rod failure through dissolution of the guide tubes by molten steel is predicted to occur at
t = 7434 s.
Oxidation runaway with a sudden temperature excursion at the top of the core occurs at about 7700 s(Figs. 16-8 and 16-11). Shortly afterwards first fuel rod clad burst is predicted to occur due to overpressurization.
Clad melt and relocation starts in the central core ring at about 7800 s in the alternative calculation,slightly earlier than for the standard one, due to the 50 K lower temperature criterion for cladding
failure. From this time point on, results of the standard and of the alternative calculation begin todiverge. Due to more metallic Zr available above the core mid-plane, the standard calculation predicts ahigher hydrogen production (Fig. 16-11), leading to higher heat generation and higher temperatures (Fig.
16-9) in comparison to the alternative calculation.

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In both calculations the metallic melt, mainly formed in the core inner rings, flows downwards andrefreezes at an elevation slightly above the liquid surface. The amount of metallic crust, together withthe molten control rod materials, is enough to cause a partial channel flow blockage in Ring 1(Fig. 16-5).
The melt temperature of the UO2-ZrO2 eutectic mixture is reached at 8686 s in the standard calculation,
160 s later than in the alternative one. At this time point ca. 7 ton of metallic melt has been generated
and dislocated to the lower core regions. Afterwards ceramic melt becomes the main contributor to thetotal melt mass (Figs. 16-12 and 16-13). The hot ceramic melt can be only partially cooled in the lower
core regions, leading to the inception of a liquid melt pool within the innermost core channel(Fig. 16-14).
At the end of this phase, before the start of ECC injection, the calculated total molten mass amounts to
41 ton (standard calculation) and 31 ton (alternative calculation). The generated H2 mass is predicted tobe 294 kg and 258 kg, respectively. These global values agree satisfactorily with the results of otherparticipants.
However, significant deviations can be observed not only with respect to the melt composition(ATHLET-CD predicts a relatively low fraction of Zr melt, ca. 7 to 8 kg) but also concerning its
distribution within the core region.
Both calculations predict a large extent of fuel rod melting mainly in the innermost core channel,
creating a cavity between the elevations 1.7 m and 3.0 m above the core bottom, with a maximum
diameter of 1.6 m (alternative calculation) and 2.2 m (standard calculation, at elevation 2.3 m),
respectively. This fuel melt relocates completely to the lower inner core regions, displacing the water
there (Fig. 16-5) and leading to a flow channel blockage. The outer fuel rods (Ring 4) do not melt
significantly.
The main parameters concerning core degradation are summarized in Table 16-4 for both calculations.
Synthetic views of the core status at the end of the degradation phase are depicted in Fig. 16-21.
Phase 3: Core reflood phase (t = 9676 s to t = 11500 s)
High pressure (HP) injection in the cold legs at a constant, total rate of 60 kg/s is initiated at t = 9676 s,
5000 s after pump trip according to the benchmark specifications. At this time point, the collapsed waterlevel in the outer core rings (Rings 3 to 5) is about 60 cm above core bottom in both calculations (Fig.
16-5). Due to melt accumulation, the inner core channels are blocked.
At reflood initiation ECC water penetrates the core region mainly through the outer channels andthrough the core bypass. Due to the strong steam formation as well as hydrogen production, primarypressure increases considerably (Fig. 16-2). The pressure peak is higher in the standard calculation (10.4
MPa) as in the alternative one (9.9 MPa) due to the larger amount of hydrogen generated during reflood(Fig. 16-11).
In this phase, the liquid melt pool could only be cooled from the sides, leading to the formation of acrust of refrozen material (Figs. 16-14 and 16-15), which in turn acts as an additional resistance tofurther melt quenching. At about 10300 s the core water level is high enough to allow water flow into
the cavity, partly from the core bypass and upper plenum, partly due to cross-flow from the outer corerings. This water flow partially quenches the melt pool from the top, leading to an enhanced steam
generation and to a second primary pressure peak. This behaviour was predicted in both calculations(Figs. 16-2, 16-14 and 16-15), with a considerably slower melt cooling rate in the standard calculation
due to the higher temperatures and larger amount of melt. Once again, crust formation inhibits further
melt cooling.

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Table 16-4 and Fig. 16-22 present the core status at t = 11500 s, 30 min after start of HP injection. At
this time point the reactor core was completely refilled. Both calculations show melt accumulation in thelower, central core regions, which could only be partially quenched. The hydrogen generated during
reflood amounts to 57 kg (standard calculation) and 19 kg (alternative), respectively. The increase of thetotal mass of molten materials was about 2 ton during this phase. Fig. 16-16 depicts the average axial
mass distribution in the core for both calculations.
For completeness, Figs. 16-17 and 16-18 show the pressure and the water level at the secondary side of
the steam generators. It was only possible to control the secondary pressure according to the benchmark
specifications as long as the primary pressure remained higher than the secondary pressure, about 7200 s
after break opening. Afterwards, heat transfer changed its direction, leading to steam condensation at thesecondary side, and to a slight increase of the secondary water level.

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Time (sec)Main Events

Standard

Alternative

Break opening and loss of main feedwater

00

PORV opens

7

7

Reactor scram (primary pressure > 16.35 MPa)

10

10

PORV closes

15

15

Pressurizer empty

200

200

Pumps shutdown (primary inventory reaches 85 ton)

4669

4669

Onset of core heat-up (TUP > Tsat)

~5100

~5100

Control rod failure

7434

7434

Onset of oxidation runaway (TCLAD ~ 1850 K)

~7700

~7700

First clad perforation/burst

7781

7781

First cladding dislocation

7810

7805

First ceramic melt dislocation

8686

8524

Start of HPI

9669

9669

End of calculation (~ 30 min after start of HPI)

11500

11500

Tab. 16-3: Chronology of main events

Parameter

Start of reflooding

End of calculation(5000 sec after pump trip)(ca. 1800 sec after start ofreflooding)Standard

Alternative

Standard

Alternative

Primary mass inventory (ton)

55.5

55.8

128.7

131.8

Generated H2 mass (ton)

0.29

0.26

0.35

0.28

Mass of molten Zr (ton)

8.2

7.2

9.0

7.2

Mass of molten ZrO2 (ton)

2.0

1.2

2.0

1.3

Mass of molten UO2 (ton)

28.0

20.1

29.5

21.7

Mass of molten contr. rods (ton)

2.9

2.7

2.9

2.8

Total magma mass (ton)

41.2

31.2

43.3

33.0

Tab. 16-4: Core status at reflooding phase95

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Fig. 16-2: Pressurizer pressure

Fig. 16-3: Break mass flow rate

Fig. 16-4: Total primary liquid inventory96

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Fig. 16-5: Core collapsed liquid level - inner ring 1

Fig. 16-6: Core collapsed liquid level - outer ring 4

Fig. 16-7: Coolant temperatures in upper plenum97

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Fig. 16-8: Fuel cladding temperatures at top of core (elevation 3.0 m)

Fig. 16-9: Fuel cladding temperatures at mid-core (elevation 1.7 m)

Fig. 16-10: Fuel cladding temperatures at bottom of core (elevation 0.6 m)

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Fig. 16-11: Total hydrogen production

Fig. 16-12: Total mass of molten materials in the coreFig. 16-13: Total mass of molten Zr in the core99

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Fig. 16-14: Total mass of liquid melt in the coreFig. 16-15: Total mass of refrozen materials in the coreFig. 16-16: Average axial mass distribution at t = 11500 s

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Fig. 16-17: Secondary pressure

Fig. 16-18: Secondary level

16.5 Synthetic Views of the Core at Selected InstantsFigs. 16-19 to 16-22 show synthetic views of the core status at selected instants for both calculations,obtained with the help of the plant analyzer ATLAS. The pictures show the void distribution (VOID) inthe lower plenum (PV-LP), core rings (PV-CORE1…5), core bypass (PV-REF) and upper plenum (PVUP-M1 and -M2), the cladding temperatures (TCSU) as well as the dislocations of the fuel rods (ROD1
to ROD4) and control rods (at the left hand side of the corresponding fuel rods).

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Standard Calculation

Alternative Calculation
Fig. 16-19: TMI core at beginning of oxidation excursion (T~1850 K)

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Standard Calculation

Alternative Calculation
Fig. 16-20: TMI core at time of first ceramic melt relocation

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Standard Calculation

Alternative Calculation
Fig 16-21: TMI core at start of reflooding
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Standard Calculation

Alternative Calculation
Fig. 16-22: TMI core at the end of calculation (30 min of reflooding)105

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16.6 Additional calculations
Some additional parametric studies have been performed in the frame of this benchmark. One of themaddressed the modelling of Zr oxidation. Using the same set of input parameters as for the “alternative
calculation”, two additional calculations have been carried out with other correlations available in
ATHLET-CD: one with the correlations of Cathcart (low temperature range) and Prater/Courtright (hightemperature range), and one with the correlations of Leistikow and Prater/Courtright.
While the results concerning hydrogen production in the low temperature range (T < 1800 K) either with
Cathcart or with the Leistikow correlation are very similar, the use of Prater/Courtright in the hightemperature range leads to about 20% lower hydrogen production than with the Urbanic-Heidrickcorrelation, for a similar extent of core degradation.
As shown in Section 16.4, the “standard calculation” predicted about 40% higher mass of UO2-ZrO2
melt than the “alternative calculation”, although the temperature criterion for oxide melting was higher(2600 K instead of 2500 K). The reason was the larger extent of oxidation during the degradation phaseresulting from the higher temperature criterion for cladding failure (more Zr in metallic form wasavailable at higher elevations) in comparison to the “alternative calculation”. An additional calculationshows that, for the same cladding failure temperature criterion, an increase of the ceramic melt
temperature from 2500 K to 2600 K leads to a reduction of about 20% of the total ceramic melt mass(ca. 15% of the total melt mass). Thus, analyses of code results should not be based only on single input
parameters, but rather take into account the combined influence of them.
Although outside the scope of this benchmark, the ATHLET-CD calculations have been complemented
by an estimation of the release of fission products and absorber materials with the code module
FIPREM. A surprising outcome is that the relatively small changes on input parameters from the“standard” to the “alternative” calculations yield a considerable reduction of the release rates: about 30%
less noble gases and volatiles, 35% less actinides and up to 60% lower Cd release. This illustrates thedegree of uncertainty related to source term determination.
16.7 Final remarks
The ATHLET-CD results for the TMI-benchmark are physically reasonable and comparable to theresults of other participants, mainly with respect to the thermal-hydraulics of primary system up to thestart of ECC injection and to the predicted total mass of hydrogen produced and of molten materials.
The main observed discrepancies are related to core degradation evolution during the reflood phase.
ATHLET-CD (without the debris bed module MEWA) was the only code to predict the formation of amelt pool in the lower, central core region, which could not be completely cooled during quenchingphase. It is however supposed that the ICARE/CATHARE simulation would also show a similar
behaviour, if the HP injection was delayed enough to allow considerable melt relocation and flow
blockage. Further investigations should also address the influence of the chosen nodalization on melt
progression.
With respect to code performance, the calculations have been performed on a CRAY XD1 Linux Clusterwith AMD Opteron 2.4 GHz processors and took about 44000 sec. CPU time, with a ratio CPU to real
about 4:1 and an average time step size of 0.1 s. The calculations up to the start of ECC injection were
faster than real time, with a considerable increase of processing time during the reflood phase. Thesefigures indicate a rather good and robust code performance and show the adequacy of the couplingtechnique between the thermal-hydraulics and core degradation models.
Furthermore, one calculation has been repeated in three different platforms, with different operating
systems and FORTRAN compilers. The results were practically identical, confirming the good code
portability.
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Finally, the benchmark calculations showed that the simulation of core degradation phenomena isstrongly dependent on the choice of code input parameters like melting temperature criteria or Zr
oxidation correlation. Therefore uncertainty evaluations, taking into account the influence of the
combined variation of input parameters and modelling options, are recommended to support code
applications for deterministic plant analyses.

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17. APPENDIX : UPI CALCULATION WITH MELCOR 1.8.5

17.1 TMI-2 Plant Modelling
The same nodalisation used to simulate the first two phase of the TMI-2 accident has been used for thisbenchmark, with proper boundary conditions. Moreover, minor changes were performed with respect tothe first benchmark phase, in order to reduce stratification in the primary side and improve heat transfer
through steam generators.
The nodalization of the TMI-2 primary system for the MELCOR 1.8.5 code is presented in the Fig 1.
Both primary loops have been modelled simulating each SGs and considering a single equivalent main
pump and cold leg. The secondary system is modelled only up to the main feed and isolation valves,
while the SGs were modelled with a high level of detail. The secondary side is connected to an imposedpressure control volume and to a “spill&feed” level control, which assure the desired imposed boundary
conditions.
Particular care has also been devoted to core simulation (Fig. 2). The core schematization is constituted
by five radial rings and twelve axial levels; four thermal-hydraulic levels are used in each ring of thecore region, with three core cells axially in each thermal-hydraulic control volume. Radial and axial
flow paths in the core region allow for the prediction of 2-D flow patterns. Heat structures representativeof the control rod guide tubes and upper tie plate in the upper plenum have been added to the MELCOR
model to permit condensation heat transfer and coolant recirculation inside the vessel. The input deck
was developed using standard default MELCOR modeling parameters as long as possible, and allows
for a complete description of an eventual severe progression of an accidental transient.
The VENT valves are also explicitly modeled between the cold and hot collectors. The plant geometry,the boundary conditions and the accident scenario have been strictly defined according to alternative
TMI-2 scenario benchmark specifications provided by IRSN, as well as recommended values of code
physical parameters for the sensitivity study.

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Fig. 1: TMI-2 plant nodalization scheme.

Fig.2: TMI-2 core simulation with MELCOR.
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17.2 Updated Initial Steady-State Conditions
The TMI2 initial conditions are obtained by a steady-state code run lasting 2000 s and starting fromplant thermal-hydraulic parameter values close to the ones specified for TMI-2. During the steady-statecalculation, some regulations in the primary and secondary sides are activated to facilitate theachievement of stable conditions, as close as possible to the TMI2 plant state at t = 0 (turbin trip in
TMI2 accident).
The regulated parameters are:
 The pressurizer pressure by turning on the heaters when the pressure is lower than the nominal value; The pressurizer liquid mass, by water injection or draining, in order to obtain the precise liquid level.
The TMI2 plant initial conditions calculated by the MELCOR code are compared with TMI-2 accident
data at turbine trip in the table below. The primary circuit conditions are very well reproduced by thecode, the larger mismatch being the mass core flowrate which resulted about 3% lower with respect toreference plant parameters.

Parameter
Core power
Cold leg temperature
Hot leg temperature
Lower plenum pressure
Outlet plenum pressure
Total RCS flow rate
Core flow rate
Bypass flow rate
Pressuriser level
Feed-water/steam flow
OTSG outlet pressure
OTSG outlet temperature
Feed-water temperature
Primary mass

Units
MWth
K
K
MPa
MPa
kg/s
kg/s
kg/s
cm
kg/s
MPa
K
K
kg

Reference
value
2772
563.76591.4315.36
15.17
17602.2
16052.4
1549.8558.8
761.596.41
572.63510.93

UPI MELCOR
calculation
2770
563.6
592.0
15.35
15.20
17040
15596
1504
568.2
761.596.41
569.5
510.93230850Difference
%
0.07
0.03
0.08
0.07
0.2
3.4
2.8
2.9
2.0
IMPOSED
IMPOSED
0.5
IMPOSED-

Table 1: TMI-2 initial conditions as calculated by MELCOR for the alternative scenario.

17.3 Standard Physical Parameters of the CodeMain core degradation physical parameters used in the standard calculation and alternative physicalparameters scenario with MELCOR are summarized in the Table 2. The physical parameters used in thestandard calculation are the MELCOR default recommended values, while in the sensitivity case they
were based on the conclusions of ISP46 (PHEBUS-FPT1) and from the latest findings on coredegradation phenomena.

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Reference caseSensitivity case

Zr oxidation correlation

Urbanic-Heidrick

Urbanic-Heidrick

Cladding failure criteria

ZrO2 thickness < 10 mm
and Tclad > 2400 K

ZrO2 thickness < 200 mm
and Tclad > 2280 K

UO2-ZrO2
temperature

melting UO2 Tm = 3113 K
ZrO2 Tm = 2990 K

Debris formation criteria

Rod failure when:
Tclad> 2500 K and
Zr thickness < 100 mm
TUO2 > 3100 K

UO2 Tm = 2650 K
ZrO2 Tm = 2650 K

Rod failure when:
Tclad> 2500 K and
Zr thickness < 100 mm
TUO2 > 2650 K

Table 2: Main assumptions on main degradation models for the reference and sensitivity calculations.
For Zircalloy oxidation, the solid-state diffusion of oxygen through an oxide layer to unoxidized metal isrepresented by a parabolic rate equation, whose rate constant is evaluate using the Urbanic-Heidrickcorrelation. For very low oxidant concentrations, gaseous diffusion may limit the reaction rate, and amass transfer coefficient is calculated via a heat-mass transfer analogy from the heat transfer
correlations.
Candling, that is the downward flow of molten core materials and the subsequent refreezing of thesematerials as they transfer latent heat to cooler structures below, is addressed with a semi-mechanistic
model, based on fundamental thermal/hydraulic principles. Relocation of core materials may result in a
reduction of area and increase of flow resistance, or even total blocking of flow, within various parts ofthe core. A model is also implemented for an oxide shell to hold up molten material until the shell isbreached. Molten material is hold up within a component if the oxide thickness is greater than a criticalvalue hold, if the component temperature is less than a critical value, and if no candling from thecomponent in that cell has yet taken place.
MELCOR contains several simple models that consider the structural integrity and support of intact
components, and convert them to particulate debris when either is lost. Most are logical models rather
than structural models; no stress calculations are performed for any component other than supporting
structure. All components other than fuel rods are immediately converted to particulate debris whenever
the unoxidized metal thickness is reduced below a user defined minimum value. The thickness criterion
is also used for cladding, which is assumed to support fuel pellets, but other criteria are also considered
for fuel rods. Oxidized rods are assumed to retain their identity until the cladding reaches 2500 K, and tocollapse unconditionally if the fuel temperature reaches 3100 K (the approximate melting temperature of
UO2).
17.4 Chronology of Major Events
The chronology of major events calculated by MELCOR is presented in the table below and discussed in
the following section 5. Minimum differences are shown only for the first melt relocation that is slightly
anticipated in the sensitivity calculation. The other significant difference is related to hydrogenproduction, as described in the following.

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Event
Break opening and loss of feedwater
Reactor scram
Pressurizer is empty
Primary pump void fraction > 0.6
Primary pump shutdown (M < 85000 kg)
Onset of core uncovery
Onset of core heatup
Beginning of oxidation
First fuel rod clad burst
First melt relocation
Onset of reflooding
End of calculation

Time (s)Standard case
0
16
215
4150
4360
4635
5300
6800
7450
7510
9360
12000

Sensitivity case
0
16
215
4150
4360
4635
5950
6800
7450
7470
9360
12000

Table 3: Chronology of major events in the two calculations.

17.5 Result Analysis
The behaviour of the main thermal-hydraulic variables is quite similar in the two MELCOR
calculations. Boundary conditions on the SGs secondary side (Figures 3 to 4) are maintained during thewhole transient, by connecting each steam generator secondary side to a control volume with imposedpressure and using a “spill&feed” level control. So, changes in pressure and in liquid level due to steam
evaporation/condensation taking place during the transient are continuously corrected byinjecting/spilling saturated steam and water from/to the two controlled source. The results of this control
are particular clear in the long-term comparison of the SGs levels, where the condensation of the steam
induced by HPI actuation is compensated by spilling away the condensate water.
The initial feedwater trip and the consequent heat removal loss through the two steam generators causes
the primary pressure to rapidly reach the opening PORV set-point and after 16 s the SCRAM is actuated.
The primary pressure (Fig. 5) rapidly approaches saturation as a consequence of the break on the hot leg
and it remains almost constant at a value of 7.5 MPa (that is a slightly higher value respect to the othercodes predictions) up to the trip of the pumps which takes place after 4400 s. Almost at the same time
the break uncovers and pressure starts increasing due to reduced removed power through the steam
generators and the break. After 10 minutes the pressure decreases again reaching a minimum value ofabout 7.1 MPa. Finally, core reflood by HPI injection is started at 9400 s, 5000 s after pump trip
according to benchmark specifications. By this time, the collapsed core water level is at the core bottom
in both calculations. The reflood of the core results in sudden primary pressure increase due to the steep
increase in convective heat transfer and steam formation. Maximum pressure peak is about 128 MPa.
Formation of a void fraction is still delayed respect to other code predictions (Fig. 6), due to the slightly
higher primary system pressure that delays saturation. This is probably due to the slight underestimation of heat transfer through the two steam generators (Fig. 7). The removed heat is directly
correlated to the primary system saturation temperature up to the trip of the main pumps, when heatremoval steeply decreases. The following pressurisation and temperature increase causes a weak heattransfer increment, even though it almost goes to zero at about 8000 s.
The mass of water in the primary system (Fig. 8) decreases with an almost constant rate during liquidflow-rate through the break. The calculation of break mass flow-rate (Fig. 9) is performed using
RETRAN correlations for choked flow, that is the Moody model for saturated water and the Henry113

NEA/CSNI/R(2009)3

Fauske model for the subcooled phase. The higher primary pressure delays the reaching of saturated
conditions in the primary system and this could explain the almost constant break mass flow-rate,compared to the decreasing values of the break flow-rate calculated by the other codes. The larger mass
flow-rate through the break leads to a slightly anticipated pump trip, as the total primary mass decreases
down below 85000 kg at about 4400 s. After break uncovering the mass inventory decreases more
slowly reaching a minimum value of about 68 tons. The actuation of HPI finally causes a steep increaseof the primary water mass.
The mass flow-rate through the two primary loops (Fig. 10) is predicted to be rapidly reduced due tovoid formation inside the primary system. A void degradation factor has been also applied in order tosimulate the decrease of pumps head as a function of the void fraction.
The core starts to uncover at 4800 s (Fig. 11), while core heatup starts at 5950 s at the core top, because
the decay power is no more removed by natural circulation in the primary circuit. Core level is predicted
to reach the bottom of the active fuel after 7800 s and it remains to this level, as also indicated by the
analysis of the by-pass, down-comer and lower plenum levels, up to the actuation of the HPI.
The temperature in the cold leg (Fig. 12) remains at saturation during almost the whole transient andundergoes a steep decrease after HPI actuation. The behaviour of the temperature inside the hot leg (Fig.
13) is similar to other code predictions, even though higher values of steam superheating are evaluatedduring core heatup phase.
Also the fuel rod temperature, as evaluated by MELCOR at three different core elevations (Figures from
14 to 16), shows a good agreement with other code calculations. At the bottom of the core, a weaker fuelrod heat-up is anyway predicted during the phase where the core level goes to zero. The core quenching
behaviour is strongly affected by the input value of the hydraulic diameter of the particulate debris (inthis calculation the debris is only the refrozen zircalloy in the middle-lower parts of the core and the fuelpellets are still intact). The data presented in the following plots have been obtained using a value of 0.5m, respect to the previous value of 3.0 m that led to an increase of the core temperatures after the refloodphase, due to a reduction in the predicted heat removal. A further reduction of this value leads to theprediction of a larger core degradation before the reflooding.
Fuel rod melting starts just before the main oxidation phase. Melting of Zircalloy starts at about 7500 s
also causing a partial dissolution of the fuel. The total mass of molten metals is predicted to be about 12
tons (Fig. 17), while about 18 tons of material debris (Fig. 18) are calculated by MELCOR.
The main oxidation phase takes place (Figures 19 and 20) between 6700 s and 8500 s since thebeginning of the transient. The hydrogen production is different for the reference and alternative
calculations (Fig. 21). In the sensitivity calculation cladding failure is anticipated (larger minimum
thickness and lower maximum temperature); this causes a larger release and relocation of molten Zr
towards lower and colder elevations, reducing available metal for the oxidation process. The final
amount of hydrogen is therefore larger (335 kg) during the reference case with respect to the alternative
parameter calculation (275 kg). No hydrogen production is calculated during and after core reflooding
phase.

114

NEA/CSNI/R(2009)3

SG A Level
2.00E+00
1.80E+00
1.60E+00

Level (m)1.40E+00
1.20E+00
1.00E+00

ENEA

8.00E-01

IVS

6.00E-01

IKE

GRS
UPI

4.00E-01

SNL
IRSN

2.00E-01

IRSN-dSNU

0.00E+00
0

2000

4000

6000

8000

10000

12000

14000

10000

12000

14000

Time (s)Fig. 3: Steam generator A level

SG A Pressure
8.00E+06
7.00E+06

Pressure (Pa)6.00E+06
5.00E+06
4.00E+06

ENEA
IVS

3.00E+06

GRS
IKE
UPI

2.00E+06

SNL
IRSN

1.00E+06

IRSN-dSNU

0.00E+00
0

2000

4000

6000

8000

Time (s)Fig. 4: Steam generator A pressure

115

NEA/CSNI/R(2009)3

Pressurizer Pressure
1.80E+07

ENEA
IVS

1.60E+07

GRS
IKE

1.40E+07

UPI
SNL

Pressure (Pa)

1.20E+07

IRSN
IRSN-d

1.00E+07

SNU

8.00E+06
6.00E+06
4.00E+06
2.00E+06
0.00E+00
0

2000

4000

6000

8000

10000

12000

14000

10000

12000

14000

Time (s)Fig. 5: Pressuriser pressure

Pump A Void Fraction
1.00E+00

ENEA
IVS

9.00E-01

GRS
8.00E-01

IKE
UPI

Void fraction

7.00E-01

SNL
IRSN

6.00E-01

IRSN-dSNU

5.00E-01
4.00E-01
3.00E-01
2.00E-01
1.00E-01
0.00E+00
0

2000

4000

6000

8000

Time (s)Fig. 6: Main pump A void fraction

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NEA/CSNI/R(2009)3

SG A Power Exchanged
2.00E+08

ENEA
IVS

1.80E+08

GRS
1.60E+08

IKE
UPI

1.40E+08

SNL
IRSN

Power (W)

1.20E+08

IRSN-d

1.00E+08

SNU

8.00E+07
6.00E+07
4.00E+07
2.00E+07
0.00E+00
0

2000

4000

6000

8000

10000

12000

14000

-2.00E+07

Time (s)Fig. 7: Steam generator A removed powerPrimary Mass
2.50E+05

ENEA
IVS
GRS

2.00E+05

IKE
UPI

Mass (kg)

SNL
IRSN

1.50E+05

IRSN-dSNU

1.00E+05

5.00E+04

0.00E+00
0

2000

4000

6000

8000

Time (s)Fig. 8: Total primary water mass

117

10000

12000

14000

NEA/CSNI/R(2009)3

Break Flow Rate
1.00E+02

ENEA
IVS
GRS

8.00E+01

IKE

Mass flow rate (kg/s)

UPI
SNL
IRSN

6.00E+01

IRSN-dSNU

4.00E+01

2.00E+01

0.00E+00
0

2000

4000

6000

8000

10000

12000

14000

Time (s)Fig. 9: Break mass flowrate

Loop A Flow Rate
1.00E+04
ENEA
IVS

8.00E+03

GRS
IKE
UPI

Flow rate (kg/s)

6.00E+03

SNL
IRSN

4.00E+03

IRSN-dSNU

2.00E+03

0.00E+00
0

2000

4000

6000

8000

-2.00E+03-4.00E+03

Time (s)Fig. 10 Loop A mass flowrate

118

10000

12000

14000

NEA/CSNI/R(2009)3

Swollen Water level
5.00E+00
ENEA
IVS
GRS

4.00E+00

IKE
UPI
SNL

Water level (m)3.00E+00

IRSN
IRSN-dSNU

2.00E+00

1.00E+00

0.00E+00
0

2000

4000

6000

8000

10000

12000

14000

10000

12000

14000

-1.00E+00

Time (s)Fig. 11 Core swollen water level

Cold Leg A Temperature
1.00E+03

ENEA
IVS

9.00E+02

GRS
IKE
UPI

Temperature (K)8.00E+02

SNL
IRSN
IRSN-d

7.00E+02

SNU
6.00E+02

5.00E+02

4.00E+02

3.00E+02
0

2000

4000

6000

8000

Time (s)Fig. 12 Cold leg A temperature

119

NEA/CSNI/R(2009)3

Hot Leg A Temperature
2.50E+03

ENEA
IVS
GRS

2.10E+03

IKE
UPI

Temperature (K)

SNL
IRSN

1.70E+03

IRSN-dSNU

1.30E+03

9.00E+02

5.00E+02
0

2000

4000

6000

8000

10000

12000

14000

12000

14000

Time (s)Fig. 13: Hot leg A temperature

Temperature at the Top of Core
3.00E+03
ENEA
IVS
GRS

2.50E+03

IKE

Temperature (K)

UPI
SNL

2.00E+03

IRSN
IRSN-dSNU

1.50E+03

1.00E+03

5.00E+02

0.00E+00
0

2000

4000

6000

8000

Time (s)Fig. 14: Fuel rod temperature at the top of the core

120

10000

NEA/CSNI/R(2009)3

Temperature at mid-Core
4.50E+03
ENEA
IVS

4.00E+03

GRS
IKE

3.50E+03

Temperature (K)

UPI
SNL

3.00E+03

IRSN
IRSN-d

2.50E+03

SNU
2.00E+03
1.50E+03
1.00E+03
5.00E+02
0.00E+00
0

2000

4000

6000

8000

10000

12000

14000

12000

14000

Time (s)Fig. 15: Fuel rod temperature at the middle of the core

Temperature at the Bottom of Core
3.50E+03
ENEA
IVS
3.00E+03

GRS
IKE
UPI

Temperature (K)2.50E+03

SNL
IRSN

2.00E+03

IRSN-dSNU

1.50E+03

1.00E+03

5.00E+02

0.00E+00
0

2000

4000

6000

8000

Time (s)Fig. 16 Fuel rod temperature at the bottom of the core

121

10000

NEA/CSNI/R(2009)3

Total Mass of Molten Metal
3.00E+04
ENEA
IVS
GRS

2.50E+04

IKE
UPI
SNL

2.00E+04

Mass (kg)IRSN
IRSN-dSNU

1.50E+04

1.00E+04

5.00E+03

0.00E+00
0

2000

4000

6000

8000

10000

12000

14000

10000

12000

14000

Time (s)Fig. 17: total mass of molten metals

Mass of Debris
6.00E+04
ENEA
IVS
GRS

5.00E+04

IKE
UPI
SNL

4.00E+04

Mass (kg)IRSN
IRSN-dSNU

3.00E+04

2.00E+04

1.00E+04

0.00E+00
0

2000

4000

6000

8000

Time (s)Fig. 18: total mass of debris

122

NEA/CSNI/R(2009)3

Fraction of non Oxidized Zr
1.00E+00
9.00E-01
8.00E-01

Mass Fraction

7.00E-01
6.00E-01
5.00E-01

ENEA
IVS

4.00E-01

GRS

3.00E-01

IKE

2.00E-01

SNL

UPI
IRSN

1.00E-01

IRSN-dSNU

0.00E+00
0

2000

4000

6000

8000

10000

12000

14000

12000

14000

Time (s)Fig. 19 Fraction of non oxidised Zr

Cumulated Hydrogen Production
4.50E+02
ENEA
IVS

4.00E+02

GRS
IKE

3.50E+02

UPI
SNL

H2 Mass (kg)3.00E+02

IRSN
2.50E+02

IRSN-dSNU

2.00E+02
1.50E+02
1.00E+02
5.00E+01
0.00E+00
0

2000

4000

6000

8000

Time (s)Fig. 20 Cumulated mass of Hydrogen

123

10000

NEA/CSNI/R(2009)3

400
350
Reference calculation

Mass [kg]300
Sensitivity calculation

250
200
150
100
50
0
-2000

0

2000

4000

6000

time [sec]Fig. 21 Hydrogen production comparison

124

8000

10000

12000

14000

NEA/CSNI/R(2009)3

Synthetic Views of the Core at Selected Instants
Standard case
Cladding temperature maps
3.50

3.00
1500
1400

2.50

1300
1200
2.00

1100
1000
900

1.50

800
700
1.00
600
500
0.50

-1.50

-1.00

-0.50

0.00

0.50

1.00

1.50

Fig. 22: Cladding temperature (t = 6600. s)
3.50

3.00
2400
2300
2.50

2200
2100
2000

2.00
1900
1800
1700

1.50

1600
1500
1400

1.00

1300
1200
0.50

-1.50

-1.00

-0.50

0.00

0.50

Fig. 23: Cladding temperature (t = 8200. s)

125

1.00

1.50

NEA/CSNI/R(2009)3
3.50

3.00
1600
1500
1400

2.50

1300
1200
1100
1000

2.00

900
800
700
1.50

600
500
400
300

1.00

200
100
0

0.50

-1.50

-1.00

-0.50

0.00

0.50

1.00

1.50

Fig. 24: Cladding temperature (t = 10000. s)
3.50

3.00
1600
1500
1400

2.50

1300
1200
1100
1000

2.00

900
800
700
1.50

600
500
400
300

1.00

200
100
0

0.50

-1.50

-1.00

-0.50

0.00

Fig. 25: Cladding temperature (t = 12000. s)

126

0.50

1.00

1.50

NEA/CSNI/R(2009)3
Linear mass of oxidised Zircalloy
3.50

3.00
0.00
0.00
2.50

0.00
0.00
0.00
0.00

2.00

0.00
0.00
0.00
1.50

0.00
0.00
0.00
0.00

1.00

0.00
-0.00
0.50

-1.50

-1.00

-0.50

0.00

0.50

1.00

1.50

Fig. 26: ZrO2 linar mass (t = 6600. s)

3.50

3.00
0.12
0.11
2.50

0.10
0.09
0.08

2.00

0.07
0.06
0.05
0.04

1.50

0.03
0.02
0.01

1.00

0.00
-0.01
0.50

-1.50

-1.00

-0.50

0.00

0.50

Fig. 27: ZrO2 linar mass (t = 8200. s)

127

1.00

1.50

NEA/CSNI/R(2009)3
3.50

3.00
0.24
0.22
2.50

0.20
0.18
0.16

2.00
0.14
0.12
0.10

1.50

0.08
0.06
0.04

1.00

0.02
0.00
0.50

-1.50

-1.00

-0.50

0.00

0.50

1.00

1.50

Fig. 28: ZrO2 linar mass (t = 10000. s)
3.50

3.00
0.24
0.22
2.50

0.20
0.18
0.16

2.00
0.14
0.12
0.10

1.50

0.08
0.06
0.04

1.00

0.02
0.00
0.50

-1.50

-1.00

-0.50

0.00

Fig. 29: ZrO2 linar mass (t = 12000. s)

128

0.50

1.00

1.50

NEA/CSNI/R(2009)3
Molten Zircalloy mass
3.50

3.00
220.00200.002.50

180.00160.00140.002.00
120.00100.00
80.00

1.50

60.00
40.00
20.00

1.00

0.00
-20.00
0.50

-1.50

-1.00

-0.50

0.00

0.50

1.00

1.50

Fig. 30: Molten Zircalloy mass (t = 7900. s)
3.50

3.00
800.00750.00700.00650.00600.00550.00500.00450.00400.00350.00300.00250.00200.00150.00100.0050.00
0.00-50.00

2.50

2.00

1.50

1.00

0.50

-1.50

-1.00

-0.50

0.00

0.50

Fig. 31: Molten Zircalloy mass (t = 8200. s)

129

1.00

1.50

NEA/CSNI/R(2009)3
3.50

3.00
1000.00
900.002.50

800.00700.00600.002.00

500.00400.001.50

300.00200.00100.001.00
0.00-100.00
0.50

-1.50

-1.00

-0.50

0.00

Fig. 32: Molten Zircalloy mass (t = 12000. s)

130

0.50

1.00

1.50

NEA/CSNI/R(2009)3

Sensitivity case
Cladding temperature maps
3.50

3.00
1550.00
1450.00

2.50

1350.00
1250.00
2.00

1150.00
1050.00
950.001.50

850.00750.001.00
650.00550.00
0.50

-1.50

-1.00

-0.50

0.00

0.50

1.00

1.50

Fig. 33: Cladding temperature (t = 6600. s)
3.50

3.00
2400.00
2300.00
2.50

2200.00
2100.00
2000.00
1900.00

2.00

1800.00
1700.00
1600.00
1.50

1500.00
1400.00
1300.00
1200.00

1.00

1100.00
1000.00
0.50

-1.50

-1.00

-0.50

0.00

0.50

Fig. 34: Cladding temperature (t = 8200. s)

131

1.00

1.50

NEA/CSNI/R(2009)3
3.50

3.00
1600.00
1500.00
1400.00

2.50

1300.00
1200.00
1100.00
1000.00

2.00

900.00800.00700.001.50

600.00500.00400.00300.001.00

200.00100.00
0.00

0.50

-1.50

-1.00

-0.50

0.00

0.50

1.00

1.50

Fig. 35: Cladding temperature (t = 10000. s)
3.50

3.00
1600.00
1500.00
1400.00
2.50

1300.00
1200.00
1100.00
1000.00

2.00

900.00800.00700.00600.001.50

500.00400.00300.001.00

200.00100.00
0.00

0.50

-1.50

-1.00

-0.50

0.00

Fig. 36: Cladding temperature (t = 12000. s)

132

0.50

1.00

1.50

NEA/CSNI/R(2009)3
Linear mass of oxidised Zircalloy
3.50

3.00
0.00
0.00
2.50

0.00
0.00
0.00
0.00

2.00

0.00
0.00
0.00
1.50

0.00
0.00
0.00
0.00

1.00

0.00
-0.00
0.50

-1.50

-1.00

-0.50

0.00

0.50

1.00

1.50

Fig. 37: ZrO2 linar mass (t = 6600. s)

3.50

3.00
0.12
0.11
2.50

0.10
0.09
0.08

2.00

0.07
0.06
0.05
0.04

1.50

0.03
0.02
0.01

1.00

0.00
-0.01
0.50

-1.50

-1.00

-0.50

0.00

0.50

Fig. 38: ZrO2 linar mass (t = 8200. s)

133

1.00

1.50

NEA/CSNI/R(2009)3
3.50

3.00
0.16
0.15
0.14

2.50

0.13
0.12
0.11
0.10

2.00

0.09
0.08
0.07
1.50

0.06
0.05
0.04
0.03

1.00

0.02
0.01
0.00

0.50

-1.50

-1.00

-0.50

0.00

0.50

1.00

1.50

Fig. 39: ZrO2 linar mass (t = 10000. s)
3.50

3.00
0.16
0.15
0.14

2.50

0.13
0.12
0.11
0.10

2.00

0.09
0.08
0.07
1.50

0.06
0.05
0.04
0.03

1.00

0.02
0.01
0.00

0.50

-1.50

-1.00

-0.50

0.00

Fig. 40: ZrO2 linar mass (t = 12000. s)

134

0.50

1.00

1.50

NEA/CSNI/R(2009)3
Molten Zircalloy mass
3.50

3.00
800.00750.00700.00650.00600.00550.00500.00450.00400.00350.00300.00250.00200.00150.00100.0050.00
0.00-50.00

2.50

2.00

1.50

1.00

0.50

-1.50

-1.00

-0.50

0.00

0.50

1.00

1.50

Fig. 41: Molten Zircalloy mass (t = 7900. s)
3.50

3.00
1500.00
1400.00
1300.00

2.50

1200.00
1100.00
1000.00
900.002.00

800.00700.00600.001.50

500.00400.00300.00200.001.00

100.000.00-100.00

0.50

-1.50

-1.00

-0.50

0.00

0.50

Fig. 42: Molten Zircalloy mass (t = 8200. s)

135

1.00

1.50

NEA/CSNI/R(2009)3
3.50

3.00
1600.00
1500.00
1400.00
1300.00
1200.00
1100.00
1000.00
900.00800.00700.00600.00500.00400.00300.00200.00100.000.00-100.00

2.50

2.00

1.50

1.00

0.50

-1.50

-1.00

-0.50

0.00

Fig. 43: Molten Zircalloy mass (t = 12000. s)

136

0.50

1.00

1.50

NEA/CSNI/R(2009)3

18. APPENDIX : IVS CALCULATION WITH ASTEC V1.3
18.1 TMI-2 Plant Modelling
The two previously described ASTEC modules were used in the coupled mode. The front-end thermalhydraulic part of the transient was performed by CESAR module and the DIVA module was activatedlater, when the core heat-up takes place. The TMI-2 plant nodalization scheme is shortly described
below.
18.1.1 Reactor vessel:








PRIMARY data (core, bypass, lower plenum and downcomer) automatically created fromVESSEL_D data by ASTEC;VENT valve was modelled between the reactor downcomer and upper plenum;
Double junctions to reactor vessel upper head with slightly different elevations were used inorder to prevent the flow stagnation in this volume;
All junctions were of type 0 (standard two-phase flow junction);
Volumes of type 0 (standard homogeneous volumes) were used preferably. Volumes of type 1(volumes with swollen level) were used in upper part of the reactor vessel. Volumes of type –1(axial elements) were used to model fuel assemblies;
All characteristic features of vessel (e.g. vessel vent valves) modelled;
Default value of tau_long parameter (time return to thermal equilibrium) for volumes of type 1
were used.

volume type 0volume type 1
volume type -1
volume type 0volume type1
1!1

Fig. 1. Nodalization of reactor vessel
137

NEA/CSNI/R(2009)3
18.1.2 Circulation loops:







Two identical loops were modelled (loop with pressurizer and break is shown);
Two parallel cold legs were modelled in each loop;
Volumes of type 0 (standard volumes) were used preferably. Volumes of type 1 (volumes withswollen level) were used only in pressurizer and SG cold collector, volumes of type –1 (axial
elements) used only in SG tubes;
Junctions of type 0 (standard two-phase flow junction) preferred, some junctions of type 1 usedin order to enable proper phase separation;
SG tubing was split into 2 parallel bundles (90% - inner channel and 10% - outer channel cooled
by EFW spray) in accordance with nodalisation of secondary side;
Reactor coolant pumps were modelled using shaft equation, respecting the pump inertia,
friction, motor and hydraulic torque.

18.1.3 Pressurizer and surge line:Pressurizer was modelled using 3 volumes of type 1 (volumes with swollen level);
RRZ surge line was modelled by 2 volumes of type 0 (standard volumes) connected by junctionin their bottom part (loop seal shape).

PORV

From
reactorTo reactor
Make-up(loop 2B)

To reactor
BreakFig. 2. Nodalization of loop with pressurizer
138

NEA/CSNI/R(2009)3
18.1.4 Steam generator – secondary side:







Two identical SGs modelled;
SG tubes bundle was split into 2 parallel nodes (90% - inner channel and 10% - outer channel);Volumes of type 0 (standard volumes) were used preferably. Volumes of type 1 (volumes withswollen level) were used only in parallel bundle channel (10% of SG tubes) for EFW injection;
High value of tau_long parameter was used for volumes of type 1 (20000 s);
All junctions are of type 0;
EFW injected into upper part of outer channel.

Emergency
feedwater

Turbine (BC)Feedwater

Outer
Inner
channel channel

Fig. 3. Nodalization of SG secondary side

18.1.5 Core, lower plenum and downcomer model:









Meshing: 6 parallel channels (core rings) + bypass, downcomer and lower plenum;
Fuel assemblies split into 24 axial meshes;
Core is assumed only as a region containing fuel;
The following solid macro components were defined:
Fuel and cladding;Control rods and gain tubes;
Core baffle and shield;
Reactor vessel with thermal insulation;

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

Lower plenum;
Model of the molten pool formation used (=> macro components DEBRIS and MAGMA were
defined).
. . .
. . .
..

Fig. 4. Nodalisation of reactor core, lower plenum and downcomer

18.2 Initial Steady-State Conditions
The initial “steady-state” conditions were obtained running the calculation (with t < 0) on nominal TMI2 plant conditions using the simplified form of most important plant controllers. This includes simple
FW controllers, cycling of PRZ heaters and turbine modelled as pressure boundary condition. The
obtained stabilised parameters and their comparison with nominal TMI-2 values at the start of transient(t = 0) is given in the table below.
TABLE I: Comparison of ASTEC steady-state conditions with TMI-2 data
Parameter

Unit

ASTEC V1.3

TMI-2

Reactor Power

W

2700

2700

Primary Pressure

MPa

15.32

15.2

Temperature Hot Leg A

K

592.9

592

Temperature Hot Leg B

K

592.9

592

Temperature Cold Leg A

K

565.1

548-561

Temperature Cold Leg B

K

565.1

565

Mass Flow Rate – Loop A

kg/s

8501

8280

Mass Flow Rate – Loop B

kg/s

8498

8560

Pressurizer Level

M

5.69

5.77

Total Primary Masston

221515-Pressure SG A

MPa

6.247*

7.31

Pressure SG B

MPa

6.246*

7.24

Steam Temperature SG A

K

578.8

586

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Steam Temperature SG B

K

578.0

586

Collapsed Level SG A

M

3.81

-

Collapsed Level SG B

M

3.65

-

Liquid Mass SG A

Kg

15303

-

Liquid Mass SG B

Kg

14409

-

Feedwater Flow SG A

Kg/s

746.5

723

Feedwater Flow SG B

Kg/s

737.2

717

SG Feedwater Temperature

K

513

-

* according to benchmark scenario, pressure boundary condition 6.2 MPa was used on the secondaryside .
18.3 Physical Parameters Used in the Analysis
Main core degradation physical parameters used in the base case calculation with ASTEC are given in
the table below. Only the base case calculation with Urbanic-Heidric Zr oxidation kinetic model was
performed for final benchmark scenario. The sensitivity analysis with FZK correlation for oxidation
kinetics was performed only in previous stage of benchmark exercise.
TABLE II: Main physical parameters used in the base case and sensitivity analysis
Parameter

ASTEC V1.3 Rev. 0

Zr Oxidation correlation

Urbanic-Heidric
T > 2260K and e(ZrO2) < 160 μm
T > 2280K and e(ZrO2) < 200 μm
T > 2340K and e(ZrO2) < 220 μm
T > 2380K and e(ZrO2) < 240 μm
T > 2450K and e(ZrO2) < 300 μm

Cladding failure criteria

1

UO2-ZrO2
temperature

melting default values from material database of the code

Debris formation criteria

code default criteria, DEBRIS macro component
and reflooding model in ASTEC are still underdevelopment

Debris diameter

5 % of 2 mm diameter
15 % of 4 mm diameter
60 % of 6 mm diameter
15 % of 8 mm diameter
5 % of 10 mm diameter

Debris porosity 2

0.7 / 0.6

Remarks:
1. This is not cladding failure criterion (cladding failure is function of temperature, internalpressure, ...) but criteria for change of the status from COMPACT to DISLOCAT(i.e. relocation).
2. The first value is porosity used for collapsing of free debris, the second parameter is porosity forcollapsing of debris bed.

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18.4 The results of the analysis
The chronology of major events calculated by ASTEC is presented in the Table III below and illustrated
on the attached figures.
After reactor scram the primary pressure starts to decrease. The primary pressure first reached saturationpressure and later further decreases and approaches to the secondary pressure. The core is cooled via
two-phase forced circulation in primary system. The core decay power and RCPs heat are removed by
the steam generators. The total heat removed by both SGs is clearly higher than the decay heat. The
mass-flow-rate through the loops is monotonously decreasing due to increased void fraction in primarycoolant.
The void primary coolant mass drops below 85 t at t = 4789 s. At this time all RCPs were tripped. Since
this time, phase separation occurs in primary system and the core decay heat is removed in natural
circulation boiler-condenser mode through loops A and B (in loop A the steam is also escaping throughthe break into containment). However, the total heat transfer from primary to secondary side is about
50% of the decay heat and later decreases due to decrease of primary pressure and even heat flux
reversal occurs in both SGs when the primary pressure drops below the secondary pressure.
TABLE III: Chronology of main events during the accident
Event

Time
[s]

Break opening and loss of feedwater0

SG B level drops below 1 m

12.6

PRZ spray valve opening

13.5

SG A level drops below 1 m

14.0

PRZ PORV opening

15.2

Reactor scram

19.4

PRZ spray valve closing

23.3

PRZ PORV closing

23.3

SGs are nearly dry (water mass in SG B/ SG A < 100 kg)

26.6/29.9

Pressurizer is nearly empty (water level < 0.5 m)

150

Primary coolant mass (water + steam) < 85 t=> all RCPs are tripped

4314

Onset of core heatup5149

Max. cladding temperature > 550°C (DIVA start)

5759

First fuel rod clad burst

6306

First debris and magma in the core

6402

Melting pool formation in the core

7925

Start of HP injection

9789

End of calculation (code crash in DIVA module)

10478

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NEA/CSNI/R(2009)3
The break opening and FW trip was considered as initiating event at time t = 0. The loss of FW results in
rapid decrease of secondary side SG water level followed by increase of primary pressure (neither the
break flow nor PRZ spray and PORV opening were sufficient to prevent further primary pressure
increase). The pressure rise activated reactor scram (P > 16.34 MPa) at t = 19.4 s.
The initial coolant leak through the break exceeds 90 kg/s. After the reactor scram dropped below 40
kg/s and later further slowly decreases below 30 kg/s. Up to the RCPs trip the leak is prevailingly in thewater phase. After the RCPs are tripped, phase separation occurs in primary system, practically puresteam is escaping from primary system and mass-flow-rate drops below 10 kg/s and further decreases in
the later course of the accident. Shortly after RCP trip the leaking steam is saturated. Later, when thecore heatup takes place, the steam is overheated.
When the maximum core temperature exceeds 550 °C at t = 5759 s, the DIVA module of ASTEC was
activated to compute in-vessel core degradation and melt relocation. Since this time the oxidation of thecladdings is calculated by DIVA module. First fuel rod clad burst due to overpressure is calculated by
ASTEC at 6306 s. Since the t ~ 6500 s an intensive oxidation take place and the maximum chemical
power for certain period even exceeded the decay heat. Since t ~ 7500 s the oxidation rate slows down
significantly and the chemical power is only fraction of decay heat.
At t = 9789 high-pressure injection into horizontal part of all cold legs starts. The primary pressure, leakflow-rate and heat transfer from primary to secondary side increase due to core quenching and intensive
steam production here. However, this injection was followed only with minor hydrogen production.
Since this time primary coolant inventory and core water level start to rise. At t = 10478 s the calculation
was terminated due to error in DIVA module. The total mass of hydrogen produced till the end of
calculation is 383.5 kg.

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NEA/CSNI/R(2009)3
18.

TMI-2 benchmark, final

ASTEC V1.3 Rev. 0

primary
secondary A

16.

Saturation in primarycircuit reached

Pressure [MPa]14.

secondary B

RCP-s stop

12.

Start of DIVA module

10.

Start of HPI

8.
6.

Secondary circuit iscooled by primary4.
2.

Remark: Primary pressure is considered as pressure in the upper part of hot
leg of loop A in accordance with pressure measurement in TMI -2 plant.

0.
0

1000

2000

3000

4000

5000

6000

7000

8000

9000

10000

Time [s]Fig 5. Pressures.
18.

TMI-2 benchmark, final

ASTEC V1.3 Rev. 0

Reactor scram

16.

primary
secondary A

PORV closing
PRZ spray valve closing

14.

secondary B

Saturation in primarycircuit reached

Pressure [MPa]

12.
10.

PORV opening
PRZ spray valve opening

8.
6.
4.

Start of EFW injection

SG-s near empty

2.

EFW mass-flow reachmaximum value (30 kg/s)

Remark: Primary pressure is considered as pressure in the upper part of hot
leg of loop A in accordance with pressure measurement in TMI -2 plant.

0.
0

25

50

75

100

125

150

Time [s]Fig 6. Pressures (fine scale).

144

175

200

225

250

NEA/CSNI/R(2009)3
4500. TMI-2 benchmark, final

ASTEC V1.3 Rev. 0

CL 1A

4000.

CL 2A
CL 1B

3500.

CL 2B

3000.

RCP-s stop

Flow [kg/s]

2500.
2000.

Start of HPI, injection
modelled to the downcomer

1500.
1000.
500.
0.
-

500.
0

1000

2000

3000

4000

5000

6000

7000

8000

9000

10000

Time [s]Fig 7. Cold leg mass-flows.
100.

TMI-2 benchmark, final

ASTEC V1.3 Rev. 0

90.

water
steam

80.

total

Flow [kg/s]

70.

RCP-s stop

60.
50.

Start of core heatupStart of HPI

40.

Start of intensivehydrogen production

30.
20.
10.
0.
0

1000

2000

3000

4000

5000

6000

Time [s]Fig 8. Break mass-flow.

145

7000

8000

9000

10000

NEA/CSNI/R(2009)3
250.

TMI-2 benchmark, final

ASTEC V1.3 Rev. 0

water
steam

Saturation in primarycircuit reached
200.

total

RCP-s stop

Start of HPI

Mass [t]

150.

100.

50.

0.
0

1000

2000

3000

4000

5000

6000

7000

8000

9000

10000

Time [s]Fig 9. Primary masses.
1800. TMI-2 benchmark, final

ASTEC V1.3 Rev. 0

hot leg A
hot leg B

1600.

Start of HPI

Temperature [°C]

1400.
1200.

Start of core heatup

1000.
800.
600.
400.
200.
0

1000

2000

3000

4000

5000

6000

7000

Time [s]Fig 10. Reactor outlet temperatures.

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NEA/CSNI/R(2009)3
80.

TMI-2 benchmark, final

ASTEC V1.3 Rev. 0

Max. EFW mass-flow (30 kg/s).
All water from EFW is evaporated

70.

SG A
SG B

SG water level formation
60.

Fluctuation of the SG-s level
due to controller behavior.

Flow [kg/s]

50.

Start of core heat-up

40.
30.

Start of DIVA module.
Beginning of oxidationand cumulation ofhydrogen in SG-s

RCP-s stop

Start of HPI

20.
10.
0.
0

1000

2000

3000

4000

5000

6000

7000

8000

9000

10000

Time [s]Fig 11. SG-s outlet.
4.

TMI-2 benchmark, final

ASTEC V1.3 Rev. 0

3.5

SG A
SG B

Fluctuation of the
SG-s level due tocontroller behavior

3.

Level [m]2.5

RCP-s stop
2.

Secondary circuit is cooled byprimary. Increasing of level
due to steam condensation.

1.5
1.
0.5

Start of HPI

0.
0

1000

2000

3000

4000

5000

6000

Time [s]Fig 12. SG-s levels.

147

7000

8000

9000

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NEA/CSNI/R(2009)3
7.

Fig. 8. Pressurizer level.

TMI-2 benchmark, final

ASTEC V1.3 Rev. 0

Fig. 8. Pressurizer level.
7.

6.

TMI-2 benchmark, final

ASTEC V1.3 Rev. 0

6.

5.

Reactor scram

4.

Level [m]Level [m]

5.

3.

4.

3.

2.

2.

1.

Time [s] (fine scale)

0.

1.

0

50

100

150

200

250

300

Time [s]

0.
0

1000

2000

3000

4000

5000

6000

7000

8000

9000

10000

Time [s]Fig 13. Pressurizer level.
1.

TMI-2 benchmark, final

ASTEC V1.3 Rev. 0

0.9
0.8
0.7

RCP-s stop

Void [-]

0.6

Start of HPI, injection
modelled to the downcomer

0.5
0.4
0.3
0.2

RCP 1A
RCP 1B

0.1
0.
0

1000

2000

3000

4000

5000

6000

Time [s]Fig 14. RCP void fractions.

148

7000

8000

9000

10000

NEA/CSNI/R(2009)3
TMI-2 benchmark, final

4.5

ASTEC V1.3 Rev. 0

CESAR
DIVA

4.

Start of core heatup
3.5

Level [m]3.
2.5

Start of HPI

2.
1.5
1.

RCP-s stop

0.5
0.
0

1000

2000

3000

4000

5000

6000

7000

8000

9000

10000

Time [s]Fig 15. Core collapsed level.
4000. TMI-2 benchmark, final

ASTEC V1.3 Rev. 0

Start of HPI
cladding (CESAR)
core (DIVA)

3500.

Temperature [°C]3000.

Very intensivehydrogen production

2500.
2000.

Start of core heatup
1500.

Start of intensivehydrogen production

1000.
500.
0.
0

1000

2000

3000

4000

5000

6000

7000

Time [s]Fig 16. Maximum temperatures.

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400.

Fig. 12. Hydrogen production.

TMI-2 benchmark, final

ASTEC V1.3 Rev. 0

350.
300.

Mass [kg]

250.
200.
150.
100.
50.
0.
5500

6000

6500

7000

7500

8000

8500

9000

9500

10000

10500

Time [s]Fig.Fig13.17.
HydrogenHydrogenproduction
production.rate.
0.6

TMI-2 benchmark, final

ASTEC V1.3 Rev. 0

0.5

Rate [kg/s]

0.4

0.3

Start of HPI
0.2

0.1

0.
5500

6000

6500

7000

7500

8000

8500

Time [s]Fig 18. Hydrogen production rate.

150

9000

9500

10000

10500

NEA/CSNI/R(2009)3
TMI-2 benchmark, final

70.

ASTEC V1.3 Rev. 0

Start of HPI
debris
magma

60.

50.

Mass [t]

40.

30.

20.

10.

Start of HPI
0.
5500

6000

6500

7000

7500

8000

8500

9000

9500

10000

10500

Time [s]Fig 19. Masses in the core.
200.

TMI-2 benchmark, final

ASTEC V1.3 Rev. 0

decay heat (total)

180.

SG water level formation

SGA
SGB

160.
140.

SG total
Zr oxidation

Start of core heat-up

Power [MW]120.

Very intensivehydrogen production

RCP-s stop

100.

Start of HPI

80.
60.
40.
20.
0.

Secondary circuit is cooled by primary
-

20.
0

1000

2000

3000

4000

5000

6000

Time [s]Fig 20. Powers.

151

7000

8000

9000

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5769 s6564 s8164 s9769 s10124 s10465 sFig. 21. Core temperature field [K].

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5769 s6564 s8164 s9769 s10124 s10465 sFig. 22. Debris saturation in the core [-].

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5769 s6564 s8164 s9769 s10124 s10465 sFig. 23. Magma saturation in the core [-].

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5769 s6564 s8164 s9769 s10124 s10465 sFig. 24. Gas temperature in the core [K].

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NEA/CSNI/R(2009)3

5769 s6564 s8164 s9769 s10124 s10465 sFig. 25. Water level in the core [m].

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NEA/CSNI/R(2009)3

5769 s6564 s8164 s9769 s10124 s10465 sFig. 26. Gas porositry in the core [-].

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19. APPENDIX : NRC-SNL CALCULATION WITH MELCOR 1.8.6

19.1 TMI-2 Plant Modeling
MELCOR version 1.8.6 (YS) was used in this calculation. The TMI-2 plant nodalization scheme used inthis calculation is depicted in the figure below. It models the reactor pressure volumes and internals, theprimary loops with pressurizer, and main emergency and control systems. The model uses a total of 143
control volumes (see table), 213 flow paths, and 134 heat structures. Of the 143 control volumes, 123
are in the primary (88 in the RPV, six in the pressurizer, and 29 in the piping, pumps, and steam
generator tubes) and 16 control volumes in the secondary. The remaining 4 control volumes serve asboundary conditions. Most of the heat structures, 64, are in contact with the RPV volumes, 38 are incontact with the secondary.
In the primary loop, the coolant flow and heat transfer through the A and B loops are represented as
distinct flow circuits. However, the division of primary coolant flow at the exit of each SG into twoparallel cold leg piping systems is neglected, and the pair of RCPs in each loop are represented as asingle composite pump. The pressurizer and surge line are modeled with six control volumes and six
heat structures.
Modeling of the secondary systems is limited to the secondary side of steam generators, the steam lines
with isolation valves, and main feedwater and auxiliary feedwater injections.
MELCOR TMI-2 plant nodalization within the active core, fuel behavior is modeled by subdividing thecore into five (plus the by-pass) radial rings and 12 axial levels. The twelve axial levels are equal inheight, representing roughly 1-foot increments of the core. Two additional axial levels are defined torepresent the volume below the active core, where damaged fuel can relocate. The core baffle, the barrel
and the thermal shield at the core periphery are also represented. Core debris can relocate outside thebaffle and down through the formers. The control rod guide tubes and drive assemblies in the upper
plenum region and a flow distributor and in-core instrumentation tubes in the lower head region are alsomodeled.

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Control Volumes
Flow Paths
Heat Structures

Total
143
213
134

Primary
123
154
114

RPV
88
121
64

Pressurizer
6
8
6

etc.
29
33
44

Secondary
16
16
38

19.2 Geometry
The geometry assumed for the MELCOR calculation is summarized in the table provided below. The
input deck used in this calculation was derived from the standard TMI-2 deck used for MELCOR
assessments where boundary/initial conditions were modified to conform to the problem specifications.
A notable variance from the problem specification is in the axial nodalization used for the core. This
calculation used 14 meshes over the total core length compared to 20 in the problem specification. Also,
the total primary volume is somewhat larger than the specification. This reported difference was not
resolved.

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NEA/CSNI/R(2009)3

ProblemSpecification

MELCOR

Primary volume

337.6

351

Primary system volumes out of pressurizer (m3)Pressurizer volume (m3)
RPV volume (m3)Secondary side free volume of 1 SG

294.6
43
116.5

308
43
127
65

12302.5
15530

12303.
15531

177
3.66
4
square0.01443208
0.01090.00470.000673

177
3.66
4
Square0.01443208
0.01090.00470.000673

93650
23050

95382.96
23026

1011
2750

1010
2818.4064

Core barrel inside diameter [m]Core barrel outside diameter [m]

3.584
3.683

3.58
3.683

Radial rings
Axial meshes (over total core length)

5 or 620

514.1778, .1335,
12x.3048

Free volume

Heat transfer with the secondary side
SG tubes surface (secondary side) / SG
Number of tubes / SGInitial Core Geometry
Core Characteristics
Number of fuel bundles of type 15x15
Active cor length
Total core length
Type of fuel lattice
Rod pitch
Number of fuel rods/assembly
External diameter of fuel rod
Fuel pellet radius
Fuel rod cladding thickness
Initial core material inventory
UO2 mass
Zircaloy mass
(H2 total mass if converted from total zircaloy(mZr/22.8))
AIC mass (Ag+In+Cd) [kg]Core Meshing

Mesh size (m)

0.2

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19.2.1 Vent valves
These valves, between the upper plenum and the top of the downcomer, are designed to avoid a directloss of water by the hot leg. They were modeled, as they were shown to have an influence on thetransient, following the problem specifications, i.e.,

If Δp < 414 Pa, the valves are closed.
If Δp > 1724 Pa, the valves are fully open, which corresponds to a total section of 0.794 m2.
If 414 < Δp < 1724, the valves are considered half-open, with a section increasing linearly withΔp.

19.3 Updated Initial Steady-State Conditions
The TMI2 plant state at transient initiation is obtained by a steady-state precursor calculationlasting 300
s and starting from plant thermal-hydraulic parameter values close to the ones specified for TMI-2.
Following the initialization, the break is initiated at time t=0. The TMI2 plant initial conditions
calculated by MELCOR 1.8.6 are compared with TMI2 accident data at turbine trip in the table below.
A noticeable point of discrepancy between the MELCOR calculation and other contributed calculations
is the total primary mass, which is ~2% higher for the MELCOR calculation. This is in part explained by
the available free volume, which is larger than the problem specification by about 4%, as discussed
previously.
Parameter

Unit

MELCOR 1.8.6

TMI-2

Reactor Power

W

2700

2700

Primary Pressure

MPa

15.0

15.2

Temperature Hot Leg A

K

592

592

Temperature Hot Leg B

K

592

592

Temperature Cold Leg A

K

564

548-561

Temperature Cold Leg B

K

564

565

Mass Flow Rate – Loop A

kg/s

8737

8280

Mass Flow Rate – Loop B

kg/s

8828

8560

Pressurizer Level

M

5.77

5.77

Total Primary Masston

231273-Pressure SG A

MPa

6.42

7.31

Pressure SG B

MPa

6.27

7.24

Steam Temperature SG A

K

568

586

Steam Temperature SG B

K

567

586

Collapsed Level SG A

M

5.6

-

Collapsed Level SG B

M

5.1

-

Liquid Mass SG A

Kg

18700

-

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Liquid Mass SG B

Kg

16900

-

Feedwater Flow SG A

Kg/s

750

723

Feedwater Flow SG B

Kg/s

770

717

SG Feedwater Temperature

K

513

-

19.4 Standard Physical Parameters of the Code
Oxidation of Zircaloy and stainless steel (released from control rods) is modeled with parabolic rateequation. The Urbanic-Heidrick correlation is used in MELCOR to evaluate rate constants for Zircaloy
in steam. New to MELCOR 1.8.6, oxidation of unquenched Zircaloy and steel surfaces below the poolsurface is also calculated. The necessary steam is assumed to come from the gas film between the hot
surface and the pool and is therefore limited only by the pool mass.
Debris bed and molten pool models are modeled in version 1.8.6. Debris beds formation occurs when
intact material fails. Cladding fails and fuel rods collapse when clad temperature exceeds 2300 K andunoxidized clad thickness exceed .3mm or unconditionally when clad thickness exceeds 2500K. For
control rods, failure occurs when either the unoxidized thickness exceeds a criterion or the steel melt
temperature is reached.
The standard best practice used by MELCOR is to define the effective UO2-ZrO2 melting temperature
as 2800 K to simulate the reduction in liquefaction temperatures that results from materials interactions.
Molten pools form as molten material that candles on surfaces is blocked and forms contiguous molten
regions in either the core or in the lower plenum. By default, MELCOR 1.8.6 assumes that molten pools
will separate into a denser oxide pool with a metallic pool above. Radionuclides are partitioned between
the molten pools and convective heat transfer is calculated from Rayleigh correlations for convecting
pools. Heat transfer from the molten pool to the supporting substrate is calculated using an integral
Stefan model.
For the sensitivity calculation, the liquefaction temperature for UO2-ZrO2 has been reduced from thestandard 2800 K to 2550 K. In addition, a new damage criterion for fuel rod collapse was implemented.
This model can lead to debris bed formation at temperatures as low as 2100K and ensures rod failure attemperatures that exceed 2500 K.

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NEA/CSNI/R(2009)3

Parameter

Standard Case

Sensitivity

Zr Oxidation
correlation

Urbanich Heidrich

Urbanich Heidrich
In addition to the standard failure
criterion, failure based damage function.
 Damage accumulates when un-oxidized
clad thickness < 1 mm

Cladding Failure(rod collapse) criteria

 T > 2500 K and e(ZrO2) < 100 μm
 T > 2800 K

UO2-ZrO2 melting
temperature

2800 K

2550 K

Debris Formation

Debris forms when clad fails

Debris forms when clad fails

Debris diameterCore - 0.01 m
Lower Plenum - 0.002 m

Core - 0.01 m
Lower Plenum - 0.002 m

Debris porosity

0.4

0.4

Temperature (K)< 10001000T<2100
2100  T < 2500
T>2500Time (sec)
1.0e9
3600.01800.060.0

19.5 Chronology of Major Events
The chronology of major events calculated by MELCOR is presented in the table below.
Time (s)EventStandard caseSensitivity case

Break opening and loss of feedwater0

0

Reactor scram

25

25

Pressurizer is empty

144

144

Primary pump void fraction > 0.6

4770

4770

Primary pump shutdown (M < 85000 kg)

4840

4840

Onset of core uncovery

4900

4900

Onset of core heatup

4910

4910

Beginning of oxidation

6500

6500

First fuel rod clad burst

6690

6710

First melt relocation (U-Zr-O)

7580

7580

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NEA/CSNI/R(2009)3

First ceramic melting (UO2-ZrO2)

8200

9200

Onset of reflooding

9840

9840

End of calculation

14000

13920

19.6 Result Analysis
Thermal hydraulics
Overall, the results of the MELCOR 1.8.6 calculation are in good agreement with the other calculationsin this benchmark activity. Loss of primary coolant inventory through the break was well simulated
though it was noted that careful modeling of flow path junction heights in the system was necessary to
avoid stratification of flow throughout the system. This stratification results from MELCOR‟s lack of aflow regime map and improper modeling of the flow path opening heights. Also, flow rates through theloops is significantly higher in the MELCOR calculation probably due to differences in pump models.
Core Degradation
Heat up of the core occurs at roughly the same time as other codes, though onset of core uncovery may
be a bit earlier than other calculations. At about 8000 seconds, fuel rods lose integrity just above themiddle of the core and the MELCOR support logic leads to failure of intact rods from the level of initialrod failure to the top of the core. As a result, the upper core slumps and there is no longer a temperaturerecorded for intact rods at the top of the core, though other codes continue to calculate temperature
increase of fuel rods at the top of the core. Oxidation signatures predicted by MELCOR initially look
similar to other codes with oxidation beginning at 6500 seconds though the total hydrogen generation
predicted by MELCOR is less than other codes as oxidation diminishes when the intact rods form debris
and slumps at 8000 seconds.
The standard default calculation predicts only 20 tons total of molten material, much less than the 40
tons predicted by other codes. This is explained by the relatively high default UO2/ZrO2 melting
temperature of 2800 K. The sensitivity calculation predicts significantly more melted material, 33 tons,
which compares more closely to the other codes. It also appears that the debris rests lower, near the coreplate than predicted by other codes. Fifty-five tons of debris are predicted by the standard calculation(65 tons in the sensitivity calculation) of which about 6 tons of debris, composed mostly of steel andcontrol rod material (with trace quantities of UO2 and ZrO2) relocates onto the lower head.
The flow blockage model has been activated for axial flow paths in the core region. However, radial
blockages were not activated. Adding radial flow blockages may reduce heat removal from the debris
bed and lead to additional melted material.
It is also of interest to point out that the standard MELCOR calculation predicted that the outer fuel rods
remain standing throughout the calculation. However, the sensitivity calculation predicts failure of theouter fuel rods during the reflood. This may be due to the time-at-temperature criterion of failure thathas become a MELCOR „best practice‟ in recent time.
Lessons Learned
Important lessons learned that are being considered by the MELCOR development team as a result of
this exercise include perhaps the need to reduce the default UO2/ZrO2 melting temperature to be morein line with Phebus FPT observations and the default settings for flow path opening heights in verticalflow paths.

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The original input deck also specified multiple radial flow paths connecting a single control volume withan adjacent control volume in the core. This was done because a single control volume was associatedwith multiple COR cells and the flow paths were specified to capture flow blockage within each COR
cell. This nodalization scheme led to non-physical convection cells within the core that redistributed
heat from the center of the core to outer rings. This practice is highly discouraged and the final
specification associated a single control volume with each COR cell and removed this nodalization
problem.

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19.7 Synthetic Views of the Core at Selected Times
400

CaseID

CaseID

400

450
500
550
600
650
700
750
800
850
900

450
500
550
600
650
700
750
800
850
900

950

950

1000
1050
1100
1150
1200
1250
1300
1350
1400
1450

1000
1050
1100
1150
1200
1250
1300
1350
1400
1450

1500

1500

1550
1600
1650
1700
1750
1800
1850
1900
1950

1550
1600
1650
1700
1750
1800
1850
1900
1950

2000

2000

2050
2100
2150
2200
2250
2300
2350
2400
2450
2500

2050
2100
2150
2200
2250
2300
2350
2400
2450
2500

2550

2550

2600
2650
2700
2750
2800
2850
2900
2950

3000

2600
2650
2700
2750
2800
2850
2900
2950
3000

6500 (sec)

6500 (sec)Core State at onset of oxidation for default case Core State at onset of oxidation for sensitivity(6500 sec)case (6500 sec)

CaseID

400

400

450
500
550
600
650
700
750
800
850
900

450
500
550
600
650
700
750
800
850
900

950

950

1000
1050
1100
1150
1200
1250
1300
1350
1400
1450

1000
1050
1100
1150
1200
1250
1300
1350
1400
1450

1500

1500

1550
1600
1650
1700
1750
1800
1850
1900
1950

1550
1600
1650
1700
1750
1800
1850
1900
1950

2000

2000

2050
2100
2150
2200
2250
2300
2350
2400
2450
2500

2050
2100
2150
2200
2250
2300
2350
2400
2450
2500

2550

2550

2600
2650
2700
2750
2800
2850
2900
2950

2600
2650
2700
2750
2800
2850
2900
2950

3000

3000

7580 (sec)

CaseID

7580 (sec)Core State at First melt relocation for default Core State at First melt relocation for sensitivity
case (7580 sec)case (7580 sec)

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400

CaseID

CaseID

400

450
500
550
600
650
700
750
800
850
900

450
500
550
600
650
700
750
800
850
900

950

950

1000
1050
1100
1150
1200
1250
1300
1350
1400
1450

1000
1050
1100
1150
1200
1250
1300
1350
1400
1450

1500

1500

1550
1600
1650
1700
1750
1800
1850
1900
1950

1550
1600
1650
1700
1750
1800
1850
1900
1950

2000

2000

2050
2100
2150
2200
2250
2300
2350
2400
2450
2500

2050
2100
2150
2200
2250
2300
2350
2400
2450
2500

2550

2550

2600
2650
2700
2750
2800
2850
2900
2950

3000

2600
2650
2700
2750
2800
2850
2900
2950

9840 (sec)

3000

9840 (sec)Core State at time of reflooding for default case Core State at time of reflooding for sensitivity case(9840 sec)
(9840 sec)

400

CaseID

CaseID

400

450
500
550
600
650
700
750
800
850
900

450
500
550
600
650
700
750
800
850
900

950

950

1000
1050
1100
1150
1200
1250
1300
1350
1400
1450

1000
1050
1100
1150
1200
1250
1300
1350
1400
1450

1500

1500

1550
1600
1650
1700
1750
1800
1850
1900
1950

1550
1600
1650
1700
1750
1800
1850
1900
1950

2000

2000

2050
2100
2150
2200
2250
2300
2350
2400
2450
2500

2050
2100
2150
2200
2250
2300
2350
2400
2450
2500

2550

2550

2600
2650
2700
2750
2800
2850
2900
2950

2600
2650
2700
2750
2800
2850
2900
2950

3000

3000

0 (sec)9840 (sec)Core State at end of calculation for default case Core State at end of calculation for sensitivity case(13920 sec)
(14000 sec)

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20. APPENDIX : IRSN CALCULATION WITH ICARE/CATHARE V2

20.1 TMI-2 Plant Modeling
The TMI-2 plant nodalization used in the Icare/Cathare benchmark calculation is shown on figure 1.
This modeling includes a detailed nodalization of the primary circuit with:The reactor pressure vessel,
The two primary coolant loops (with one hot leg and two cold legs for each loop) with the oncethrough steam generators,
The pressurizer (with the surge-line, the heaters, the spray-line and the pilot operated relief
valve),All the main emergency and control systems are also represented.

Concerning the secondary circuit, the modeling is limited to the secondary side of the steam generators,
the steam lines with the isolation valves and of course the main and auxiliary feed waters injections.
The reactor vessel is modeled as an Icare/Cathare full 2D module (see figure 2) with the core and all the
major internal structures. A total of 8x45 meshes are used. 6 radial rings and 25 axial meshes model thecore. Another ring is devoted to the core bypass and another one for the vessel downcomer.
All other elements of the plant are modeled as standard Cathare 0D or 1D modules with their external
walls.
Finally, all the boundaries conditions are been modeled according to the prescribed scenario:With a constant make-up flow,
With regulations of the secondary side of the steam generators.

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Figure 1: Reactor Plant Nodalization

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Figure 2: Reactor Vessel Meshing171

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20.2 Physical Parameters of the CodeThe benchmark calculation has been performed with the last version of Icare/Cathare, which is theversion V2.1-rev1 released in July 2007. The main core degradation physical parameters used in
Icare/Cathare calculation are given in table 1. All the parameters requested in the benchmark
specifications, actually correspond to the recommended values by Icare/Cathare team for reactor
applications. Therefore, no sensitivity study was necessary.
The most advanced models concerning the degradation phase have been used:

2D molten material movement model,
Advanced 2D radiative model in core region,
Molten pool formation and evolution model,
Corium oxidation model.

Finally, the Cathare 3D thermal hydraulics model has been used to calculate the thermal hydraulics inthe whole reactor vessel.
Remark: According to the benchmark specifications, an alternative calculation should have been
performed with different parameters. It was decided to enable the specific debris model and porous
media model, but unfortunately the current Icare/Cathare version was not able to perform suchcalculation due to numerical instabilities in the transition phase from intact geometry to debris
geometry and in the transition phase from standard hydraulic model to porous hydraulic model. Thisshould be solved in the next Icare/Cathare version (V2.2-rev1), which is planned to be released in June
2008.Such results could be available in the nextsteps of this benchmark activity, if it continues.
Parameters

Zr Oxidation correlation

Cladding Failure criteria

Standard Case
Correlations proposed by Schanz et al. (2001)
 Cathcart correlation if T < 1800 K
 Prater correlation if T > 1900 K
 Linear interpolation between 1800 – 1900
K


T > 2300 K and e(ZrO2) < 300 μm
T > 2500 K

UO2-ZrO2 melting temperature

2550-2600 K

2D Corium Movement Model

enableCorium Oxidation Model

enableDebris model

disableTable 1: Mains Icare/Cathare Code Parameters

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20.3 Initial Steady State Conditions
The TMI2 plant state at the beginning of the transient is obtained by performing an Icare/Cathare steadystate calculation of the reactor following by a transient calculation at nominal conditions with standard
regulation processes. This transient lasts about 5000 s and, at the end, all the main plant parameters arestable. All those parameters are given in the following table and compared with TMI-2 accident data at
turbine trip.

Parameter

Unit

ICARE/CATHARE V2

TMI-2

Reactor Power

W

2700

2700

Primary Pressure

MPa

15.02

15.2

Temperature Hot Leg A

K

592.3

592

Temperature Hot Leg B

K

592.3

592

Temperature Cold Leg A

K

566.3

548-561

Temperature Cold Leg B

K

564.7

565

Mass Flow Rate – Loop A

kg/s

8805

8280

Mass Flow Rate – Loop B

kg/s

8846

8560

Pressurizer Level

M

5.60

5.77

Total Primary Masston

230389-Pressure SG A

MPa

6.38

7.31

Pressure SG B

MPa

6.24

7.24

Steam Temperature SG A

K

584.9

586

Steam Temperature SG B

K

583.6

586

Collapsed Level SG A

M

5.26

-

Collapsed Level SG B

M

5.39

-

Liquid Mass SG A

Kg

18425

-

Liquid Mass SG B

Kg

18989

-

Feedwater Flow SG A

Kg/s

700

723

Feedwater Flow SG B

Kg/s

742

717

SG Feedwater Temperature

K

513.0

-

Table 2: Steady State Conditions Parameters

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Since the conditions in the TMI-2 plant just before the turbine trip are not very well known, it notpossible to reach exactly the TMI-2 data, especially for the secondary circuit.
20.4 Analysis of code results
The chronology of major events as calculated by Icare/Cathare is presented in the table 3. Fours
synthetic views of the reactor vessel state at different times are shown on figures 3 to 6.

Time (s)EventStandard Case
Break opening and loss of feedwater

0

Reactor scram

16

Pressurizer is empty

240

Primary pump shutdown (M < 85000 kg)

4831

Primary pump void fraction > 0.6

4855

Onset of core uncovery

5788

Onset of core heatup

6250

Beginning of oxidation

7550

First fuel rod clad burst or perforation

8462

First melt relocation (U-Zr-O)

9272

Onset of reflooding

9831

End of calculation

15000
Table 3: Chronology of Main Events

The scenario calculation can be divided into three phases: the thermal hydraulics phase, the coredegradation phase and the core reflood phase.
20.4.1 The thermal hydraulics phase (from t = 0 s to t = 6250 s)
The scenario starts with the opening of a small break on hot leg A and with the loss of main feedwater at
the same initial time. The loss of feedwater induces a strong reduction of heat transfer between primaryand secondary circuit. The resulting loss of heat removal causes the primary pressure to reach rapidly
the high-pressure reactor trip point and the reactor scram occurs at about 16 s. After this scram, theprimary pressure decreases progressively to reach a stable value near the secondary pressure value atabout 450 s.
During this phase the mass inventory in the primary circuit continuously decreases due to the break on
hot leg A. But as the primary pumps are still working, a forced circulation of a liquid water and steam

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mixture is established. In those conditions, the core decay power is removed totally from the primarycircuit by the steam generators.
According to the benchmark scenario, this behavior continues until the primary mass inventory falls
below 85000 kg. This occurs at 4831 s. At this time, the primary pumps are shutdown.
After pumps stop, the forced convective circulation in primary circuit is replaced by a natural
circulation. This induces a large increase of primary pressure. The heat removal from the core is
progressively reduced and the core decay power is no more removed at about 5450 s. From this moment
the vessel water level decreases continuously and the core uncovery starts at about 5800 s.
The degradation process starts after this moment.
20.4.2 The core degradation phase (from t = 6250 s to t = 9831 s)
Shortly after the start of core uncovery the fuel rod temperatures increase rapidly in the central zone of
the upper part of the core (at about 6250 s). The temperature escalations expand radially and downwards
to the bottom of the core.
The oxidation runaway occurs at about 7550 s and increases the core heat-up rate (the state of the core at
this time is shown on figure 3). It is followed by the first clad failure at about 8460 s and by the firstmelt relocation at about 9270 s (see figure 4).
The evolution of the degradation process can be seen on figures 7. The degradation zone is limited to theupper part of the core because the total dry-out of the core is never reached.
At the time 9831 s, a total of 258 kg of hydrogen has been released and 36577 kg of corium has been
formed. The state of the core is presented on figure 5).
Remark: Between t = 8000 s and t = 9831 s, it is not possible to impose the right boundary conditions
on the steam generators. There is an increase of the steam generators level and at the same time adecrease of steam generators pressure because of steam condensation in the secondary side of the steam
generators.
20.4.3 The core reflood phase (from t = 9831 s to 15000 s)
According to the benchmark specifications, the High Pressure Injection is enabled 5000 s after primary
pumps shutdown. This occurs at 9831 s. The total flowrate injection is 60 kg/s.
The reflood of the core causes a large primary pressure increase (about 9.0 MPa) due to steam
generation in the core. This reflood is very efficient and the vessel is totally refilled and cooled down in
about 1200 s. During this quenching phase, the hydrogen production and the corium formation are very
rapidly stopped. The hydrogen mass release is limited to 17 kg (compared to 258 kg during thedegradation phase) and the corium mass production is only 3368 kg (compared to 36577 kg during theprevious phase).
The evolution of reflood process is presented on figures 8 to 10. This process can be divided into 4 main
phases:

A core bypass phase (during about 300 s)
A top core quenching phase (during about 500 s)

175NEA/CSNI/R(2009)3

A bottom core quenching phase (during about 400 s)
A final steady state phase (from t = 11000 s to t = 15000 s)

Remark: After t = 12500 s, it is not possible to impose the right boundary conditions on the steam
generators. There is an increase of the steam generators level and at the same time a decrease of steam
generators pressure because of steam condensation in the secondary side of the steam generators.

20.5 Comparisons to other code results
In general there is a good agreement between Icare/Cathare and other codes results:The thermal-hydraulic behavior as calculated by Icare/Cathare is very consistent with other code
results.
There is no large discrepancy concerning total hydrogen release and total molten material
formation.
As other codes, Icare/Cathare predicts a very efficient reflood process with a rapid vessel
refilling and a very limited amount of hydrogen release and corium formation during quenchingphase.

The main discrepancies with other codes results are the following:The primary pressure evolution after pumps stop differs.
The starts of core uncovery and core heat-up rate are different.
The degradation process seems to be rather different between different codes.

It seems that the thermal hydraulic behavior in the steam generators just after the pump shutdown are
not the same in the different calculations and it should have a non negligible impact on the beginning ofdegradation process and then on the conditions just before reflooding. This is not yet clearly explained.

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Figure 3: View of the Reactor Vessel at the beginning of oxidation

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Figure 4: View of the Reactor Vessel at time of first melt relocation

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Figure 5: View of the Reactor Vessel at the start of the reflooding

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Figure 6: View of the Reactor Vessel at the end of the calculation

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Figure 7: Core Degradation Phase: Relocation Process (t = 9250 s  t = 9750 s)

t = 9250 s

t = 9350 s

t = 9450 s

t = 9550 s

t = 9650 s

t = 9750 s

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Figure 8: Core Reflood Phase: Core By-Pass Phase (t = 9800 s  t = 10100 s)

t = 9800 s

t = 9900 s

t = 10000 s

t = 10100 s

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Figure 9: Core Reflood Phase: Top Core Quenching Phase (t = 10100 s  t = 10600 s)

t = 10100 s

t = 10200 s

t = 10300 s

t = 10400 s

t = 10500 s

t = 10600 s

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Figure 10: Core Reflood Phase: Bottom Core Quenching Phase (t = 10600 s  t = 11000 s)

t = 10600 s

t = 10700 s

t = 10900 s

t = 11000 s

t = 10800 s

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21. APPENDIX : SNU CALCULATION WITH MAAP4

21.1 TMI-2 Plant Modelling
The MAAP4.03 primary system nodalization for tracking these quantities in a Babcock &Wilcox type
plant is shown in Fig. 1. The reactor consists of four volumes: core, downcomer, upper plenum, andreactor dome. All loops except one, the broken loop, are lumped together, and the broken loop is treatedseparately. The broken loop refers to the loop that can contain a primary system break. The user selects
whether the pressurizer is in the broken or unbroken loop. The A-loop is taken as the broken loop for
analysis of the TMI-2 accident.
The reactor vessel is nodalized in the form of heat sinks and control volumes. Fig. 2 specifies the coreregion with the number of radial rings and axial rows. A radial peaking factor and volume fraction are
fixed for each ring, and an axial peaking factor is assigned for each row. Seven rings and thirteen rows
are used to nodalize the TMI-2 core for simulation of the accident.
Two input files are required to simulate the TMI-2 accident with MAAP4: the parameter file and an
input deck. The user inputs various parameters for each component such as volumes or masses. The
parameter file has the information on TMI-2 relevant to the plant geometry, system performance,
controls and initial conditions used to perform these benchmark calculations. The input deck defines theoperator actions and boundary conditions during the course of the accident.
Mass and energy rates of change for core materials are calculated for each core node. Steam andhydrogen are assumed to flow along the uncovered and unblocked flow channels, and the mass flow
rates and enthalpies in channel are determined by accounting for the generation and consumption at eachaxial level. The core water pool is treated as a lumped mass and energy volume.

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Unbroken
Loop

BrokenLoop

Candy Cane⑩
Candy Cane
( Hot Leg Tubes)④

Hot Leg③
Hot Leg⑨

SG Tubes⑪

Reactor Dome⑭
SG Tubes(Cold Leg Tubes)⑤

Upper Plenum②

Break
LocationCore①
Cold Leg⑬

Cold Leg⑦
Downcomer⑧

Intermediate Leg⑫

Intermediate Leg⑥

Figure 1 Primary System Nodalization

Core Shroud/Baffle

7X23

Upper Tie Plate +Upper Gas Plenum

23
22
21
20
19
18
17
16
15
14
13
12
11
10
9
8
7
6
5
4

Row

3
Lower Tie Plate +Lower Gas PlenumCore Support Plate

2
1
1

2

3

4

5

6 7

Ring

Figure 2 Primary System Nodalization

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21.2 Initial Steady State Conditions and Scenarios
The accident is initiated by a small break at the bottom of hot leg A together with a loss of main
feedwater for the TMI-2 nuclear power plant. The main assumptions include stopping of the primary
pumps when the primary mass void fraction exceeds 0.6 and starting of the High Pressure Injection(HPI) 5000 s thence forth. There is no letdown flow nor Power Operated Relief Valve (PORV) failure.
This alternative scenario was chosen because it describes a standard small-break scenario with core
uncovery and core degradation and then reflood so that it could lead to good comparison without a great
deal of uncertainties involved with the actual accident.
Table 1 lists the TMI-2 plant initial conditions for MAAP4. Table 2 presents the core heat flux profile
according to radial ring and axial row.
Table 1. Initial Conditions
Main Parameters

Unit

MAAP4.03

Reactor PowerPrimary pressure
Temperature Hot Leg A
Temperature Hot Leg B
Temperature Cold Leg A
Temperature Cold Leg B
Pressurizer level
Total primary mass
Pressure SG A
Pressure SG B
Steam temperature SG A (inlet)
Steam temperature SG B (inlet)
Collapsed level SG A
Collapsed level SG B
Liquid mass SG A
Liquid mass SG B
Feedwater flow SG A
Feedwater flow SG B
SG feedwater temp.

MW
MPa
K
K
K
K
m
kg
MPa
MPa
K
K
m
m
kg
kg
kg/s
kg/s
K

2700
15.2
579
579
559
559
6.78
211170
6.38
6.38
508
508
3.28
3.28
17506
17506
756
756
508

Table 2. Core Heat Flux Profile
Radial profile
Ring
1
2
3
4
5
6
7

Power factor
1.162
1.160
1.155
1.150
1.130
0.803
0.480

187

Area factor
0.09370.11230.16800.16160.15480.15480.1548

NEA/CSNI/R(2009)3

Axial profile
Row
23
22
21
20
19
18
17
16
15
14
13
12
11
10
9
8
7
6
5
4
3
2
1

Z (m)
4.56
3.99
3.81
3.62
3.44
3.26
3.08
2.89
2.71
2.53
2.34
2.16
1.98
1.80
1.61
1.43
1.25
1.06
0.88
0.70
0.52
0.33
0.08

Power factor
00.57019
0.81171
0.95979
1.02413
1.04428
1.05070
1.05660
1.06329
1.06838
1.07127
1.07435
1.08075
1.091171.10233
1.10843
1.10482
1.08932
1.05178
0.93659
0.555060021.3 Sensitivity Study
The main core degradation physiscal parameters used in the standard calculation with MAAP4 arecompared against alternative physical parameters for PHEBUS-FPT1 in Table 3.

Table 3. Core Degradation Physiscal Parameters
PHEBUS FPT1

MAAP4
Baker-Just model (Tclad > 1875 K)
Cathcart model (Tclad < 1850 K)
Interpolated inbetween

Zr
Oxidationcorrelation
Urbanic-Heidrick model

Cadding
failure T > 2280 K and e(ZrO2) < 200 mm
T > 2500 K
criteria
T > 2450 K and e(ZrO2) < 300 mm
Fraction of clad reacted > 0.675
UO2-ZrO2 melting T > 2550K
T > 2500 K
temperature
Debrisformation No debris bed and molten pool formation Melt progression model
criteria
modelling
Debris diameter
Debris porosity
0.4 (debris bed)

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21.4 Chronology of Major Events
The chronology of major events calculated by MAAP4 is presented in Table 4 and discussed in
Section 6.

Table 4. Major Events Squence
Time (s)EventBreak opening and loss of feedwater
Reactor scram
Pressurizer is empty
FP models on
Primary pump shutdown
Onset of core uncovery
Beginning of oxidation
Maximum core temperature has exeeded 2499KFirst fuel rod clad collapses
First melt relocation
HPI swich no forced off
Onset of reflooding
End of calculation

Standard

Alternative

0
18
221
1497
5342
5682
6251
7638
7789
7794
10342
10342
10422

0
18
221
1797
5342
5682
6251
7528
7775
7775
10342
10342
10422

21.5 Results Analysis
The overall system behavior was reasonably predicted according to the hot leg break opening at 0 s, thecontemporary feedwater trip with consequent loss of heat removal by the secondary side results in the
primary pressure increase. Reactor scrammed due to the pressure in the primary system is higher than a
MAAP4 specified high pressurizer-pressure trip point at 18 s. The primary system starts to depressurize
after the reactor scram. The steam generator secondary side is dry due to loss of feedwater. The primarypressure decrease is accelerated when the pressurizer gets empty by 221 s.
Before the pump shutdown a steam-water mixture circulates in the primary circuit and through the coreat the saturated condition, while the void fraction in the primary system increases due to loss of fluid(mostly liquid) from the hot leg break. During this time the core decay power and pump heat areremoved by the steam generators. The primary pumps are shut down when the primary pump void
fraction exceeds 0.6 after 5342 s.
After the pump shutdown, the liquid water starts to settle down in the primary circuit and inside thevessel. By this time the break flow rate quickly reduces, changing from mostly liquid to pure steam flow
due to liquid stratification in the primary system with hot leg draining. The reactor core starts to uncoverat 5682 s because the decay power is no longer removed by natural circulation in the primary circuit.

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Cladding oxidation in core begins at 6251 s. The maximum core temperature exceeds 2499K by 7638 s
in the stnadard calculation. In the sensitivity calulation, on the other hand, the maximum coretemperature exceeds 2499K at 7528 s. It is followed after few seconds by clad failure and first melt
relocation. First clad collapse occurs a bit earlier at 7775 s due to higher oxidation rate below 1200 K
with the Urbanic-Heidrick correlation. Melt relocation is calculated later on at 7775 s in the sensitivity
calculation due to HPI operation. Table 5 presents sensitivity results according to main events.
Table 5. Comparison of Sensitivity Results
Baker-Just Model
Hydrogen
Generation [kg]Clad Reacted
Fraction [%]3.297E+11 2.50E+04

7.47E-05

07789

3.716 E+11 2.45E+10

84

18.18

Time of first melt relocation

7794

3.762E+11 2.49E+10

85

18.47

Time reflooding

10342

4.373E+11 6.57E+10

225

48.77

End of the calculation

10422

4.39E+11

225

48.81

Main Events

Time [s]DecayEnergy [J]Oxidation
Energy [J]Hydrogen
Generation [kg]Clad Reacted
Fraction [%]Beginning of oxidation

6251

3.297E+11

8.08E+04

2.42E-040

Beginning of node collapses

7775

3.755E+11

3.20E+10

110

23.79

Time of first melt relocation

7775

3.755E+11

3.20E+10

110

23.79

Time reflooding

10342

4.289E+11

6.81E+10

238

51.70

End of the calculation

10422

4.38E+11

7.78E+10

292

64.30

Main Events

Time [s]DecayEnergy [J]Beginning of oxidation

6251

Beginning of node collapses

Oxidation
Energy [J]6.58E+10

Urbanic-Heidrick Model

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21.6 Synthetic Views of the Core
21.6.1 Standard Case

- At the beginning of oxidation

191

NEA/CSNI/R(2009)3- At the time of first melt relocation

192

NEA/CSNI/R(2009)3- At the time of reflooding

193

NEA/CSNI/R(2009)3- At the end of the caculation

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21.6.2 Sensitivity Case

- At the beiginning of oxidation

195

NEA/CSNI/R(2009)3- At the time of first melt relocation

196

NEA/CSNI/R(2009)3- At the time of reflooding

197

NEA/CSNI/R(2009)3- At the end of the caculation

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22. APPENDIX : IKE CALCULATION WITH ATHLET/MEWA

The results of the calculations of IKE could not be provided in time to be included in the general
presentation of the benchmark results. Therefore, major results of the calculations with ATHLET-CD
including the porous media module MEWA are briefly described here.
22.1 TMI-2 Plant Modelling
For the calculations, ATHLET-CD was run with essentially the same input that was used by GRS for the
standard calculations (made available by courtesy of Henrique Austregesilo). Thus, the plant modelling
ist he same as described in the GRS contribution.
22.2 Main modelling options
Up to the onset of quenching by high pressure injection 5000s after switch-off of the main coolant
pump, the modelling options were the same as in the GRS calculation with ATHLET-CD for the
standard case.
In the calculations of IKE, the description for the core region was switched to the MEWA model, whenthe quenching of the core is starting (at about 9700 s). Switching to the MEWA model was done at once,
i.e. the formation of a particulate debris bed was assumed to take place instantaneously, notprogressively. Temperatures, porosity and melt fractions were derived from the results of the ATHLETCD calculation with representative rod modelling at this time. A representative particle diameter of 3
mm assumed.
22.3 Discussion of Results
The calculation of IKE with ATHLET-CD only differs from the calculation of GRS for the standard
case after onset of quenching, when the MEWA module was activated. For the results up to this point, itis therefore referred to the results of GRS.
Figs. 22-1 to 22-5 show synthetic views of the core status at selected instants. It can be seen that the coreis quenched with water rising in lateral, colder and more intact core regions (higher porosity), whichthen penetrates from the sides into the degraded parts, driven by the outer water head. This clearydemonstrates that multi-dimensional aspects decide about the success of cooling.
The quenching takes quite a long time, due to the slow water penetration into the debris bed. At the endof the calculation, still some parts of liquid metallic melt remain in the compacted region at the bottomof the core, but the temperatures in this part remain below the melting temperature of ceramic
corium.Also, the melt region is not increasing in extent any more.
Fig 22-6 shows the development of the pressure in the pressurizer. Up to the onset of quenching it
coincides with the respective GRS calculation for the standard case. The pressure increase after about

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9700 s during quenching is less pronounced, since the steam production is limited by the water access tothe hot regions of the core, which is much slower in the case of the debris bed, due to the large friction.
Concerning hydrogen production shown in Fig 22-7, the ATHLET-CD/MEWA calculation only predicts
a very small additional amount of about 25 kg during the quenching phase. According to the calculation,hydrogen production in this phase is limited by steam availability, not by temperature. The low steam
availability is again caused by the large friction and thus slow water penetration into the debris bed.
Fig 22-8 and Fig 22-9 show that further degradation and melting of the core is practically stopped after
onset of flooding. Only some limited additional melting takes place, while the melt that was present
prior to flooding is partly being solidified. The ceramic melt refreezes completely, while parts ofmetallic melt in a compacted region in the lower part of the core remain liquid still at the end of thecalculation.

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4.0

Time t = 9780s

3.5

3.5

3.0

3.0

2.5

2.5

Height [m]Height [m]

4.0

2.0

1.5

Time t = 9870.9s

2.0

1.5

Liquid Volume Fraction [-]

1.0

0.80
0.70
0.60
0.50
0.40
0.30
0.20
0.10

0.5

Liquid Volume Fraction [-]

1.0

0.80
0.70
0.60
0.50
0.40
0.30
0.20
0.10

0.5

0.0

0.0
-2

-1

0

1

2

-2

-1

Radius [m]
4.0

1

4.0

Time t = 9907.2s

3.5

3.5

3.0

3.0

2.5

2.5

Height [m]Height [m]02Radius [m]2.0Time t = 10036s

2.0

1.5

1.5

Liquid Volume Fraction [-]

1.0

0.80
0.70
0.60
0.50
0.40
0.30
0.20
0.10

0.5

Liquid Volume Fraction [-]

1.0

0.80
0.70
0.60
0.50
0.40
0.30
0.20
0.10

0.5

0.0

0.0
-2

-1

0

1

-2

2

-1

0

Radius [m]Radius [m]Fig 22-1: Water volume fraction in TMI core during reflooding

201

1

2

NEA/CSNI/R(2009)3

4.0

Time t = 10299s

3.5

3.5

3.0

3.0

2.5

2.5

Height [m]Height [m]

4.0

2.0

1.5

Time t = 10437s

2.0

1.5

Liquid Volume Fraction [-]

1.0

0.5

Liquid Volume Fraction [-]

1.0

0.80
0.70
0.60
0.50
0.40
0.30
0.20
0.10

0.80
0.70
0.60
0.50
0.40
0.30
0.20
0.10

0.5

0.0

0.0
-2

-1

0

1

2

-2

-1

Radius [m]
4.0

1

4.0

Time t = 10584s

3.5

3.5

3.0

3.0

2.5

2.5

Height [m]Height [m]02Radius [m]

2.0

1.5

Time t = 11536s

2.0

1.5

Liquid Volume Fraction [-]

1.0

0.5

Liquid Volume Fraction [-]

1.0

0.80
0.70
0.60
0.50
0.40
0.30
0.20
0.10

0.80
0.70
0.60
0.50
0.40
0.30
0.20
0.10

0.5

0.0

0.0
-2

-1

0

1

2

-2

Radius [m]-10Radius [m]Fig 22-2: Water volume fraction in TMI core during reflooding (continued).

202

1

2

NEA/CSNI/R(2009)3

4.0

Time t = 9870.9s

3.5

3.5

3.0

3.0

2.5

2.5

2.0
Solid Temperature [K]1.5

Height [m]Height [m]

4.0

0.5

2.0
Solid Temperature [K]

1.5

2'800
2'600
2'400
2'200
2'000
1'800
1'600
1'400
1'200
1'000
800
600
400
200

1.0

Time t = 9780s

2'800
2'600
2'400
2'200
2'000
1'800
1'600
1'400
1'200
1'000
800
600
400
200

1.0

0.5

0.0

0.0
-2

-1

0

1

2

-2

-1

Radius [m]
4.0

1

4.0

Time t = 9907.2s

3.5

3.5

3.0

3.0

2.5

2.5

2.0
Solid Temperature [K]1.5

Height [m]Height [m]02Radius [m]

0.5

2.0
Solid Temperature [K]

1.5

2'800
2'600
2'400
2'200
2'000
1'800
1'600
1'400
1'200
1'000
800
600
400
200

1.0

Time t = 10036s

2'800
2'600
2'400
2'200
2'000
1'800
1'600
1'400
1'200
1'000
800
600
400
200

1.0

0.5

0.0

0.0
-2

-1

0

1

2

-2

-1

Radius [m]0Radius [m]Fig 22-3: Temperature distribution in TMI core during reflooding.

203

1

2

NEA/CSNI/R(2009)3

4.0

Time t = 10299s

3.5

3.5

3.0

3.0

2.5

2.5

2.0
Solid Temperature [K]1.5

Height [m]Height [m]

4.0

0.5

2.0
Solid Temperature [K]

1.5

2'800
2'600
2'400
2'200
2'000
1'800
1'600
1'400
1'200
1'000
800
600
400
200

1.0

Time t = 10437s

2'800
2'600
2'400
2'200
2'000
1'800
1'600
1'400
1'200
1'000
800
600
400
200

1.0

0.5

0.0

0.0
-2

-1

0

1

-2

2

-1

4.0

1

4.0

Time t = 10584s

3.5

3.5

3.0

3.0

2.5

2.5

2.0
Solid Temperature [K]1.5

Height [m]Height [m]02Radius [m]Radius [m]

0.5

2.0
Solid Temperature [K]

1.5

2'800
2'600
2'400
2'200
2'000
1'800
1'600
1'400
1'200
1'000
800
600
400
200

1.0

Time t = 11536s

2'800
2'600
2'400
2'200
2'000
1'800
1'600
1'400
1'200
1'000
800
600
400
200

1.0

0.5

0.0

0.0
-2

-1

0

1

-2

2

-1

0

Radius [m]Radius [m]Fig 22-4: Temperature distribution in TMI core during reflooding (continued).

204

1

2

NEA/CSNI/R(2009)3

4.0

Time t = 9780s

3.5

3.5

3.0

3.0

2.5

2.5

Height [m]Height [m]

4.0

2.0

1.5

Time t = 11536s

2.0

1.5
Porosity [-]1.0

Porosity [-]

1.0

0.9
0.8
0.7
0.6
0.5
0.4
0.3
0.2
0.1

0.5

0.9
0.8
0.7
0.6
0.5
0.4
0.3
0.2
0.1

0.5

0.0

0.0
-2

-1

0

1

2

-2

Radius [m]

-1

0

1

2

Radius [m]Fig 22-5: Porosity distribution in TMI core during reflooding.
Pressurizer Pressure
1.80E+07
1.60E+07
1.40E+07

Pressure (Pa)

1.20E+07
1.00E+07
8.00E+06
6.00E+06
4.00E+06
2.00E+06
0.00E+00
0

2000

4000

6000

8000

Time (s)Fig 22-6: Development of pressure in the pressurizer

205

10000

12000

14000

NEA/CSNI/R(2009)3

Cumulated Hydrogen Production
3.50E+02

3.00E+02

H2 Mass (kg)2.50E+02

2.00E+02

1.50E+02

1.00E+02

5.00E+01

0.00E+00
0

2000

4000

6000

8000

10000

12000

14000

12000

14000

Time (s)Fig 22-7: Cumulated hydrogen production
Total Mass of Molten Materials
4.50E+04

4.00E+04

3.50E+04

Mass (kg)3.00E+04

2.50E+04

2.00E+04

1.50E+04

1.00E+04

5.00E+03

0.00E+00
0

2000

4000

6000

8000

10000

Time (s)Fig 22-8: Development of total mass of molten materials (i.e. sum of liquid and refrozen melt)

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Mass of Molten Pool
2.50E+04

2.00E+04

Mass (kg)1.50E+04

1.00E+04

5.00E+03

0.00E+00
0

2000

4000

6000

8000

10000

12000

14000

Time (s)Fig 22-9: Mass of molten pool (liquid melt)

22.4 Additional calculations
In additional calculations, it was tried to evaluate the limits of coolability of the degraded core, i.e. thepossibility of reflooding the core before melting further progresses such that an un-coolable situation isreached, due to too large friction in the debris bed (which impedes water access resp. vapour removal).
For this, the representative particle diameter was varied. All other parameters were as in the previous
calculation. For particles size smaller than 1 mm, complete quenching was not possible any more (seeFig 22-10). The heat-up of the dry parts and melting continues and a larger molten pool is obtained at
the end of the calculation (Fig 22-11).
In general, the possibility to quench a particulate debris bed in the core is quite good according to thepresent calculations. Only in the case of very finely fragmented core (very large friction, representative
particle diameter of 1 mm or below) a non-coolable situation is reached. Such core configurations are
considered as relatively un-realistic. Of course, in the present case the the coolability is strongly
supported by the high pressure. In similar calculations done with MEWA for reactor cases at lower
pressures (2-5 bars), limits of coolability were rather obtained for average particle sizes around 2-3 mm.
But, also in these cases, it was shown that 2D/3D features strongly help to reflood a strongly degradedcore and to establish coolable configurations.

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4.0

4.0

Time t = 10433s

3.5Time t = 11782s

3.5

Liquid Velocity U [m/s]Liquid Velocity U [m/s]
0.03

3.0

3.0

2.5

2.5

Height [m]Height [m]

0.03

2.0

1.5

2.0

1.5

Liquid Volume Fraction [-]

1.0

0.5

Liquid Volume Fraction [-]

1.0

0.80
0.70
0.60
0.50
0.40
0.30
0.20
0.10

0.80
0.70
0.60
0.50
0.40
0.30
0.20
0.10

0.5

0.0

0.0
-2

-1

0

1

2

-2

-1

Radius [m]

0

1

2

Radius [m]Fig 22-10: Development of water fraction in during reflooding for the case with 1 mm representative
particle diameter
4.0

Time t = 11782s

3.5

3.5

3.0

3.0

2.5

2.5

2.0
Solid Temperature [K]1.5

Height [m]Height [m]

4.0

0.5

2.0

1.5

2'800
2'600
2'400
2'200
2'000
1'800
1'600
1'400
1'200
1'000
800
600
400
200

1.0

Time t = 11782s

Porosity [-]

1.0

0.9
0.8
0.7
0.6
0.5
0.4
0.3
0.2
0.1

0.5

0.0

0.0
-2

-1

0

1

-2

2

-1

0

1

2

Radius [m]Radius [m]Fig 22-11: Temperature (left) and porosity distribution (right) at the end of the calculation for the casewith 1 mm representative particle diameter

208