Light Water Reactor Safety: The Development of Advanced Models and Codes for Light Water Reactor Safety Analysis

Light Water Reactor Safety The Development of Advanced Models and Codes for Light Water Reactor Safety Analysis

Light Water Reactor Safety The Development of Advanced

Models and Codes for Light Water Reactor Safety Analysis

J.N. LILLINGTON

AEA Technology Winfrith Technology Centre Dorchester, Dorset, U.K.

1995 ELSEVIER Amsterdam - Lausanne - New York - Oxford - Shannon - Tokyo

ELSEVIER SCIENCE B.V. Sara Burgerhartstraat 25 P.O. Box 2II, 1000 AE Amsterdam, The Netherlands

L1brary of Congress Catalog1ng-1n-Pub11cat1on Data

L1111ngton,

J.

N.

Llght water reactor safety

the development of advanced models

and codes for 11ght water reactor safety analysls

/ J.

N.

L1111ngton. p.

cm.

Inc l udes blb l l ograph1ca l

ISBN 0-444-89741-0 1.

references and

1ndex.

(acld-free paper)

Llght water reactors--Safety measures.

TK9203.L45L55

I.

Tltle.

1995

621.48'35--dc20

95-2322 CIP

ISBN: 0444 8974 1 0

© 1995 Elsevier Science B.V. All rights reserved. No part of this publication may be reproduced, stored in a retrieval system or transmitted in any form or by any means, electronic, mechanical, photocopying, recording or otherwise, without the prior written permission of the publisher, Elsevier Science B.V., Copyright & Permissions Department, P.O. Box 521, 1000 AM Amsterdam, The Netherlands.

Special regulations for readers in the U.S.A. - This publication has been registered with the Copyright Clearance Center Inc. (Ccq, 222 Rosewood Drive, Danvers, MA, 01923. Information can be obtained from the CCC about conditions under which photocopies of parts of this publication may be made in the U.S.A. All other copyright questions, including photocopying outside of the U.S.A., should be referred to the publisher. No responsibility is assumed by the publisher for any injury and/or damage to persons or property as a matter of products liability, negligence or otherwise, or from any use or operation of any methods, products, instructions or ideas contained in the material herein. This book is printed on acid-free paper. Printed in The Netherlands.

v

PREFACE A large proportion of the nuclear reactors in operation worldwide are Light Water Reactors (LWRs). There are two principal designs, Pressurised Water Reactors (PWRs) and Boiling Water Reactors (BWRs). Over the past 10-20 years, much experimental and theoretical research has been carried out in Europe, the USA and elsewhere to ensure safe operation and to understand the behaviour of these plants under various accident conditions. This has included a significant level of severe accident related research, particularly since 1979 when a core melt-down occurred at the Three Mile Island - Unit 2 PWR in the USA. It is of paramount importance that the core in a nuclear reactor should remain adequately cooled at all times. Even after a successful shut down, heat continues to be produced from the decay heat of unstable isotopes in the fuel. One of the main objectives of reactor safety research is to consider ways in which core cooling can be achieved and the consequences if it is not achieved. Thus, in the case of a LWR, the physics of water/steam flows, associated heat transfer and other phenomena (e.g. hydrogen production and fuel degradation, in severe accidents) that could occur under accident conditions need to be adequatel y understood. Many complex and expensive experiments have been perfonned and sophisticated analysis techniques (large computer codes etc.) have been developed to provide the necessary infonnation. The main purpose of this book is to provide a review of and references to the main activities that have been carried out towards the development of advanced mechanistic models/codes for this LWR Safety Analysis. The book will describe the state of the art and discuss the likel y future direction of research in this area. This objective will be achieved by summarising the basic features of LWRs that impact on safety, the key accident scenarios and physical phenomena, the major experiments, the resulting models/codes developed and their validation together with some representative plant calculations that have been perfonned with the codes. The book opens with a brief historical review of the key events from the inception of civil nuclear power soon after the Second World War up to the present time. The prominant safety studies in support of Light Water Reactor design and licensing, the major areas of technical research and nuclear safety research objectives are among the main items summarised. These latter objectives have helped to define modelling requirements which have then driven research to provide experimental data and tools for analysis. The principal features of the designs of both types of Western LWRs i.e. PWRs and B WRs, relevant to the safety issues presented in this book, are given. These include the elements of the reactor coolant system (PWR), steam cycle (BWR), the core and pressure vessel, the containment and the main protection systems. The main emphasis of the book is on the PWR but where appropriate some mention is made of the BWR. Many of the important modelling requirements are common to both designs.

VI

Preface

Accident conditions are usually classified within various categories. These categories range from classes of relatively trivial faults expected within the working life of the plant, through to design basis accidents and the very low probability beyond design basis (severe) accidents. Within these categories there are various types of accident condition. A range is described which is intended to span the scope of relevant phenomena. Integral experiments are performed for two main reasons: (a) to identify important phenom­ ena expected to exist within a certain accident scenario and (b) for code validation. The main integral programmes addressing both thermal-hydraulic and severe accident issues are briefly reviewed. Separate�ffects experiments are referenced where appropriate in the chapters devoted to individual modelling areas. Thermal-hydraulic model development has been a major activity during the last twenty years. As understanding and analysis techniques have improved the trend has been to produce "best estimate" models/codes with some justified estimate of uncertainty. These are replacing the conservative or pessimistic models/codes which supported safety cases hitherto. Some of the important models available in present best estimate system codes are described. The response of a nuclear power plant under many forms of accident condition will depend on its response to abnormal heat and pressure loads. Models for component (fuel rods, structures etc.) heat conduction and heat transfer coupling to the thermal-hydraulics are reviewed. Mechanical response models in response to heat loads are also swnmarised. Under severe accident conditions there are a large number of additional phenomena( compared with the phenomena present in design basis accidents) which also require modelling. Core melting, reactor coolant circuit failure and the threat to containment are possible events that need to be considered. It is these events that have attracted the most attention in recent years. A large proportion and the emphasis of the book therefore concentrates on severe accident modelling. At sufficiently high temperatures, oxidation of core components, particularly Zircaloy fuel rod cladding results in the production of significant hydrogen. This is an extremely important reaction in relation to the progression of many severe accident conditions. Considerable heat is evolved during the oxidation process resulting in strong positive feedback and further increase in temperature. Nwnerous materials reactions occur, resulting in loss of geometry through the production of low melting point eutectics. Melting points of indi vidual constitu­ ents present in the plant are much higher. A chapter is devoted to all these various interactions. Both mechanistic and parametric models have been produced for predicting the meltdown and degradation of the reactor core under these extreme conditions. The sequence of events is affected by the formation of the different eutectics particularly those formed between the metallic Zircaloy and other core components. During a melt progression sequence there is the potential for molten material to fall into water giving rise to an energetic interaction. Steam explosion phenomena and research are briefly summarised for both in-vessel and ex-vessel melt/water interactions.

Preface

vii

If the debris contacts water varying degrees of fragmentation may occur. Whether such debris is coolable or not is a key safety issue. If the volume to surface area ratio of the debris is too large this may not be the case. If the debris becomes too finely fragmented the potential for steam explosions increases. During an in-vessel melt progression phase, debris may interact with the vessel and the vessel internals. There is a particular interest in the way instrument penetrations may be attacked in the lower vessel head of a PWR, providing a mechanism for vessel failure. The release and transport of fission products in the primary circuit is briefly covered. While thermal-hydraulics, fluids and thermal transfer are the main concerns of the book , some discussion of fission product issues both in the primary circuit and containment is included for completeness. Material behaviour in the cavity is discussed. The extent to which debris might be swept out of a cavity in a high pressure release from the vessel determines the threat of direct containment heating and potential for early failure of the containment. Core/concrete interactions provide an important source mechanism for fission product release to the containment. Fission products would be carried along with copious production of gases e.g. steam, carbon dioxide and aerosols. These gases may also be reduced by metallic corium resulting in the production of further flammable hydrogen (and carbon monoxide). Containment thermal-hydraulics and the transport of fission products within the containment are the subjects reviewed in the final chapters concerned with modelling. New phenomena associated with the operation of engineered safety features such as ice condensers and sprays are also addressed. A chapter is given to thermo-physical models. Densities, thermal conductivities, specific heats etc. are required for the important materials e.g. Zircaloy, fuel, stainless steel and concrete structures. These properties are also required for the material compounds e.g. oxides, eutectics and for water/steam and various gas properties. Many codes for LWR safety analysis are now being developed. Emphasis here is given to the system codes developed by the United States Regulatory Commission (USNRC} and its contractors. These codes are widely used worldwide and provide a state-of-the-art capability for primary circuit thermal-hydraulics, in-vessel core degradation and primary circuit severe accident modelling, fission product behaviour and containment thermal-hydraulics. Some representative simulations of various integral experiments are shown in a later chapter to give an overall impression of the adequacy of current predictions, compared with experimental data. The TMI 2 accident is judged to be a sufficiently important subject to merit a chapter by itself. This accident provides unique data at full plant scale on the nature of core degradation and melt phase progression. -

Vlll

Preface

Towards the end of the book selected code calculations of certain representative accident sequences are briefly presented. Plant studies are the ultimate goal of the model development and the experimental research. Accident management modelling requirements are briefly discussed. These are providing new challenges to the system codes at the present time. Boundary conditions may be complex: there are also difficulties in modelling certain new phenomena e.g. flooding of a degraded core. The final chapter is concerned with Advanced LWRs. These put greater emphasis on passi ve safety systems and new phenomena are present. New experimental programmes are planned and underway and further systems code development will be influenced by Advanced LWR safety concerns.

IX

ACKNOWLEDGEM ENTS This book includes reference to the work of many scientists and engineers, worldwide. The author would like to acknowledge this material and in particular technical work carried out by other colleagues in AEA Technology. Individual contributions are referenced where appropriate: apologies are given for any significant omissions. Particular names to mention include Dr A T D Butland, currently Division Director, Consultancy Services, AEA Technology, who led a major review of UK modelling re­ quirements in the 1 980s, Dr S R Kinnersly, Reactor Safety Studies Department Manager, Dr J C Birchley, Dr W M Bryce, Dr P Ellicott, Dr T J Haste, Dr B J Holmes, Dr G R Kimber, Mr A J Lyons, Dr B W Morris, Mr R O'Mahoney, Dr P N Smith, Mr D W Sweet, Dr B D rurland and Dr D A Williams, also of Consultancy Services, AEA Technology. The author would like to thank other colleagues in AEA Technology: Mr P L Holden, Senior Manager, for granting permission to write this book, Mr B Hallett, Graphics Department Manager and Mr D M Burden for technical production advice. Particular gratitude is expressed to Mrs J M Ramsden, for carrying out the large proportion of the word -processing and also to the late Mr J Hale for the early work on this book. Thanks are also due to Mrs L Dade for preparation of the illustrations. Appreciation is expressed to Dr G Jones of Technical Communications for accepting the original proposal to write the book and to Mrs van der Heide of Elsevier for advice on the camera-ready production. A final acknowledgement is to my family, my wife Marilyn and children Mark, James, Hannah and Joseph, for their support and encouragement

xi

C ONTENTS

PREFACE ACKNOWLEDGEMENTS

1

INTRODUCTION

1.1 1 .2 1 .3 1 .4 1 .5 1 .6 2

BASIC FEATURES OF LIGHT WATER REACTORS

2. 1 2.2

2.3

3

Nuclear Power Light Water Reactors Prominant Safety Studies Major Technical Areas Nuclear Safety Research Objectives Modelling Requirements and Capabilities

Introduction Pressurised Water Reactors 2.2. 1 Reactor Coolant System 2.2.2 Reactor Core and Pressure Vessel 2.2.3 Containment 2.2.4 Safety Systems Boiling Water Reactors 2.3 . 1 BWR Steam Cycle 2.3.2 Reactor Core and Vessel 2.3.3 Containment 2.3.4 Safety Systems

ACCIDENT SCENARIOS

3.1 3.2 3.3

3.4

3.5

Introduction Accident Classification Intact Circuit Faults 3.3 . 1 Pressurised Water Reactors 3.3.2 Boiling Water Reactors Loss of Coolant Accidents (LOCAs) 3.4. 1 Pressurised Water Reactors 3.4 . 1 . 1 Large Break LOCAs 3.4. 1 .2 Small Break LOCAs 3.4.2 Boiling Water Reactors 3.4.2. 1 Large Break LOCAs 3.4.2.2 Small Break LOCAs Severe Accidents 3.5 . 1 Pressurised Water Reactors 3.5. 1 . 1 Intact Circuit 3.5.1 .2 LOCAs 3.5. 1 .3 Containment Bypass Sequences 3.5.2 Boiling Water Reactors

(v) (ix) 1 1 2 2 3 4 5 9 9 9 9 11 13 13 14 16 16 16 17 21 21 21 21 21 23 24 24 24 24 25 25 26 26 26 26 28 30 31

Contents

XU

3.5 .2. 1 Intact Circuit Faults 3.5.2.2 LOCAs 4

INTEGRAL EXPERIMENTS

4. 1 4.2 4.3 4.4 4.5 4.6 4.7 4.8 4.9 4.10 4.1 1 4.12 S

Introduction Thermal Hydraulics Fuel and Cladding Behaviour Materials and Structural Behaviour Core Melt Programmes Natural Circulation Fission Product Release and Transport Debris Beds Melt/Water Interactions Core Concrete Interaction High Pressure Melt Ejection and Direct Containment Heating Experiments Containment Phenomena

THERMAL HYDRAULIC MODELS

5.1 5.2 5.3 5.4

5 .5

5.6

Introduction Code Categories Detailed Physical Phenomena Classification of Two-Phase Flow Models Homogeneous Flow (a) Homogeneous Model with Slip (b) Drift Flux Model (c) Disequilibrium Two-Phase Fluid Models (d) Disequilibrium Fluid Models including Incondensable Gases (e) Derivation of Two-Phase Flow Equations 5.5 . 1 Field Equations Additional Field Equations 5.5.2 Constitutive Relations 5.6. 1 Flow Regime Maps Vertical Flow Maps (a) Horizontal Flow Maps (b) Other Flow Maps (c) Interphase Drag 5.6.2 5.6.3 Wall Friction Wall Heat Transfer 5.6.4 Interphase Mass Transfer 5.6.5 5.6.6 Reflood Heat Transfer Turbulence Modelling 5.6.7 Choked Flow 5.6.8 Subcooled Flow (a) Two-Phase Flow (b) Single-Phase Vapour (c) Transition Regime (d)

31 31 33 33 33 35 36 36 39 39 40 41 41 42 43 47 47 47 48 49 49 49 49 49 49 50 51 53 54 54 54 54 54 54 56 56 57 57 58

60 60 60 60 60

Contents 6

COMPONENT HEAT TRANSFER

6.1 6.2 6.3 6.4

6.5 7

MECHANICAL MODELS

7.1 7.2 7.3 7.4 7.5 7.6 7.7 8

Introduction Clad Ballooning Models Mechanical Modelling of Structures Design Characteristics Modelling Criteria Detailed Models

MATERIALS INTERACTIONS

8. 1 8.2

8.3

8.4

8.5

8.6 8.7 9

Introduction Heat Conduction in Structures Heat Conduction Under Reflood Conditions Fuel and Cladding Energy Transfer Heat Generation in the Fuel 6.4. 1 6.4.2 Gap Conductance 6.4.3 Thennophysical Properties Cladding Oxidation Heating 6.4.4 6.4.5 Fuel Rod Boundary Conditions Radiation Models

Introduction Zircaloy Oxidation 8.2. 1 Phenomena 8.2.2 Kinetics Fuel and Cladding Interactions 8.3 . 1 Zircaloy/Uranium Dioxide Interactions 8.3.2 Molten Zircaloy/Solid Zirconium Dioxide Dissolution Control Rod Materials PWR Absorber Rods 8.4.1 8.4.2 BWR Absorber Rods Other Zircaloy Interactions 8.5. 1 Inconel Grids 8.5.2 Burnable Poison Rods Steel and Structures Oxidation Oxidation of Fuel

MELT PROGRESSION MODELS

9. 1 9.2 9.3 9.4

Introduction Parametric Core Meltdown Models Mechanistic Models: General Approaches Liquefaction of the Intact Core and the Fonnation of B lockages PWR Control Rods 9.4 . 1 9.4.2 BWR Control Rods 9.4.3 Fuel Rods

XU1

63 63 63 65 67 67 68 70 70 70 70 73 73 74 75 76 77 78 78 83 83 83 83 85 86 86 87 88 88 88 90 90 90 90 91 95 95 95 97 99 99 99 100

xiv

Contents

9.5 9.6 9.7 9.8 10

STEAM EXPLOSIONS

10. 1 10.2 10.3 10.4 10.5 10.6 10.7 10.8 1 0.9

11

Introduction Phases of a Steam Explosion Boundary Conditions Mixing of Molten Debris with Water Triggering Mechanical Energy from Steam Explosions Damage Potential Experiments Calculational Models 10.9. 1 Heat Transfer from Particles Following a Steam Explosion 10.9.2 Relief and Safety Valves

DEBRIS COOLABILITY MODELS

1 1.1 1 1 .2 1 1 .3 1 1 .4

1 1 .5

1 1 .6

12

Core Fragmentation and Blockage Formation Molten Pools B lockage Melt Release Models Melt Relocation from Blockages and Melt Progression Paths

Introduction Phenomena Experimental Programmes Bottom Reflood Models 1 1 .4.1 Lumped Parameter Models 1 1 .4.2 Quenching Models 1 1 .4.3 Models for Fluidisation 1 1 .4.4 More Sophisticated Models Top Reflood Models 1 1 .5.1 Critical Heat Flux Models 1 1 .5.2 Upper Bed Quenching Rate Limitation Models 1 1 .5.3 Quench Front Limitation Models Status of Modelling 1 1 .6. 1 Bottom Reflood 1 1 .6.2 Top Reflood

DEBRIS INTERACTIONS WITHIN THE VESSEL

12. 1 12.2 12.3 12.4 12.5 12.6 12.7 12.8 12.9

In trod uction Release of Material from the Core Region Debris Interactions with the Lower Vessel Internals Debris Interactions with Water Debris Interactions with the Vessel Vessel Failure Potential Vessel Response at Elevated Temperatures Behaviour of Penetrations Debris Behaviour at Vessel Failure

101 103 103 104 107 107 1 09 111 1 12 1 15 1 16 1 19 120 121 121 122 127 127 128 128 131 131 131 132 132 1 32 1 34 1 34 135 1 36 136 1 37 14 1 14 1 142 142 144 145 146 146 147 149

Contents 13

FISSION PRODUCTS 13.1 Introduction 13.2 13.3

13.4 13.5

13.6 13.7

13.8

14

Convective Transport Chemistry 13.5.1 Equilibrium Chemistry 13.5.2 Non-Equilibrium Chemistry Fission Product Effects on Decay Heating Aerosols 13.7.1 Deposition Mechanisms 13.7.2 Agglomeration Mechanisms Fission Product Release from Debris

155 157 158 159 161 162 162 162 163 163 164 165

CAVITY PHENOMENA

169 169

14.3 14.4 14.5 14.6 14.7 14.8 14.9

Introduction Debris Formation and Mixing Hydrogen Production Steam Explosions in the Cavity Debris Transport within the Cavity Debris Coolability Uncoolable Debris Uncertainties in Debris Behaviour in the Cavity Summary

169 170 171 173 174 175 176 178

CORE DEBRIS INTERACTIONS WITH CONCRETE 15.1 Phenomenology

181 181

15.2

181

Relevant System Components Heat Generation 15.4 Pool Heat Transfer 15.5 Surface Heat Transfer 15.6 Heat Transfer Between the Melt and the Concrete 15.7 Crust Behaviour 15.8 Concrete Ablation 15.9 Chemical Interactions 15.10 Mass and Energy Transfer 15.11 Energy Conservation 15.12 Bubble Behaviour 15.3

16

153 153 154

14.1 14.2

15

Stages of Severe Accident Fission Product Release Fission Product Behaviour in the Fuel 13.3.1 Intra-Granular Processes 13.3.2 Extra-granular Processes

xv

183 183 185 185 186 187 188 189 190 190

AEROSOL PRODUCTION FROM CORE-CONCRETE INTERACTIONS

193

16.1 16.2

193

16.3 16.4

Introduction Important Mechanisms Aerosol Production Models Material Entrainment

193 194 194

Contents

XVI

16.5 16.6 16.7 16.8 16.9 16.10 16.11 16.12 16.13

17

Aerosol Production via Bubble Collapse Vaporisation Condensation Aerosol Particle Size Gas Composition at Equilibrium Over the Melt Bubble Rise Phenomena Mass Transport Departure from Equilibrium Conditions Mechanical Aerosol Production

196 196 197 198 200 200

CONTAINMENT THERMAL-HYDRAULICS

203 203 204 204 205 207 208 208 209 211 211 213

CONTAINMENT AEROSOL AND FISSION PRODUCT MODELS 18.1 Introduction

215 215

18.2

216

18.3 18.4 18.5 18.6 18.7 18.8 18.9 18.10 18.11 18.12 18.13 18.14

19

195 195

17.1

Introduction 17.2 Major Phenomena Affecting Thennal Hydraulics 17.3 Flow Between Compartments 17.4 Fundamental Thennal-Hydraulics 17.5 BWR Specific Processes 17.6 Material Properties 17.7 Gas Burning Models 17.8 Energy and Mass Transfer 17.9 Heat Conduction in the Structures 17.10 Engineered Safety Features 17.11 Summary

18

195

Size Distribution Agglomeration Aerosol Condensation and Evaporation Deposition Aerosol Sources Scrubbing Radionuclide Behaviour Modelling Decay Chains Decay Heating Models Transfer Rates in the Atmosphere Transfer in the Liquid Engineered Safety Features Containment Sprays

THERMOPHYSICAL PROPERTIES 19.1 19.2 19.3 19.4 19.5 19.6 19.7

Introduction Fuel Requirements Fuel Pin Cladding Control Rod Material Control Rod Cladding Fuel Cladding Eutectic Control Rod Eutectic

216 218 218 219 219 220 221 221 222 222 222 224 227 227 227 231 234 235 237 238

Contents 19.8 19.9 19.10 19. 1 1 19. 12

20

COMPUTER CODES

20. 1 20.2

20.3

21

Introduction Heat Transfer and Hydraulics 20.2. 1 Thennal-Hydraulics System Codes 20.2.2 Thennal-Hydraulic Reactor Coolant System Codes 20.2.3 Fuel and Cladding Behaviour 20.2.4 Containment Severe Accident Codes 20.3 . 1 Integrated Codes 20.3.2 Mechanistic System Codes 20.3.3 Separate Effects Codes

CODE VALIDATION

21.1 2 1 .2

2 1 .3

22

Structural Material Core-Concrete Material Water/S team Properties Non-condensable Gases Other Gases

Introduction Thennal-Hydraulics Experiments 2 1 .2. 1 Intact Circuit Faults 2 1 .2.2 Loss of Coolant Accidents Severe Accidents 2 1 .3 . 1 Early Phase Core Degradation 2 1 .3.2 Natural Circulation 2 1 .3.3 Fission Product Transport in the Reactor Coolant System 2 1 .3.4 Core - Concrete Interaction 2 1 .3 .5 Containment Phenomena

XVll

239 240 24 1 243 243 249 249 249 249 249 252 252 253 253 255 258 265 265 265 266 268 273 273 276 277 278 280

THE ACCIDENT AT THREE MILE ISLAND UNIT 2: IMPLICATIONS FOR MODEL DEVELOPMENT

22. 1 22.2 22.3 22.4 22.5 22.6 22.7 22.8 22.9 22. 10 22. 1 1

Introduction Accident Scenario Thennal-Hydraulic System Analysis Upper Vessel Structural Temperatures Final State of the Plant Melt Progression Debris in the Lower head State of the Lower Head Lower Head Failure Analysis Fission Product Release and Transport in TMI-2 Implications for Modelling and Conclusions arising from the TMI-2 Accident

283 283 284 285 287 288 290 29 1 292 293 294 294

xviii

Contents

23

PLANT STUDIES 23 . 1 Introduction 23.2 Intact Circuit Faults 23 .2. 1 Loss of Feed Water 23 .2.2 Steam Line Break 23 .2.3 Reactivity Transients 23.2.4 Anticipated Transient Without Scram (ATWS) 23 .3 Loss of Coolant Accidents 23 .3 . 1 Steam Generation Tube Rupture (SGlR) 23 .3.2 Large Break LOCA 23.4 Core Degradation and Thermal Hydraulics 23.4 . 1 Station Blackout 23 .4.2 Small Break Loss of Cooling Severe Accident 23.4 .3 Containment Bypass: Interfacing LOCA 23 .4.4 Shutdown Accidents 23.5 Fission Product Transport under Severe Accident Conditions 23.5. 1 Station Blackout 23.6 Containment Behaviour Under Severe Accident Conditions 23 .6. 1 Station Blackout

297 297 298 298 300 300 301 301 301 303 304 305 308 309 309 3 10 3 10 3 12 3 12

24

ACCIDENT MANAGEMENT 24. 1 Introduction 24.2 Preventative Accident Management 24.2. 1 Total Loss of Feed Water 24.3 Computer Code Requirements 24.4 Code Assessments 24.5 Summary - Preventative Accident Management 24.6 Mitigative Accident Management 24.7 Water Addition to a Degraded Core 24.8 Primary Circuit Integrity 24.9 Late Phase Melt Progression 24. 10 Melt-Water Interactions in the Reactor Vessel 24. 1 1 Steam Explosions 24. 1 2 Failure of the Reactor Vessel 24. 1 3 Threat to the Containment 24. 14 Ex-Vessel Melt/Water Interactions 24. 1 5 Molten Core/Concrete Interactions 24. 16 Summary - Mitigative Accident Management

3 17 317 318 3 19 320 321 322 322 324 324 325 325 326 326 326 327 327 327

2S

ADVANCED REACTORS 25 . 1 Introduction 25.2 Design Concepts 25.2. 1 Evolutionary Designs 25.2.2 Passi ve Designs 25.2.3 Innovative Designs 25.3 New Phenomena

33 1 33 1 33 1 33 1 333 335 336

Contents

25.4 25.5 25.6

Experiments Accident Assessments Computer Code Requirements

INDEX

XIX

338 340 342 345

1

Chapter 1 INTRODUCTION

1.1

Nuclear Power

During the Second World War it became apparent that the nuclear fission of uranium and plutonium could provide a means for the production of both atomic bombs and for peaceful uses such as the production of electricity. Noclear fission involves the splitting of a heavy atom e.g. uranium-235 or plutonium-239 following collision with a neutron, arompanied by a rapid large release of energy and further neutrons. One kilogram ofuranium-235 totally fissioned produces about 8x 1 013 jollies: the energy JXltential of plutonium is several orders of magnitude higher. The challenge for producing continuous energy release is to engineer a controlled self-sustaining chain reaction. This was frrstachieved by Enrico Fermi in 1 942 in an experimental reactor at the University of Chicago. Civilian nuclear power generation research and development programmes started in the 1 950s in the technologically advanced countries including USA, the former USSR, UK and Western Europe. To promulgate the peaceful uses of nuclear energy, a conference was held in Geneva in 1 955. Wide ranges of designs were considered involving many possible fuels, moderators, control materials, coolants and structural materials. From the early days the potential hazards of radionuclide release and the consequences of failure to control the chain reaction were realised. The nuclear industry invested heavily in safety related work to ensure the safety of its workers and the general public. Of all the designs, Light Water Reactors (LWRs) are the most widely used in the world for electricity generation. Of these, Pressurised Water Reactors (PWRs) and Boiling Water Reactors (BWRs) are the most common in the USA, Western Europe and the Pacific Rim . Another type of pressurised water reactor, the VVER, i s being used i n Russia and Eastern Europe. The PWR owes its origins from submarine reactor designs. The first commercial scale civilian reactor was commissioned at Shippingport, Pennsylvania in 1 957. This was small (60MWe) by present day standards: a large modem PWR generates typically -12001300MWe. The first commercial-scale BWR was commissioned at Dresden, Illinois in 1 960. This was also small {l80MWe). Modem BWRs have much greater capacity than the early plants.

Introduction

2

The first VVER was constructed in Novoronezh (210MWe) and went into operation in 1964. The reactor concept of VVERs is similar to PWRs in the West but there are significant and extensive differences in safety related features between VVERs and PWRs. VVERs have been designed in two principal sizes 440 MWe and 1000 MWe.

1.2

Light Water Reactors

In Light Water Reactors (LWRs) ordinary water serves as both coolant and moderator. The coolant passes through the reactor core, extracts the energy to ultimately drive the turbine and produce the electricity. All Western LWRs use fuel rods in an open lattice framework supported by grids. PWRs nonnally incorporate a two-coolant system. The primary circuit transfers heat from the core to steam generators. The number of circuits varies between designs, as do the steam generators. The secondary circuit produces the steam for the turbine. The primary side pressure is high enough to prevent boiling. BWRs commonly used today nonnally utilise a single coolant system. Steam is produced as the coolant passes through the core and is passed directly to the turbine. A disadvantage with this direct cycle system is that the turbine will tend to become contaminated since the steam passing through will be radioactive. Since PWRs and BWRs have similar materials and coolants they do exhibit some common safety characteristics. Differences result due to differences in design within either category. There are major differences in design between VVERs and Western LWRs. VVERs have horizontal steam generators: Western LWRs have vertical steam generators. However, like the PWRs, VVER- l000s have an open lattice fuel assembly in a 4 loop system. VVER440s have a triangular fuel lattice with 6 primary coolant loops. VVER- l000s also have improved safety systems compared with VVER-44 Os. The discussion in this book is primarily concerned with advanced modelling techniques for Western LWRs (i.e. PWR and BWR). Unless otherwise stated LWR refers to 'Modem Western LWR ' and does not include VVER. The development of advanced safety analysis techniques for VVERs is beginning but is still in its infancy. It is likely to expand in the next decade.

1.3

Prominant Safety Studies

One of the first studies on the consequences of major accidents in nuclear power plants (WASH -740 [ 1 . 1]) was carried out by Brookhaven National Laboratory in 1957 for the then United States Atomic Energy Commission (USAEC). The aim was to estimate the consequences of fission product release from an intennediate size LWR of the day. The first major study on reactor safety (WASH- I400 [1 .2]), also known as the Rasmussen study, was

Introduction

3

carried out in 1975 by the USAEC successor, the US Nuclear Regulatory Commission (USNRC). This reactor safety study (RSS) was a wide ranging review of the consequences of fission product releases from many accident sequences. This �as followed by the German Risk Safety Study [ 1 .3] in 1980, later updated. Following the major research effort during the 1980s a new reactor safety study was initiated by the USNRC (NUREG 1 150 [l .4]). The aim was to review the WASH- 1400 conclusions taking account of improved understanding and better calculational tools than were available previously and to extend the number of plants considered. In the UK the main focus ofreactor safety research in the last 10-20 years has been to provide support for the Sizewell B PWR. Major safety assessments were carried out both for the Pre­ Construction Safety Report (PCSR) in the early 1980s and later for the Pre-Operational Safety Report (POSR) in the early 1 990s.

1.4

Major Technical Areas

During the 1950s the main interests in research were concentrated around the core design. Initially modelling techniques relied on analytical solutions derived for idealised geometries. Simple experiments were carried out to provide data for new materials and designs. The drive towards optimisation of the cores for commercial plants required more powerful tools for analysis. The core geometries were becoming more complex. Many and varied materials were being considered. Computers started to become available during this decade. Computer models for the cores were therefore developed and reactor physics codes were produced. These were improved as new data became available. There was considerable emphasis during this time on the validation of these reactor physics codes to provide the tools required by core designers and safety analysts. Fuel performance issues attracted attention in the quest to produce fuels with increased power density and more prolonged burn-up. Much work was carr ied out in Canada and Norway. Experiments were conducted to investigate fuel behaviour both under normal operation and under fault conditions. The main issues were concerned with the behaviour of the fuel and cladding under irradiation. Irradiation causes pellet cracking and re-sintering. Cladding may also contract down on to the fuel due to the high external pressure in PWRs leading to failure and consequent fission product release into the primary circuit In parallel with improving fuel performance the thermal-hydraulics of the core needed to be addressed to determine safety margins. Heat transfer between the cladding and water is good but the critical heat flux at which the surface of the pin drys out, leading to rise in surface temperature and possible cladding failure, sets limits on fuel pin performance. With the design of reactors with a large thermal output a major issue in LWR safety research became the need to ensure that the systems response to Loss of Cooling Accidents (LOCAs) were adequate to ensure that core temperatures do not rise above an acceptable level.

4

Introduction

Emergency Core Cooling Systems (BCCS) were installed to protect the plant against large pipe breaks involving complete depressurisation (e.g. to the containment pressure of a few atmospheres). LOCAs attracted major modelling development activity. The structural integrity of the reactor vessel and primary circuit was another important safety issue in PWRs. A particular concern was whether the vessel could stand the thermal shock from ECCS operation in the event of a large break LOCA. Much materials property research was carried out into the fracture and corrosion characteristics of the materials. Thermal-hydraulic system codes were therefore developed in the US and elsewhere by both utilities and licensing bodies to be used to demonstrate the plants are safe under a very wide range of intact circuit and LOCA conditions. The system response under many accident conditions is very complicated and this necessitated the development of sophisticated computer models over many years. The aim of the later more advanced codes was to include mechanistic or physics based modelling as far as possible but even present generation codes incorporate various assumptions and empirical data. The developmentand validation needs of these codes spawned many experimental programmes in both the US and Western Europe. These experiments were in several catagories: 'separate effects' and 'integral'. Separate-effects experiments aim to address single phenomenon or single system component effects: integral experiments are concerned with overall system response. The accident in Unit 2 of the Three Mile Island PWR (TMI-2) [l.S] on 28 March, 1 979 had important ramifications for reactor safety research. Up to that time much of the research had concentrated on accidents within the 'design basis ' defined later and in particular on the double ended guillotine fracture of a main coolant pipe. After the TMI-2 accident it could no longer be claimed that severe accidents could not occur. The accident also focused more attention on the need for training of operators and on accident management. It took many years for the full extent of the damage to the TMJ-2 reactor to be realised. A large amount of work to understand severe accident phenomena has been carried out during this time.

1.5

Nuclear Safety Research Objectives

The main aim of reactor safety is to ensure that the release of radionuclides is prevented. During normal operation the releases are kept as low as reasonably practical: this is known as the ALARP principle. There are various barriers to radionuclide release. 1 . Fuel pellets. Modem LWRs use ceramic fuels. Ceramic fuels have been shown to have good fission product retention properties compared with metallic fuels, which tended to be used in some early reactor designs. 2. Fuel pin cladding. The fuel rods are clad is a ductile material, Zircaloy, an alloy of zirconium. (Different designs incorporated slightly different compositions of alloy).

Introduction

5

3. Reactor Coolant Circuit. This may be single, 2 component or more. 4. Containment. The strength of containments is design dependent. The tendency in modem LWR designs is to incorporate very strong containments. In most designs however, the coolant circuit is not entirely contained within the containment which means that failure of certain plant components in the reactor coolant system can result in a containment bypass release.

The main threat to breaching these barriers arises if the core is allowed to uncover resulting in rise of the core and reactor coolant circuit temperatures due to decay and later Zircaloy oxidation heat The situation is aggravated further if large pressure gradients exist across vulnerable components. Broadly the higher the temperatures and pressures, the more likely it is that one or more of these boundaries could be breached. A note of comfort however is that, despite very high temperatures attained during the TMI-2 accident, the primary circuit remained intact. For public acceptability it is likely to be necessary to ensure that the operation of nuclear power plants does not increase the risk to society (compared with other risks) and this is a principal goal of nuclear safety research [ 1 .6]. 1.6

Modelling Requirements and Capabilities

During the past 10-20 years there has been considerable activity in the development of codes for LWR safety analysis. This book aims to provide a concise review of the major activities and a broad overview of current modelling capabilities. The book commences with a brief description of the various types of Western LWR (pWR, BWR) currently in operation. Potential accident conditions are classed in various categories, broadly intact circuit faults and loss of coolant accidents. These can be of various degrees of severity. Some representative ' design basis' and severe accident' scenarios are described to provide an indication of the scope of important phenomena that need to be modelled. Safety systems are installed to protect the plant against a wide range of accident conditions. Such accidents are termed to be within the 'design basis' . Severe Accidents are more extreme accidents that are beyond the design basis. Experiments have been carried out in most areas to provide data for model development (Separate-Effects Experiments) and for code validation (Integral Experiments). The major integral programmes addressing thermal-hydraulics, in-vessel core melt and containment related issues are summarised in the book . The major portion of the work is concerned with advanced modelling techniques in the following areas: thermal-hydraulics -

structural heat transfer

6

Introduction

mechanical properties materials interactions melt progression steam explosions debris interactions in the lower vessel fission product transport cavity phenomena core/concrete interactions core/concrete fission product release containment thennal-hydraulics containment vapours/aerosols thermophysical properties. Mathematical models for these phenomena have been integrated in advanced computer codes and some of the major codes employing state-of-the-art capabilities are reviewed. While order of magnitude predictions can often be made using quite simple calculations, the codes enable the coupling of the various phenomena to be investigated and more reliable results to be obtained. Such a capability is particularly important in severe accident analysis where the coupling effects between the various phenomena are very important. Some representative code predictions are compared with integral experiments, to demonstrate the perfonnance of the models and codes discussed. The comparisons cover thennal­ hydraulics and melt progression, and containment related areas. A complete chapter is devoted to TMI-2 which provides unique full-scale data on melt progression and system response under severe accident conditions. The core uncovered, principally due to a sequence of operator errors, and post accident examination of the debris showed that very high temperatures were attained. Two chapters are concerned with advanced code predictions for some of the accident scenarios described earlier in the book and with accident managemenl A capability to carry out safety analysis for a complete set of accident scenarios including possible accident management actions is the ultimate objective of the modeller and code developer.

Introduction

7

Advanced reactor designs are under development by various vendors (Westinghouse, GE, ABB-CE and others). These designs fall into different categories. Some of the new designs are basically similar to curre n t plant but with certain weaknesses improved. From a modelling perspective the phenomena are similar between the new and old designs. Other designs include more radical systems, different phenomena are important, and new code modelling and data are required. These designs are providing the focus for current modelling development

REFERENCES

1.1

1.2

1.3 1 .4 1.5 1.6

US Atomic Energy Commission "Theoretical Possibilities and Consequences of Major Accidents in Large Nuclear Power Plants" USAEC Report WASH-740, March 1957. US Nuclear Regulatory Commission "Reactor Safety Study, an Assessment of Accident Risks in US Commercial Nuclear Power Plants" USAEC Report W ASH1400, October 1 975. Federal Minister for Research and Technology "The German Risk S tudy Nuclear Power Plants" Verlag TUB Rheinland, 1980. US Nuclear Regulatory Commission, Reactor Risk Reference Document, NUREG- 1 1 50, Washington DC, USA, 1987. Three Mile Island Unit 2: Nuclear Technology "A Journal of the American Nuclear Society", Vol 87, No 1 NUTYBB 87( 1) 1 -334 (1989), August 1989. D Hicks, Nuclear Safety Research on Thennal Reactors, Paper presented at the 5th Vernon Clancey Memorial Lecture given at City University, London, UK on 17 March 1 993, published by Institution of Chemical Engineers, 1 993.

9

Chapter 2 BASIC FEATURES OF LIGHT WATER REACTORS

2.1

Introduction

The aim of this chapter is to describe the main feanu-es of Western style LWRs, especiall y those that are relevant to reactor safety considerations. There are two types of LWR cwrently in operntion, Presswised Water Reactors (PWRs) and Boiling Water Reactors (BWRs).

2.2

Pressurised Water Reactors

Pressurised Water Reactors are characterised by a high pressure coolant system in order to prevent boiling. They generally have smaller cores than BWRs and higher power density. PWRs have been manufactured in the US, various countries in Western Europe (notably France and Germany), Japan and elsewhere. The description below is for a typical large modem four loop Westinghouse PWR. Different manufacturers have produced some differences in design, but the basic principles are similar: only the more significant differences in the various designs are mentioned below.

A large modem PWR station includes a 3425 MWt Nuclear Steam Supply System which supplies steam to turbine generators with a combined gross output of 1 1 50- 1250 MWe. The system comprises the reactor itself, the reactor coolant system together with its various components (e.g. pressuriser and steam generators) and various auxiliary and safety systems. Representative PWR Design Data are given in Table 2. 1 [2. 1 ] . 2.2.1 Reactor Coolant System

The reactor coolant system (RCS), Figure 2. 1 , consists of the reactor pressure vessel and cooling loops connecting the reactor vessel to the steam generators. The principal thermal­ hydraulic regions of the vessel are shown in Figure 2.2. Two, three and four primary cooling loops have been utilised in various designs. Water is circulated through the reactor and steam generators via connecting pipework. The coolant temperature is basically about 298°C at the inlet to the reactor core; the temperature rise across the core region being about 28°C. The primary reactor coolant water is at high pressure, 1 5.5 MPa. Heat is removed from the reactor via the steam generators; steam is produced at about 250°C. The flow rate through each coolant loop is about 3975 kg/so N.B. There are small variations in the nominal core temperature rise and flow rate across various designs. During normal operation (and fault conditions) temperature and pressure changes are experienced in the reactor coolant system. The pressure is controlled by the pressuriser,

Basic Features of Light Water Reactors

10

STEAH NOZZLE STEAH GEN ERAT OR

REACTOR COOLANT

LEG

REACTOR PRESSURE VESSEL (CONTAINS COREl

FIGURE 2.1

PWR FOUR LOOP NUCLEAR STEAM SUPPLY SYSTEM

Figure 2. 1 connected to a hot leg via a surge line. This has a volume of -5 1cum, about one eighth of the total reactor coolant system volume. The pressuriser is partly filled with water and partly with steam. Primary circuit pressure is maintained within desired limits via pressuriser heaters and sprays. The pressuriser heaters serve to increase the steam quality if it is desired to increase the pressure of the primary circuit Sprays are turned on to condense steam in order to

Basic Features of Light Water Reactors

11

decrease the pressure. Protection against over-pressurisation is provided by relief valves. Some modem designs e.g. Sizewell-B [2.2] utilise both power-operated relief valves and three passive spring loaded safety valves making use of the French Sebim valve system [2.3] . Any steam discharged through the valves i s condensed i n the pressure relief tank. This i n turn is protected against over-pressurisation by discs that can burst and release water and steam to the containment The Sizewell-B design is based on SNUPPS (Standardised N ucIear Unit Power Plant System) [2.4] with some improvements, taking advantage of the feedback of experience from the large number of operating PWRs including the lessons learned from TMI-2 [2.5] . The steam generators transfer heat from the reactor coolant water to feed water from the turbine condensate system. A Westinghouse U-tube steam generator is shown in Figure 2. 1 . It consists of three components, a bottom head carrying the primary coolant inlet and outlet nozzles, the evaporator section enclosing the U tube bundle and an upper section containing the moisture separators. The tubes are made from Inconel. Some other designs (e.g. General Electric) employ once-through steam generators [2. 1]. Once a reactor has been tripped, heat associated with the radioactive decay of fission products (referred to as decay heat) is transported away initially via the steam generators and then via a residual heat removal system (RHRS). In the latter case heat is extracted from the coolant by passing the coolant through water cooled heat exchangers. In the RHRS, pumps draw water from two of the hot legs, pass the water through the heat exchangers and return it to the cold legs through the low head safety injection system. The RHRS is operated once the coolant temperatures are sufficiently low - 1 77°C and the pressure is less than about 3.2 MPa. Such conditions are achieved at about 4 hours after shutdown. 2.2.2 Reactor Core and Pressure Vessel

The reactor core consists of 193 fuel assemblies in a square lattice arrangement, Figure 2.3. Each fuel assembly contains 264 fuel rods containing low enrichment U02 fuel clad inZircaloy tubes. Core reactivity is controlled by S ilver-Indium-Cadmium control rods inserted from above and borated water. The coolant water serves both as a moderator and as a medium for conducting the heat away from the core. The active length of the fuel rods is 3.7 m long and the cladding is 9.5 mm in outer diameter. The fuel rods are pressurised with helium and supported at intervals by grids. The 264 fuel rods are held in a 17 x 17 square array. The centre position is in-core instrumentation: the remaining positions being occupied by Zircaloy guide tubes which are used for control rod assemblies, burnable poison rods, neutron source assemblies or are simply plugged. Drawings of a modem fuel assembly can be found in [2.2] . There are a large number of vessel internal structures to support the core, direct the coolant flow etc. These can be divided into the upper core structure 60 tonnes and lower core support structures -127 tonnes, Figure 2.2.

Basic Features of Light Water Reactors

12

Each rod cluster control assembly can insert, hold and withdraw the control rods within the core. Under accident conditions control rods can fall rapidly into the core, thus tripping the reactor. The reactor vessel itself is a steel cy linder with a domed upper and lower head. The cy lindrical section is about 4.4 m internal diameter surrounded by a thick wall, a lower plain portion about 219 mm thick and an upper portion about 267 mm thick containing four inlet and outlet nozzles (of diameters about 698 mm and 736 mm respectively). These nozzles are

- - - - - - S I G N I F I C A N T M E TA L S T R U CT U R E S

CO NT R O L GUI DE

ROD

TUBES

UPPER

PLENUM

HOT

LEG

C O L D L EG

SUPPORT COLUMN

OOW N CO M E R

A CT I V E

CO R E

_--t-tlr--

CORE BYPASS

CORE

INLET REGION

L O W E R P LE N U M

FIGURE 2.2 THERMAL-HYDRAULIC REGIONS OF THE PWR VESSEL

Basic Features of Light Water Reactors

13

symmetrically arranged around the vessel and connect the vessel to the hot and cold legs of the primary circuit. The upper head is removable: the lower head is integral with the cylinder. The lower head contains penetration nozzles to admit the in-core instrumentation. 2.2.3

Containment

The reactor and reactor coolant system are surrounded by a large pre-stressed and reinforced containment building. A i m thick wall is made of reinforced concrete to provide mechanical strength: a 64 mm liner provides a leak tight membrane against escape of gases. The main purpose of the containment is to house the reactor components and also to contain any large releases of steam and water which might occur under accident conditions and also to provide a final barrier preventing release of radioactive fission products to the environment. Containment buildings vary in design. Different classes of containment design include reinforced concrete. ice-condenser and spherical steel designs. 2.2.4 Safety Systems

Pressurised water reactors have elaborate reactor protection systems to initiate the required safety functions. Control rods can fall into the core under gravity. Reactor shutdown can also be achieved by varying boron concentration in the reactor coolant. The primary circuit is protected from over-pressurisation by the safety relief valves at the top of the pressuriser. Steam generators must be supplied by feedwater to be effective in removing heat If the main feed fails then an auxiliary feed system is available to transport the heat away. Under loss of coolant accident conditions the plant is protected by an emergency core cooling system (ECCS). This system is composed of three components. the accumulators. the high head and low head injection systems. In a large LOCA. the core is uncovered very rapidly and there is a need to deliver a large quantity of water to the core region. This function is discharged by the accumulators. The accumulators discharge at about 4.5 MPa RCS pressure and are connected to the cold legs. The high head injection system consists of four pumPS. one connected to each cold leg. These deliver water from the refuelling water storage tank once the pressures fall below about 12.5 MPa. The high head pumPs ensure sufficient injection to protect against small breaks or loss of coolant system inventory where the pressure in the primary circuit remains relatively high.

The low head safety injection system is required to provide adequate cooling in the longer tenn in the event of a large LOCA. This system uses the two pumps and heat exchangers in the RHRS to deliver water from the refuelling water storage tank to the four cold legs. The low head system can deliver once pressures fall beneath about 1 .7 MPa. There are various systems provided to ensure the effectiveness of the containment following an accident. These include containment isolation systems. containment spray systems. ice condenser compartments. containment fan coolers and combustible gas control systems.

Basic Features of Light Water Reactors

14

Fuel Assembly

Reactor Vessel Control Rod Assemblies

Neutron S h ield Pad Irradiation Specimen Guide

FIGURE 2.3 PWR CORE CROSS SECTION

The containment isolation system ensures if radioactivity is released to the containment that all penetrations are closed except those necessary to the safety of the reactor. The other systems are primarily to reduce the pressure load in the event of a serious accident. Under LOCA conditions or other accident conditions whereby water and steam are released to the containment, the containment spray system exists to reduce the temperature and pressure of the containment and also to scrub fission products from the atmosphere. This system draws water from the refuelling water storage tank or from the containment sumps. Containment fan coolers can provide additional cooling under LOCA or steam line break conditions. Cooling is provided by the component cooling water system. The fan coolers promote good mixing within the containment atmosphere. Under various severe accident conditions hydrogen may be released following steam interaction with Zircaloy cladding of the fuel rods. To avoid the build-up of flammable concentrations of hydrogen a combustible gas control system exists to provide control either by mixing or recombination of hydrogen with air in the containment to form water.

2.3

Boiling Water Reactors

Boiling Water reactors are based on a direct steam cycle and therefore steam generators are not required. Steam is generated within the core region and enters the turbine directly. Thus,

Basic Features of Light Water Reactors

15

TABLE 2.1 REPRESENTATIVE PWR DESIGN DATA [2.1] Plant Manufacturer

Westinghouse

Thennal power

3425 MWt

Electric output (grosslnet)

1 1 50/1 1 00 MWe

Efficiency

33%

Core Active core (height/diameter)

3.7/3.4m

Fuel inventory

1 0 1 -t U02

Nwnber of fuel assemblies

1 93

Assembly pitch

30.4 cm

Rod pitch

l . 26 cm

Average core power density

1 04.5 kWmtre

Fuel Fuel material

U02

Enrichment

Three regions with 2. 1 , 2.6, 3 . 1 %

Pellet dimensions (diameternength)

0.82/1 .35

Assembly array

17

x

em

17

Total number of fuel rods

50,952

Cladding material

Zircaloy-4

Cladding (outer diameter/thickness)

9.5/0.6mm

Control Nwnber of control clusters

53

Nwnber of control rods per cluster

20

Absorber material

Ag-In-Cd

Absorber rod cladding

304 stainless steel

Control rod type

Cylindrical rods

Other control system (first core)

Burnable poison rods, borosilicate glass

Vessel Material

SA533, Mg-Mo-Ni steel with inner cladding

Wall thickness

2 1 .9cm

Vessel dimensions (diameterlheight

4.4/12.6m

Coolant Material

(H20)-liquid phase

System pressure

15.5 MPa

Nwnber of loops/steam generators

4

Mass flow

1 5.9 Mg/sec

Core temperature (inlet/outlet) 298/326°C

Fueling Type

Off-load, radial shuffling

Refueling sequence

1 /3 of core every 12 months

Shutdown period

30 days

Annual spent fuel discharge

30.4 t

Design fuel burnup

3 3000 MWd/t

Basic Features ofLight Water Reactors

16

from a safety point of view, any release of radioactivity is not contained within a secure primary system and therefore the problem of controlling dose rate to operators is more demanding than for other plants. Another important difference from PWRs is that control rods are inserted from the bottom and therefore the effects of gravity cannot be used for scramming the reactor. 2.3.1

BWR Steam Cycle

BWRs have been manufactured in various countries including the US, Japan, Germany and Sweden. A typical modem US General Electric Co. manufactured plant has a thermal core power of 3579 MWt with a total gross electric output of 1269 MWe. Typical design data are given in Table 2.2 [2. 1 ] . Information for Swedish plants is given in [2.6] . The pressure in the loop is maintained at about 7 MPa. Water boils as it passes through the core, at the core exit the steam fraction is about 1 0% by weight The core inlet temperature is about 277°C and the temperature rise over the core region is about 1 1 °C. Water is circulated to the bottom of the core via circulation pumps, Figure 2.4. The flow rate through the core is about 1 3240 kg! s: about one third of the total flow goes through the recirculating pumps. The steam passes to the turbine and thence to a condenser, after which it is returned to the reactor vessel. 2.3.2

Reactor Core and Vessel

The fuel rods in BWRs are similar to those in PWRs except they are thicker (outer diameter 12.3 mm) with a cladding thickness of 0.8 1 mm and active fuel length 3.8m. The cladding material is Zircaloy-2. However there are significant differences in the fuel assemblies. The more recent General Electric manufactured BWRs have assemblies each containing an 8 x 8 array of fuel rods. The assemblies are enclosed in fuel channels made of Zircaloy-4. A typical large BWR incorporates 748 fuel assemblies and 1 77 control rods. Four fuel bundles surround a control rod (or control blade) at the centre. Each control blade contains compacted boron carbide in stainless steel tubes. The core power densities in BWRs are typically about half those of a PWR and occupy a larger volume. The BWR vessel is larger than a PWR vessel because of this and also because it has to contain other large additional pieces of equipment, e.g. the jet pump assemblies and steam separators and driers. A typical vessel is over 22m high and about 6.5m in diameter. However the wall thicknesses are somewhat less than for a PWR since the reactor coolant system pressure is about half that of a PWR. 2.3.3

Containment

Containment design in BWRs has evolved through a number of stages. In the US these have been designated Mark I, II and III.

Basic Features of Light Water Reactors

17

Mark I containments were also known as inverted light bulb containments. The vessel is enclosed in a dry well which is connected to a large torus (the wetwell) and somewhat less than half filled with water. This water reservoir acts as a heat sink therefore limiting pressure increase. BWR containments are based on a pressure suppression principle.

The Mark II containment was similar to that of the Mark I except that the inverted light bulb was replaced by the frustwn of a cone on top of a cylinder and a floor separated the drywell from the wetwell. The Mark III design consists of a concrete dry well, inside a steel containment structure. The pressure suppression pool is an annular channel between the drywell and the containment. The steel containment encompasses all the equipment of the reactor building. It is designed to withstand the temperature and pressure of a large loss of coolant accident. 2.3.4

Safety Systems

As for PWRs. BWRs have elaborate protection systems to ensure the core remains cooled at all times and that radioacti ve release to the environment is minimal. Control rods are inserted

Wate r poo l

Automatic D e p ressu risation System

Steam Feedwater

r+--H>t:r---

Condensate S u p p ly

FIGURE 2.4 BWR STEAM CYCLE WITH EMERGENCY CORE COOLING SYSTEMS

Basic Features of Light Water Reactors

18

TABLE 2.2 REPRESENTATIVE BWR DESIGN DATA [2.1] Manufacturer Core thennal power Electric output (net) Plant efficiency Active core (height/diameter) Fuel inventory Number of fuel assemblies Assembly pitch Rod pitch Average power density Fuel material Enrichment Pellet dimensions (diameterlheight) Assembly array Total number of fuel rods Cladding material Cladding (outer diameter/thickness) Number of control rods Materi al Other control systems Material Wall thickness Vessel dimensions (height/inner diameter) Vessel weight (including head) Material Pressure Number of recirculation loops Core coolant flow Core coolant temperature (outlet/inlet) Feedwater flow rate Feedwater temperature Average coolant exit quality Type Refueling sequence Shutdown for refueling Annual spent fuel discharge Design fuel burnup

Plant

General Electric 3579 MWt 1269/1233 MWe 33.5%

Core

3 .76/4.65m 1 38-t U02 748 1 5.2 cm 1 .63 cm 56kW/litre

Fuel

U02 Average 2.8 mU (initial core 1 .77-2. 1%) 1 .06/1 .Ocm 8 x 8 with fuel channel around fuel rods 46,376 Zircaloy-2 12.5,u.86mm

Control

1 77 Boron carbide (B4C) "Crucifonn" blades Use of burnable poison

Vessel

SA533 (or 533B) manganese molybdenum nickel steel, inner layer of cladding l/8in of austenitic stainless steel 16.4 cm 2 1.6/6.Om 885t

Coolant

(H20)-two phase 7 MPa 2 1 3.2 Mg/hr 288{277°C 1 .94Mg/sec 2 16°C 14.7% steam by weight

Fueling

Off-load, radial shuffling 1/3 of core every 18 months or 1/4 of core every 1 2 months 60 days 32 t/yr 28,400 MWd/t at equilibrium

Basic Features of Light Water Reactors

19

by hydraulic systems or motored into the core. Boron injection systems are available in the event of failure of the control rods. Pressure relief systems are available to safeguard the reactor from over-pressure. An auxiliary feed water system exists if the nonnal feed water system becomes unavailable. There are a number of systems to safeguard the core under loss of coolant accident conditions. There are both high and low pressure core spray systems and also a low pressure coolant injection system, Figure 2.4. The pressure suppression pool is designed to condense the steam under loss of coolant accident conditions. Spray systems are incorporated to cool the pool . Spray systems also exist in the drywell in some designs. REFERENCES

2.1 2.2 2.3 2.4 2.5 2.6

F J Rahn, A G Adamantiades, J E Kenton, C Braun "A Guide to Nuclear Power Technology, Wiley and Sons, 1984. B V George, J A Board "The Sizewell-B Design Nuclear Energy, 1987, 26, No 3, June, pp 1 33- 148. G Olivon et al, Improving French PWR Overpressure Protection with Sebim Valves, Nuclear Energy International, 1984, 29 May, 40-43. W A Petrick, SNUPPS The Multiple Utility Standardisation Project, Nuclear Energy International, 1 975, 20 November, 939-937. J Kirk, J R Harrison, The Approach to Safety for Sizewell B, Nuclear Energy, 1987, 26, No 3, June, 161- 174. B Pershagen, Light Water Reactor Safety, Pergamon Press, 1989.

21

Chapter 3 ACCIDENT S C ENARIOS

3.1

Introduction

The aim o f this chapter i s to describe briefly some o f the classes o f accident scenarios that are typically considered in reactor safety analysis. In order to assess whether a plant is completely safe, a wide spectrum of accident conditions need to be studied. These range from minor faults which might be expected to occur during plant life time through to improbable events, not expected, but catered for in the design (Design Basis Accidents), and extremely improbable events not catered for in the design (Beyond Design Basis or Severe Accidents), for which mitigative action procedures need to be developed.

3.2

Accident Classification

Events are usually classified into five categories, as shown in Table 3 . 1 . These events fall broadly into two classes: Intact circuit faults or transients. Loss of coolant accidents.

Of the events in Table 3.1, the events in category 1 are not relevant from the viewpoint of safety. Events in Categories 2-4 are those included for in the design: events in category 5 are outside the design basis and these are referred to as severe accidents. Examples of particular events in various categories are shown in Table 3.2. A good summary of events is given in [3. 1] and [3.2] .

3.3

Intact Circuit Faults

Some examples of intact circuit faults are given below. 3.3.1 Pressurised Water Reactors

Typical PWR intact circuit faults considered in reactor safety analysis include: Flow reduction; this might be due to pump failure, requiring corresponding reduction in reactor power in order to avoid critical heat flux.

Accident Scenarios

22

Reactivity control malfunction; such transients could result from uncontrolled withdrawal of control rods at power, resulting in a potential mismatch between the power and flow. The reactor is scrammed. A worse transient might involve control rod ejection, resulting in rapid reactivity excursion. Steam flow malfunction; this might result from a main steam line break requiring safety injection, reactor scram, closure of isolation valves and the start-up of auxiliary pumps. Feedwater system faults; necessitating scram and initiation of the auxiliary feedwater system. Load rejection; the reactor power is regulated and if necessary the reactor is scrammed. Auxiliary power loss; power is normally guaranteed by start-up of diesel generators. An example of an intact circuit severe accident (assuming total loss of power supplies and back-up diesels) is considered in Section 3.5.

TABLE 3.1 EVENT CATEGORIES Category

Description

FrequencylReactor year

Conditions that occur regularly in normal operation

-10

2

Faults that are expected during the life of the plant: anticipated moderately frequent events requiring safety response

-1

3

Faults not expected during the life of a particular plant: Anticipated infrequent events requiring safety response

4

Improbable events not expected to occur in the nuclear industry but provided for by the design

-10-4

5

Extremely improbable events not provided for in the design of the plant

- 10-6

2 -10-

23

Accident Scenarios 3.3.2 Boiling Water Reactors

Examples of intact circuit fault transients in boiling water reactors include: Reactivity control malfunction; such a fault could result from uncontrolled withdrawal of control rods. A more serious event is the control rod accident where a control rod is suddenly ejected. Main recirculation pump trip or failure; trip of one or two pumps would be compensated for by the remaining pumps. This transient is plant specific in tenns of the impact of pump inertia on the ensuing transient. Steam flow malfunction; events associated with either increase or decrease of steam result in rapid transients. Feedwater system faults; these include loss of feedwater, failure of feedwater control system and various other abnonnalities of the feedwater system. Failure of the decay heat removal system; in such circumstances heat is removed via the condensation pool . Auxiliary power loss.

TABLE 3.2 EXAMPLE EVENTS Events

Categories

Bringing the reactor to full power

1

Loss of external grid

2

Loss of feedwater Loss of reactor coolant pump Small LOCA

3

Valves open Large LOCA

4

Main steam line break LOCAs without ECCS Transients with total loss of on- and off-site power

5

Accident Scenarios

24 3.4

Loss or coolant accidents (LOCAs)

3.4.1

Pressurised Water Reactors

3.4.1.1 Large break LOCAs

These accidents are characterised by a sufficiently large break in the primary coolant system to cause the system to depressurise rapidly. The accident involving a large double-ended guillotine break in a cold leg has been the most widely studied since it provides the greatest challenge to the emergency core cooling systems (ECCS). This accident is characterised by a number of phases. During the first phase (blow down phase), the coolant is lost through the break, the system rapidly depressurises and the core is uncovered. After about l Os the pressure falls to a sufficiently low level to allow the high pressure injection system to operate and the accumulators to discharge, both into the cold leg. Events are summarised in Table 3.3. Initially there will be a tendency for the injected water to bypass the core region, due to significant steam flow up the downcomer. Eventually though the pressure will fall sufficiently to allow refill of the lower vessel to take place. The low pressure injection system (LPIS) will also operate by this time. Finally the core will be reflooded and cooled. Long term cooling is established by the LPIS pump maintaining the core covered. Any steam produced is condensed by the containment sprays and the resulting water is recycled. Timescales of events are shown in Table 3.3. TABLE 3.3 PWR LARGE BREAK LOCA PHASES Phase

Time(s)

Bypass Refill

20-30 30-40 40-250 250-

Reflood Long term cooling

3.4.1.2 Small break LOCAs

LOCAs involving a break in the cold legs are potentially serious and these have also been widely studied. After the initiating event the pressure falls and the reactor trips. The rate at which a reactor depressurises depends on the size of the break. The high pressure injection system OWlS) pumps are available initially but not the accumulators or the low pressure injection system. Initially the primary circuit pressure will be higher then the operating set point pressures of these systems.

Accident Scenarios

25

The sequence of events might be as follows, Table 3 .4. It is assumed that the pumps are tripped: this is the advice to operators in order to minimise the loss of water from the primary circuit. Once saturation conditions are reached, steam will be produced, the upper vessel and the steam generators become voided. There will only be condensation heat transfer from primary to secondary as long as a temperature gradient exists. In order to maintain this temperature gradient secondary side cooling or depressurisation to reduce the saturation temperature may be required. Eventually the core will begin to dry out Steam will start to be lost through the break once the water level in the loop seal in the broken loop falls sufficiently. Once this point is reached rapid depressurisation will occur leading to flashing and possible core rewetting. This depressurisation enables the accumulators to dump and cool the core. The remainder of the accident follows that as described for a large LOCA. Reference 3. 1 shows the effects of different break sizes on the timescale of depressurisation.

TABLE 3.4 PWR SMALL BREAK LOCA PHASES Ph ase

Time(s)

Initial stage

0- 10 1 0-220 220-280 280-3 1 0

Reflux condensation Potential for frrst core uncovery Loop seal clearing, potential for core uncovery Long term cooling

3 10-

3.4.2 Boiling Water Reactors 3.4.2.1 Large Break LOCAs

The design basis LOCA for the BWR starts with a rupture in a pipe connecting a recirculating pump with the reactor vessel. If this occurs the vessel is isolated from the turbine by shutting the steam line within a few seconds. Some core cooling is maintained initially as the feed pump coasts down and water continues to be circulated in the unbroken loop. From initiation of the break. the system depressurises, albeit at a slower rate than in the PWR. The relevant pipework in the BWR is smaller than the legs in the PWR. Eventually core flow stops as the suction of the jet pumps is lost The core then drys out and starts to increase in temperature approximately lOs after the start of the accident

Accident Scenarios

26

After about 30s the ECCS is triggered. The vessel pressure is lowered further by the automatic depressurisation system enabling the low-pressure coolant injection (LPCI) and low pressure core spray (LPCS) system to inject water into the vessel above the core. The steam rises rapidly upwards against the downward flow of water. Eventually the lower plenum and core reflood. The reflooding rate is limited by the rate that steam can escape. This steam binding phenomenon occurs in a similar way to the PWR case. 3.4.2.2 Small Break LOCA

Again as for the PWR a full spectrum of break: sizes are considered in accident analysis. Peak clad temperatures (PCTs) attained are not a monotone function of break size. PCTs increase with break size for very small breaks up to about 1 OOcm2 then fall with break: size but eventually reach a maximum value at the maximum possible break size. 3.5

Severe Accidents

3.5.1

Pressurised Water Reactors

3.5.1.1 Intact circuit

There are a number of severe accident transient faults that under very pessimistic assumptions could lead to core melt down. Examples include complete loss of on- and off-site power, total loss of feedwater to the steam generators plus operator failure to correctly control primary feed and bleed and other initiating events with failure of the operator to take the correct remedial action. A very commonly studied transient is the TMLB (in WASH 1400 [3.3] terminology) sequence. This terminology is summarised in Table 3.5 for PWRs [3.3] (see [3.3] for corresponding BWR Table.) It provides a means for describing in short hand, different severe accident scenarios. In this particular sequence [3.4] , [3.5] , [3.6] , it is assumed that there is a loss of off-site power with failure of the on-site diesel generators and loss of turbine driven steam generator auxiliary feeds. There is therefore a loss of heat sink which results in overheating of the primary and secondary sides. With no available ECCS the systems overpressure, water and steam inventory is lost through safety relief valves and eventually the core uncovers. Many complex thermal hydraulic phenomena occur during the early phase of any severe accident. These are similar to those described for the design basis accidents earlier. For a large modem PWR the core uncovers in about 3-4 hours after accident initiation. Timescales can be seen from Figure 3. 1 . The core heats up due to the decay heat and the Zircaloy cladding begins to react exothermically with the steam as temperatures rise, Figure 3. 1 . This reaction becomes very rapid above about 1 300K and is an important contribution to core heat-up. The Zircaloy/steam reaction results in hydrogen production, which together with steam will be released to the containment Oxidation of stainless steel control cladding and other metallic structural materials can result in further hydrogen being produced.

27

Accident Scenarios 3000 �-------,

2 500

�

+-' ttl

2000

�

Q)

E 1 500

Q) I-

1 000

500 4-----�----�--� 7 000

7 500

8000

8 500

9000

9500

TI M E (s)

F U E L MAXI M U M CLAD D I N G TE M P E RATU R E

200

-a �

1 50

-0

1 00

c Q) 01 o >­

I

ttl

�

50

O �====�====��

�

�____�__

__

7000

7 500

8000

8500

9000

TI M E (s)

C U M U LATIVE HYD ROG E N G E N E RATI O N

FIGURE 3.1 STATION BLACKOUT FOR PWR

9 500

28

Accident Scenarios

As temperatures continue to rise, a number of core melt processes begin to occur. At temperatures between 1473 and 1673K, control rod, burnable poison rod and structural materials form low temperature liquid phases and molten material can relocate to form blockages in the lower core region. At higher temperatures in the range 2033K - 2273K any unoxidised Zircaloy will melt and react with fuel. This can result in a eutectic candling process forming blockages in the lower core region. Finally at very high temperatures above 2873K ceramic materials will melt and at these temperatures complete melt down of the core materials would occur. Molten core may therefore pass through the lower core support structures. A number of key issues have been considered by safety assessors. There is the potential for core debris/water interaction raising fears about the likelihood of possible steam explosions. This likelihood though is thought to be relatively small. Other important issues concern the coherent nature or otherwise of the debris, the extent of any core debrisllower head interactions and the containment threat posed by a high pressure melt ejection from the bottom of the vessel into the containment. Two alternative scenarios have been postulated concerning the progress of this particular accident. In the first scenario the lower vessel is assumed to fail while the system is still at high pressure resulting in a rapid and fast ejection of the melt. In the second scenario it is assumed that the primary circuit fails prior to vessel rupture, due to overheating via natural circulation and deposition of fission products. This would result in accumulator dump and whilst not preventing ultimate lower head failure would result in a more quiescent release of material to the containment and reduce the short term dynamic load on the containment. The actual amount of heat transfemxl to the contairunent atmosphere will depend on the scenario. It follows that the threat to early contairunent failure will also be scenario dependent The timescale for the start of the core heat-up is relatively long in this scenario. 3.5.1.2 �()<:Jls Several severe accident scenarios resulting from pipe break conditions have been studied in detail [3.4] , [3.5] . Two particular examples include the AB Hot Leg and the S2D in WASH 1400 terminology. The AB Hot Leg event refers to a large break in the hot leg, coupled with the loss of electrical power to all engineered safety features. The loss of electrical power prevents the ECCS from working, except for the accumulators which are driven by compressed nitrogen. Water will be expelled through the break during the blowdown phase. Any remaining water together with water discharged from the accumulators will be boiled off and core uncovering will take place relatively early. Since the break is in the hot leg the steam generators will be largely bypassed and any released fission products will be released directly to the containment.

Accident Scenarios

29

Once core WlCOVery has taken place many of the phenomena described earlier including Zircaloy oxidation and core melt down will take place. Some differences in detail would be expected compued with the high pressure � due to reduced density and steam availability.

Clad ballooning would occur which impacts of oxidation and melt progression phenomena. S ince the pressures are lower there will be a reduced tendency for natural circulation. Core debris will eventually slump into any remaining water in the bottom head. Failure of the lower head would eventually ensue without re-installation of cooling. Water and steam emitted from the break during the blow down phase will be transmitted to the containment. During the oxidation phase hydrogen is produced and this would in tum pass to the containment. Due to the low pressure some potential for a steam explosion may exist as debris slumps into water fIrstly in the bottom of the lower head and secondly into any water in the cavity. Debris attack of the concrete basemat could also occur releasing more combustible gases, e.g. hydrogen and carbon monoxide and further fIssion products through aerosol production . In contrast to the transient considered previously the timescales for this very severe accident would be very short.

TABLE 3.5 KEY TO PWR ACCIDENT SEQUENCE SYMBOLS A

B BI

C

D F

G H K

L M

Q R

SI

S2

T V ex

�

X o £

mtennediate to large LoeA. Failure of electric power to ESFs. Failure to recover either onsite or offsite electric power within about 1 to 3 hours following an initiating transient which is a loss of offsite AC power. Failure of the containment spray injection system. Failure of the emergency core cooling injection system. Failure of the containment spray recirculation system. Failure of the containment heat removal system. Failure of the emergency core cooling recirculation system. Failure of the reactor protection system. Failure of the secondary system steam relief valves and the auxiliary feedwater system. Failure of the secondary system steam relief valves and the power conversion system. Failure of the primary system safety relief valves to reclose after opening. Massive rupture of the reactor vessel. A small LoeA with an equivalent diameter of about 2 to 6 inches. A small LoeA with an equivalent diameter of about 1/2 to 2 inches. Transient event. LPIS check valve failure Containment rupture due to a reactor vessel steam explosion. Containment failure resulting from inadequate isolation of containment openings and penetrations. Containment failure due to hydrogen burning. Containment failure due to overpressure. Containment vessel melt-though.

30

Accident Scenarios

In WASH 1400 tenninology the S2D accident is defined by: S2 - a small break LOCA diameter (12.5 - 50 mm) D - failure of the emergency core cooling system. In this scenario failure of the high and low ECCS injection system is assumed. Water and steam are lost through the break and the core uncovers about 1 - 2 hours after the initiating event In contrast to the AB hot leg scenario above increased fission product retention would be expected within the steam generators. Core heat-up events are quite similar except that the timescales are larger. Following failure of the lower head molten material would be expected to be ejected more energetically than in the AB hot leg, and as for the intact circuit case direct heating of the atmosphere is possible. There may also be more water present in the cavity if the accumulators discharge during the lower head blow down. This may impact on the extent of any core concrete interaction. There are significant differences in the containment for this S2D case where operation of the containment fan coolers and sprays is not discounted. These devices can be used to mitigate pressure spikes. Operation of these devices could have a worsening effect on the accident since they could result in steam condensation and increase the possibility of a hydrogen deflagration (or even detonation). From a severe accident viewpoint; a positive feature is that the route to the exit will be via the steam generators. This is advantageous because the plate-out of Fission Products will be more effective in the cooler sa tube regions. 3.5.1 .3 Containment Bypass Sequences

The final PWR �o discussed is the interfacing LOCA, designated event V in WASH 1400 and subsequent studies. The scenario is usually known as the V-sequence. It occurs when a direct path is opened from a PWR primary circuit to a residual heat removal (RHR) system outside the primary containment while the reactor is in normal operation. It is initiated by failure of the check valves �arating the high pressure circuit from the RHR system. Since the RHR system is not nonnally designed to accommodate the primary operating pressure 1 5.5 MPa, it is assumed to rupture and to allow blowdown of the primary coolant into the auxiliary building (outside the containment). The principal routes through which this blowdown might occur are the RHR return flow pipes. The accident is initially like a LOCA with an intermediate break size but with one LPI system effectively disabled. Successful operation of the other LPI system(s) and the high pressure injection system and also secondary side cooling could delay core melt for

Accident Scenarios

31

several hours. However, when the ECCS water supply runs out, a switch to recirculation from the containment sumps is not possible, because the ECCS water has been lost outside the containment, and the sumps remain empty. Thus there is a potential for core melt in a situation with a large leak route direct from the primary circuit to a point in the auxiliary building (outside the containment). A mitigating phenomenon is the plate-out of FPs in the RHR lines. 3.5.2 Boiling Water Reactors 3. 5. 2.1 Intact circuit faults

Station blackout transients have been studied in Swedish power plants [3. 1 ] . The transient results through a prolonged loss of on- and off-site power, including back-up diesels. The sequence of events follows a similar pattern for thesetransients. Primary coolant is boiled off via the decay heat through relief valves. The primary circuit pressure remains at the safety relief valve pressure, typically at about 7 MPa . The core then uncovers. As the water level drops, a possible operator action is to deliberately depressure in order to avoid the problem of high pressure melt ejection. In either case the core will degrade and core debris will fonn on the core support plate and eventually relocate downwards to the lower plenum of the pressure vessel. Eventually if the accident continues the pressure vessel may fail and molten material will be ejected energeticall y on to the floor of the drywell (if the system is at pressure) or merely fall under gravity (if the system has already been depressured). There is then a potential threat to the containment. Steam generation due to melt interaction with water in the condensation pool or direct heating and hydrogen gas fonnation in the containment atmosphere are all main contributors to increase in containment pressure. Although there are some differences many of the phenomena associated with core melt down, fuel water interaction and containment behaviour are similar in both BWR and PWR postulated severe accident scenarios. 3.5.2.2 LOCAs

The phenomena that occur during core degradation and later phase melt progression have simularities with those discussed for the PWR. Some differences in materials and their oxides and the consequent effect on events are commented upon in Chapter 8. REFERENCES 3.1

3.2 3.3

B Pershagen, LIght Water Reactor Safety, Pergamon Press, 1989. J G Collier, G F Hewitt, Introduction to Nuclear Power, Hemisphere, 1987. US Nuclear Regulatory Commission, Reactor Safety Study: An Assessment of Accident Risks in US Commercial Nuclear Power Plants, USAEC Report WASH1400, October 1975.

32 3 .4

Accident Scenarios

A T D Butland et al, PWR Severe Accident Containment Study, Phase 1 , AEEW­

R I 842, December 1984. 3.5

A T D Butland et al, PWR Severe Accident Containment Study, Phase 2, AEEW­

R I964, April 1986. 3.6

I N Lillington et aI, SCDAP/RELAP5 Severe Accident Predictions, Presentation to the 1993 RELAP5 International Users Seminar, Boston, Massachusetts, Iuly 6-9 1993.

33

Chapter 4 INTEG RAL EXPERIMENTS

4.1

Introduction

A wide range of experimental programmes have been carried out in the field of Light Water Reactor Safety. This chapter summarises the main features addressed in some of the key integral experimental programmes within the US , Western European countries and Japan. Both separate effects and integral experiments have been perfonned. Separate effects tests are usually required for understanding of particular phenomena and for model development Integral tests are required to understand overall phenomenology, including coupling effects and also to provide data for code model validation. The emphasis in early reactor safety research was on large scale thennal hydraulics programmes. The main purpose of these experiments was to verify design and safety criteria for licensing purposes. Following the Three Mile Island accident there was increased emphasis on severe accident safety research and a number of major severe accident programmes were initiated. These have been primarily concerned with understanding core melt phenomenology, detennining the consequences of such events on the containment and with detennining the ultimate source tenn to the containment in the event of a severe accident The emphasis of this chapter will be in summarising the main objectives of the more recent significant integral tests in the various areas of reactor safety research. Considerable emphasis is therefore placed on severe accident related programmes. Finally this chapter does not claim to provide exhaustive coverage of all the tests performed towards reactor safety, but aims to describe some of the principal test programmes that have been used for validation of the state of the art codes discussed later in the book . The experimental programmes referred to in Sections 4.2 - 4.4 relate mainly to nonnal operating and design basis safety research: the later Sections are concerned principally with severe accident research. 4.2

Thermal Hydraulics

Two series of programmes of experiments were carried out in the Loss of Fluid Test (LOm Facility at Idaho Falls, USA. The facility was a USO scale model of a commercial 4 loop PWR. These experiments relate to pressurised water reactor safety concerns.

34

Integral Experiments

The fIrst series [4. 1] sponsored by the United States Regulatory Commission (US NRC) included both LOCA and transient tests. It was followed by a second series [4.2] funded by OECD. This series contained additional thennal hydraulic experiments, but in addition, the last two experiments addressed fission product release and fuel damage. Another series of tests carried out at the Idaho National Engineering Laboratory (lNEL) for pressurised water reactors was the SEMISCALE tests [4.3] . The SEMISCALE facility was a two-loop, full height, 1/2000 volume scale facility using electrically-heated fuel rod simulators. Tests included LOCAs and other sequences where natural circulation is an important phenomenon. The USNRCs LOCA programme for boiling water reactors included the TLTA (Two Loop Test Apparatus) [4.4] and FIST (Full Integral Simulation Tests) [4.5] operated by General Electric. In Europe, the LOBI (LWR Off-Nonnal Behaviour Investigation) test facility [4.6] was a high pressure integral system test facility operated by the Joint Research Centre of the Commission of the European Communities at Ispra in Italy. The facility was scaled (volumetric scale I n 1 2) to represent a 4 loop 3800 MWt Gennan PWR. The facility simulated full power and pressure. It consisted of 2 loops, a broken and intact loop simulating one or three of the real loops in the reference design). There were two LOBI experimental programmes, designated MOD 1 and MOD2. MODI was primarily concerned with the investigation of large break LOCA phenomenologies. The MOD2 research priorities were directed towards the investigation of small break LOCA and special transients. LOBI - MOD2 was a full height electrically heated model of a 4 loop PWR. Also in Italy is the SPES (Simulators per Esperienze di Sicurezza) integral test facility [4.7] operated at SIET, Piacenza. This 3 loop facility at 1/427 volumetric scale was representative of a \Vestinghouse 3 loop Plant. This facility has full power and pressure modelling capabilities. The PKL loop in Gennany is a three loop, full height integral test facility with a volume scale 1/134 . A main aim o f the PKL programme has been to investigate natural circulation in reflux condenser mode and other characteristic phenomena under small break LOCA and transient conditions [4.2]. The ROSA-III programme in Japan aimed to address small LOCAs in boiling water reactors [4.9] . This programme, managed by JAERI, was carried out in the LSTF (Large Scale Test Facility) which is a 1/48 volumetric scale model of a Westinghouse 3420 MW 4-loop PWR. The Upper Plenum Test Facility (UP1F) is a full scale model of a 4-loop PWR vessel operated by Siemens at Mannheim in Germany. Although not strictly an integral test facility, UPTF has enabled the full scale investigation of both cold leg and hot leg breaks,

35

Integral Experiments

particularly large breaks [4. 10] . Coupled phenomena occur in some of the tests, for example, in tests where incondensable gases have been introduced. UPTF provides unique data for code validation for large breaks. The BETHSY (Bouches d'Etudes Thermohydrauliques Systeme) programme is an on­ going programme [4. 1 1 ] operated by CEA at Grenoble, France. BETHSY is a 1/100 volumetric scale, 3-loop integral facility which is representative of a Framatome 2775 MWt PWR. A schematic is shown in Figure 4. 1 . BETHSY has the capacity for modelling a wide range of LOCA and non-LOCA transients. Later tests in the series will be concerned with accident management related issues and with accidents under shut­ down or mid-loop operating conditions.

PRESS U RISE R

n SG3

W I

SG I

STEAM GENER A T OR SG2

D O WNC OMER

FIGURE 4.1 BETHSY FACILITY 4.3

Fuel and Cladding Behaviour

This Section considers briefly experimental programmes concerned with fuel and cladding behaviour under design basis accident conditions. The corresponding behaviour under severe accident conditions is considered in Section 4.5.

Integral Experiments

36

Cladding creep under LOCA conditions has been studied in both inert and oxidising atmospheres by UKAEA at Springfields [4. 1 2] and other programmes [4. 13], [4. 14] . Cladding oxidisation kinetics have also been measured i n a nwnber o f separate effects programmes in temperature ranges relevant to both LOCA and severe accident conditions. These are specifically considered in Chapter 8. Fuel material may contact cladding through various mechanisms including fuel swelling due to irradiation, fuel relocation etc. This can lead to the well known phenomenon of PCI (pellet Clad Interaction). This phenomenon has also been widely studied for example at Studsvik, [4. 15]. Heat generation due to fission product decay has been studied in the US and various decay curves have been produced. There have been many research programmes to determine the quantity and composition of fission product release from damaged rods under normal operating and LOCA accident conditions. 4.4

Materials and Structural Behaviour

Material and thermophysical properties have been derived in numerous USNRC sponsored programmes in the US e.g. [4. 14]. The database is currently being extended to cover severe accident conditions. Fracture mechanics experiments e.g. the HSST programme have been carried out. Other programmes to address corrosion and water chemistry and fatigue stress corrosion issues were set up by the USNRC and EPRI during the 1970s. 4.5

Core Melt Programmes

This Section gives a summary of the major integral scale experimental programmes performed to investigate the behaviour of PWR and BWR cores under severe accident conditions e.g. extreme loss of cooling conditions. The principal aim of these tests is to improve understanding of the behaviour of fuel rods, control rods and other materials present in the core, under high temperature, steam oxidizing conditions. The accident at Three Mile Island - Unit 2, considered in detail in Chapter 22 provided the initial impetus for these tests. Various programmes are sponsored by the United S tates Nuclear Regulatory Commission (USNRC). Early phase melt progression data were obtained from the Severe Fuel Damage (SFD) tests in the Power Burst Facility (PBF) at the Idaho National Engineering Laboratory (lNEL) [4. 16] . These in-pile 32 rod bundle tests provided some of the first data available on fuel rod damage following severe cladding oxidation, melt relocation and fuel rod fragmentation.

Integral Experiments

37

Damaged Fuel (DF) tests were also perfonned in the Annular Core Research Reactor (ACRR) at Sandia National Laboratories (SNL) [4. 1 7]. These in-pile tests comprised short length 9 rod bundles. The main aim was to investigate fundamental fuel damage phenomena at various heat-up rates in an undercooled controlled steam environment. The Full-Length High Temperature (FLHT) series of tests was carried out in the National Research Universal (NRU) reactor at Chalk River Nuclear laboratories in Canada. The programme [4. 1 8] aimed to investigate fuel damage in full-length bundles. These tests also consisted of small numbers of fuel rods (up to about 12 rods in a bundle). The Loss of Fluid Tests (LOFT) at INEL have already been referred to earlier in the Section on thennal-hydraulics. The final tests in the series. LP-FP- l and LP-FP-2. were concerned with severe accident phenomena. In LP-FP- l . the main concern was with fission product release: in the LP-FP-2 test. significant melt progression occurred. In Europe. a Severe Fuel Damage (CSD) series of in-pile tests was perfonned in the PHEBUS reactor at Cadarache. France [4. 19] . These tests aimed to address specific phenomena under well controlled steam inert gas atmosphere conditions. The NIELS out-of-pile single rod and small (3x3) bundle experiments at [4.20] Kemforschungszentrum Karlsruhe (KfK) in the Federal Republic of Gennany were the first experiments performed world-wide addressing fuel rod behaviour under severe fuel damage conditions. The series was started before the TMI-2 accident and the NIELS results gave some early insights into what might have happened during the accident. The following CORA programme is a wide ranging series of out of pile tests providing information on the behaviour of light water reactor (LWR) core materials (both PWR and BWR) under severe accident conditions [4.21]. Cross-sections of CORA bundles are shown in Figure 4.2. The aim of these tests is to investigate the impact on phenomenology of a systematic variation of various parameters. maximum temperature. system pressure. internal heat-up rate. rod internal pressure. steam supply. conditions of tennination of the test. bundle size and chemical conditions of the bundle. The thennal hydraulic conditions have been generally chosen to simulate the low flows associated with core boil-down. and to produce representative temperature excursion rates. The above test programmes provide a database on fuel damage for bundles with various numbers of fuel rods and control rods (from 1- 1(>0). of various lengths (from O.5m to full length) and also including grids. Some tests have been quenched once the bundle is severely degraded. others have been slow cooled. The above experiments are restricted to early phase melt down phenomena i.e. up to initial blockage fonnation. A late phase experiment. ACRR-MPI has been perfonned in the ACRR at Sandia. This takes as its initial geometry a simulated blockage and aims to investigate subsequent melting behaviour.

38

Integral Experiments

(Ag, In, (d}-Abso rb er ro ds

1\

rod

sim l.! !ato r

U nhe�tec r o c S h ro u d

Shroud insu l ation (Zr02fi bre)

PWR

F u e ! rod

SS-b l e: d e

B ite powder

Ab s o r b e r rod

Zirca loy-cha nnel (box) wa l l

B WR FIGURE 4.2 CORA BUNDLE ARRANGEMENTS

Integral Experiments

39

Reactivity Initiated Accidents (RIA) have been addressed in a JAERI programme [4. 1 8] in the Nuclear Safety Research Reactor (NSRR). An RIA is characterised by an abnormal increase in reactivity caused by an unusual withdrawal or ejection of control rods leading to a rapid reactor power and fuel temperature increase. This can lead to adverse effects on fuel rod integrity.

4.6

Natural Circulation

If the pumps have been tripped, natural circulation can occur at various times during many of the accident scenarios of interest The phenomenon is of particular interest as a mechanism for providing loop circulation in PWRs under intact circuit or small break LOCA accident conditions. Under severe accident conditions it provides a mechanism to transfer heat from an overheated PWR core to other parts of the primary circuit and possibly threaten the circuit integrity. In the latter case, few data have been published in the literature. A programme [4.22] was conducted by Westinghouse, sponsored by EPRI to investigate natural circulation phenomena in a one-seventh scale replica of a pressurised water reactor system. Integral tests were carried out with water and sulphur hexaflouride both at low and high pressures. The latter fluid was chosen to provide more realistic scaling to plant conditions.

There have been some experimental investigations in porous media consisting of particle beds, but few in the rod bundle configurations relevant to LWR severe accident safety research. However natural circulation in a rod bundle configuration was simulated in a rig [4.23] at the Institut fur Kemenergetik und Energiesysteme (IKE), Germany. This provides basic data for natural circulation in the core region simulating conditions when the core is uncovered.

4.7

Fission Product Release and Transport

Fission product release from simulants for molten corium was investigated in the SASCHA Facili ty at Karlsruhe [4.24] . The main purpose was to investigate the transient fission product release behaviour from fuel. A series of tests at Oak Ridge National Laboratories [4 .25] were performed to investigate fission product release from intact LWR fuel. This series included tests with high bum-up fuel, but all the tests were limited to atmospheric pressure. At the integral test scale, fuel release data have been obtained from the last two tests in the OECD LOFT' programme and from the last PBF fuel damage test, PBF 1 .4. Marviken Aerosol Transport Tests (ATT) [4.26] have produced a large data-base on the behaviour of aerosols in a reactor primary coolant circuit These tests used simulant fission products and core materials and were close to full scale.

Integral Experiments

40

A series of large scale aerosol experiments have also been performed at the Hanford Engineering Development Laboratory, USA. These addressed the behaviour of aerosols under realistic severe accident conditions. The Winfrith FALCON experiments [4.27] are providing a current source of data relevant to a wide range of important issues. In particular, the experiments are designed to ensure that realistic fission product chemistry interactions (with simulant materials) occur and can be taken into account. The experiments are small scale and performed at atmospheric pressure.

4.8

Debris Beds

Various experimental programmes have addressed the issue of whether a debris bed (to simulate the materials present in a degraded core) can be quenched or not Depending on the circumstances reflood from either the top or the bottom of the bed may need to be considered. Top reflood is relevant to ex-vessel events where e.g. debris in a reactor cavity might be flooded by water from on top. In some reactor designs ECCS systems inject into the hot legs and here top reflood could occur in-vessel. Top reflood experiments [4.22] were carried out at Argonne National Laboratories in 1982. The main purpose of these experiments was to investigate the coolability margins of a degraded LWR core, the steam production rate and to understand the essential phenomenology. Experiments [4.29] designed to investigate more prototypical beds have been carried out at the University of California, Los Angeles (UCLA). The experiments were performed to determine the extent of the effects of internal heating and gas injection on the quenching of core debris and also to address both axial and radial stratification in top reflood. Transient core debris heat removal tests were carried out at Brookhaven National Laboratories in a programme [4.30] , [4.3 1 ] sponsored by the USNRC, where hot beds of stainless steel particles were quenched by overlying pool s of water. The coolability and dryout limits for volumetrically heated beds of particles were investigated at Kemforschungszentrum Karlsruhe, Germany [4.32] . Some transient tests, as well as steady-state tests, were also carried out in this series. The USNRC sponsored a programme [4 .33] at Sandia National Laboratories to investigate degraded core coolability. These in-pile tests, performed in the Annular Core Research Reactor (ACRR) aimed to investigate prototypical materials over a large pressure range, i.e. in parameter ranges, that had previously not been investigated. Bottom reflood of debris beds is relevant to the issue of in-vessel core flooding e.g. via ECCS injection into the cold leg/downcomer of a PWR.

Integral Experiments

41

The phenomenology of in-vessel debris coolability by bottom reflooding was investigated at Berkeley Nuclear Laboratories of Nuclear Electric (formerly CEGB) [4.34] . Bottom reflood was also investigated at Brookhaven [4.35] as a follow-up to previous work on top reflooding cited earlier. The aim of this work was directed toward late phase model development in the SCDAP/RELAP5 code. More realistic debris beds with bottom reflood were also studied by the experimental group at UCLA [4.29] . Their programme included fluidisation and radial and vertical stratification experiments. 4.9

MeltlWater Interactions

If a severe accident in a PWR developed to the stage of partial fuel melting, molten fuel could come into contact with water, heating it rapidly and causing it to expand explosively. Such an energetic interaction is termed a Molten Fuel-Coolant Interaction (MFCI). MFCIs have been studied in the UK and US [4.36] since the 1970s. Experiments were carried out at Winfrith [4.37] in which high temperature molten uranium dioxide fuel was produced using a thermite process and dropped into water. Several series of experiments were carried out in which the relative volumes of corium and water, and also the over gas volume, were varied. Experiment conditions were relevant to the case of melt falling into the bottom of a pressure vessel and either falling or being ejected through the lower head into a partially flooded vault

4.10

Core Concrete Interaction

The SURC series of tests [4.38] were performed at Sandia National Laboratories. The main purpose of the SURC series was to investigate the interaction of molten steel with concrete. Melt concrete interactions were also investigated in the BETA Test facility [4.39] performed at Kernforschungszenttum Karlsruhe (KfK). Some of these tests investigated the effects of high zirconium metal content in the steel melt in order to provide more prototypical gas generation. This phenomenon was a main topic of interest in the ACE programme. This programme included experiments which were concerned with prototypical oxide melt interacting with concrete. The ACE programme was conducted in the USA. The SWISS series of experiments [4.40] also conducted in the USA, included tests (SWISS-2) designed to investigate the sustained simultaneous interaction of molten stainless steel with concrete with an overlying water poo l.

Integral Experiments

42

MAI N O F F-GAS LI N E

T U N G ST E N-LI N E D LI D

COOLI N G PAN E LS D E PLET E D U 0 2 POWD E R

T U N G STEN E LECTRO D ES

--tt-ESgW2d

TH E R M O C O U PLES AT VAR I O U S D E PT H S

CONCR ETE BAS E M AT

FIGURE 4.3 ADVANCED CONTAINMENT EXPERIMENTS INVOL VING MOLTEN CORE CONCRETE INTERACTION AT ARG ONNE NATIONAL LABORATORY 4.1 1

High Pressure Melt Ejection and Direct Containment Heating Experiments

In the event of a failure in the lower head of a reactor vessel, molten debris would be ejected from a hole in the bottom of the pressure vessel and material transported up into the containment. A number of experimental programmes have been carried out to address the associated phenomena. Integral tests at l/lOth scale [4.4 1 ] have been perfonned in the Surtsey Facility at Sandia National Laboratories in the USA. The rig consisted of a pressurised thermite melt generator, a I/lOth linear scale model of a large modem PWR cavity, and a single volume containment vessel at the same scale. An iron/alumina melt was discharged into the cavity with the potential to simulate the thennal and chemical behaviour of a melt generated during a reactor severe accident. The main purpose of Surtsey was to provide data on gas discharge, hole ablation, heat loss to structures and plume effects. Data on jet break up have been obtained from the HIPS and JETA experiments [4.42] . The HIPS experiments used similar materials and apparatus as in the Surtsey experiments. The JETA experiments were carried out in the SPIT experimental apparatus. (The SPIT experiments were similar to HIPS but at V20th scale.) Depending on the accident sequence water may or may not be present in the cavity. Corium Water Thennal Interaction (CWTI) experiments, similar to the Surtsey experiments

Integral Experiments

43

except at smaller scale and using corium melt, provided data on cavity melt phenomena, the effects of subcompartments and the effect of airborne water. Experiments to address the flow of air and water through a model of a PWR cavity have been carried out at Winfrith [4.43] . The purpose of these tests was to provide data (including visual data) on flow phenomena and liquid step fonnation. Other dispersal experiments, similar to those at Winfrith have been carried out at Berkeley Nuclear Laboratories of Nuclear Electric. These provide more integral scale data: two phase transient discharge at higher pressure than the Winfrith experiments was simulated at two different scales (I/lOOth and l/25th scales). Tests have also been conducted in the 1/IOOth scale model cavity at Berkeley to extend the database on fuel coolant interactions in wet cavities.

4.12

Containment Phenomena

Thi s section summarises the experimental programmes which address thermal hydraulics, aerosol behaviour and fission product transport in the containment for various severe accident conditions. The containment building of the disused HDR reactor has provided detailed thcnnal­ hydraulic data, [4.44] , [4.45] for conditions relevant to both design basis and severe accident conditions. The facility has the same height as a modem PWR containment but is only about 1/3 of the diameter. The Battelle Model Containment (BMC) was a 640m3 concrete model containment Thennal-hydraulic and aerosol effects were investigated in the DEMONA series of experiments. [4.46] , [4.47]. Most of the experiments in DEMONA were single cell except for the final experiment in the series. The LACE series of experiments [4.48] co-ordinated and managed by EPRI, was conducted in the 852m3 Containment Systems Test Facility (CSlF) at the Hanford Engineering Development Laboratories (HEDL) in the USA. In the LACE experiments, more complex thermal-hydraulic conditions existed, compared with DEMONA. More transient and also complex aerosol phenomena were investigated. The ACE/MACE programme, also conducted and managed by EPRI , addresses further wide ranging topics associated with containment loading under severe accident conditions. The programmes include assessment of filtration systems for vented containment concepts, in-containment iodine behaviour, core concrete interactions, Figure 4.3, and core debris coolability experiments. The FALCON Facility at Winfrith includes a small scale model of the containment (O.3m3). The main attribute of the FALCON experiments is that materials are chosen to ensure prototype aerosols and vapours are provided as a source term.

44

Integral Experiments

REFERENCES

4.1 4.2 4.3 4.4

4.5 4.6 4.7

4 .8

4.9 4.10 4. 1 1

4.12

4.13

4.14 4.15 4.16

4.17

4.18

D L Reeder, LOFT System an d Test Description NUREG/CR-0247 (TREE- 1208) July 1978. J Fell, S M Modro, An Account of the OECD LOFT Project, OECD LOFT-T-3907, May 1 990. L J Ball, K A Dietz, D J Manson, D J Olson, SEMISCALE Program Description TREE-NUREG 12 10. W S Hwang, BWR Small Break Simulation Tests with and without Degraded ECC Systems, BWR Blowdown!Emergency Core Cooling Program NUREG/ Cr-2230, EPRI NP- 1 782, GEAP-24963, January 1 982 A G Stephens, FIST Facility Description Report, NUREG/CR-2576, EPRI NP-23 14, GEAP-22054, December 1982. C Addabbo, LOBI Seminar, 3 1 March-2 April 1992, Arona, Italy, EUP 1 4 1 74 EN, 1992. A Annunziato (ENEA), L Mazzocc hi (CISE), G Palazzi (ENEA), R Ravetta (SIET). SPES: The Italian Integral Test Facility for PWR Safety Research Journal "Energia Nucleare" N. 1 December 1984. D Hein, K Watzinger, Small-Break LOCA. Analysis, Control and Experimental Results, Paper lAEA-CN-39/A-7-30, at Inl Conf. on Current Nuclear Power Plant Safety Issues, Stockholm, 20-24 October 1980. Y Anoda, K Tasaka, H Kumamaru, M Shiba ROSA-III System Description for Fuel Assembly No 4 JAERI-M 9363, February 1 98 1 . 2D/3 D-Projekt, Planung der UPTF (Upper Plenum Test Facility) KWu Abschlussbericht R9 14/026/80, September 1980. OECD/NEA/CSNI International Standard Problem No 27, BETHSY Experiment 9.1 B.2 "Cold Leg Break Without HPSI and with Delayed Ultimate Procedure, NEN CSNI/R (R92) 20, November 1992. E P Hindle, C A Mann and A E Reynolds. The Deformation of Zircaloy PWR Cladding with Low Internal Pressure, under Mainly Convective Cooling by Steam. ND-R-478(S), August 198 1 . A T Donaldson, G Knowles. Primary an d Steady State Creep o f Westinghouse PWR Fuel Tube Material; Constant Stress Tests between 973 and 1073K. TRPD/B/0267/ N83, June 1983. J K Hohorst, D L Hagrman et aI, MATPRO - A Library of Materials Properties for Light Water Reactor Accident Analysis, NUREG/CR-5273, January 1 990. H Mogard ''The Studsvik Materials Testing Reactor in Domestic and International Fuel Research and Development, Studsvik Energiteknik AB, 1982. D J Osetek "Results of the Four PBF Severe Fuel Damage Tests", Transactions of the Fifteenth Water Reactor Safety Information Meeting, Gaithersburg, MD, October 2630, 1987, NUREG/CP-009 0. R 0 Gauntt et al "Results o f the ACRR-DFR Experiments", Proceedings o f the International ANS/ENS Topical Meeting on Thermal Reactor Safety, Vol. 3, San Diego, C A, February 2-6, 1986. In-Vessel Core Degradation i n LWR Severe Accidents. A State o f the Art Report to CSNI, NENCSNI/R/9 1) 12, January 1 99 1 .

Integral Experiments

4.19

45

C Gonnier, et al "In-Pile Investigation at PHEBUS-SFD Facility of the Behaviour of PWR Type Fuel Bundles in Severe Accident Conditions beyond the Design Criteria", International Conference on Thennal Reactor Safety, A vignon, France, October 1988. 4.20 S Hagen, P Hofmann ''PWR Fuel Element Behaviour at Temperatures up to 22S00C" , IAEA Specialists; Meeting on "Water Reactor Fuel Behaviour and FISSion Products Relt"$e in Off-Normal and Accident Conditions" Vienna, 1(}'13 November 1986. 4.21 S Hagen, L Sepold, P Hofmann and P Schanz "Results of the CORA Experiments on Severe Fuel Damage with and without Absorber Material", Proceedings of the Heat Transfer Conference, AIChE, Vol. 85 ( 1989) 135- 140, Philadelphia 1989. 4.22 W A Steward, A T Pieczyuski, V Srinivas, Experiments on Natural Circulation Flows in Steam Generators During Severe Accidents, International ANS/ENS Topical Melting or Thennal Reactor Safety, San Diego, USA February 2-6, 1 986. 4.23 R Kulenovic, S RosIer, M Groll, J Unfried and M Burger, Experimental Investigation of Natural Convection in a Vessel with Rod Bundle Configuration and Comparison with Results from 2D Thennal-Hydraulics Computer-Code Frecon. Institut fur Kernenergetik & Energiesysteme (IKE) Univ. Stuttgart, Pfaffenwaldring 3 1 , D7000 Stuttgart 80, F.R.G. 1 990. 4.24 H Albrecht, H Wild, Review of the Main Results of the SASCHA Program on Fission Product Release under Core Melting Conditions, ANS Meeting on Fission Product Behaviour and Source Tenn Research, Snow Bird, Utah, 15- 19 July 1984. 4.25 M F Osborne, J L Collins, R A Lorenz, R V Strain, Fission Product Behaviour in Tests of LWR Fuel Under Accident Conditions, Proc. Int. Symp. on Source Tenn Evaluation of Accident Conditions, Columbus, Ohio, Oct 28-Nov I , 1985, IAEA-SM-28 1 . 4.26 J Collen, H Unneberg, D Mecham, Overview of Marviken Experimental Procedures, Proc. ANS Meeting on Fission Product Behaviour and Source Term Research, Snowbird, Utah, July 15- 19, 1984. 4.27 A M Beard, C G Benson, B R Bowsher, Fission Product Vapour-Aerosol Interactions in the Containment Simulant Fuel Studies, AEEW-R 2449, December 1988. 4.28 D H Cho, D R Annstrong, S H Chan, On the Pattern of Water Penetration into a Hot Particle Bed, Nuclear Technology 65, 23-3 1 (1984). 4.29 V X Tung, V K Dhir, Quenching of Debris Beds Having Variable Penneability in the Axial and Radial Directions, Nucl. Eng. and Des. 99, 275-284 ( 1 987). 4.30 T Ginsberg, N K Tutu, J Klages, J Klein, C E Schwartz, Y Sanborn, Transient Core Debris Heat Removal Experiments and Analysis, Proceedings of the International meeting on Thennal Nuclear Reactor Safety, Chicago, IL (August 29-September 2, 1982). 4.3 1 T Ginsberg, J Klein, J Klages, Y Sanborn, C E Schwartz, J C Chan, L Wei, An Experimental and Analytical Investigation of Quenching of Superheated Debris Beds Under Top-Reflood Conditions: Final Report, Brookhaven National Laboratory BNL-NUREG-5 195 1 , NUREG/CR-449 3 ( 1986). 4.32 L Barloon, K Thomauske, H Werle, Extended Dryout and Rewetting of Small­ Particle Core Debris, Nucl. Eng. and Des. 102, 59-69 ( 1987).

46 4.33

4.34

4 .35

4.36

4.37 4.38

4.39 4.40

4.4 1 4.42

4.43 4.44 4.45

4.46 4.47

4.48

Integral Experiments

A W Reed, K R Boldt, T R Schmidt, Coolability of LWR Debris: a Summary of the DCC Experiments, Paper VI.5, Proceedings of the International ANSIENS Topical Meeting on Thennal Reactor Safety, San Diego (1986). P C Hall, C M Hall, Quenching of Heated Particulate Beds by Bottom Flooding Preliminary Results and Analysis, presented at the European Two-Phase Flow Group Meeting, Eindhoven, Holland (198 1). N K Tutu, T Ginsberg, Debris Bed Quench Characteristics Under Bottom-Flood Conditions, Topical Meeting on Thennal Reactor Safety, ANS/ENS, San Diego, CA (2-6 February 1986). M L Corradini, Vapor Explosions: An Experimental Review for Accident Analysis, University of Wisconsin, Nuclear Engineering Department, Madison, WI537061 687, 1 990. M J Bird, An Experimental Study of Scaling in Core Melt/Water Interactions, 22nd National Heat Transfer Conference, Niagara Falls, August 1 984. E R Copus et al. Core-Concrete Interactions Using Molten Steel with Zirconium on a Basaltic Basemat: The SURC-4 Experiment. NUREG/CR-4994 SAND87-2008 (1 989). H Alsmeyer. Beta Experiments in Verification of the Wechsl Code: Experimental Results on the Melt-Concrete Interaction. Nucl. Engrg. Des. 103 1 987 p1 1 5. R E Blose, J E Gronager, A J Suo-Anttila and J E Brockmann. SWISS : Sustained Heated Metallic Melt/Concrete Interactions with Overlying Water Pools. NUREG/ CR-4727, SAND85- 1 546, July 1987. W W Tarbell et al, Direct Containment Heating: Surtsey Test Results and Models, Proc. Int ENS/ANS Conf. on Thennal Reactor Safety, Avignon, 2-7 October 1988. W W Tarbell et al, Melt Expulsion and Direct Containment Heating in Realistic Plant Geometries, Proc. Int. Meeting Thennal Reactor Safety, San Diego February 2-6 1986. P W Rose, Experimental Modelling of Core Debris Dispersion from the Vault Under a PWR Pressure Vessel, AEEW-R 2143 (December 1987). H Karwat, ISP-23: Rupture of a Large Diameter Pipe in the HDR Containment, CSNI Report No 160, Volumes 1 and 2 (1989). P N Smith, P Ellicott, A UK Analysis of Light Gas Distribution Experiment E l 1 .2 in the HDR Facility, Proc. Seminar on Containment of Nuclear Reactors, Shanghai, 141 5 August 199 1 , SMIRT 1 1 ( 1 99 1). W Shock, Post Test Calculations of Aerosol Behaviour in DEMONA Experiment B3 with Various Computer Codes used in CEC Member States, EUR 1 1374 EN (1988) .. J Gauvain, Post-test Calculation of Thennal-hydraulic Behaviour in DEMONA Experiment B3 with Various Computer Codes used in EC Member States, EUR 12 197 EN (1989). R Ricchena, CEC Comparative Exercise on the Perfonnances of the Computer Codes Simulating the LACE-4 Experiment, Paper in Proc. CSNI Workshop on Water­ Cooled Reactor Aerosol Code Evaluation and Uncertainty Assessment, Brussels, 91 1 September 1987, EUR 1 1 351 EN (1988).

47

Chapter 5 THERMAL HYDRAULIC MODELS

5.1

Introduction

The purpose of this chapter i s to summarise the important thennal-hydraulic models and methods that are used in modem transient system analysis computer codes. A number of such codes have been developed and a brief description of some of the more advanced mechanistic codes is given in Chapter 20. Examples of system two fluid codes include RELAP5t TRACt CATHARE and ATHLET. COBRA is an example of an advanced two-phase subchannel code. The discussion as far as possible in this chapter will not be centered on specific codes but on generic principles and models. The more modem codes differ from the earlier codes in that they have (or will have in the near future): 1.

Multi-dimensional capabilities for modelling the vessel in comparison with more one-dimensional lumped parameter models.

2. Increased number of conservation equations to model disequilibrium two phase effects. 3. Improved closure relations resulting from increased validation against experiment. 4. Improved numerical techniques. These are required to accommodate the increased complexity arising from points 1-3.

5.2

Code Categories

Two basic categories of codes have been developed for transient system analysis. These are: Evaluation Model (EM) codes Best Estimate (BE) codes. The purpose of EM codes was largely to provide conservative calculations for licensing. The codes included conservative models. The purpose of BE codes was to model the essential physics and provide a satisfactory description of the phenomena. Methodology concerning computer validation is

48

Thermal Hydraulic Models

discussed in Fabic et al [5. 1 ] Currently BE codes are also used for licensing via imposition of conservative boundary conditions. They may also be used to detennine uncertainties in best estimate predictions. BE code validation against experiments is necessary to provide confidence in extrapolating computer code predictions to plant scale. BE codes are also used for the design and analysis of experiments. The discussion in this chapter concentrates almost exclusively on best estimate methods and models. The fluid flow and heat transfer phenomena encountered in reactor accident analysis are complicated since in many accident conditions the flows are transient and two­ phase. This has led to the development of a wide range of two-phase flow models for the best estimate codes.

5.3

Detailed Physical Phenomena

Figure 5. 1 shows some of the various heat transfer regimes that can exist in flows through vertical fuel pin arrays. The fluid drag and heat transfer associated with the different flow regimes is a complicated function of quality and temperature. In BE codes there are generally various selection criteria for defining the flow topology and heat transfer regimes for all the various rod bundle configurations, pipes and other components that occur in the water cooling circuits in LWRs. In the large break LOCA scenario for example, the partitioning between the liquid and vapour (in tenns of possibly both velocity and temperature) the break-up of liquid columns, the entrainment of droplets etc may all be modelled locally within various pipes or components. In the case of reflood there may be in the core region counter current flow associated with liquid fall-back from the upper plenum. Multi-dimensional effects may occur during the blow-down phase due to core bypass, once accumulator discharge has commenced. Also within the core region the coupling between the thennal-hydraulics and heat conduction within the fuel pellets and cladding (with possibly changes in geometry associated with clad ballooning) may need to be modelled. Thus to model the whole range of possible thennal-hydraulic conditions a broad combination of one up to three dimensional models for modelling disequilibrium two phase flows is required. Examples of accidents that require such advanced models include not only large break LOCAs as described above but also other transients such as (Anticipated Transients Without Scram) ATWS and various reactivity insertion scenarios.

Thermal Hydraulic Models S.4

49

Classification of Two-Phase Flow Models

Multi-phase flow models are usually derived by performing spatial and temporal averaging of the instantaneous Navier-Stokes equations. Further assumptions are then required concerning the velocities and temperatures of each phase which depend on the physics of the two-phase flows being modelled. Models (in the order of increasing complexity) that have been derived (Wallis [5.2] , Hughes [5.3]) include the following: (a)

Homogeneous Flow

In this model, it is assumed that the phases are in thennal and dynamic equilibrium i.e. have equal temperatures and velocities. There are the same basic number of equations to be solved (3 for mass, momentum and energy) as for single phase flows. The usual correlations that are required for closure of such single phase flows e.g. for frictional drag usually need to be modified. (b)

Homogeneous Model with Slip

This model allows the velocities of each phase to be different but the velocity ratio is correlated as a function of flow field parameters. Here again three conservation equations are solved. (c)

Drift Flux Model

This consists of separate mass conservation equations, but usually, still a single momentum and energy equation. The system is closed via a correlation for the relative velocity between the phases. (d)

Disequilibrium Two-Phase Fluid Models

In principle, separate conservation models for each phase can be fonnulated and this is the basis of the now commonly used 6-equation model. The energy and momentum transfer between the phases and with the solid surfaces have to be correlated to close the system of equations. All the modem system codes include this level of modelling. (e)

Disequilibrium Fluid Models including Incondensable Gases

The six equation model has been extended to allow for incondensable gases such as nitrogen. The gas and steam phases are usually assumed to be in equilibrium and therefore an additional continuity equation is required for each additional species that needs to be modelled. Typically the partial pressures of the gaseous phase are assumed to obey Dalton's Law. Homogeneous models only take account of the behaviour of the two phase mixture as a whole. Even if slip or drift flux correlations are included, these are usually limited to steady state. The use of separate momentum equations enables the transient

Thermal Hydraulic Models

50

temporal and spatial effects of the individual phases to be modelled in a more physical manner. One of the disadvantages of more separated conservation equations is that more constitutive relations are required to close the system of equations. These inevitably rely on further correlations with their inherent limitations or complexity. Mass, momentum and energy equations may be fonnulated for the two individual phases or indeed for the mixture as described. Clearly various combinations of equations can be derived and some examples are listed in Table 5. 1 .

TABLE 5.1 TWO PHASE FLOW MODELS

No of Conservation Equations MODEL

MASS

MOMENTUM ENERGY

Homogeneous

1

1

1

Homogeneous Model With Slip

1

1

1

Drift Flux

2

1

1

Disequilibrium Two Phase Fluid

2

2

2

Disequilibrium Including m Scalar Fields

2+m

2

2

Derivation or Two-Phase Flow Equations

5.5

For practical modelling purposes, the macroscopic aspects of the flow are usually the most important. Starting from the instantaneous flow equations for the individual phases, e.g. Truesdell [5 .5] : Instantaneous Flow Equations Mass

op/ot + V .PY... 0 =

Momentum

opyJot + V.PIT

=

Pi. - Vp+ Vt

Thermal Hydraulic Models

51

Energy

ope(c)t + V ,pye = -p V.y + V.g where p, y, g, p, 't, e, g denote density, velocity, gravity, pressure, shear stress, internal energy and heat flux, respectively, it is necessary to apply both temporal and spatial averaging techniques. Temporal averaging is required to average fluctuations (perhaps turbulent fluctuations) on very small time scales. Spatial averaging is required to average out the effects of complex moving boundaries or interfaces between the two phases. The most commonly used techniques are the Eulerian time and spatial averaging techniques due to Ishii [5.4] . Time averaging the function f(tv over a time interval �T:

( I f(t, [}dt J ,'iT is one dimensional and indeed is commonly used for modelling single phase turbulent flows. Spatial averaging may be 1, 2 and 3 dimensional corresponding to line, area and volume averaging:

(

I

f(t, [}dr

J I\r

The procedure adopted will depend on whether the flow is primarily one dimensional e.g. as in a pipe or perhaps three dimensional as in a vessel. Applying these techniques to the instantaneous form of mass, momentum and energy equations for each phase a typical set of two phase conservation equations may be derived as follows. 5.5.1 Field Equations

After averaging a set of field equations is obtained: Conservation of Mass

52

Thermal Hydraulic Models

Conservation of Momentum

v

IV

VI

VII

VIII

Conservation of Energy

IV

VII

VIII

V

VI

IX

where � denotes the void fraction for the kth phase. The first tenns in these equations represent the rate of change of mass, momentwn and energy. The second tenns are the convective changes of these quantities due to bulk fluid flow. The third tenn in the mass conservation equation,rk, denotes the rate of mass transfer to phase k from the other phases. The third tenns in the momentum equations are the pressure gradients that drive the flow. In the above fonnulation it is assumed that pressures in each phase are in equilibrium. The third tenns in the energy equations denote the rate at which internal energy changes due to compressibility. The fourth tenns in the momentum equations represent the gravitational force. The fourth tenn in the energy equation denotes a work tenn associated with expansion or contraction of the phases. The fifth tenns in the momentum equation denote viscous stresses associated with wall shear, and fluid flow shear. In most practical situations involving reactor safety applications the latter may be neglected. Similarly the fifth tenns in the energy equation represent a wall and fluid-fluid conduction heat flux. The wall shear and wall conduction heat flux tenns are denoted by Fwk and Q k respectively in Sections 5.6.3 and 5.6.4. W

The sixth tenns in the momentum and energy equations denote turbulent stresses and heat flux. The seventh tenns in these equations, Frt' �, denote momentum and energy exchange due to mass transfer to phase k. The eighth tenns in these equations, Felk ' QeIk' denote interphase momentum and energy transfer between the phases.

Thermal Hydraulic Models

53

The ninth term in the energy equation. QI)t' denotes energy dissipation due to frictional effects. The aim in modelling is to close the above set of equations by means of various closure assumptions and to form a set of partial differential equations in a relatively small number of principal flow parameters. Twelve unknowns that are commonly chosen include the void fraction a. the pressure p. the vapour and liquid densities Pv and Pl' the three components of each velocity � and li and the specific internal energies e and el . v The remainder of the terms are usually modelled as functions of these variables and also of the temperatures Tv and Tr The total system of equations is finally closed by recognising the existence of equations of state:

Equations similar to these have been discussed in Harlow and Amsden [5.6] . 5.5.2 Additonal Field Equations

Incondensable gases are modelled assuming that the gas e.g. nitrogen is in thermal and dynamic equilibrium with the vapour phase. For each additional incondensable added an additional mass continuity equation has to be solved:

where P. denotes the mass of incondensable per unit volume. Similarly in the liquid phase a solute field for tracking boron concentration has been added to some models. The amount of dissolved or separated boron could affect reactivity feedback. If the solute concentration exceeds the orthoboric acid solubility at the fluid temperature of interest models must allow for the boron to plate out A solute moving with the liquid field may be represented by: a/iJt ( l -cx)mp/at + V.[( l -a)mPlli] = rb

where m denotes the boron concentration per unit mass of liquid. Additional conservation equations such as these are found in the TRAC code [5.7] .

Thermal Hydraulic Models

54

5.6

Constitutive Relations

In two-phase flows, constitutive relations are required to approximate the infonnation that has been lost through the averaging. Approximations are required for wall frictional forces, heat transfer and also interfacial transfers between the phases. These approximations will necessarily have to be in tenns of averaged variables and will also be flow regime dependent. Separate effects experiments will generally be required to supply the empirical infonn ation. For practical engineering purposes, the two phases may be in a large number of fonns and in different geometries. The constitutive relations are likely to depend on the thickness of liquid and vapour films, the pressure drops and the distribution of droplet size, the proportion of bubbles in water etc. Flow regime maps are produced where different regimes, Ishii [5.8] , may be characterised in tenns of mass flow rates and vapour fractions. Kelly [5.9] uses a different approach by attempting to fonnulate a model that applies over different flow regimes. Some example constitutive relations that are used in system codes are discussed below. 5.6.1 Flow Regime Maps

Flow regime maps have been produced for vertical flows in vertical pipes, horizontal flows in horizontal pipes and other circumstances. (a) Vertical Flow Maps Two-phase vertical flow maps pre- Critical Heat Flux (CHF) consist of three or four regimes, bubble, slug, annular or annular mist Such regimes have been identified in Taitel and Dukler [5. 10] , [5. 1 1] and Ishii [5. 12] , [5. 1 3] . In the post-CHF region inverted annular, inverted slug and mist (dispersed droplet) regimes have been identified by Ishii. Figure 5 . 1 shows the vertical flow regime map for the RELAP5 code. (b) Horizontal Flow Maps Criteria for horizontal stratified flow have been identified by Taitel and Dukler [5. 1 0] . (c) Other Flow Maps Other maps have been produced in special circumstances e.g. for flows in pumps. 5.6.2 Interphase Drag

Interphase drag is frequently modelled via expressions of the form: Felk = where

Cd

-Cd

I vII; -Vj I (VII; -v)

is the interfacial drag coefficient.

Cd

is flow regime dependent.

Thermal Hydraulic Models

55

The dispersed bubbles and droplets in bubbly or annular mist flow are typically assumed to be spherical particles with some size distribution e.g. [5. 14] . The drag then depends on standard expressions for spheres. The maximum diameters of bubbles are usually detennined by a critical Weber number criterion. Slug flow in pipes is nonnally modelled by assuming that large and long bubbles are separated by liquid slugs with small bubbles. A drag coefficient has been given by Ishii and Chowla [5 . 1 5] . Annular or annular mist flow is characterised by a liquid film along the wall with a core of vapour with entrained droplets. Correlations have been given by Bharathan [5. 16] . In horizontal stratified flows i t i s assumed that the liquid and vapour phases are separated by a well defined horizontal boundary. The interface Reynolds number is usually defined in tenns of vapour properties, assuming that liquid is the continuous phase. Inverted flow regimes are treated in a similar fashion to the corresponding flow regimes where the phases are not inverted, with vapour and liquid interchanged.

7-

BOilin g Reg ime

Posl-d-yout

t

Inc:reosi-q Mixture Velocity

Increasing vo id fraction

FIGURE 5.1 VERTICAL FLOW REGIME MAP

56

Thermal Hydraulic Models

5.6.3 Wall Friction

Wall friction tenns for the k- phase Fwk ' are usually given by: Fwit = cwk

I

vIt I vIt

where c wk are wall shear coefficients for the kth phase respectively. A standard approach to two-phase wall friction is to use a known two-phase wall friction correlation e.g. the modified Baroczy correlation [5. 1 7] and to apportion it to the liquid and vapour phases. One way of doing this is via a technique derived from the Lockhart-Martinelli model [5. 1 8] . Chisholm [5. 1 9] developed a theoretical basis for this model. In order to describe the wall shear stress, the surface area wetted by each phase must be known, this may be derived from the wall film volume fractions for each phase. There are numerous correlations for friction factors e.g. Moody, Colebrook etc. Losses associated with abrupt area changes, e.g. spacer grids, are usually accommodated via mechanistic fonn loss models. 5.6.4

Wall Heat Transfer

Wall heat transfer tenns for the kth phase, QWIt are given by:

where AWk denotes the heated wall areas, hwk are wall heat transfer coefficients of the kill phase. These are flow regime dependent To define the wall heat transfer various boiling heat transfer regimes are usually defined. These fall into three main regimes, pre-CHF, CHF and post-CHF regimes. The pre-CHF regime consists of single phase liquid, subcooled and saturated nucleate boiling sub-regimes. Heat transfer to the vapour is essentially zero. Heat transfer to the liquid may be either by forced convection (e.g. Dittus Boelter [5.20] or by natural convection, for example depending on a Grashof-Reynolds criterion. Nucleate boiling is frequently modelled using the Chen correlation [5.2 1 ] . The post-CHF regime can be broken down into transition film boiling, film boiling and single phase vapour. A mechanistic model has been developed by Tong and Young [5 .22] which together with a liquid contact area model due to Chen can be used to describe transition film boiling and the film boiling regime. Other models in the transition boiling regime are the Biasi CHF correlation [5.23] and for film boiling the Bromley and

Thermal Hydraulic Models

57

Dougall-Rosenow correlations [5.24] . Dittus Boelter may be used for the single phase vapour regime as for the single phase liquid. In the condensation regime, heat transfer to the wall from the liquid and the vapour are dependent on the flow regime. Heat transfer from liquid to the wall is often modelled by a weighting based on void fraction between a low void volume value and a condensation heat transfer coefficient in the high void volume. 5.6.5 Interphase Mass Transfer Interphase mass transfer is frequently modelled as consisting of two components, mass transfer at the wall and mass transfer in the bulk fluid:

Models for the mass transfer at the wall have been given by Chen. For boiling and condensation processes the interfacial mass transfer is calculated from the interfacial transfer. Interphase heat transfers may be typically defined as:

where Ad is the interfacial area, T. is the saturation temperature, and the heat transfer coefficients are flow regime dependent. The interfacial mass transfer rate is then calculated by a thermal-energy jump relation:

where A is the latent heat. The interfacial mass transfer in the bulk fluid can be modelled according to the flow regime. For example, for the bubble flow regime, Piesset and Zwick [5.25] have developed a bubble growth model. Other correlations are used for other flow regimes. For condensation processes the Unal [5.27] bubble collapse model has been developed for the bubbly flow regime and Theofanous [5.28] has given a model in the annular mist flow regime. 5.6.6 Reflood Heat Transfer During the reflood phase e.g. in a Large Break LOCA, the pressures and mass flow rates relatively low. This has led to special treatment of the reflood phase in many computer codes, particularly in the transition and nucleate boiling regimes. The heat flux at the incipience of boiling may be derived from models developed by Bergles and Rohsenow [5 .27] . For wall temperatures higher than the incipience of boiling the heat flux may be computed by a composite relationship involving convective heat transfer to fluid and are

58

Thermal Hydraulic Models

a nucleate boiling component e.g. McAdams. The modified Zuber correlation may be used to determine the critical heat flux. For wall temperatures higher than the critical heat flux temperature the heat flux can then be computed by a composite relationship involving convective heat transfer from the wall to the gas phase e.g. Dougall-Rohsenow [5.24] and a transition boiling heat flux e.g. Weisman et al. The quench temperature between the transition and film boiling regimes has been given in a model due to Arrieta and Yadigaroglu. Heat transfer in the film boiling phase may be calculated via a composite of the Weisman et al and modified Bromley corrrelations [5.30] . At higher wall temperatures radiation heat transfer may have a significant effect. The Sun et al model assumes that the vapour-droplet mixture is a transparent medium and describes the radiation heat flux between the wall droplets and vapour.

5.6.7 Turbulence Modelling In a lumped parameter approach, it is assumed for the temperature and spatial averaging, that the variables are constant within a control volume. The lumped parameter approach is suitable if, for example, the lateral scale is commensurate with the scale of many lateral subchannels, Figure 5.2. When the control volume of interest is small the effect of turbulence at the boundaries in the mass, momentum and energy transfer is significant. This would be true if the control volume represented part of a subchannel and the ratio of boundary surfaces to internal surfaces is high. Turbulent mixing has been well studied in single phase flows. In two phase flows the more open areas tend to have a higher vapour concentration, i.e. vapour tends to diffuse to unobstructed areas. This phenomenon usually has to be modelled separately in two phase flows. Another complicating factor of two-phase flow turbulent transfer is that no net mass transfer takes place due to turbulence. However in two-phase flows this is not so and the concept of equal mass exchange must be replaced by an equal volume exchange model. An approach for two-phase flow models has been to develop models which if either the liquid or vapour phase is solely present then the two-phase mixing model reduces to the appropriate single-phase mixing model. There a number of possibilities for apportioning the mixing between the phases. The momentum exchange rate, often formulated as a Reynolds shear stress in single-phase flow, has been extended to two phase flows by Lakey on the basis of experimental work

59

Thermal Hydraulic Models

CORE -

-

AXIAL V L.,.oo� I,; -- LATE

ff

F U E L ASS E M B LY

I NT E G RAL COD E C E LL

CAN B E I G N O R E D

\

S U B C H AN N E L

T U RB U LANT M IX I N G CONTROL VO LU M E

o 0 �IJ-----f"'"I

I M PORTANT

FIGURE S.2 TURBULENCE MIXING AND MODELLING SCALES

60

Thermal Hydraulic Models

Gonzales-SantaIo and Griffith [5. 3 1 ] . The influence of the flow regime has been considered by Beus [5.32].

5.6.8 Choked Flow (a) Subcooled Flow During a subcooled blowdown phase, the water will change phase at the break since the primary circuit pressure is much higher than the containment pressure. This change of phase was discussed by Lakey and leads to a discontinuity in the sound speed at the break. The homogeneous sound speed can be calculated by maximising the mass flux using a method of characteristics.

(b) Two-Phase Flow Models have been developed by Ransom and Trapp which are based on the assumption of thermal equilibrium between the phases. These have been extended to incorporate an additional inert gas phase e.g. as in the TRAC code. Choking can be calculated using similar methods as above. (c) Single-Phase Vapour The choking velocity can be calculated by assuming that the vapour expands isentropically and then by maximising the mass flux. (d) Transition Regime Since there is a discontinuity in the sound speed during the transition from liquid to two­ phase flow the flow during the transition regime must be carefully analysed. However the transition from the two-phase regime to the vapour regime is smooth since the two­ phase solution approaches the homogeneous-equilibrium limit as the void fraction approaches unity.

REFERENCES 5.1 5.2 5.3 5 .4

5 .5

S Fabic, P S Anderson: Plan for Assessment of Best-Estimate LWR Systems Codes, NUREG-0676, 198 1 G B Wallis: One-Dimensional Two-Phase Flow, McGraw-Mill, 1 969 E D Hughes: Macroscopic Balance Equations for Two-Phase Flow Models, Nuel Engng Des 54, 1979, 239-259 M Ishii: Thermo-Fluid Dynamic Theory of Two-Phase Flow, Collection de la Direction des Etudes et Recherches D'Electricite de France (Eyrolles, Paris, 1975) C Truesdell: Mechanical Basis of Diffusions, J Chern Phys, 37, 1 962, 2336

Thermal Hydraulic Models 5.6 5.7

5.8 5.9

5.10

5.1 1

5.12 5.13 5 . 14 5. 15 5.16 5.17 5.18

5.19

5.20

5.21 5.22

5.23

61

F M Harlow. AA Amsden: Numerical Calculation of Multiphase Huid How. J Comp Phys. 1 7. 1975. 19-52 "TRAC-PFI/MOD I : An Advanced Best-Estimate Computer Program for PWR Thermal-Hydraulic Analysis". LA- I O I 57-MS . NUREG/CR-3858, Los Alamos National Laboratory. ( 1 986). M Ishii, K Mishima: ''Two-Fluid Model and Analysis of Interfacial Area", ANL80- 1 1 1 , ( 1 98 1). J E Kelly, M S Kazimi: Interfacial Exchange Relations for Two Huid Vapour­ Liquid Flow: A Simplified Regime Map Approach, Nuclear Science and Engineering, 8 1 , ( 1982), 305-3 1 8. Y Taitel and A E Dukler, "A Model of Predicting Flow Regime Transitions in Horizontal and Near Horizontal Gas-Liquid Flow", AIChE J, Vol 26, pp 47-55, (1 976). Y Taitel, D Bornea, and A E Dukler. "Modelling How Pattern Transitions for Steady Upward Gas-Liquid How in Vertical Tubes". AIChE J. Vol 26. pp 345354. (1 980). M Ishii and T C Chawla. Local Drag Laws in Dispersed Two-Phase How. NUREG/CR- 1 230. ANL-79- 105, ( 1 979). M Ishii and K Mishima. Study of Two-Fluid Model and Interfacial Area. NUREG/CR-1873, ANL-80-111. (1 980). G B Wallis, One-Dimensional Two-Phase How, McGraw-Hill Book Company. New York. ( 1 969). V H Ransom et al. RELAP5/MOD I Code Manual, Volume I . NUREG/CR- 1 826. EGG-2070, (March 1982). D Bharathan. H T Richter and G B Wallis, Air-Water Counter-Current Annular Flow in Vertical Tubes. EPRI. NP-786. ( 1 978). K T Chaxton, J G Collier. A J War, HlFS. Correlation for Two-Phase Pressure Drop and Void Fraction in Tubes. AERE-R7 162. ( 1972). R W Lockhart an d R C Martinelli, "Proposed Correlation o f Data for Isothermal Two Phase, Two-Component How in Pipes". Chern Engr Prog. Vol 45(1). pp3948. ( 1 949). D Chisholm. "A Theoretical Basis for the Lockhart-Martinelli Correlation for Two-Phase How". J Heat-Mass Transfer. Vol 10. pp 1 767- 1 778. Pergamon Press Ltd. ( 1967). Great Britain. F W Dittus and L M K Boelter. "Heat Transfer in Automobile Radiators of the Tubular Type", Publications in Engineering. University of California. Berkeley, 2. pp 443-46 1 . ( 1 930). J C Chen. "A Correlation for Boiling Heat Transfer to Saturated Huids in Convective How". Process Design Development, 5. pp 322-327. ( 1 966). L S Tong. and J D Young. S "A Phenomenological Transition and Film Boiling Correlation". Proceedings of the 5th International Heat Transfer Conference. Vol IV. B 3 .9. Tokyo. ( 1 974). L Biasi, G C Clerici, S Garribba, R Sala and A Tozzi "Studies on Burnout Part 3," Energ. Nucl. (Milan) 14, 530-536 (1967).

62 5.24

5 .25 5.26 5.27

5 .28

5 .29 5.30 5.3 1 5.32

Thermal Hydraulic Models R S Dougal and W M Rohsenow, "Film Boiling on the Inside of Vertical Tubes with Upward Flow of the Fluid at Low Qualities," Massachusetts Institute of Technology Mechanical Engineering report 9079-26 9 1 963). M S Plesset and S A Zwick, "Growth of Vapour Bubbles in Superheated Liquids", J Applied Physics 25 (4), 493-500 , (1954) F Krieth, Principles of Heat Transfer, Intext Press Inc, ( 1973). H C Unal, "Maximum Bubble Diameter, Maximum Bubble-Growth Time and Bubble Growth Rate During the Subcooled Nucleate Flow Boiling of Water Up to 17 MN/m1" Int J Heat Mass Transfer, Vol 19, pp 643-649, ( 1976). T G Theofanous, "Modelling of Basic Condensation Processes' The Water Reactor Safety Research Workshop on Condensation, Silver Springs. MD, May 24-25, ( 1 979). A E Bergles and W M Rohsenow, ''The Detennination of Forced-Convection S urface Boiling Heat Transfer", J Heat Transfer, Trans ASME, 86, 365, ( 1964). L A Bromley, "Heat Transfer in Stable Film Boiling", Chern Eng Progr 46, pp 22 1-227. ( 1950). Gonzalez-Santalo, J M; Griffith, P: Two-Phase Flow Mixing in Rod Bundle Subchannels, ASME, 72-WA/NE- 19, ( 1 972). Beus, S G: Two-Phase Turbulent Mixing Model for Flow in Rod Bundles, WAPDT-2538, ( 1 970).

63

Chapter 6 COMPONENT HEAT TRANSFER

6.1

Introduction

This chapter considers the modelling of heat transfer within and between the various structures present in a nuclear reactor under accident conditions.

6.2

Heat Conduction in Structures

Various structures are provided for in most computer codes to permit calculation of the heat transferred across solid boundaries that interface with hydrodynamic volumes. Most codes accommodate generic heat transfer models for these structures which include fuel pins or plates with nuclear or electrical heating, heat transfer across steam generator tubes, and heat transfer from pipe and vessel walls. In many situations structures can be represented by one-dimensional heat conduction in rectangular, cylindrical, or spherical geometry. Geometry dependent multipliers may be used to link the unit surface of the one-dimensional calculation to the actual surface of the heat structure. Thermophysical properties for the various materials present in the core are required. These properties include temperature-dependent thermal conductivities, specific heat capacities and densities. These are considered in Chapter 19. The spatial dependence of the materials may be accounted for by using finite difference or finite element methods to advance the heat conduction solutions. The mesh or grid may contain differently spaced intervals with different materials. Internal heat sources which may vary spatially can also be easily modelled by this means. The heat source may be obtained from reactor kinetics, from a decay heat curve or by other means e.g. input directly for the purposes of modelling an experiment. Boundary conditions are required to link the structural heat conduction equation to hydrodynamic volumes or to model symmetry or insulated boundaries. For heat structure surfaces connected to hydrodynamic volumes, a heat transfer package containing correlations for forced or natural convection, nucleate boiling, transition boiling, and film heat transfer etc between the wall and the fluid will be needed. This has already been discussed in the previous chapter. The general form of the heat conduction equation is: a/at pC T - VlC VT = S p

where p is the density, C is the specific heat, lC is the thermal conductivity, S is the internal p

64

Component Heat Transfer

heat source, and T is the temperature. The thermophysical properties are generally temperature and spatially dependent Boundary conditions at an external surface may be written in the form: aT + bOT/an = c where a, b, c are temperature dependent coefficients and n denotes the unit normal vector away from the boundary surface. In particular either the boundary heat flux may be described e.g. -l(

aT/an = h(T-T ) .

where h is the heat transfer coefficient or the surface temperature T= T

•

where T and T , denote the surface and sink: temperatures respectively. . The prescription of valid boundary conditions is dependent on the dimensionally of the model, the topology, and whether the physical situation is steady-state or transient. For example; in one dimensional problems, boundary conditions have to be applied on the left and right surfaces. In steady-state problems a valid physical problem requires that the temperature is prescribed on at least one boundary surface. In both transient or steady­ state, in cylindrical and spherical geometry there is a singularity at the centre of zero radius which requires special treatment. In this case aT/an is usually set to zero. Reference 6. 1 shows a typical mesh layout. In between left and right boundaries zones are typically occupied by different materials; these will in turn contain different numbers of mesh points of the finite difference grid. In deriving the numerical approximations in many of the cells it may be assumed that there is continuity of temperature and heat flow across all interfaces. However, a contact resistance interface condition must be modelled with care since the temperature gradient will no longer be continuous across the interface, although the heat flux will continue to be continuous. Once assembled the difference approximations for the mesh points lead to a sparse matrix: e.g. a central difference approximation gives a tri­ diagonal set of equations. In the case of heat flux boundary conditions the coefficient matrix is symmetric. The solution to the above equations may be obtained by a number of efficient matrix equation solution algorithms, many of which exploit the sparcity of the coefficient matrix. The system of equations can be solved using either direct methods e.g. Gaussian Elimination or by direct iterative methods. The latter is possible since the off-diagonal elements are

Component Heat Transfer

65

negative and the magnitude of each diagonal element is greater than the sum of the magnitudes of the off-diagonal elements. Since the topology is invariant these conditions are satisfied for any values of the mesh point spacing, time step, and thermal properties. The thermal conductivity, K, and the volumetric heat capacity, C , are generally functions of p temperature and space. Spatial variations can be accommodated in numerical models via appropriate interpolation. However, when these parameters are temperature-dependent, iterations may be necessary to resolve the difficulty of obtaining thermal parameters as a function of temperature when the temperatures are unknown. The Finite Element Method (FEM) for heat conduction has been used in the SCDAP/ RELAP5 code [6. 1 ] . This method involves approximating temperature by dividing the solution region into a finite number of regions called (elements), Huebner [6.2] , Zienkiewicz e.g. [6.3] and Cook [6.4]. Temperature is expressed in terms of an assumed approximating function for each element. The nodal temperatures are the desired unknowns and, using the FEM, the entire temperature distribution can be determined by assembling the contributions of each element. The Galerkin Method of Weighted Residuals is used in SCDAP/RELAP5. This method assumes a general functional behaviour for the temperature w ithin an element. Substitution of the approximation into the differential equation governing heat conduction and into the boundary conditions results in residual error. The product of this residual and the weighting function is required to vanish in an average sense over the solution region. Mter the element contributions are determined, the element contributions are assembled to obtain a set of simultaneous linear equations for the entire solution region. These equations are then solved yielding the temperatures at each node point.

6.3

Heat Conduction Under Reflood Conditions

Two-dimensional heat conduction schemes are used in reflood models. Axial and radial heat conduction is required to be modelled in fuel rods i.e. cylindrical structures and also other rectangular heat structures. The resulting difference equations may be solved by a variety of methods: the coefficient matrices are diagonally dominant for standard central difference schemes in the same manner as described for the one-dimensional heat conduction case described earlier. Alternative-Directional Implicit (ADI) solution methods can be used. Some codes [6.5] employ a fine mesh-rezoning scheme to efficiently use the two­ dimensional conduction solution for reflood calculations. The schemes are intended to resolve the large axial variation of wall temperatures and heat fluxes. The number of axial nodes in the heat structures may be varied in such a way that the fine nodes exist in regions where there are marked temperature variations e.g. in the nucleate boiling and transition boiling regions.

66

Component Heat Transfer

Figure 6. 1 shows a typical heat-structure geometry with one fluid control volume connected to each heat structure. The dots denote radial mesh points. At the initiation of the reflood model, each heat structure is subdivided into two or more axial intervals. A two-dimensional array of mesh points is thus fonned. Afterwards, the number of axial intervals may be doubled, halved or unchanged at each time step according to a prescribed set of rules.

H E AT STR U CT U R E S

n

NODES

/ �

n -1

F LU I D CE LLS

I I I I I

� �

I---� - - - - - - - -

-�

.. .. L-_�

-- - ------

I I I I

-4

I I I I I I I I i------, - - - - - - - - -t I I I I I I I I I-----f - - - - - - - -t I I I I I I

. I I

I

I I

I

���� -------- � I I I I I

I I I

���� --- - - - - - � I I I I I I I I ���� - -- - -- - - � I

I I I I

I

...J

e---4....� ... ... --- - - - _ -J

P R I O R TO R E F LO O D

D U R I N G R E F LOOD

'--_---'

______ _

FIGURE 6.1 HEAT STRUCTURE RE-ZONING FOR REFLOOD

In codes such as RELAP5 the number of axial mesh intervals in a heat structure depends on the heat transferregimes in the heat structure. Where the mesh may be subdivided is typically dependent on the vicinity of TCHF, the wall temperature where the critical heat flux occurs, TQ, the quench or reweuing temperature, and TIB . the wall temperature at the incipience of boiling.

Component Heat Transfer 6.4

67

Fuel and Cladding Energy Transfer

Fuel rods are special structures and some of their special features are considered below. In most codes the temperature distribution throughout the fuel and cladding is calculated at a number of different axial levels or nodes. In some circumstances, e.g. to model clad ballooning it may be necessary to calculate a two-dimensional (radial plus azimuthal) temperature distribution for each axial node. It may also be necessary to perform calculations of radiative heat transfer, both from fuel pellet to fuel cladding and from the surface of pins to neighbouring pins and other structures. Heat transfer coefficients and subchannel fluid temperatures obtained from the thermal hydraulics calculations supply the boundary conditions for the conduction (or radiation) solution. Heat generation in the fuel, and also possibly from the cladding will also need to be prescribed. The thermal properties of the fuel and the gap conductance need to be calculated. In codes such as MABEL, the equations are linearised and depend on the appropriate node temperatures from the previous iteration or timestep. The gap conductance calculation also depends on the most recently calculated value of gap size. All of the variables required for the solution of the temperature distribution can then be determined, and the solution obtained using standard numerical methods. The radiative heat transfer across the pellet-to-cladding gap forms one part of the total across-gap heat transfer. The effective radiative heat transfer coefficient can be evaluated from the radiative heat flux and this heat transfer coefficient forms one component of the gap conductance. The radiative interchange at the outer surface of the fuel rod can be regarded as a source or sink term in the heat balance performed on the cladding nodes. Heat generated by the reaction of Zircaloy cladding with steam may also be included. This phenomenon is considered later. In order to calculate the correct gap conductance it may be necessary to take account of thermal expansion of the fuel. The models used in the fuel and cladding temperature determination generally involve a number of assumptions and limitations, the most important of which are usually: (i) no convective mode of heat transfer occurs across the pellet-to-cladding gap; (ii) no 'solid conductance' term is included in the calculation of gap conductance. Some of the commonly used models included in the fuel and cladding heat transfer codes considered below.

are

6.4.1

Heat Generation in the Fuel

A basic requirement in all computer models is to prescribe heat generation distribution in the fuel as a function of time. For a reference fuel rod this is typically comprised of various components:

68

Component Heat Transfer

(a)

the power conditions at the start of the transient;

(b)

a model for the fuel rod power history e.g. from a reactor physics code or from a decay curve for the case of a transient or LOCA where the reactor is scrammed;

(c)

axial and radial power distributions in the fuel rod;

(d)

information on the power tilt: for a reactor at power, the power tilt in a nuclear fuel rod arises from the flux tilt.

The total energy released for U235 fission is the sum of the following components: (i)

� decay energy

(ii)

fission fragment kinetic energy

(iii)

y decay energy

(iv)

y prompt energy

(v)

neutron energy

(vi)

U238 neutron capture energy.

The first two comprise about 90% of the total energy per fission and these energies are produced at the site of fission. Under shut down conditions, y heating is a major contribution to the total heat generation accounting for about 50 % of the decay heat. Gamma energy is distributed over a wide area and in these circumstances therefore, the power tilt is only about 50% of the original flux tilt This change must be taken into account in modelling short timescale accident scenarios such as large break LOCAs.

6.4.2

Gap Conductance

The conductance between the fuel pellets and the cladding in a fuel rod is important in determining centre fuel rod temperature. It is an important model in fuel performance codes such as HOTROD [6.6] and MINIPAT [6.7] under nominal power conditions and at short timescales following reactor scram . The conductance depends upon: gap width the gas composition and temperature in the gap.

69

Component Heat Transfer

Models have been proposed by Ross and Stoute [6.8] and used in codes such as MABEL [6.9] and HOmOD [6.6]. Most of the codes assume that the mechanism of heat transfer across the gap is only via conduction through the gap and by radiation. If the gaps are sufficiently large e.g. under cladding ballooning conditions then gas natural circulation may also be an important mechanism though as stated earlier this is not normally included. The conductance is then modelled as the sum of conductive and radiative components:

can be found from the literature see for example Mason and Saxena [6. 10] . Radiative gap conductance can be calculated using radiative heat transfer models appropriate for heat transfer between concentric cylinders [6.9] .

Gas mixture thermal conductivities

-

ZIRCALOY C LADD I NG

./

/ :. '

.:

REG I ON

'

.

. . ..

:

.

STEAM

: : r FUEL RODS ..

.

::- / !

:/ : : . . /' : - : 0 °

.

/ :.

. • •• •

BUN D L E WA LL

STE A M

B U B B L ES

. 0 o •

SUBCO O L E O REGION

INLET

F LO W

FIGURE 6.2 FUEL BUNDLE IN THE PROCESS OF BOILING DRY

70

6.4.3

Component Heat Transfer Tbermopbysical Properties

Data for the thermal conductivity, heat capacity and density of the fuel rod materials (U02' Zr) are described in Chapter 19. These physical properties are usually correlated with temperature. The calculation of the thermal conductivity of the fuel pellets may be modified to allow for the effects of porosity on thermal conductivity. The effects of both microporosity and macroporosity are included in fuel performance codes. Thermophysical properties for all the main fuel rod constituents are discussed in Chapter 19. In general temperature-dependent thermal properties introduce non-linearities into the heat conduction equation. These are usually evaluated using the temperatures obtained in the previous iteration or timestep.

6.4.4

Cladding Oxidation Heating

At higher temperatures e.g. if the fuel rods become uncovered, Figure 6.2, the oxidation of Zircaloy cladding by stearn must be accounted for in the conduction model. The kinetics of the phenomenon are discussed in Chapter 8. The heat generated by the Zircaloy-steam reaction (approximately 6.45 x I ()6J/kg zirconium) is a major component of the total heat generation rate at higher core temperatures under severe accident conditions. Oxidation of stainless steel cladding of control rods also takes place once the temperature rises sufficiently high. This phenomenon is also considered in Chapter 8. According to thermodynamic considerations, this reaction provides insignificant heat compared with the Zircaloy/steam reaction and therefore has little impact on core heat-up. The phenomenon though does have importance from the point of view of hydrogen produc tion .

6.4.5

Fuel Rod Boundary Conditions

The convective heat transfer mechanisms by which heat is convected away from the fuel rod cladding boundary have already been discussed in the thermal-hydraulics Chapter 5 . The calculation of radiative heat losses requires the specification of surface emissivities, view factors and absorptivities in the fluid medium surrounding the pins. Core radiation heat transfer models are discussed below.

6.5

Radiation Models

A feature of heat transfer in the core region is the increasing importance of radiation heat transfer between rods if temperatures start to exceed l 000K . This radiation heat transfer should be and is included in most severe fuel damage codes. To model radiation in the complex geometry of rod bundles in an exact mechanistic manner is very difficult.

71

Component Heat Transfer

The absorption spectrum for steam varies with temperature and i t i s thus very difficult to determine the absorption through bundles of widely varying temperature. Even for isothermal conditions, view factors and optical path lengths are difficult to define particularly in three dimensions and during core degradation away from nominal geometry. Radiation is modelled on a two-dimensional basis using the net method [6. 1 1 ] in many codes. This has been used in a number of applications [6. 1 2] - [6. 16] in the context of nuclear reactor heat transfer. In this method, each component (fuel rod, control rod, or shroud) surface fonns one side of an enclosure with a finite number of sides and the enclosure is filled with the medium (see Figure 6.3). The equation of radiation heat transfer is derived for each surface and describes radiation exchange between all the surfaces (including itself if it radiates to itself), and absorption and emittance by the enclosed coolant. The resulting equations are solved simultaneously by a matrix inversion method to obtain the radiosity (the sum of emitted and reflected radiation energy rates) of each surface. The difference between the radiosity and incident energy from the surroundings gives the net heat flux to or from a surface. The algebraic sum of net heat flux corresponding to each surface gives the total radiation heat absorbed by the coolant. However two-dimensional models in a plane normal to the centre axis of the vessel neglect the important heat transfer out of the top of the core to upper plenum internals and down to the water pool . The former are important in detennining such issues as fission product retention in the above core regions. The latter, radiation heat transfer to the water pool , affects the steam generation rate; particularly important during the latter stages of core heat-up when steam starvation will be reducing the oxidation and therefore core heat-up rate.

REFLECTION CF I N C I DENT RAY E M I S S I ON

t

RAD IANT

ENERGY

SU RFACE

K

D I R ECT EM ISSION FROM Oi"HER SURFACES

!

L E AV I N G

EM I S S I O N

oY VAPOUR

PHASE

I

,

EM ISS I O N BY L I Q U I D PHASE

�

R AD I ANT E N ERGY I N C I OErn G N S U R FA C E K ENCLOSURE

OF SU RFACe:S

FIGURE 6.3 RADIATION NET ENCLOSURE MODEL

Component Heat Transfer

72

These axial effects are treated in some codes e.g. MARCH2 and also the MELPROG code. The latter model is probably the most advanced among models envisaged for the severe fuel damage codes. It is based on a combination of the net enclosure method and a diffuse treatment between computational cells. Each cell in a 2-D grid is treated as a net enclosure with internal structures. To detennine net radiation fluxes the codes usually average over the whole wavelength spectrum assuming that the surfaces are grey and assuming an idealised spectrum for steam absorption.

REFERENCES

6. 1 6.2 6.3 6.4 6.5 6.6 6.7

6.8 6.9 6.10 6. 1 1 6. 12

6. 13 6. 14 6. 1 5

6.16

SCDAP/RELAP5 Code Manual, NUREG/CR-5273, June 1 989. K H Huebner, The Finite Element Method for Engineers, John Wiley & Sons, 1975. 0 C Zienkiewicz and Y K Cheung, The Finite Element Method in S tructural and Continuum Mechanics, McGraw-Hill Publishing Company Ltd, 1 967. R D Cook, Concepts and Applications of Finite Element Analysis, John Wiley & Sons, 1974. J M Kelly, Quench Front Modelling and Reflood Heat Transfer in COBRA-TF, ASME Winter Annual Meeting, 79-WA/HT-63, New York, 1979. L F A Raven, Comparison of HOTROD Code Predictions with PIE Data, ANS Meeting on Water Reactor Fuel Perfonnance, Chicago, May 1 977. R Hargreaves and D A Collins, A Quantitive Model for Fission Gas Release and Swelling in Irradiated Uranium Oxide, Journal British Nuclear Engineering Society, Vol 1 5, pp 3 1 1 -3 1 8, 1 976. A M Ross and R L Stoute, Heat Transfer Coefficient Between UO, and Zircaloy-2, AECL- 1552, 1962. R W Bowring and C A Cooper,MABEL2: A Code to Analyse Cladding Defonnation in a Loss of Coolant Accident, AEEW-R I 530, June 1983. E A Mason and S C Saxena, Approximate Fonnula for the Thennal Conductivity of Gas Mixtures, Physics of Fluids 1 , 361 , 1958. R Siegel and J R Howell, Thennal Radiation Heat Transfer, McGraw-Hill Publishing Company., pp 550-558. J W Spore, M M Giles, and R W Shumway, A Best Estimate Radiation Heat Transfer Model Developed for TRAC-BD1 , ASME Paper NO 8 1-HT-68, 20th Joint ASME/AIChE National Heat Transfer Conference, Milwaukee, August 198 1 . C J Shaffer, Importance ofThennal Radiation to Steam in Rod Bundles, Topical Meeting on Water-Reactor Safety, Salt Lake City, Utah, March �28, 1973, pp 371-379. D A Mandell, A Radiative Heat Transfer Model for the TRAC Code, Los Alamos Scientific Laboratory, Report No NUREG/CR-0994 , LA-7965-MS , November 1979. J G M Andersen and C L Tien, Radiation Heat Transfer in a BWR Fuel Bundle Under LOCA Conditions, Fluid Flow and Heat Transfer Over Rod or Tube Bundles, S C Yao and P A Pfund eds, ASME, New York, NY , 1979, pp 197-207. K H Sun, J M Gonzales-Santalo, and C L Tien, Calculations of Combined Radiation and Convection Heat Transfer in Rod Bundle Under Emergency Cooling Conditions, J Heat Transfer, 98, 1976, pp 4 14420.

73

Chapter 7 MECHANICAL MODELS

7.1

Introduction

Under certain accident conditions, temperatures may rise to well in excess of normal operating levels. There will be an increasing tendency for high temperature components to creep, rupture, fail etc, particularly if they are under stress or when temperatures approach the material melting point This chapter reviews the range and scope of some of the mechanical models that have been developed for accident analysis. Much attention has focused on the behaviour of fuel rods and their response to changes in system pressure and temperature under accident conditions. Fuel rods are pre-pressurised with helium and under LOCA conditions there is a potential for the rods to deform and rupture, a phenomenon known as clad ballooning. This mechanical behaviour depends primarily on the stress, strength and temperature of the materials concerned [7. 1]. Since the fuel rod cladding material Zircaloy becomes ductile as temperatures increase, clad ballooning will OCCID' once the tensile stress across the clad induced by the pressure gradient exceeds the yield strength of the cladding. Fuel rod behaviour under accident conditions depends on the balance between the heat generated by nuclear reaction, heat generated by Zircaloy oxidation and the heat transfer mechanisms at the surface of the fuel rod. These, together with the thermophysical properties of the rods, enable the spatial temperature distribution to be determined as discussed in the previous Chapter 6. In general the mechanical strength, structural integrity and the susceptibility of the materials decrease and the potential for oxidation increases with increasing temperature. The behaviour of the fuel also depends on the irradiation. The cladding behaviour will be dependent on the cladding constituents: steel clad control rods will behave quite differently from Zircaloy clad fuel rods. However fuel rod cladding deformation will also be dependent on the exact constituents of the trace materials in the Zircaloy. Many experiments have been performed to determine clad deformation and rupture in both single rod and multi-rod bundle systems. The main objective of most of these tests has been to investigate the relationship between burst temperatures and pressures on the final strain. Corresponding codes have also been developed. The effects of cladding oxidation and resulting embrittlement have also been investigated. In the case of embrittlement, criteria have been developed which depend on the extent

74

Mechanical Models

of oxidation and the thennal shock loading, for example by a sudden core reflood. Under these conditions fuel fragmentation will also occur. There are fewer data on the response of other reactor components to external loadings. To determine the component response requires a knowledge of the ductility, toughness and strength of its material under the relevant accident conditions. This is particularly important under severe accident conditions where it is necessary to assess the integrity of vessel and primary circuit components to severe conditions of temperature and pressure. Nevertheless models are now being developed and these are also discussed briefly. 7.2

Clad Ballooning

Under nonnal operation PWR fuel rods are pressurised to a cold fill pressure -2.5 MPa. In the case of a system depressurisation the rods may, depending on the temperature transient, balloon outwards as the clad temperatures rise beyond 1 025K as the internal rod pressure exceeds the coolant pressure. The extent of the ballooning will depend on the detailed conditions including when clad failure occurs. ....

A number of consequences result from clad ballooning. The increased void in the fuel rod provides space for fuel relocation, either by collapse of solid material or through

120 100

o

I

D<·�B..

I I

80

z'

�

� z is o «

d

: "

.. q

I

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o

0

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ll.

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0

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O A-

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0

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[I]

•

1

S

10 30

K/s

o o

60

:;(

0:: � VI

f · ···0 ..30

6

·?:�b�

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�

I

TEMPWATURE RAMP RATE

40 20 aI

600

700

800

J3 900

1000

1 1 00

1 2 00

BOO

1400

D(

FIGURE 7.1 DEPENDENCE OF RUPTURE STRAIN ON TEMPERATURE IN A STEAM ATMOSPHERE

Mechanical Models

75

relocation as a liquefied melt: in the case of a severe accident it can also affect the extent of fueVclad mechanical or chemical interaction. Clad ballooning causes the subchannel areas for coolant flow to be reduced. Rod cooling may improve at modest diametrical clad strains because of the increased cladding surface area accessible to the coolant, but rod cooling at high clad strains may be adversely affected because the flow area reduces leading to a reduction of the flow rate particularly following rod-to-rod contact. Over the temperature range 1 100 - 1250K, Zircaloy changes from the hexagonal a-phase to the cubic, �-phase and the creep properties change. For large break conditions ballooning usually occurs in the high a or low a-� phase. Forpressurised sequences ballooning will take place, if at all. at higher temperatures i.e. in the higher mixed phase or possibly �-phase. In the a-� region rupture strains are reduced - Raff [7.3] , Figure 7 . 1 . - at the strain rates relevant to these transients making it unlikely that significant coplanar flow blockage would occur. In the �-phase. although large balloons have been observed in inert atmospheres. oxidation is likely to reduce the strains in the reactor so that again severe blockage is unlikely. Furthermore. at these higher temperatures, modest oxidation of the Zircaloy will have a strengthening effect on the cladding. Mechanical restraint from neighbouring rods will take place at higher strains i.e. > 33%. 7.3

Models

A state-of-the-art clad ballooning model has to satisfy a range of requirements. It first relies on the availability of an adequate database to supply the Zircaloy creep properties. These databases exist [7.4]. Additionally the model must take account of initial clad thickness perturbations, azimuthal temperature variation. mechanical restraint by neighbouring rods and the effects of oxide strengthening at the higher temperatures of interest in the pressuri sed accident sequences. The feed back on pin internal pressure during ballooning needs to be accounted for: as a fuel rod balloons. its pressure decreases. Thermal hydraulics boundary conditions must allow for blockage bypass effects to be included. for if not then the mass flows through ballooned regions remain unchanged, rather than in reality reduced. and the clad temperatures are under­ predicted. This is even more important for long balloons where cooling is worse than for short balloons. Under severe accident conditions. if clad temperatures are underestimated. then the position of the oxidation driven temperature excursion will be changed and the core heat-up rate and subsequent core melt behaviour will be affected. A state-of-the-art ballooning model capability is provided for by the CANSWEL-2 code. Thiscode [7.8] uses a two dimensional (r. 9) membrane theory model ata11 times during

the deformation history. There is a detailed treaunent of the effect of the mechanical restraint of neighbouring ballooned but unburst rods in slowing down the rate of defonnation. Azimuthal temperature variations and initial non-unifonnities in wall thickness are also treated. In early versions of the code. the effects of oxidation were taken into account using the model of Malen and Tarkpea [7.9] in which the strong external oxide layer bears most of the cladding stress. and hence reduces the strain rate. until the oxide cracks at a strain of

76

Mechanical Models

2 %; the oxide is then assumed to have no further effect This was found to be inadequate, particularly in the alpha-phase region, and furthermore could not be used to calculate the effect of oxidation in reducing rupture strains. A more detailed model has therefore been developed which calculates separately the strengthening effects of the outer oxygen-rich oxide and retained alpha layers. takes account of the experimentally observed cracking of such layers as strain proceeds (in the range 2- 10%), the stress and strain concentrations which occur at such sites, and the subsequent re-oxidation of the underlying metal. The early code included a local ductility rupture criterion, the MATPRO-II stress criterion, and a strain rate criterion. The last two have been found to be inconsistent with the results of Donaldson [7. 10] . who noted that in constant stress creep tests the rupture ductility was similar to that in constant pressure tests (where the stress increases rapidly with strain), in the alpha-phase region. In a constant pressure test both internal and external pin pressures are held constant whereas in a constant stress test the internal pressure is adjusted to keep the stress constant. Raff [7.3] also had demonstrated that burst strains are independent of strain rate (and therefore stress) in the same temperature range. The local ductility criterion has been validated [7. 1 1] in the alpha-phase region. but cannot yet account for the dependence of rupture strain on strain rate and temperature found in the phase change region [7.3], particularly at low strain rates when superplasticity is important The modified strain fraction rule of Raff [7.3] has therefore been introduced into CANSWEL-2. A coupled ballooning thermal-hydraulic capability has been obtained by coupling CANSWEL to a thermal-hydraulic code MABEL [7.2]. Other ballooning models in the literature include the localised deformation model BALON2 [7.6] and the FAR I probabilistic model [7.7] . In general the models for the severe accident codes are less well developed. The BALON2 model specifies a small variation in temperature in the axial and azimuthal directions near the ballooning location, and carries out a detailed membrane theory calculation taking into account axial radii of curvature. No account is taken of mechanical restraint of neighbouring rods in calculating this deformation. The FAR I model [7.7] is used to calculate flow blockage. 7.4

Mechanical Modelling or Structures

The need to analyse the effects of severe accidents has led to the development of structural mechanics models for the severe accident codes, see for example [7. 1 7] . The main aim of such models is to predict the failure potential of large in-vessel structures under extreme conditions of temperature and mechanical loading. The structures present in a representative large PWR are shown in Figure 7.2. These fall into various basic categories. These include cylindrical walls. plates with or without perforations. supported at regular intervals or around their periphery. There may also be special structures for different reactor types.

77

Mechanical Models

Upper VelSel Head

Head Flange

Vessel Wall

Core Barrel

Lower Core Grid Plate

Support Columns

Lower Vessel Head

FIGURE 7.2 PRINCIPAL STRUCTURES IN A WESTINGHOUSE PWR VESSEL 7.5

Design Characteristics

For structures under relatively low stress, high temperatures will be required for failure to occur. The mechanism will probably be via plastic or creep failure at temperatures close to the melting point of the structure. In a core melt down scenario structures might be heated and also ablated by molten debris, Figure 7.3. Failure at low temperatures could occur if a structure was subjected to high stress, resulting in a fracture failure. Under normal operating conditions, most internal structures in a nuclear reactor are under relatively little stress, albeit that they possibly support considerable weight. High stresses could be exerted on these structures following some event during the accident

78

Mechanical Models

such as an energetic steam explosion. The vessel itself, particularly in a PWR, contains the high internal pressure of the reactor coolant circuit. Therefore certain components are under high stress, such as the restraining head bolts, which might be particularly vulnerable under transient accident events [7. 1 8] . 7.6

Modelling Criteria

Mechanical behaviour models have been developed including detailed finite element and finite difference representations of the mechanics equations. These determine detailed stress-strain distributions in the structures, but are time consuming in terms of computer time. For the purposes of severe accident modelling [7. 17], it is typically assumed that little structural deformation takes place under the given loading conditions for most of the accident sequence. This means that for most of the time the stresses can be calculated assuming elastic deformation. The structures remain close to their initial geometry until failure is about to occur. In [7. 1 7] the stresses are calculated analytically at known peak stress locations: once failure is imminent these are then analysed assuming fully plastic conditions, thereby eliminating the need for an intermediate elastic-plastic calculation. Under these kind of assumptions the Larson-Miller Parameter (LMP) correlation provides a criterion for relating time-ta-failure to the stress and temperature of the structure. This failure criterion considers the stress under both plastic behaviour and high temperature creep conditions but is independent of strain. 7.7

Detailed Models

This section considers typical component models, that have been developed for the different types of structures present in a representative Westinghouse PWR. In such a plant, the core rests on a lower core grid plate which is supported by support columns, which are themselves supported by the lower core support plate. This is attached at its upper edge to the upper head flange. The bottom of the vessel is supported by the vessel wall. The upper core grid plate is attached to support columns from the upper core support plate which is supported at its periphery by the upper head flange. In [7. 1 7] the lower core grid plate is modelled by using analysis methods for a solid plate [7. 19] , [7.20] , taking account of the presence of holes by modifying the mechanical properties of the plate. The vessel top and bottom and the core support plate are modelled using an elliptical plate model (7.23), again taking account of perforations by modifying the mechanical properties. Cylindrical components are modelled via thick wall equations to provide stresses on the inner surface.

79

Mechanical Models

WALL AB LATION

PLATE

PLATE

MELT

PLATE ABLATION

LOWER H E AD

FlGURE 7.3 ABLATION OF STRUCTURES

Mechanical Models

80

Columns may be in either compression or tension. For compressive loads, the Euler buckling criteria has been used [7.24] . For tensile loads the stress is compared to the yield stress. An Egg Crate Grill model has been used for the upper core support plate in modelling a Westinghouse PWR and also for modelling the reinforced reactor support plate. The model assumes a plate supported on its periphery: the effect of the re-enforced grill is taken account of by increasing the plate rigidity.

REFERENCES

7.1 7.2 7.3 7.4 7.5 7.6 7.7 7.8 7.9 7.10

7. 1 1 7.12 7.13 7.14 7.15

F J Erbacher, A Review of Significant Safety Research Results o n Zircaloy Fuel Cladding Defonnation and Coolability of Defonned Rod Bundles in a LOCA, IAEA Meeting, Vienna, 1 0- 1 3 November 1986. R W Bowring, C A Cooper, MABEL 2: A Code to Analyse Cladding Defonnation in a Loss-of-Coolant Accident. AEEW-R I 530, June 1983. S Raff, Entwickl ung eines Defonnations und Versangsmodells fur Zircaloy im Hochtemperaturbereich zur Anwendung bei Kuhlmittelverlustofalluntersuchungen an Leichwasserreaktoren, KfK-3 1 84, November 1982. D L Hagrman et al, MATPRO: A Library of Material Properties for Light Water Reactor Accident Analysis, NUREG/CR-5273, January 1990. D A Powers and R 0 Meyer, Cladding Swelling and Rupture Models for LOCA Analysis, NUREG-0630, 1980. D L Hagrman, Zircaloy Cladding Shape at Failure (BALON2), EGG-CDAP5379, July 198 1. E R Carlson, Probabilistic Flow Area Reduction Model FAR 1 , EGG-CDD5567, November 198 1 T J Haste, CANSWEL-2: A Computer Model of the Creep Defonnation of Zircaloy under Loss-of-Coolant Accident Conditions, ND-R-8 14(S), Parts 1 -3, July 1982 April 1984. K Maten and P Tarkpea, Evaluation of Creep in Oxidising Zircaloy during Temperature Transients, CSNI Specialists' Meeting, Spatind, Norway, September 1976. A T Donaldson and G Knowles, Primary and Steady-State Creep of Westinghouse PWR Fuel Tube Material; Constant Stress Tests between 973 and 1 073K, TRPD/B/0267/N83, June 1983. T J Haste, Validation of CANSWEL-2 for Zircaloy 4 in the Alpha Phase Region using BNL Single Rod Burst Data, to be published, 1985. C L Mohr. LOCA Simulation i n NRU; Summary Materials Test 4, Battelle/PNWL, December 1982. E P Hindle, C A Mann, an d A E Reynolds, The Defonnation o f Zircaloy PWR Cladding with Low Internal Pressure, under Mainly Convective Cooling by Steam, ND-R478(S), August 198 1 . A T Donaldson, R A Horwood, and T Healey, Biaxial Creep Defonnation ofZirclaoy4 in the High Alpha-Phase Temperature Region, TPRD/B/Ol00/N82. June 1982. R F Cameron, Applications of the FRAP-T Codes to PWR Loss-of-Coolant Accident Analysis, and Comparisons with the MABEL-2 Model Part 1 . The FRAP-T5 Code, NDR-896(S,X) Part 1 , January 1984.

Mechanical Models

7.16

81

T J Haste, Modelling the Effects ofDifferent Mechanical Restraints on PWR Fuel Rod Defonnation under Conditions Relevant to the NRU MT-3 Experiment, Using the MABEL-2 Code, IAEA Meeting on Water Reactor Fuel Element Perfonnance Computer Modelling, Bowness-on-Windennere, UK, April 1984. 7.17 S S Dosanjh et aI, MELPROG-PWR/MODI :A Two-Dimensional, Mechanistic Code for Analysis of Reactor Core Melt Progression and Vessel Attack Under Severe Accident conditions, NUREG/CR-5 193, May 1 989. 7. 18 D V Swenson and M L Corradini, Monte Carlo Analysis of LWR Steam Explosions, NUREG/CR-2307, SAND8 1-1092, Sandia National Laboratories, Albuquerque, NM, October 198 1 . 7.19 W J O'Donnell and B F Langer, Design of Perforated Plates, Trans of the ASME J of Eng for Industry, p 307, August 1962. 7.20 J S Porowski and W J 0' Donnell, Elastic Design Methods for Perforated Plates, Trans ASME, Vol 100, pp 356-362, April 1978. 7.21 S Timoshenko and S Woinowsky-Krieger, Theory of Plates and Shells, McGraw­ Hill, 1959. 7.22 A Mendelson, Plasticity: Theory and Application, The MacMillan Company, New York, NY, 1983. 7.23 J F Harvey, Pressure Component Construction, Van Nostrand Reinhold, 1 980. 7.24 T H Lin, Theory of In-elastic Structures, John Wiley and Sons, New York, NY, 1968. 7.25 Zion Nuclear Station, Final Safety Analysis Report, Commonwealth Edison Co,

83

Chapter 8 MATERIALS INTERACTIONS

8.1

Introduction

Many accidents of interest in LWR safety analysis are characterised by loss of cooling in the core region. This chapter considers the various chemical interactions that occur and therefore need to be modelled once temperatures start to rise. Different phenomena occur during different phases in an accident sequence: in general the potential for interactions increases as temperatures rise. Once the core becomes uncovered by whatever mechanism, the core components, fuel rods, control rods, steel structures etc become exposed to a steam environment, and as temperatures rise there is an increased potential for the components to oxidise. Oxidation models for the principal materials, e.g. Zircaloy (the fuel rod cladding material), various stainless steels (control rod cladding and structural material), V02 (in the event of cladding failure by some mechanism) and various other materials are considered in this chapter. At higher temperatures e.g. under severe accident conditions there are a wide range of possible interactions. These are summarised in Table 8. 1 and discussed in the sequel. In addition to oxidation reaction with steam, chemical interactions take place between Zircaloy and fuel, control rod materials and structural materials. Binary, tertiary or multicomponent systems may be formed. The main interactions and associated models are considered in this chapter.

8.2

Zircaloy Oxidation

8.2.1 Phenomenon This is an important phenomenon to be considered in many accident sequences. In the large break Loss of Cooling Accident, oxidation of Zircaloy-4 cladding affects the ballooning of rods during the refill and reflood phases 970- 1 1 70K. The integrity of the cladding is also affected at temperatures in excess of 1 270K which could result in an increased likelihood of rod failure, and release of fission products to the primary circuit. The oxidation behaviour may also depend on the detailed alloy composition. Under severe accident conditions, Zircaloy oxidation is extremely important for a number of

Materials Interactions

84

reasons. At high temperatures the reaction is very rapid resulting in considerable additional heat and hydrogen release. The fonner is a positive feedback mechanism causing the rate of core heating to escalate: the latter provides a possible potential for a damaging energy release resulting from ignition later on in the accident. The oxidation of Zircaloy at high temperatures has been extensively reviewed [8. 1 ]-[8.4] . TE M P E RAT U R E 1 700

1 3 00

1 500

(o()

1 1 00

--

1 00 0

BAK E R - J U ST

. - ORNL

, Vl

N

-

N

1 0-5

- - - - - U RB A N I C - H E IDRICK

,

,

,

,

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v

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0

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, I

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-

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,

,

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\ ,

,

.

, . , \

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.

,\ ,

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'"

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al

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C U B IC

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M O NOCLI N IC

Z rO]

ZrO]

ZrO]

--··�I��r----.�I�.r--

1 0-9

--L.-----r---""'"T"""---.,...--r---I

5

6

7

T E M PE RATU RE-1

8 X

9

1 04 (K- 1 )

FIGURE 8 . 1 PARABOLIC RATE CONSTANTS FOR ZIRCALOY OXIDATION IN STEAM

Materials Interactions

85

8.2.2 Kinetics

Cladding oxidation occurs during normal light water reactor operation viz 570-670K. At these temperatures a thin oxide film develops over the alpha phase Zircaloy. Well characterised data are available from numerous out of pile experiments. Empirical models based on out of pile data have been published by Van de Linde [8.5] but in-reactor oxidation shows enhancement compared with the predictions of models developed from out of pile data. The enhancement is different in B WR environments and PWR environments and more pronounced in the lower temperature range. Past studies have concentrated on the effects of dissolved oxygen in the water [8.6] , [8.7], fast neutron flux [8. 1 ] , fast neutron fluence [8.7] and y - irradiation [8.9] . The oxidation of Zircaloy-4 in stream and helium/oxygen mixtures has been investigated at Springfields in a temperature range 970- 1270K relevant to large leak LOCA. This work was carried out to investigate the impact of oxidation on ballooning under large leak LOCA conditions. At these temperatures the rods reveal an oxide layer and an underlying ox ygen enriched layer, the oxygen stabilised alpha layer. Oxidation of Zircaloy at high temperatures (> 1253K) in steam produces a three-layer structure: an outer brittle layer of zirconium dioxide, an underlying brittle layer of alpha­ phase metal stabilised against transformation into beta-phase by the presence of dissolved oxygen, and a beta-phase core, containing a smaller amount of dissolved oxygen. The cladding may oxidise similarly on the inner surface, near the rupture site in ballooned cladding. The reaction is highly exothermic, (- 586 kl/mol), and this may lead to uncontrolled temperature escalation under conditions of poor heat transfer; typically at temperatures above about 1 500K. Generally the data can be represented by a correlation in the form dN= Ke -Qff t for steady state oxidation where d is the oxide or oxygen stabilised alpha thickness, T is the absolute temperature, K is a constant and N is an integer describing the time dependence (= 1 for linear, 2 for parabolic, 3 for cubic kinetics). Below 1 1 70K approximately cubic kinetics are observed for the oxide layer [8. 10]-[8. 12] . However at normal operating temperatures reactor oxidation i s not so well predicted for reasons that are not particularly clear. The growth of the stabilised alpha layer is parabolic. This behaviour has been associated [8. 12] with the formation of monoclinic rather than tetragonal oxide. At long times (>30 mins) at temperatures - 1 220K, transition to breakaway oxidation with linear kinetics is observed. At temperatures relevant to large break LOCA analysis the oxidation response of the Zircaloy-4 cladding is dependent on the composition of the alloy and also on the oxidising atmosphere [8. 13].

86

Materials Interactions

The rate-controlling mechanism at high temperatures is believed to be diffusion of oxygen anions through the anion-deficient zirconia structure. The parabolic kinetics expected from this theory are obeyed above 1 323K for weight gain, and for the thickness of the oxide and oxygen-stabilised alpha layer. The oxide layer data show a smooth transition from cubic to parabolic kinetics in the temperature range 1 170- 1320K. The data [8. 14]-[8 . 1 5] in this intermediate range have been approximated by logarithmic interpolation between the equations fitted at the upper and lower ranges. A step increase in oxide layer growth rate, but not alpha-layer growth rate, is seen at 1 850K [8. 1 6] , this has again been associated with a change in the oxide phase, from tetragonal to cubic. There is little evidence at temperatures relevant to severe accidents of any dependence on pressure, material type, small amounts « 1 0%) of steam contaminates (e.g. nitrogen) on the reaction rate, although some inhibition of the reaction at high hydrogen concentrations (>50%) partial pressure) has been observed [8. 1 7]. However hydriding does not occur at high temperatures. Examples of measured parabolic rate constants expressed as a function of temperature are given in [8. 12] , [8. 16] and [8. 18], Figure 8. 1 . The ORNL correlation [8. 1 2] is commonly used [8.4] for best estimate calculations of severe accident transients, with the UrbanicIHeidrick correlation [8. 1 6] being used at high temperatures. The AEA correlation [8. 10] is consistent with this approach, in the cubic/parabolic region. The Baker-Just correlation [8. 1 8] based on early experiments employing fine zirconium wires heated in water, has been extensively used in licensing work. 8.3

Fuel and Cladding Interactions

8.3.1

Zircaloy/Uranium Dioxide Interactions

Under severe accident conditions, as core temperatures rise, various chemical reactions can take place in vol ving the Zircaloy cladding and the U02 fuel. The external interactions between Zircaloy and the steam environment have been described earlier in this chapter. These result in oxygen stabilised a-Zr(O) and zr02 phases at the outer surface. At the inner surface there results the fonnation of oxygen stabilised a-Zr(o) phases and also the formation of (U, Zr) see e.g. [8. 19]-[8.20]. The kinetics of these interactions have been studied under steady state conditions and also under transient conditions. Figure 8.2 shows the various phases that exist in a fuel rod undergoing external oxidation by steam. The extent of the internal cladding reaction depends critically on whether there exists good contact between the fuel and cladding. The UOjil interaction is initially parabolic. However the oxygen potential on the outside is greater than the inside. It follows that the extent of interaction is determined by the rate of oxygen diffusion in from the outside and that eventually the UOjZr growth rates slow down. The final states are zr02 and ceramic (U, Zr)02 crusts. The PECLOX computer model has been developed to describe the chemical interactions between solid Zircaloy and U02 fuel. The model predicts, the formation, growth and disappearance of the various material layers. The numerical model is based on the Fick

Materials Interactions

87

Zr02

- Zr(O)

PRIOR B-Zry ( R E LATIV E LY U N REACTE D Zry-4)

Zr(O) b

- Zr(O)a+(U, Zr)

FIGURE 8.2 FUEL ROD UNDERGOING EXTERNAL STEAM OXIDATION

and Stefan diffusion models. The solution of these enables the detailed oxygen profile within the clad to be calculated as a function of temperature and time. The melting point of VOz fuel is about 3 1 OOK: however once VOz is in contact with molten Zircaloy the fuel may be liquefied at much lower temperatures. A metallic (Zr, V, 0) melt is formed. The reaction kinetics have been investigated, see for example [8.2 1], [8.22]. The chemical interaction between VOl and molten Zircaloy is parabolic: the dissolution rate has an Arrenhius dependence on temperature. 8.3.2 Molten Zircaloy/Solid Zirconium Dioxide Dissolution

As shown in Chapter 9, it is quite possible for molten Zircaloy to be enclosed by the ZrOz crust. In these circumstances, the oxidised cladding can be dissolved by the liquid Zircaloy at temperatures up to l000K lower than the oxide melting point. This has been investigated by Hofmann. Again the dissolution obeys a parabolic rate law: however the dissolution rate ofZrOz by molten Zircaloy is about five times slower than the corresponding VOz dissolution rate.

Materials Interactions

88 8.4

Control Rod Materials

8.4.1 PWR Absorber Rods

PWR control rods consist of a Silver-Indium-Cadmium (Ag, In, Cd) alloy. This material becomes molten at about 1037K. Zircaloy can be chemically dissolved by molten absorber rod alloy. Reaction rates are fast at temperatures higher than 1473K: the dissolution rates are parabolic. Eutectic reactions can take place between stainless steel and Zircaloy once temperatures reach 1273K and the reaction rate becomes fast above 1 523K. The interaction rates are also parabolic. Studies of these reactions have been investigated with various experiments. Under severe accident conditions, there are several mechanisms by which control rods can fail. The stainless steel cladding may fail because of internal cadmium vapour pressurisation: cadmium is the most volatile of the control rod constituents. Alternatively if stainless steel cladding contacts the Zircaloy guide tube, e.g. due to bowing or other mechanism then the control rod could fail due to a Zircaloy/stainless steel eutectic reaction. In either mechanism, the molten alloy is released possibly evoking further dissolution of the Zircaloy guide tube or other Zircaloy components of the core. These reactions are somewhat delayed by the fonnation of thin oxide layers on the Zircaloy surfaces. However, dissolution of the oxide layers by the unreacted Zircaloy must take place before the other reactions can proceed. The presence of additional oxygen in the Zircaloy tends to reduce the reaction rates and push liquefaction temperatures slightly higher. The molten behaviour of (Ag, In, Cd) control rods has an important influence on the early melt down behaviour under severe accident conditions. The mechanism of material release, i.e. control rod failure may be sequence dependent For low pressure sequences the control rods may fail by a stainless steelfZircaloy guide tube interaction. At high pressures, the control rod may fail earlier due to internal pressuration. In summary control rod failure may be expected if temperatures rise beyond 1 523K. Since the resulting melt has the potential to dissolve Zircaloy at these temperatures, control rod failure will have a strong impact on early melt progression and also fission product release. 8.4.2 BWR Absorber Rods

In BWRs, the absorber blades contain boron carbide (B4C) pellets in stainless steel tubes. In each control element, there are four blades fonned in the shape of a cross. The interaction of B4C with stainless steel (AISI 3 16) and with Zircaloy-4 has been studied by Hofmann [8.23] et al for the temperature range 1073-1523K. In both reactions liquefaction occurs below the melting points of the constituents.

89

Materials Interactions

B4C reacts with stainless steel i n a manner described b y a parabolic rate law. Th e kinetics become very rapid above 1473K, as liquid phases form and above 1 523K complete liquefaction takes place. B4C also reacts with Zircaloy but the kinetics are several orders of magnitude below the B"C/stainless steel interaction rates. Complete liquefaction doesn ' t take place until about 1923K.

TABLE 8.1 TEMPERATURE REGIMES FOR MATERIALS INTERACTIONS IN A SEVERE LWR ACCIDENT

Temperature (K) Total core destruction

3273 3 123

Melting of V02 , zr02 and (U,Zr)02 reaction products

2873

I

Complete meltdown of all core materials and formation of a ceramic melt Extended core damage

2273

2023

1673

1473

Melting of metallic Zircaloy cladding and n-Zr(O), V02 dissolution. Eutectic interactions of A 1 203 with Zr02 and/or V02

Eutectic interactions between Fe-Zr, Ni-Zr, B "C-stainless steel, B"C-Zircaloy, Al (AI203) -Zr Ag-Zr interactions

Formation of liquid and metallic phases that result in a local fuel rod cladding damage

Localised core damage

1273

1073

I

Formation of metallic and ceramic melts that relocate and form large blockages on solidification

Melting of Ag-In-Cd absorber alloy

90

Materials Interactions

8.5

Other Zircaloy Interactions

8.5.1

Inconel Grids

Zircaloy reacts with Inconel grids in PWRs at temperatures below the melting JX>int of Inconel -1723K. Experiments to sbJdy the interaction between Zircaloy and Inconel 718 have been carried out by Hofmann et aI. These studies also a:ldressed the situation when the Zircaloy is covered by a thin oxide fIlm . With unoxidised Zircaloy, interactions between the Inconel 718 � grid and Zircaloy start at aroWld 1273K, when liquid phases fIrst fonn. Fast and complete liquefaction occurs at 1523K. The kinetics are parnbolic. The presence of these oxide layers delays the rate of reaction but does not prevent it. As for the case of stainless steel liquefaction, the oxide layers must be dissolved before the Inconel reaction can take place, resulting in the formation of oxygen stabilised a-Zr(O) which tends to shift liquefaction to higher temperatures.

8.5.2

Burnable Poison Rods

Some nuclear power plants use a mixture of alumina (A�03) and a small amount of boron carbide (B"C) as a burnable poison. The interaction rates between �03 and Zircaloy-4 can be described by parabolic rate laws. These interactions have been studied by Hofmann et al [8.24] in the temperature range 1473K- 1 773K. First liquid phases form about 1473K, above 1 773K the kinetics are very rapid. The chemical reactions can be described by Arrenhius rate laws. Under severe accident conditions, low temperature eutectic melt would be released on failure ofcladding. This would also give increased liquefaction ofUOl, since the melt can react with the UOl, at temperatures below the melting point of the Zircaloy (2033K).

8.6

Steel and Structures Oxidation

The stainless steel oxidation reaction is complex and several species are present in the oxide layer. These also depend on whether the external environment is oxidising or reducing (e.g. in the case of hydrogen presence and the absence of steam). The heat of reaction of steel with steam is about an order of magnitude less than the corresponding Zircaloy/steam interaction. The feedback in core heat-up is therefore much less positive. At short times above 1 273K the reaction kinetics are linear [8.25] : at longer times the kinetics are parabolic. At temperatures above 1673K there is considerable foaming of the steel [8.25]­ [8.28] . The rate of reaction is faster than that for Zircaloy above 1425K, as temperatures approach the steel melting point.

Materials Interactions 8.7

91

Oxidation or Fuel

The oxidation of VOz in steam atmospheres has been studied by Bittel et al [8.29] . The kinetics were found to be approximately parabolic and the parabolic rate constants could be correlated by Arrenhius laws. The oxidation of VOz in steam has also been studied by Cox et al [8.30]. The temperature ranges studied were 1 273 - 1 923K. Cox 's analysis indicated that the kinetics were only linear during the initial period of oxidation. The rate constants during this period agreed with those of Bittel. Cox also found that increased exposure to steam produced significant weight losses due to volatilisation. The kinetics for this process were linear. The oxidation of VOz in other environments has also been studied. The oxidation of VOl in air and oxygen was studied many years ago because of interest in the behaviour of dry storage for fuel. Cox has recently performed some measurements. He showed a growth ofV409 with parabolic kinetics. The kinetics in air (e.g. at 1 273K) were much faster than for steam. At lower temperatures in steam , V30S is formed [8.3 1 ] . Carter and Lay [8.32] investigated the behaviour of VOl i n COl an d CO atmospheres. They developed an Arrenhius expression for the rate of stoichiometry increase.

REFERENCES 8. 1 8.2

8.3

8.4

8.5 8.6

8.7 8.8

C A Mann, E 0 Hindle and P 0 Parsons, ''The Deformation of PWR fuel in a LOCA", ND-R-70 1 (S), April 1 992. P 0 Parsons and W G B urns, "Oxidation of Zircaloy Fuel Cladding and Hydrogen Evolution", in J H Gittus (ed.), PWR Degraded Core Analysis, ND-R-6 1 0(S), April 1982. P 0 Parsons, E 0 Hindle and C A Mann, "The Deformation, Oxidation and Embrittlement of PWR Cladding in a Loss-Of-Coolant Accident" CSNI State­ of-the-Art Report ND-R- 135 1 (S), September 1986. N L Hampton and 0 L Hagrman, "Zircaloy Oxidation in Water and Steam", in J K Hohorst (cd.) SCDAP/RELAP5/MOD2 Code Manual, Volume 4: MATPRO - A Library of Materials Properties for Light Water Reactor Accident Analysis, NUREG/ CR-5273 Vol. 4, February 1990. A. Van der Linde, Calculation of the Safe Life Time Expectancy of Zirconium Alloy Canning in the Fuel-Elements of the Nero Reactor, RCN-4 1 (July 1 965). A. B. Johnson, Jr., "Effects of Nuclear Radiation on the Corrosion, Hydriding, and Oxide Properties of Six Zirconium Alloys", Applications-Related Phenomena for Zirconium and its Alloys, ASTM-STP-458 ( 1 969) pp. 301 -324. W. A. Burns, Effects of Fast Neutron Irradiation, Fabrication History, and Water/ Oxygen on the Environmental Behaviour of Zirconium Alloys, BNWL-88 ( 1 965). A. B. Johnson, Jr., and J. E. lrvin, Radiation-Enhanced Oxidation ofZircaloy-2 in pH10 LiOH and pH- I0 NH40H, BNWL-463 (1967).

92 8.9

8.10

8. 1 1

8. 12 8. 1 3

8.14

8. 15 8. 16 8. 1 7

8. 1 8

8. 1 9 8.20

8.2 1

8.22 8.23

8.24

8.25

Materials Interactions L. Lundeand K. Videm, "Effects of Surface Treatment on the Irradiation Enhancement of Corrosion of Zircaloy-2 in HBWR", Zirconium in Nuclear Applications, ASTM­ STP-55 1 (1 974) pp 5 14-526. T J Haste, W R Harrison and E D Hindle, ''Zircaloy Oxidation Kinetics in the Temperature Range 700- 1 300·C", lAEA Technical Committee Meeting on Water Reactor Fuel Element Modelling in Steady-State Transient and Accident Conditions, Preston, September 1 988. S Leistikow, G Schanz, H V Berg and A E Aly, "Comprehensive Presentation ofExtended Zircaloy-Steam OxidationResults600- 100°C", OECD-NEA-CSNI/lAEAMeeting, Riso, Denmark, May 1983, IAEA Summary Report IWGFPT/16, 1983. R E Pawel, J V Cathcart and R A McKee, 'The Kinetics of Oxidation of Zircaloy-4 in Steam at High Temperatures", J. Electrochem. Soc. 126 1 105- 1 1 1 1 , July 1 979. P D Parsons, E D Hindle and C A Mann. The Defonnation, Oxidation and Embrittlement of PWR Cladding in a Loss-Of-Coolant Accident. ND-R1 3 5 1 (S), September 1986. S Leistikow, G Schanz, H V Berg and A E Aly Comprehensive Presentation of ExtendedZircaloy-4 Steam OxidationResults 600- 1600°C. OECD-NEA-CSNI/lAEA Meeting, Riso, Denmark, May 1983, lAEA Summary Report IWGFPT/l6, 1 983. S Leistikow and G Schanz. High Tem� Oxidation of Zircaloy4 in Steam (� 16(X)OC). Ninth Int Congress on Metallic Corrosion, Toronto, Canada, June 1984. V F Urbanic and T R Heidrick, "High Temperature Oxidation of Zircaloy-2 and Zircaloy-4 in Steam", J. Nucl. Mat 75(2), 25 1 -26 1 , August 1 978. H M Chung and G R Thomas, "Rate-Limiting Effects o n the Effect o f Gaseous Hydrogen on Zircaloy Oxidation" NRC Workshop on the Impact of Hydrogen on Water Reactor Safety, Albuquerque, USA, January 198 1 . L Baker and L C Just, "S tudies of Metal-Water Reactions a t High Temperatures. III. Experimental and Theoretical S tudies of the Zirconium-Water Reaction", ANL-6548, May 1 962. P Hofmann et aI . , "Reac tor Core Materials Interactions at very H igh Temperatures", Nuclear Technology, Vol. 87 ( 1 990) 146- 1 86. P Hofmann, D W Kerwin-Peck, "UOfZircaloy-4 Chemical Interactions and Reaction Kinetics from 1 000 to 1 700·C under Isothermal and Transien t Conditions", J. Nucl. Mater. 1 24 ( 1 984) 80- 1 -5. P Hofmann, H Uetsuka, A N Wilhelm, E A Garcia, "Dissolution of Solid U02 by Molten Zircaloy and its Modelling", Int. Sympos. on Severe Accidents in Nuclear Power Plants, Sorrento, Italy, 2 1 -25 March 1 988 (IAEA-SM-2986/l). K T Kim, D R Olander, "Dissolution of Uranium Dioxide by Molten Zircaloy", J. Nucl. Mater 1 54 (1 988) 85- 1 12. P Hofmann, M Markiewicz, J Spino, "Reaction Behaviour of B4C Absorber Material with S tainless Steel and Zircaloy in Severe LWR Accidents", Nuclear Technology, Vol 90 ( 1 990) 226-244. P Hofmann, M Markiewicz, J Spino, "Chemical Interactions Between A 1 203, which is used in Burnable Poison Rods, and Zircaloy-4 up to 1 500·C", J. Nucl. Mater. 166 ( 1989) 287-299. J T Bittel, L H Sjodahl and J F White, "Oxidation of 304L Stainless Steel by Steam and Air", Corrosion 25( 1), 7- 14 January 1 969.

Materials Interactions 8.26 8.27

8.28 8.29 8.30

8.3 1

93

R E Wilson, C Bames, L Baker and R 0 Ivins, "Chemical Engineering Division Semi­ Annual Report, Reactor Safety Metal-Water Reactions", ANL-7 125, May 1 966 . H C Brassfield, J F White, L Sjodahl and T J Bittel, ''Recommended Property and Reaction Kinetic Data for use in Evaluating a Light-Water-Cooled Reactor Loss-Of-Coolant Accident Involving Zircaloy4 or 304 SS Clad V02", GEMP 482, 1968. J C Hesson et aI., "Laboratory S imulation of Cladding-Steam Reactions Following Loss-Of-Coolant Accidents in Water-Cooled Reactors". ANL-7609, 1 970. J T Bittel, L H Sjodahl, J F White, Amer. Ceram. Soc. 52 ( 1969) 446. D S Cox, F C Iglesias, C E L Hunt, N A Keller, R D Barrand, J R Mitchell, R F O'Connor, "Oxidation of V02 in Air and Steam with Relevance to Fission Product Releases", Proceedings of the American Chemical Society Symposium on Chemical Phenomena Associated with Radioactivity Releases During Severe Nuclear Plant Accidents, Anaheim C. A., 7- 1 2 September 1986, NUREG/CP-0078. D S Cox, C E L Hunt, R F O'Connor, R D Barrand, and F C Iglesias, "High Temperature Oxidation Behaviour of V02 in Air and Steam", Proceedings of the International Symposium on High Temperature Oxidation and S ulphidation Processes, Hamilton, Ontario, Canada, 1 990, August 26-30. Published by Pergamon Press, ISBN 0- 1 8-0404 1 5-4.

95

Chapter 9 MELT PROGRESSION M ODELS

9.1

Introduction

This chapter describes some of the methods that have been developed to predict melt progression in Light Water Reactors under severe accident conditions. Initially models were largely parametric: the quantity of melt was effectively detennined by input parameters and the impact of various assumptions was explored. An example of such a model is described in the Section 9.2. Later the TMI-2 accident [9. 1] and various severe fuel damage experiments [9.2] -[9.7] have shown that core damage is likely to proceed through various states before the core slumps into the lower head. Mechanistic models were fonnulated which attempted to recognise the various events of melt progression. The available data imply that various states of core damage occur which include: (i)

the liquefaction and solidification of the fuel rod cladding, including the dissolution of fuel and leading to the formation of largely non-porous blockages,

(ii)

fragmentation of fuel, perhaps as a consequence of rapid cooling, leading to the formation of porous debris blockage,

(iii)

the fonnation of a molten pool supported by non-porous crust,

(iv)

melt-through of the crust and slumping into the upper head.

A selection of models are considered in Section 9.3 onwards. 9.2

Parametric Core Meltdown Models

The early models developed for core melt down were both simple and parametric. The MARCH code [9.8] is an example of a code that incorporates such methodology. The models assume that at some stage of core meltdown, molten homogeneous core material will begin to fall on to the grid plates. A number of input variables control when dropping of melt from the core begins and ends. In such a simple approach the quantity of melt relocating is dependent on the numerical representation of the geometry or the finite difference nodalisation. The criteria for melt relocation in the MARCH code is based on certain key variables which include, the temperature of the fuel rods which must exceed that required to overcome the latent heat of fusion before the material in a particular node can drop and the fraction of the

96

Melt Progression Models

core which must be melted before material dropping from that node can occur. Additional criteria concern the number of the node which must be melted before the melted nodes in a particular radial region can drop and the fraction of the core melted. If all these criteria are satisfied then total collapse of the remaining core into the head is deemed to occur. To scope the effects of core meltdown on core heatup. various core meltdown models were developed. The models were parametric in the sense that criteria for slumping were not based on calculations of stress levels. creep rates. or flow rates of molten materials. Some of the meltdown models assumed the molten fuel is retained in the core as a continuous region; other models assumed the molten fuel falls as drops into the bottom head. Calculations indicate that these different model assumptions could give rise to significantly different core heat-up history primarily because of the influence of the meltdown model on the boil-off rate and the cladding water reaction. For the case when melt is held up in the core region various assumptions are made in the MARCH code concerning the heat transfer from the blockage (or molten pool). One option is to assume that the excess heat in the molten poo l (above that required to just keep the pool molten) is transferred downward. There is no convection of heat to the top and sides of the pool. Such a model would be physically consistent with a meltdown situation in which the molten region tends to cover and mix (downward) with the solid region at such a rate that the homogenized molten region remains just at the melting temperature. This approach maximises the downward movement of the molten poo l . With this option the code assumes that no solid core material falls into the molten pool from above. Another option is to assume that the excess heat in the molten pool is transferred upward, and none is transferred downward. Within the molten region, heat may be transferred radially if the average temperature of a radial power region exceeds the melting temperature. The heat transferred upward is used to melt solid core material, which is assumed to fall into the molten pool . The amount of solid material falling into the molten pool is sufficient to keep the homogenised temperature of the molten pool at the core melting point. When the top nodes in the core are melted, it is assumed the decay heat in the pool may be radiated to the support structures above the core. The two models yield very similar results for core meltdown fractions of up to about 50 per cent. However. for larger meltdown fractions the first model results in faster core heatup and a more rapid downward progression of the molten region. If it is assumed that during the melting period. a small part of the molten core continually falls into the water. the results for the two models are similar. The other approach i.e. when a fuel node melts, it immediately falls to the first grid plate below the bottom of the core is also an option in the MARCH code. In this case. the decay heat of the node. its heat associated with its thermal capacity and heat from the metal­ water reaction are added to the water. The large boiloff rates obtained under these assumptions result in very high heatup rates. due to the cladding-steam reaction,

97

Melt Progression Models

between the time when core melting first starts and all of the water in the bottom of the pressure vessel is boiled off. This model is not intended to give a complete picture of core meltdown into the lower head. It was developed to illustrate the effect of molten fuel dropping out of the core rather than being retained in a molten zone supported by a crust within the core region.

9.3

Mechanistic Models: General Approaches

The exothennic Zircaloy/steam reaction causes the fuel rod temperatures to rise rapidly to more than 2000K leading to clad melting and fuel dissolution. This process in which the Zircaloy interacts with the fuel to fonn a (0, Zr, 0) melt is an important mechanism in degrading the fuel geometry at temperatures far below the fuel melting point 3000K . Considerable work to characterise these phenomena has been carried out by Hofmann, as described in Chapter 8. Control rods, clad in stainless steel enclosed in Zircaloy guide tubes and grid spacers, constructed of Zircaloy or Inconel will also provide sources of molten material at much lower temperatures. -

As the melt flows into cooler regions it may freeze. There are uncertainties on the mechanism of freezing even in the 'simple case' of no chemical reactions. Two extreme cases are generally considered in mechanistic models: (a)

(b)

conduction controlled freezing in which a crust is deposited on the rod surface and bulk freezing in which heat is removed from the bulk flow and any melted material is entrained.

In practice, experiments tend to show an intennediate mechanism since although a crust may fonn, it may become unstable particularly if the substrate melts. Crust entrainment and heat transfer then become important. The addition of chemistry poses a further complication. It is not clear that for conduction controlled freezing there will be access to the heated clad since this will be covered with a corium crust For a bulk freezing model the effect of oxidation of the (0, Zr, 0) melt also has to be considered. The effect of these phenomena in detennining the manner of the melt progression and possible blockage fonnation is currently uncertain as although there are experimental data available on blockage fonnation, it is not clear whether conclusions extrapolate to plant scale. A summary of the liquefaction and relocation models in the principal severe accident codes is given in [9.9] . The main impact on the thennal-hydraulics is concerned with the possible contact of hot material with water and change in axial power distribution that results from clad liquefaction and relocation and also from the change in flow area. The effects of liquefied fuel and fuel relocation are treated in some models, as extensions of the pin geometry for moderate levels of core degradation. For more severe core loss of

98

Melt Progression Models

Fuel rod centre l i ne

•

,.

t•

•

:

. 0. •

-----�

B reach rel e a s i ng m ixtu re of Z i rca l oy I fuel e utectic (Z r . U .O.) M o lten m ixtu re of Zr. U . O Zr

So l i dified Zr. U . O crust form ation M ixtu re of f l ow i ng Zr. U . O F u rt her Zr.U .O. crust form ati o n

FIGURE 9. 1 FUEL ROD MELTDOWN

Melt Progression Models

99

geometry, it may be more appropriate to include the relocation of such material as a liquid phase in the thennal-hydraulics. In this approach, corium is treated as a high density liquid phase in the thennal-hydraulic equations. The corium phase may include fuel, Zircaloy, zirconium dioxide, steel, control rod material and eutectic. This approach, see for example [9. 10] , requires an advanced fluids description in tenns of the coupled mass, momentum and energy transfer between the water, vapour and corium phases. The numerical methods will need to be robust to model these exchange processes between such a wide range of materials.

9.4

Liquefaction of the Intact Core and the Formation of Blockages

9.4.1 PWR Control Rods Control rod temperatures may be obtained using similar heat transfer models as for the fuel rods. Once temperatures exceed a certain threshold control rod failure accompanied by release and relocation of control rod materials will occur. As described in Chapter 8 there is a potential for failure as soon as the core exceeds the eutectic melting temperature for contacting stainless steel control rod cladding with silver-indium­ cadmium (Ag-In-Cd) absorber material. This temperature is approximately 1 500K. Once the guide tube melts, or is breached by some mechanism , the absorber material will move through the breach and relocate downwards. The lowest boiling point component Cd will be released in vapour or aerosol. Ag will be lost as it traverses downwards to a cooler part of the core, where it will freeze. Existing models, for control rod relocation, see example [9. 1 1 ] , tend to be rather simplistic compared with fuel rod relocation models considered in the next section. For in general momentum and energy equations for the melt relocation are not fonnulated. In [9. 1 1 ] the frozen material is assumed to come to a halt when the temperature difference between the molten flowing material and its contact surface becomes sufficiently large. The more sophisticated models that are available for fuel rod relocation are described in Section 9.4.3.

9.4.2 BWR Control Rods Models have been developed to calculate the melting and relocation of B4C control rods. One such [9. 1 1 ] is based on the concept of a well developed incompressible viscous flow over a cylindrical geometry. The internal energy associated with the relocated molten material is assumed to instantly exchange with the internal energy of material at lower elevations. Should liquefied material at this new elevation still appear, downward moving calculation of the liquefied material continues. If the middle region of a control rod becomes depleted then the unsupported control rod segment is assumed to slump.

1 00

9.4.3

Melt Progression Models Fuel Rods

Widespread melting can occur once temperatures reach the cladding melting temperature -2200K. Models have been developed to describe the dripping of the liquefied fuel and cladding down the outside of fuel rods and to calculate where the material solidifies and blockage fonns. In order for melt to flow, the protective oxide crust which will have fonned during the heat-up under steam conditions must be broken or breached, Figure 9. 1 . Whether the cladding oxide protective shell breaches or not is dependent on the thickness of the oxide cladding once Zircaloy melting temperatures have been reached. Typical criteria that are used are that breach occurs if the oxide thickness is less than 50% oxidised once the melting temperatures have been reached: otherwise breach cannot occur until the ceramic oxide melting temperatures are reached, about 2950K. If axial temperature distributions are modelled then breach sites can in principle be predicted at a number of axial locations. The available models for the flow of melt contain various simplifying assumptions. Axi­ symmetric slugs are assumed to relocate down the outside of fuel rods. The fonner assumption may not be justified in all cases. Relocation could in principle occur down the inside of fuel rods, if clad ballooning has taken place. Other assumptions e.g. [9. 1 1 ] are that the slugs are of fixed length, well mixed and at uniform temperature. Heat is transferred to the fuel rod on the basis of a temperature front which propagates through the rod as the melt progresses downwards. Heat transfer to neighbouring rods and the steam is neglected in some models. The velocity of the melt can be calculated from a force balance between gravity and frictional losses, at the melt/fuel rod interface. The criteria for freezing is typically when the slug temperature falls below the melt solidus temperature. As core material relocates downwards a blockage can grow both radially and axially. A non-porous debris bed will fonn and coolant flow paths through the blockage region may be virtually cut off. S ince the blockage may contain unoxidised Zircaloy which can produce heat and is generating decay heat, there is a potential for a re-melting of the blockage at a later phase in the accident. Models have also been developed to describe the relocation of previously solidified melt. Inevitably melt relocation implies a change in geometry and since melt progression events are largely controlled by temperature, correct prediction of relocation events will be dependent on correct modelling of convection and radiation to the coolant and neighbouring structures. At blockage temperatures in excess of 2000K , radiation heat transfer will be particularly important.

Melt Progression Models

101

The rate of temperature increase reduces markedly during melt relocation. Oxidation heating is dominant during the peak temperature excursion phase and the removal of Zircaloy from a particular part of the core implies the removal of a large heat source. If the melt falls down into water, the additional convective cooling of the steam may involve a further temperature decrease. However under steam starved conditions, additonal steam could promote further Zircaloy oxidation in the bundle which would tend to counteract this effect. The blockage size and quantity of melt relocated will depend on the degree of dissolution of fuel [9. 1 2] and perhaps other materials as described in the previous chapter. The blockage composition will also be affected. A large amount of UOz will make the blockage more refractory. The modelling of the solubility of UOz in the Zircaloy will be affected by whether in the Zr-u-o ternary phase diagram, the solidus line or the liquidus line is used. The former assumption gives rise to a factor of five increased dissolution of UOz compared with the later assumption. It should be noted however that the amount of material relocated is frequently determined by the oxide breach criteria. The extent to which blockages are porous or non-porous requires consideration in modelling. Data are available from various fuel damage tests and the TMI-2 accident.

9.S

Core Fragmentation and Blockage Formation

There are various mechanisms through which fragmentation of the core can occur. One mechanism is concerned with rapid cooling of substantially oxidised fuel rods. In Chapter 8 the various layers of differing oxygen concentration in cladding undergoing oxidation were considered. As the zirconium dioxide and oxygen stabilised layers grow the clad is said to become embrittled. Under rapid cooling conditions embrittled fuel rods will tend to fragment into particles. Criteria for fragmentation typically involve the thickness of the �-phase Zircaloy remaining and cladding temperature history. A model was developed by Chung and Kassner [9. 1 3] for predicting the embrittlement and fragmentation of oxidised fuel rods. This was studied by Haggag [9. 14] and found to accuratel y predict events. Data on debris porosity and particle diameters have been obtained from TMI-2 debris [9. 1 5] , [9. 16] . Another mechanism of fragmenting the fuel occurs under reducing conditions e.g. the cladding reaches melting conditions with a thin oxide layer, and the amount of dissolved oxygen is so small that the fuel is not melted [9. 1 7] . Under these conditions the melt will just run off leaving behind fuel and a thin shell of oxidised cladding. Since the fuel will be pre-fragmented due to its power history (this would also be the case in the mechanism considered above), the fragmented fuel would be released. Typical criteria for the production of porous debris by this means are that the cladding is sufficiently thin and the temperature exceeds the melting temperature 2200K of the cladding.

1 02

Melt Progression Models

The heat up of porous debris beds depends on the decay heat, oxidation of previously unoxidised Zircaloy and convective and radiative heat transfer to the coolant flowing through the porous spaces. These thennal-hydraulic mechanisms of debris beds are considered separately in a subsequent chapter. The fIrst mechanism referred to above is relevant to accident sequences where a reflood is modelled. In most LWR accident sequences without reflood the amount of oxidation would be sufficiently large, that the oxidised clad would be sufficiently strong to contain the fragmented fuel. The fuel debris would only be released once temperatures exceeded the oxidised cladding melting point (about 2950K). The survivability of oxidised cladding under reducing conditions will depend on the extent of dissolution of the oxide layer by any unoxidised Zircaloy.

FIGURE 9.2 TMI-2 END-STATE CORE CONFIGURATION THROUGH A ROW OF FUEL ASSEMBLIES

Melt Progression Models 9.6

103

Molten Pools

Various mechanisms have been discussed by which both porous and non-porous debris may accumulate in the lower core region. Figure 9.2 shows the TMI -2 end-state core configuration with clear evidence of a blockage in this region. Concerning non-porous blockages, liquefied material may relocate, then freeze out during the early and late melt down phases. As the blockage of non-porous debris grows, the internal part of the blockage can then increase in temperature due to decay heat. Another pool is therefore formed consisting of a layer of molten material which is supported by a non-porous crust. Clearly also a porous material blockage has the potential to form a molten pool if the coolability is sufficiently poor . The liquefied debris would tend to permeate porous debris and so therefore there could be the tendency for a molten pool to grow as above, though possibly this time supported by more porous debris. In this case the melt is found to be a unified mixture of liquid and solid debris. Thus in both cases there is a potential to form a melt pool once temperatures rise above the solidus temperature. This pool , if the supporting crust fails, could dump a large amount of molten material into the lower head of the vessel. Heat is transferred by natural convection within the molten pool . Models for the heat flux to the top and bottom surfaces have been formulated in terms of Rayleigh number correlations Jabn and Reineke [9. 1 8] . These have also been applied to TMI-2 by Epstein and Fauske [9. 19]. The pool temperatures can then be determined by a heat balance between these heat fluxes and internal heating sources. Models are also necessary to allow for the growth and melting of the upper and lower crusts. Some models assume equilibration of the molten and porous debris. This is likely to over-estimate the penetration of molten material into porous debris beds and this exaggerates the growth of the melt pool . Clearly the crusts supporting the melt pool may contain a complex m ixture of fuel rod and control rod materials which affect their melting points and therefore the potential for failure and melt relocation. Correlation for these heat fluxes over hemispheres are available from the literature: correlations for the convective heat flux from a molten pool have also been derived from steady-state data [9. 1 1 ] . In reality to allow for transit effects a better approach would be to correlate the heat flux to the temperature difference between the bulk melt temperature and the boundary temperature. 9.7

Blockage Melt Release Models

Under severe accident conditions it seems likely that a molten region will form in the core region and be contained by a surrounding crust. As the temperature of the pool rises,

104

Melt Progression Models

convection will become an effective mechanism of heat transfer, the crust will thin and failure may somewhere occur. There are effectively two possibilities. Either the upper crust will fail or the lower crust, will fail. For failure of the lower crust criteria based on thickness are used and all the melt will be released (e.g. if the crust is less than 25mm thick). A second mechanism for which models have been produced is failure of the upper crust. Again the criteria used have been based on thickness, e.g. the crust fails if it is less than 0.5mm thick if it is supported underneath by liquid: otherwise it fails when it is less than 25mm thick [9.20] . Liquid spills over and is calculated from the displacement of liquid melt by solid material.

9.8

Melt Relocation from Blockages and Melt Progression Paths

Once the crust fails, the melt will fall under gravity down into the lower part of the vessel. The questions that arise are whether the melt will refreeze again lower down and also the path that the melt will take. These questions are important to define the boundary conditions for melt interaction with the lower head or the initial stages of molten fuel coolant interaction. The indications from TMI-2 are that melt relocation can take place through the core by­ pass region as well as through the lower core region. Many of the PWR severe accident codes assume that melt progresses downwards to the lower core region. Melt is likely to flow through large holes in the lower core plate and whether or not melt can reach the lower head is determined via modelling an effectively extended fuel rod. The more sophisticated codes include flow freezing models but these have not been satisfactorily validated at the present time. Turland has investigated the degree of melt penetration and flow paths using the PLUGM code [9.2 1 ] . PLUGM models the time dependent flow o f molten material through a prescribed flow channel, allowing for heat transfer between the molten material and the channel wall, and for the reduction of flow area as crusts develop. The code includes in it as options both a conduction freezing model and a bulk freezing model. With the first option a crust is allowed to form on the cool surfaces of the flow channel and the crust growth or dissolution is calculated using a Stefan condition [9.2 1 ] . Heat transfer in the crust and channel wall is calculated assuming one dimensional conduction. With the second option, i.e. the bulk freezing model no crust is assumed to form. Once the molten material cools to below its liquidus temperature, a solid particle fraction is calculated, allowing for release of latent heat. The channel is assumed to plug when this

Melt Progression Models

105

fraction reaches a critical value typically prescribed as 0.62 (the maximum spherical packing fraction). The flow of melt is assumed to be one dimensional pipe flow. Discretised mass and momentum conservation equations are solved. The velocity profile assumes the flow is turbulent. Turland applied the code to flows through an intact fuel bundle and also to model the flow through bypass regions. TMI-2 was the chosen accident scenario. In order to model the flow through intact bundles various possibilities were considered on how the real channel geometry could be mapped to a cylindrical geometrical model. The approaches produced rapid plugging for the bulk freezing assumption, whereas for the conduction freezing model the extent of melt relocation to the lower head was dependent on the length of the intact fuel rods. The prediction of melt via the core bypass requires less complicated modelling than is required for melt flow through the lower core region and the results are therefore more certain. Turland concluded that if the length of the intact fuel rods was relatively long, perhaps more than 2m, then the melt would probably relocate through the bypass region, for shorter lengths of intact fuel rods the likely route was less certain. The code has been compared with other flow freezing codes and also with experimental data for UO/Mo flow down a tube. These and other comparisons of flow freezing codes with experimental data indicate the conduction freezing model gives reasonable agreement with data from simulant experiments but the behaviour of U02 melt is intermediate between predictions based on the conduction freezing (stable crust) and bulk freezing (slurry) models. Examination of frozen U02 melts indicates that away from the leading edge, stable crusts form: bulk freezing is only appropriate for the leading edge and once the edge has passed through the system without causing blockage, the conduction model is more appropriate.

REFERENCES 9.1 9.2 9.3 9.4

9.5

E L Tolman et al., TMI-2 Accident Scenario Update, EGG-TMI-7489, December 1 986. Z R Martinson et aI. , PBF Severe Fuel Damage Test 1 - 1 Test Results Report, NUREG/ CR-4684, EGG-2463, October 1 986. D A Petti et al., PBF Severe Fuel Damage Test 1 -4 Test Results Report, NUREG/CR5 1 63 , EGG-2542, May 1989. S Hagen et a1., "Out-of-Pile Experiments on Severe Fuel Damage Behaviour ofLWR Fuel Elements (CORA Program)," International Symposium on Severe Accidents in Nuclear Power Plants, Sorrento, Italy, March 2 1 -25, 1 988, paper IAEA-SM-296126. A B Wahba and E F Hicken, "Ten Years of Experiments in the Loss of Fluid Test (LOFT) Facility," Atomkernenergie-kerntechnik, 49, No. 1/2, 1 986, pp. 68-73.

106 9.6

9.7 9.8

9.9

9.10

9. 1 1

9.12

9.13

9.14 9.15 9.16 9.17

9. 1 8

9.19 9.20 9.21

Melt Progression Models R 0 Gauntt et al., "Results and Phenomena Observed from the DF-4 BWR Control Blade Channel Box Test," Proceedings of the International ENS-ANS Conference on Thermal Reactor Safety, 5, Avignon, France, October 2-7, 1 988. M L Russell, "TMI-2 Core Geometry," Proceedings of the TMJ-2 Topical Meeting, Washington, D. C., November 1 988. R 0 Wooton, P Cybulskis and S F Quayle. MARCH 2 (Meltdown Accident Response Characteristics) Code Description and Users Manual, Battelle-Columbus Laboratories, NUREG/CR-3988, September 1 984. J N Lillington and A J Lyons. The Modelling of Vessel Heat Transfer Phenomena in PWR Severe Accident Analysis. Eighth International Heat Transfer Conference, San Francisco, August 1986. S S Dosanjh e t al, MELPROG-PWR/MOD1 : A Two-Dimensional Mechanistic Code for Analysis of Reactor Core Melt Progression and Vessel Attack Under Severe Accident Conditions, NUREG/CR-5 193, May 1 989. SCDAP/RELAP5/MOD2 Code Manual: SCDAP Code Structure, Models, and Solution Methods. (Edited by CM Allison and E C Johnson.) NUREG/CR-5273, EGG-2555, volume 2 ( 1 986). P Hoffmann an d H Ostereka, "Dissolution o f Solid UOz b y Molten Zircaloy and the Modelling," Paper IAEA-SM-296/ 1 , International Symposium of Severe Accidents in Nuclear Power Plants, Sorrento, Italy, March 2 1 -25, 1 988. H M Chung and T F Kassner, Embrittlement Criteria for Zircaloy Fuel Cladding Applicable to Accident Situations in Light Water Reactors, NUREG/CR- 1 344, ANL79-48, January 1980. F M Haggag, Zircaloy Cladding Embrittlement Criteria: Comparison of In-Pile and Out-of-Pile Results, NUREG/CR-2757, July 1 982. D W Akers e t al., TMI-2 Core Debris Grab Samples - Examination and Analysis, GEND-INF-975, September 1986. Pui Kuan, TMI-2 Upper Core Particle Bed Thermal Behaviour, EGG-TMI-7757, August 1 987. R R Hobbins et al., "PBF Severe Fuel Damage Test 1 -4 Melt Progression Scenario," Severe Accident Research Program Partner ' s Review Meeting, Sandia National Laboratories, Albuquerque, New Mexico, April 25-30, 1988. M Jahn and H Reineke, "Free Convection Heat Transfer with Internal Heat Sources, Calculations and Measurements," Proceedings of the 5th International Heat Transfer Conference, Tokyo, Japan, September 1 974. M Epstein and H K Fauske, The TMI-2 Core Relocation Heat Transfer and Mechanism, FAI/87-44 , July 1 987. Pui Kuan, Core Relocation in the TMI -2 Accident, EGG-TMI -7402, Septem ber 1 986. M Pilch and P K Mast, PLUGM: a Coupled Thermal Hydraulic Computer Model for Freezing Melt Flow in a Channel. SAND82- 1 580, NUREG/CR-3 1 90, ( 1 984).

1 07

Chapter 10 STEAM EXPLOSIONS

10.1

Introduction

In the previous chapter, the phenomena of melt progression under degraded core accident conditions were considered. Under these circumstances a pool of molten core material may arise within the core region surrounded by a hard crust. If the crust fails then molten material may fall downwards under gravity towards the lower vessel, which could well contain residual water. There is therefore a possibility for a steam explosion as the melt contacts the water: this chapter considers the m echanisms involved, some of the experiments performed to provide understanding and the models that have been developed. The consequences of a steam explosion depend on the severity or the magnitude of the energy release and the way in which this energy is dissipated. A large steam explosion could damage the reactor pressure vessel, perhaps generate missiles which might breach the containment. In these circumstances there would be a path for release of fission products to the environment. There are a number of other consequences of a steam explosion which could in principle threaten the containment and lead to early failure. Core debris might be finely fragmented and dispersed within the containment. This could give rise to a considerable pressure spike with heat being rapidly transferred from the finely fragmented melt. Steam generated via a steam explosion will directly contribute to containment pressurisation: this could arise not only from vaporisation of primary circuit water but also from steam released as a consequence of secondary side breach. The quantity of hydrogen and fission products in the containment could be increased due to increased chemical interactions between steam and molten core materials. The generation of hydrogen could also impact on the containment pressurisation, since this could lead to a hydrogen burn in the containment if the hydrogen concentration becomes sufficiently high. A steam explosion has a precise definition in the parlance of light water reactor safety. It is generally defined as a physical interaction between a hot and a cold liquid, in the case ofLWR materials, chemical interactions may also occur. A steam explosion is characterised by a rapid energy transfer from the melt to a small amount of water. For this to occur fine fragmentation is necessary in order to generate a sufficiently large contact area for heat transfer.

Steam Explosions

1 08

The rapid heat transfer from the liquid/melt mixture causes the water to vaporise with consequently high pressures within the mixture. These high pressures could damage directly neighbouring structures close to the region of vaporisation or could accelerate missiles or cause further fragmentation and increase the energy release. Parametric models/codes have been developed for analysis of core melt-down which simulate rapid generation of steam associated with a steam explosion. These codes generally do not model the dynamic effects of a steam explosion or attempt to predict the level of damage resulting from a steam explosion. Separate assessments are carried out to derive a probability of causing containment failure. Some of the available models!codes are considered in this chapter. The damage that might ensue from a large steam explosion in a reactor pressure vessel has been evaluated by comparing the thermal energy that could be stored in core debris (-1 86 GJ) [10. 1 ] with the mechanical energy that would need to be supplied to fail the lower head ( IOOO- I SOOMJ) [ 1 0.2] and the energy required to lift the vessel up to the top of the containment ( 260MJ) [ 1 0. 1 ] . The conclusion is it would be necessary to convert only a relatively small fraction of the thermal energy in the molten core debris to mechanical work to possibly cause serious damage to the vessel and the containment. -

-

(a) S I N G LE MASS

-

-

- - �- �

(b) TH I N J ET

STAG E 1 : F U E L J ET E NTRY P R I O R TO F U E L COO LANT M I XI NG FIGURE 10.1 INITIAL CONDITIONS

Steam Explosions

1 09

The phenomena associated with steam explosions are usually associated with well defined stages Figures 10. 1 10.3. These are justified by experimental investigations. These stages are briefly summarised in the next section and then considered in more detail later in the chapter.

10.2

Phases or a Steam Explosion

(I)

Initial Conditions/Core Melt Down

In a PWR there are essentially three locations where melt can come into contact with water. 1. Within the core region during possible reflood.

2. In the lower vessel due to melt relocation downwards into water that has not previously been boiled off.

3. In the cavity beneath the reactor pressure vessel due to release of molten material into any water that may be present. For interaction within the core region (Case 1), the extent of the energy release will depend on the timing of reflood relative to the stage of core heat up and therefore temperature of the core. The extent of interaction in Case 2 will be dependent of the coherency of core melt which will effect the range of possible flow rates. In Case 3 considered in a later section the rate of interaction will depend on the mode of vessel failure: possibilities include a gross failure of the lower head or possible a lesser failure associated with failure of an instrument penetration tube(s).

(II)

Mixing

Molten debris tends to disintegrate rapidly into a coarse mixture on contact with water. Steam flows also have an influence on the mixing process: some break-up can be accounted for by hydrodynamic instability. These processes are complex and few fundamental hydrodynamic models have been developed. Other simpler models have been developed which aim to provide bounds for the process.

(III)

Triggering

Triggering events can occur spontaneously or can result from external events. To initiate a steam explosion, a consensus view is that it is necessary to initiate a pressure pulse to set off a fragmentation process which then propagates through the system . Spontaneous triggering often occurs on solid surfaces but can also occur on the surface of a quantity of molten debris. These spontaneous events may arise through compression of coolant against a surface, where the expansion of trapped water is impeded, or by

1 10

Steam Explosions

shocks associated with bubble collapse. Either of these events can produce a pressure pulse. External triggering mechanisms might include steel component impacts or collisions or explosive detonation.

(IV)

Propagation

Propagation occurs due to a passage of a pressure wave driven through the mixture by energy derived from the expansion of coolant heated by fragmented melt. The propagation velocity as derived from film is typically in the range 50 to 500 mls provided the mixture characteristics are appropriate. It appears that once triggered, the interaction will proceed rapidly through the system. This process has been described by a thermal detonation model [ 10.3] which draws on the analogy between steam explosions and chemical explosions. The model assumes that the propagating pressure wave collapses any vapour blankets surrounding the relatively coarsely mixed particles, producing finer hydrodynamic fragmentation and greater area for heat transfer. This leads to rapid vapour generation and high local pressures behind the wave front. The flow behind the front is therefore accelerated and is sustained against dissipative processes. Calculational models are currently limited by several problems. The fragmentation of liquid drops due to a sudden pressure wave is not well understood and hence the rapid heat transfer to the coolant cannot be accurately predicted. Secondly, the mechanisms of melt progression and the quantities of melt and liquid interactions are uncertain, i.e. it is difficult to specify the initial conditions for the interaction.

(V)

Expansion

Thermodynamic arguments can be used to determine an upper limit for the conversion of thermal to mechanical energy. Mechanical energy is generated by the expansion of the heated steam. The optimum mixture, for the materials of interest corresponds to approximately equal volumes of melt and coolant. The upper limit for thermal to mechanical energy conversions is in the range 30 to 50%: however this is much higher than the values observed from experiments. Typical experimental values are more of the order of 1 % (with considerable scatter above and below this value) and therefore these thermodynamic considerations are of limited value in safety analysis since they produce unrealistically high and conservative values. For the analysis of experimental data the measured yield is normally defined as: Yield = Mass x Efficiency x Specific Thermal Content.

Steam Explosions

111

The first two tenns in the right hand side can be thought of in several ways. The mass referred to is nonnally the total mass available for interaction in some sense, e.g. in a PWR it could be the mass of melt that has dropped from the core into the lower head. This may be less than the total mass melted. It is quite difficult to identify the mass actually involved in the interaction. For the efficiency a typical value of 1 % is used for estimating the yield. The assessment of the severity of a steam explosion requires a knowledge of the energy yield as well as the pressure wave generated. As stated earlier the damage arising from a steam explosion can be caused either by the pressure spike generated by the explosion or by the mechanical momentum of the fluid as it is accelerated away from the explosion site. There are also difficulties in extrapolating from the relatively small experiments to reactor scale. In order to assess the contribution that steam explosions make to risk, it is necessary to consider the occurrence probability that an in-vessel steam explosion is sufficiently large to cause a breach of the containment. In-vessel steam explosions might involve much larger quantities of liquids than those in experiments: the size distribution of debris may also be different. The pressures could also be different from those considered in experiment. Assessment of the probability and the consequences of an in-vessel steam explosion will necessarily involve extrapolation from experimental correlations to conditions beyond those investigated in experiments where these correlations may not be valid. This introduces uncertainties in the assessments. There is also uncertainty in the state of the molten core prior to the steam explosion. These two uncertainties must be combined and the result is that the extrapolation of results from relatively small scale experiments is subject to considerable uncertainty. The extrapolation will be particularly difficult in the context of PWR where the pressure vessel contains complex and extensive internal structures which could effect the chain of events. These structures tend not to be included in experiments.

10.3

Boundary Conditions

The conditions at core slump are scenario dependent. For an unrecovered hot leg breach with total failure of all core cooling systems, core slump has been predicted to occur as soon as an hour after accident initiation. Pressures would be in equilibrium with the containment and quite low e.g. -O.3MPa. For an unrecovered small break accident e.g. the S2D, core slump would occur in -about 1 .5 hours from the initiating event: the pressures might be about 7MPa. For an intact circuit fault where the circuit remains intact up to vessel failure, the pressures could be as high at 1 7MPa. The quantities of melt involved are uncertain, see the discussion in the previous chapter. The temperatures of the melt are also subject to uncertainty.

1 12

Steam Explosions

The lack of understanding of core melt-down creates uncertainty in the conditions that might exist prior to a steam explosion. Melt progression could take place in such a way that a large pool of debris is supported by a crust (as discussed previously). Continued heating of the pool via decay heat and possibly some oxidation heating would mean that eventually the crust would fail and a coherent flow of molten debris would then fall downwards into the lower head. It has been suggested [ 1 004] that as much as 75% of the core could fall into the lower plenium via this mechanism although this is likely to be pessimistic: if this happened however a damaging steam explosion might ensue. On the other hand molten debris might fall continuously [ 1 0.5], but not coherently, resulting in a series of lesser events perhaps with quench and resulting in no, little or few steam explosions. However even a single steam explosion might be sufficient to cause the blockage to fail and lead in a more energetic steam explosion event

10.4

Mixing of Molten Debris with Water

The extent of coarse mixing of molten core material with water has an important impact on the quantity of melt participating in any steam explosion and therefore the energy yield. It is important to assess the extent of coarse mixing prior to an in-vessel steam explosion. Experiments have shown that molten fuel breaks up and mixes in the surrounding water. The mechanism for the break-up is likely to be both hydrodynamic instability as well as the influence of steam flows. Mixing is a complex process and various simple models have been developed to attempt to bound the process. Calculations to determine the energy required to break up and mix two liquids indicate that drag forces are very important [ 10.6] . This approach enables bounds to be placed on the extent and rate of mixing when account is taken of the energy available. For a PWR melt-down it is virtually impossible to generate a single stage steam explosion without a coarse mixing stage. The length and timescales for the mixing process where drag might be considered important or otherwise have been considered by Porter [ 1 0. 1 ] . If molten debris and coolant intermixed on a scale of O.O l m in a 0. 1 - 1 .0s timescale then little viscous dissipation would occur. However mixing on a scale of O.OO l m in a time of order O.OO l s requires large quantities of energy to overcome drag forces. When molten debris falls into water the m ixing process results in rapid steam production. Two models have been considered which propose limits for the steam flows. Henry and Fauske [ 10.7] , [ 10.8] have argued that since water flows into the mixture, the maximum steam flow can be approximately estimated by considering the critical heat flux (CHF). The model assumes steady state conditions, that the water and rising steam are in counter-current flow and uses a flat plate CHF correlation in combination with the vessel area. Corradini has proposed a model [ 10.0], [ 1 0. 1 0] , [ 10. 1 1 ] based on experimental work at Sandia National Laboratories, USA. This allows for a 3D representation with water flow into the bottom and sides of the mixture and steam flowing out of the top. The Corradini approach was based on the sweep-out of either the fuel or the water by upward

1 13

Steam Explosions

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VESSEL

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•

,

.. It

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,

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STAG E I I : COARSE M I XI NG WITH S LOW H EAT TRAN S F E R AN D N O P R E SS U R E I N CR EAS E

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..

.. •

4 ,

•

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STAG E I I I : TR I G G E R I N G PROCESS WITH H I G H LOCAL P R ESSU RE

FIGURE 10.2 MIXING AND TRIGGERING PROCESS

1 14

Steam Explosions

flowing steam. Pool depth is a critical parameter in this model. The maximum participating melt mass increases rapidly with pool depth. Other insights on mixing are given in [1004]. Henry and Fauske [10.12] suggest that steam explosions only occur when the melt particles are small « lcm in diameter) but there is conflicting evidence [ 10. 12], [ 10. 13] . The tendency in modelling have been to ignore the effect of steam generation on mixing. Indeed enhancement of mixing by steam can cause stratified configurations of water overlying the melt to explode [ 10. 14]. Corradini has shown that changing the criterion from a critical heat flux limitation to a fluidisation assumption can give between one to two orders of magnitude increase in mixed mass of melt depending on the particle size. The mass of melt is also geometry dependent. If a cylindrical configuration rather than a spherical configuration is assumed then the mass of melt is further enhanced. Corridini' s model allows for the break-up of the melt [ 1 0. 1 1 ] but ignores instabilities arising from steam flows. The small scale of most past experimental work « 10kg of melt) makes extrapolation to greater masses very speculative. Chamberlain [ 10. 15] suggests that the Fauske model [ 10.8] may underestimate the degree of mixing. At the present time the degree of mixing, particularly on the larger or plant scale, is uncertain. The experiments conducted to understand fuel-coolant mixing are fairly limited. Significant work has been carried out at Sandia National Laboratories and Argonne National Laboratories, Bennan et al 1980-1990 [ 10. 1 6] , [10. 17], Mitchell et al 1981 [ 1 0. 1 8], 1982 [ 10. 19], 1986 [ 10.20] , Marshall et al 1988 [10.21] , [ 1 0.22] and Spencer et al 1986 [10.23] , 1987 [1 0.24, 1989 [1 0.25] . In the Sandia experiments, the fuel mass was usually poured into the water coolant as a discrete mass, i.e. the pour diameter was of the order of the fuel melt diameter. In the Argonne tests, the fuel mass entered as a liquid jet, i.e. the pour diameter was small relative to the fuel mass and its diameter. Larger masses were involved in the Sandia experiments (2-50kg) compared with the Argonne experiments ( 1 - 10kg). In both cases the fuel kinetics in the coolant pool were observed, the fuel leading edge penetration rate was measured and the radial spread of the fuel. In addition to the measurement of the jet kinematics both investigations estimated the amount of fuel delivered to the coolant pool , and in some cases the pool level swell, due to fuel entry and vapour production, was measured. Both groups used the Chu et al model 1984 [1 0.26] to predict the debris size. This model has been further refined by Chu who has produced a mechanistic model [ 1 0.27] , [ 1 0.28] , [ 10.29] for dynamic fuel fragmentation.

Steam Explosions 10.5

1 15

Triggering

If a disturbance triggers a steam explosion, then the heat transfer will occur very much faster than if the melt is simply quenched. Heat transfer will automatically take place due to the temperature gradient between the two liquids. Spontaneous triggering (e.g. when the source of the disturbance is not immediately apparent), often occurs at a solid surface, although it can also occur at the melt surface. Externally triggered steam explosions usually have an easily identifiable source of disturbance. Externally triggered steam explosions have been identified in experiments e.g. due to the impact of one steel structure on another and the firing of mini-detonators and bridge wires. Multiple steam explosions have also been observed [ 10.30] where several triggering events have occurred, later ones probably as a result of earlier explosions. The conditions under which contact between melt and water occur can effect the triggering mechanism of the steam explosion. Bennan et al [ 10.4] have investigated the effect on triggering of varying the ambient conditions but the data-base is sparse at high pressures. Experiments for pressures in the range 0. 1 to 0.8 MPa [10.3 1 ] , [ 10.32] provide some evidence that triggering for small drops occurs in this pressure range. For larger melts [10.33] this was not so but efficient triggering was given if a detonator was used. Externally triggered explosions were observed at ISPRA for pressures as high as 3.0MPa [10.34] but only with difficulty. These results have led to the tentative conclusion that increase in pressure tends to discourage triggering. In addition there is some evidence from single droplet tests that non-condensable gases such as hydrogen tend to impede triggering, as also does water near its saturation temperature. However intermediate scale tests do not necessarily support these tentative observations. When a small fuel mass enters water there is a delay between initial entry and any spontaneously triggered FCI. This has been observed by most investigators. Experiments with tin/water, Frost et al 1988 [10.35] have shown that as the temperatures of either component increase, at same point, no spontaneously triggered FCIs are observed. If fuel solidifies as it falls through the liquid pool then the outer surface can prevent liquid fuel fragmentation after film collapse. This solidification boundary is important for a triggered FCI particularly when the stable film boiling temperature is far below the fuel melting temperature as in the VO/water and iron oxide experiments carried out by Nelson et aI, 1977 [10.36] , [10.37] , 1988 [ 10.40] , and Akiyoshi et al 1990 [ 10.4 1 ] . Kim and Corradini 1985 [ 10.42] have conjectured that the mechanism i s that the gas stabilises the vapour film by acting as a cushion and thus helping to resist film collapse. In various small scale tests, Kim 1989 [ 10.43] it has been noted that multiple cyclic interactions are more likely to occur as the fuel mass increases, probably because the fuel mass is not completely fragmented initially. Explosions can also occur as the fuel settles on the base of e.g. the water pool . This has been observed in some of the large scale industrial tests for the steel industry and more

1 16

Steam Explosions

recently Nelson et al 1988 [10.44] has obtained data for aluminium/water explosions. Material properties of the fuel and coolant can also influence the triggering, Nelson et al 1982 [ 10.39], 1986 [ 10.40] and more recently Kim et al 1986 [ 10.45] and 1989 [ 10.43] . Nelson showed that increase in the fuel or coolant viscosity tends to suppress the explosion. It is clearly difficult to extrapolate from small scale tests, where the observations are little understood , to the plant scale. The effects of many of the complex structures within the inner vessel, not simulated in these smaller scale experiments, are another source of uncertainty. 10.6

Mechanical Energy from Steam Explosions

The size of a steam explosion may be conveniently characterised in terms of the mechanical work done by the expanding melt/water/steam mixture. This may easily be related to experimental data but it should be noted that the work done is strongly dependent on the environment of the steam explosion. This definition of yield is not a measure of the maximum potential of a steam explosion to do work, e.g. the yield can be increased if the volume available to a steam explosion is increased. In experiments this yield can be derived from the kinetic energy of surrounding materials and from the compression of gas if the explosion takes place in a confined space. Under reactor conditions, the former is likely to be the more significant. The work yield from a steam explosion is influenced by a number of factors. These include: (a)

the extent of coarse melt/water mixing prior to the explosion usually only a fraction of the melt mixes effectively and undergoes fine fragmentation and rapid heat transfer during the explosion,

(b)

the process of fine fragmentation this directly influences the rate of heat transfer from the melt to the water,

(c)

loss of heat during the expansion stage. This can result in reduced yield if heat is lost to cooler structures such as steel structures or adjacent masses of water.

Work yields have been measured in numerous experiments for small- and intermediate­ scale steam explosions. In the SUW experiments at Winfrith [10.46], yields of 0.2% to 2.3% of the available thermal energy were reached. These small scale experiments involved releasing thermite generated urania and molybdenum melts of mass about 24 kg into 1 .5 Te of water within an enclosed vessel. Intermediate scale experiments at Sandia [ 10.33] , [ 10.47], [ 10.48] showed similar yields: the largest mechanical yield was about 4.4% which occurred in a single droplet explosion [10.3 1 ] .

Steam Explosions

FINE M I XT U R E P R E SS U R E WAVE

1 17

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COO LANT M I SS I L E

o

STAG E V: COO LANT AN D/O R M I SSI L E EJ ECTI O N

FIGURE 10.3 PROPAGATION AND EJECTION PROCESSES

118

Steam Explosions

An experimental investigation of scaling in experiments perfonned by Bird [ 10.46] at Winfrith has shown that the efficiency of thennal to mechanical energy conversion increases with decreasing water sub-cooling. The efficiency was largely independent of melt mass and was unchanged over a melt mass range 0.03- 18.0kg. These experiments also showed that the fraction of melt participating in the steam explosion increased with increasing ambient pressure. Thus the conversion efficiency referring to mass of melt released into the water increased with rising pressure. It must be emphasised that a steam explosion could occur in a PWR pressure vessel under conditions radically different from those conditions studied in experiments. The most significant differences could be in respect of the quantities of melt and water participating and the pressure conditions which could be much higher. Estimates of conversion ratios have been summarised in Bennan et al [ 10.4] who cites work of Mayinger [ 10.5], Theofanus [ 1 0. 1 2] , [ 10. 13], Becker, Haag and Korber and Spencer [1 0.49]. Different conversion ratios in the range 0.4 to 16% are quoted: these are also consistent with Gittus. Calculations for the Zion/lndian Point Study suggest that at the larger scale, conversion ratios in the upper range may be the more plausible. These calculations, however, made use of important simplying assumptions and highlighted the difficulties in extrapolating the present experimental evidence to the larger, plant scales of interest. It is clear that there are large differences between the conditions realised in steam explosion experiments and those that might be realised in a PWR steam explosion. Consequently calculations of conversion efficiency based on thermodynamic considerations can play a useful role in producing possible bounds on the efficiency of an in-vessel steam explosion. On approach has been adopted by Hicks and Menzies [10.50] which has been used in comparisons with the efficiencies measured in experiments. This method assumes that: (1)

the melt mixes with a prescribed mass of water and this system reaches thennal equilibrium at constant volume,

(2)

the system expands isentropically to a prescribed end state with the melt and water under thermal equilibrium conditions.

The maximum efficiencies using this method are found to be in the range 30-50% for expansion of the mixture to atmospheric pressure. These are reduced if the expansion takes place in an enclosed volume where the expansion is restricted. This method is a straightforward way of estimating the maximum magnitude of the maximum conversion efficiency. Since the Hicks Menzies approach is general it can be applied to any fuel-coolant state with given equations of state and it has been improved upon by many investigators,

Steam Explosions

1 19

Cline et al 1989 [10.5 1] and Bang et al 1990 [ 10.52] . The latter used the Hicks Menzies approach to estimate an upper bound work potential for a range of fuel coolant pairs. Board et al 1975 [10.3] used an analogy between a vapour explosion and a chemical detonation to estimate the maximum work potential. The Board et al model has been compared with the Hicks Menzies model by Bang et al 1990. The comparison indicated that the net work output from the explosion is less than or equal to the Hicks Menzies result indicating that the latter probably represents a truer upper bound for a vapour explosion. It has also been observed that fuel material has an effect on the work output. The peak conversion ratio for a corium water explosion is much larger -60% than that of an aluminium/water system -40%. Corradini et al 1988 [10.53] has reviewed various theories which have been advanced to describe the vapour explosion progress. 10.7

Damage Potential

The main concern resulting from a large in-vessel steam explosion is with damage to the reactor pressure vessel and the containment. Breach in the pressure vessel results in a release of fission products at least to the containment and possibly outside, if the containment is breached (e.g. as a consequence of the explosion). Such a breach could result from missiles set in motion by the explosion: these could be the pressure vessel itself or attachments such as control rod drive mechanisms. Corradini and Swenson [ 10.54] have considered gross vessel motion. For the vessel of a large modem PWR to rise to the top of the containment building would require a mechanical yield of the order 1600MJ. If an explosion generated this yield it is possible it would probably be mitigated by failure of the bottom of the vessel. Several studies [ 10.2] [ 10.54] of steam explosion induced failure have been made which treat the pressure vessel as a continuous elasto-plastic system. Analysis of vessel failure is complex since the mode of failure can be influenced by flows. Energetic missiles could result by fragmentation of the pressure vessel. In the Zion/Indian Point Study [ 10. 1 1] it is stated that the maximum slug energy that could be generated by a steam explosion without failing the bottom head is in the range 1 000- 1 500MJ. In a larger explosion the bottom head would fail and then the yield would be reduced by the downward material flow through the breach. Thus under such circumstances a considerably larger explosion would be required to generate an upward moving missile of sufficient energy to breach the containment. Top head vessel failure has been considered in several studies [ 10.55] [ 10.56] and [1 0.2]. A conservative approach was adopted in some of these e.g. in [10.55] and [ 10.2] the impulsive loading of the top head due to the impact of a closely fitting material slug was

Steam Explosions

1 20

considered. However in practice, the slug would not fit the top head exactly, therefore being retarded on arrival in the time it would take to travel its own length with a correspondingly reduced loading over an extended period. In the Zion/lndian Point Study [10.2] more realistic loadings resulting from slugs of material contacting the upper head at its centre were considered. With this model, a l 00Te slug of kinetic energy greater than 1200MJ would be expected to cause top head failure. It is uncertain how failure of the top head would occur and to what extent it could become a large missile or several small missiles. A mechanism for generating small missiles could be ejection of components attached to the top head, resulting from an impulsive loading. There is however considerable uncertainty on which components would in fact become missiles and their mode of failure and acceleration. The generation of such missiles was also considered in the Zion/lndian Point Study [10.2] . Even if a steam explosion does not occur as molten material contacts water, some consideration needs to be given as to whether the resulting steam spikes could overpressurise the primary circuit and cause it to fail. This might present a threat to the containment if missiles were generated, perhaps fracturing steam generator tubes which would then result in bypass of the containment. Threats following steam spikes have been considered by Baines et al. He demonstrated that for the loss of power transient (TMLB) a steady power of about 0.35 OW would have to be delivered to the coolant over a period of l00s to cause steam generator tube failure. When additional factors such as heat and mass sink effects associated with the steam generators and also hydrogen are taken into account, the steady heating rate required to cause steam generator tube failure is about 0.70W. This is hardly feasible in the case of a large PWR and although there are significant uncertainties it is thought that a steam spike is unlikely to cause primary circuit failure and containment bypass. 10.8

Experiments

The physical processes in steam explosions have been studied in several series of experiments at Winfrith. In the first series, [ 10.30] relatively small -Q.5kg quantities of thermite generated urania and molybdenum melt were released under water in closed vessels. To examine scale effects larger scale testing involving 24kg of melt was also carried out in a second series of experiments [ 10.46] . These aimed to examine the effects of pressure and reduced cover gas volume on melt dispersal. The programmes showed that reducing the compressibility of the system by decreasing the cover gas volume or by increasing the initial gas pressure had the effect of restricting the rate and extent of dispersion of the melt into the surrounding water. Increasing the

Steam Explosions

121

system compressibility tended to enhance the oscillatory nature of the pressure and resulted in coarser debris. Yields were also significantly reduced by subcooling, at least with a large excess of water. At the 24kg scale there was a significant increase in the mass of melt involved and in the yield, if the time of trigger was delayed. This was not observed in the 5kg tests where the opposite was true possibly because of melt freezing effects. The work efficiencies were between 3 and 16% of the Hicks-Menzies efficiencies. 10.9

Calculational Models

Bounding calculations have been perfonned to estimate the pressure generated if the steam and water contents in the reactor vessel are equilibriated with the molten corium. This results in extremely high estimates of pressure even neglecting any Zircaloy reaction. It is clear that there is a significant potential for high pressures provided that the heat transfer rate exceeds any dissipation mechanism via the safety valves or other heat sinks. A series of codes have been developed at Winfrith to model a range of different steam explosion mechanisms. This suite includes a calculation for the ideal Hicks - Menzies efficiency: the parameters being different melt properties and the melt/water ratios either with or without steam blanketing. The mass fraction partition of the debris is detennined by a log-nonnal distribution provided that the mean size and standard deviation are supplied. The detennination of heat transfer from the particle spectrum can be derived via the Bird model [10.30] using the methods of Carslaw and Jaegar [10.57] for heat conduction. This method was generalised by Porter [10. 1 ] for the Winfrith codes. The method have been applied to situations inside the pressure vessel, where even if subcooling is absent, there might be high ambient pressure and also to situations in the cavity where inertial and subcooling effects are important. In the application of computer codes, an outstanding uncertainty is associated with the general momentum balance since the hydrodynamic surface boundary conditions are ill-defined (e.g. the surface can be relatively easily vented). Clearly the codes require validation against experiment to provide confidence in their prediction of pressure transients and impulse loadings. However it must be appreciated that in a PWR the situation is very complex compared with experiment. The multiplicity of components and phases in the plant system give it many degrees of freedom. 10.9.1 Heat Transfer from Particles Following a Steam Explosion

The size distribution of the fragments resulting from a steam explosion have been determined from the Winfrith experiments. These indicate an average particle size of about I mm with a log-nonnal distribution and a mean deviation of 0.5 on a log 10 basis.

Steam Explosions

1 22

The Winfrith experiments tend to imply that average size may be reduced with increasing pressure. The fractional heat removal from a particle is relatively unaffected by its shape for an infinite heat transfer coefficient and for an isothennal heat sink. The expression for a sphere is tabulated in [ 10.57] . Small particles tend to give above average heat release in the early stages, large particles give above average release in the later stages. For the particles from the Winfrith experiments most of the heat would be released over about 0. 1 s: if every particle was 1 0 times bigger then the significant release would be over lOs. Particle size is therefore a critical parameter even more so than distribution in determining timescales of heat transfer. For a finite external heat transfer coefficient, the time to achieve a given fraction of release, all other parameters remaining the same is inversely proportional to the heat transfer coefficient Black body radiation alone at temperatures of 2000K would indicate 1kw/m2K ignoring interparticle radiation. With this external resistance to heat transfer the particulate (in the infinite heat transfer coefficient case) material would be expected to release its heat on a timescale of 1 s rather than O. ls. 10.9.2 Relief and Safety Valves

Methods for the release of fluids have been well reported [10.10] , [10. 1 1 ] , [ 10.33] . In the Winfrith model, the fluid is assumed to expand isentropically maintaining its total enthalpy and its entropy to the throat or exit static pressure condition. The pressure is determined to maximise the mass flow. The effective area for release is calculated typically taking the total effective area of the relief and safety valves. The calculations make no allowance for an additional release area in lifting the upper head. A complicating factor in the modelling is the uncertainty on the hydrogen concentration. Hydrogen is useful in encouraging the molar or volume release through the valves. Another factor allowed for in the calculational model is the influence of accumulators. REFERENCES

10.1 1 0.2 1 0.3 10.4 1 0.5

A T D Butland et al (1 984) "Report on Phase 1 of the PWR Severe Accident Containment Study", AEEW - R 1842. W B Murfin (ed) (1980) "Re}XXt of the Zion/Indian Point Study", NUREGICR- 1410. S J Board, R W Hall and R S Hall ( 1975) Nature 254, 3 19. M Berman, D V Swenson and A J Wickett (1984) "An Uncertainty Study of PWR Steam Explosions", NUREG/CR-3369. F Mayinger (1 982) Atomwirtschaft, Feb, P74 - English Translation: UKAEA Risley Trans. 4636.

Steam Explosions

10.6 10.7 10.8 10.9 10. 1 0 10. 1 1 10.12 10. 1 3 10.14 10.15 10. 16 10. 17 10. 1 8 10.19 10.20 10.2 1 10.22

1 23

D E Cho, H K Fauske and M A Grolmes (1976) "Some Aspects of Mixing in Large Mass Energetic Fuel-Coolant Interactions" , Proc . Int. Mtg. on Fast Reactor Safety and Related Physics, Chicago, CONF-761001 , �, 1 852. R E Henry and H K Fauske (198 1 ) "Required Initial Conditions for Energetic Steam Explosions", ASME Winter Mtg. Washington DC, 198 1 . H K Fauske ( 1982) "Scale Considerations and Vapour Explosions (Rapid Phase Transitions)", LNG Safety Workshop, MIT. M L Corradini ( 1982) "Proposed Model for Fuel-Coolant Mixing", Trans. ANS, 4 1 , 4 1 5. M L Corradini (1982) "Proposed Model for Fuel-Coolant Mixing during a Core-Melt Accident", Proc. Int. Mtg. on Thermal Nuclear Reactor Safety, Chicago, 11 1 . NUREG/CP-0027, 2 1399. M L Corradini and G R Moses ( 1 983) "A Dynamic Model for Fuel-Coolant Mixing", Proc. Int. Mtg on LWR Severe Accident Evaluation, Cambridge, MA, 6.3- 1 . T G Theofanus and M Saito ( 198 1) "An Assessment of Class 9 (Core-Melt) Accidents for PWR Dry Containment Systems", Nucl.Eng. Des. 66, 30 1 . "Preliminary Assessment of Core Melt Accidents at the Zion and Indian Point Nuclear Power Plants and Strategies for Mitigating their Effects", ( 198 1), USNRC, NUREG0850, Vol 1 . G A Greene et al (1983) "Some Observations on Simulated Molten Debris­ Coolant Layer Dynamics", Proc. Int Mtg. on LWR Severe Accident Evaluation, Cambridge, MA, 12.2- 1 . A T Chamberlain and F M Page ( 1983) "An Experimental Examination of the Henry­ Fauske Voiding Hypothesis", University of Aston, Birmingham, UK . M Berman, LWR Safety Quarterly and Semi Annual Progress Reports, Sandia National Laboratories Internal Reports. M Berman, D Beck and L S Nelson, Review of Molten Fuel-Coolant Interactions, Ex-Vessel Severe Accident Review for the Heavy Water NPR, Sandia National Laboratories, SAND 90-0234, 1990. D E Mitchell, N A Evans and M L Corradini Intermediate Scale Steam Explosion Phenomena: Experiments and Analysis, SAND 8 1-0 1 24, NUREG/CR-2 145, Sandia National Laboratories, September, 198 1b. D E Mitchell, N A Evans, The Effect of Water to Fuel Mass Radio and Geometry on the Behaviour of Molten Core-Coolant Interaction at Intermediate Scale, Proc. Int. Mtg. Thermal Nuclear Reactor Safety, Chicago, IL NUREG/CP-0027, August 1982. D E Mitchell and N A Evans, Steam Explosion Experiments at Intermediate Scale: FITSB Series, NVREG/CR-3983, SAND83- 1057, Sandia National Laboratories, Feb 1986. B W Marshall Jr, Recent Fuel-Coolant Interaction Experiments Conducted in FITS Vessel, 1988 ASME- AIChE National Heat Transfer Conference, Houston, TX, lITC-Vol 3, 265-278, July 1 988. B W Marshall, D F BECK and M BERMAN, Mixing of Isothermal and Boiling Molten-Core Jets with Water: The Initial Conditions for Energetic FCls, Proc. of the International ENS/ANS Conference on Thermal Reactor Safety, Avignon, France, Vol 1 , 1 17-127, October 2-7, 1988.

1 24

Steam Explosions

1 0.23 B W Spencer, J D Gabor and J C Cassulo, Effect of Boiling Regime on Melt Steam Breakup in Water, IntI. Symp. on Multiphase Transport, Miami, 1986. 10.24 B W Spencer, J Sienicki, and L McUmber, Hydrodynamics and Thennodynamics of Corium-Water Thennal Interactions, EPRI Report, NP-5127, 1 987. 1 0.25 B W Spencer, S K Wang, C A B lomquist, L McUmberand J P Schneider, Experimental Study of the Fragmentation and Quench Behaviour of Corium Jets in Water, ANS Winter Annual Meeting San Francisco, November 1 989. 1 0.26 C C Chu and M L Corradini, Hydrodynamic Fragmentation ofLiquid Droplets, Trans. Am. Nuc. Soc., 47: Washington, DC (Nov 1984). 10.27 C C Chu and M L Corradini, One-Dimensional Transient Model for Fuel-Coolant Fragmentation and Mixing. ANS/ENS International Meeting on Thennal Reactor Safety, February 1 986. 10.28 CC Chu and M L Corradini, One-Dimensional Transient Model for Fuel-Coolant Interaction Analysis, Nuc. Sci. Engr., 1 0 1 : 48-7 1 (January 1 989). 10.29 C C Chu, One-Dimensional Transient Fluid Model for Fuel-Coolant Interactions, Ph.D. Thesis, University of Wisconsin, UWRSR-30, May 1986a. 1 0.30 M J B ird (1981) "An Experimental Study of Thermal Interaction between Molten Uranium Dioxide and Water", AEEW-R I40 1 . 10.3 1 L S Nelson and P M Duda (to be published) "Steam Explosion Experiments with Single Drops of Iron Oxide Melted with a CO2 Laser, Part II, Parametric Studies", NUREG/CR-27 18. 10.32 L S Nelson and P M Duda ( 1 982) "S team Explosions of Molten Iron Oxide Drops: Easier Initiation at Small Pressurizations", Nature 296, 844. 10.33 D E Mitchell, M L Corradini and W W Tarbell (198 1 ) "Intermediate Scale Steam Explosion Phenomena: Experiments and Analysis", NUREG/CR-21 45. 10.34 R Hohmann, H Kottowski, H Schins and R E Henry ( 1982) "Experimental Investigations of Spontaneous and Triggered Vapour Explosions in the Molten Salt/Water Systems" , Proc. Int. Mtg. on Thermal Nuclear Reactor Safety, Chicago, NUREG/CP-0027, 2, 962. 10.35 D Frost and G Ciccarelli, Dynamics of Explosive Interactions Between Multiple Drops of Tin and Water, Dynamics of Explosions, V1I4, AIAA, Washington, DC (1988). 1 0.36 L S Nelson and L D Buxton, The Thermal Interactions of Molten LWR Core Materials with Water, Trans. Am. Nuc. Soc., 26: 397 (1 977). 10.37 L S Nelson and L D Buxton, Steam Explosions Triggering Phenomena: Stainless Steel and Corium-E Simulants Studies with a Floodable ARC Melting Apparatus, SAND 77-0998, NUREG/CR-01 22. 10.38 L S Nelson and P M DUda, Steam Explosion Experiments with Single Drops of Iron Oxide Melted with CO2 Laser, SAND 8 1 - 1 346, NUREG/CR-2295, Sandia National Laboratory, September 1 98 1 . 1 0.39 L S Nelson, and P M Duda, Steam Explosion Experiments with Single Drops of lmn Oxide Melted with a CO2 Laser, Part 2, Parametric Studies, SAND 82- 1 105, NUREG/CR-27 18 , Sandia National Laboratory, September 1 983 . 10.40 L S Nelson and K P Guay, Suppression of Steam Explosion in Tin and Fe-Al20J Melts by Increasing the Viscosity of the Coolant, High Temperatures-High Pressures, 1 8: 107- 1 1 1 ( 1986).

1 25

Steam Explosions

10.4 1 R Akiyoghi, S Nisho, I A Tanasawa Study on the Effect of Non-Condensible Gas in the Vapor Film on Vapor Explosion, In. 1 . Heat Mass Transfer, 33(4): 603-609 ( 1990). 10.42 B J Kim and M L Corradini, Recent Film Boiling Calculations: Implication on Fuel­ Coolant Interactions, Fifth Int. Mtg. on Thennal Nuclear Reactor Safety, Vol 2, Karlsruhe, FRG, October 1985. 10.43 H Kim , M L Corradini and 1 Krueger, Single Droplet Vapor Explosions Experiments: Effect of Coolant Viscosity, Proc. of NURETH-4 Conf., Karlsruhe, FRG, 1 989. 10.4 5 H Kim and M L Corradini, Single Droplet Vapor Explosion Experiments, Proc. of the International ANS/ENS Topical Meeting on Thennal Reactor Safety, San Diego, CA, February 1986. 10.46 M J Bird ( 1984) "An Experimental Study of Scaling in Core Melt/Water Interactions" , 22nd Nat. Heat Transfer Conf, Niagara Falls, 84HT 1 7 . 10.47 L D Buxton, W B Benedick an d M L Corradini ( 1 980) "Steam Explosions Efficiency Studies, Part II: Corium Experiments", NUREG/CR- 1 746. 10.48 L D Buxton and W B Benedick (1979) "Steam Explosion Efficiency Studies", NUREG/CR-0947. 10.49 D Squarer and M C Leverett ( 1 983) "Steam Explosion in Perspective", Proc. Int. Mtg. on LWR Severe Accident Evaluation, Cambridge, p6. 1 - 1 . 10.50 E P Hicks and D C Menzies ( 1 965) ''Theoretical Studies on the Fast Reactor Maximum Accident", ANL-7 1 20, p654. 10.5 1 D C Cline, L Pong, M Bennan, An Equation of State Fonnulation for Hicks-Menzies FCr Efficiencies, Proceedings 26th National Heat Transfer Conference, Philadelphia, PA, August, 1989. 10.52 K H Bang and M L Corradini, Thennodynamic Analyses of Vapor Explosions, Technical Note, Nuclear Science and Engineering (Submitted for Publication 1990b). 10.53 M L Corradini, B 1 Kim, M D Oh, Vapor Explosions in Light Water Reactors: A Review of Theory and Modelling, Progress in Nuclear Energy, 1 : 1 - 1 1 7 ( 1 988). 10.54 M L Corradini and D V Swenson ( 1 981) "Probability of Containment Failure due to Steam Explosions Following a Postulated Core Meltdown in an LWR NUREG/CR22 14. 10.55 "Reactor Safety Study - An Assessment of Accident Risks in US Commercial Reactor Plants", ( 1975), WASH- 1400, NUREG-75/0 14. 10.56 "The Gennan Risk Study", (1979), Published by the Federal Minister of Research & Technology (S ummary in English, main document in Gennan). 10.57 H S Carslaw and J C Jaeger (2nd ed, 1 959) "Conduction of Heat in Solids" publ. OUP, p.234. ",

1 27

Chapter 11 DEB RIS C O OLABILITY MODELS

11.1

Introduction

If a reactor core loses cooling for an extended period then the core will become uncovered and the temperatures rise. Examples of such scenarios, the unrecovered LOCA or the station black-out have been discussed earlier. Under these circumstances cladding oxidation will become important and there will be a temperature excursion driven by the cladding oxidation reaction. This will initiate a stage by stage melting process, involving fustly the control materials and secondly low temperature cladding eutectics. These processes cause blockages to form as the molten material relocates downwards and freezes. In many of the scenarios the lower head will still be full of water as the melt falls into it. The melt may react violently with the water by the mechanisms described in the previous chapter or more quiescently by mechanisms which are the subject of this chapter. B lockage formation has been amply demonstrated in various experiments and a significant blockage has also been found from post accident examination of Three Mile Island. The sudden shock of quench tends to result in the formation of more porous blockages: porous debris is produced due to fragmentation by thermal stresses. The porous blockage or bed will be internally heated through its own decay heat and also by possibly some oxidation heating. A key issue to the safety analyst is the extent to which the blockage is coolable. This will depend on whether flow paths are produced through the localised regions of the plant. If the blockage is sufficiently wide and coherent and if the blockage obstructs the inlet (in the case of the PWR then flow will be forced into the debris. Whereas if the blockage is higher up and surrounded by unblocked channels bypass effects would be important and the resultant blockage would be somewhat hotter. If coo lent flow through the bed is not sustained then the bed will overheat and the integri ty of a debris bed will depend on whether it can be quenched or not. Whether or not it can be quenched could ultimately determine whether the accident will be confined in-vessel or whether the vessel will be breached and the containment threatened. If the vessel is breached then debris will be ejected possibly energetically if the circuit pressure is high and debris will encounter water, thus forming an ex-vessel debris bed. If the debris is not coolable in this configuration, then again it will tend to dry out because of its own decay heat and start to attack the concrete basement. If coolant can be re-established then it may be possible to quench the debris and maintain a coolable state.

Debris Coolability Models

1 28

The relevant physics here therefore concern the mechanisms of top reflood. If the amount of superheated debris is large it is possible that the debris will still attack the concrete while remain cooled from above. However the gases evolved as a consequence of the core concrete interaction will tend to prevent the ingress of water into the debris. The quench process could be so seriously affected that the lower bed regions could be molten when the quench front finally arrives.

11.2

Phenomena

The purpose of this chapter is to review the latest understanding of quenching of overheating debris beds. The essential physical processes have been studied mainly through experiments with simulant materials for both top and bottom reflood conditions. A smaller number of studies have been carried out using more prototypical materials. The main phenomena of interest include the heat removal processes of the quenching material and the rate of coolant penetration into the debris bed. The models developed for these processes will be described below. The important processes of top and bottom reflood are different, Figure 1 1 . 1 . The thennal-hydraulics of bottom reflood is simpler than top reflood. Essentially as water flows into the bottom of the bed steam is expelled from the top and the flow is one­ dimensional. Limiting factors are likely to be the rate of the conductive heat flow out of the larger particles, heat transfer from these particles to the coolant and the resistance to flow through the bed. Top-reflood involves counter-current flow conditions. Water flowing into the bed from above will be impeded by steam flowing up out of the bed and there may also be additional gases arising from an underlying core concrete reaction.

1 1.3

Experimental Programmes

Experiments by P Hall and C Hall [ 1 1 . 1] were carried out at Berkeley to investigate the problems of hot debris quench by bottom reflooding. The experiments used spherical iron shot particles with mean particle diameters of O.3mm, 1 .3mm and 2.0mm in a 1 m high vertical steel tube o f 35mm internal diameter. The main conclusions were that larger driving heads gave rise to faster quench but for low driving heads an equilibrium could be reached whereby the steam pressure drop in the unquenched regions is sufficient to halt the inlet flow and halt the quench front This phenomena known as 'steam binding' is an unstable phenomena however and can be destroyed by small increases in head. Tutu and Ginsburg also investigated the quenching of superheated debris beds by bottom-reflooding [ 1 1 .2] , [ 1 1 .3], [ 1 1 .4] at Brookhaven National Laboratories, USA. These experiments involved 3mm stainless steel particles. The results showed that quenching was essentially one-dimensional for low temperature debris and low flow rates

Debris Coolability Models

1 29

and the bed heat flux was fairly constant: for hotter beds and larger liquid supply rates they found that the quenching process is complex and multi-dimensional and the bed heat flux is not constant with time.

VAP O U R --

T O P O F B ED

--

VAP O U R

--

BOILING

--

LI Q U I D H EATI N G

CO-C U RR E NT F LOW

-- B OnO M O F B E D

LI Q U I D

(a) BOTTO M R E F LO O D

WAT E R

� f � t

VAPO U R

CO U NTE R - CURR ENT F LOW

-- COOLED D E B R I S

-- N OT COO L E D D E B R I S

(b) TO P RE F LOO D

FIGURE 1 1 . 1 DEBRIS BED COOLING A prototypical programme of experiments has been carried out at the University of California, LA, USA. Tests were carried out with stratified beds, gas injection at the bottom of the bed and also including zirconium oxidation. The test programmes included simple bottom reflood experiments [ 1 1 .5]. fluidisation experiments. radial and vertical statification experiments [ 1 1 .6] and multi-dimensional quenching of a simulated core debris bed.

1 30

Debris Coolability Models

The later experiments were intended to be the most propotypical of all. The main conclusions were that: increased pressure difference across the bed accelerates quenching but also leads to multidimensional effects (c.f. the BNL bottom reflood experiments); The Zircaloy-steam reaction raises local temperatures which slow down the quench fro nt; Non-uniformities in the initial temperature profile can produce large changes in the quenching pattern and hence the hydrogen production rate. In the area of top reflood experiments, Cho and Armstrong [ 1 1 .7] , [ 1 1 .8] , [ 1 1 .9] provided data on the coolability margins of a degraded L WR and on the steam production rate: the concern in the latter case being the threat to the containment from over­ pressurisation. These tests involved both stainless steel and aluminium balls, about 3mm in diameter. Tung and Dhir from UCLA also carried out experiments for top reflood. They investigated the effects of internal heating and gas injection [ 1 1 . 1 0] on the quenching of super-heated core debris with stainless steel particles of small diameters in the range 0.6mm to 1 .6mm. They also considered axial and radial stratification [ 1 1 .6] . Ginsberg et al [ 1 1 . 1 2] , [ 1 1 . 13] , [ 1 1 . 14] investigated the phenomenology of steam spikes in connection with transient core debris heat removal in LWR In core experiment hot packed beds of stainless steel particles about 3 m m diameter were quenched by overlying poo ls of water. .

Ginsberg et al also carried out a thorough experimental and analytical study of quenching of superheated debris beds by top reflooding. These results showed that the bed heat flux is strongly dependent on particle diameter but only weakly dependent on particle initial temperature. Barleon, Thomanske and Werle [ 1 1 . 15], [1 1 . 16] and later Barleon and Hofmann [ 1 1 . 1 7] looked at the coolability and dryout limits for volumetically heated beds. These experiments were performed at Kernforschungzentrum Karlsruhe in Germany. Particle diameters in the range from 0. 1 5mm up to O.3mm were considered. A series of experiments (DCC) [ 1 1 . 1 8] , [ 1 1 . 19], [ 1 1 .20], [ 1 1 .2 1 ] were carried out in deep U02 fission heated beds in a pressurised water bath in the Annular Core Research Reactor (ACRR) at Sandia in the USA. These tests involved particles of diameters in the range of 0.9mm6.4mm. Figure 1 1 .2 shows a comparison of experimental and predicted temperatures of the particles [ 1 1 .20] .

Debris Coolability Models 11.4

131

Bottom Reflood Models

Models of varying sophistication and generality have been developed. The simplest approach has been to develop lumped parameter or zero-dimensional models which only treat the global processes in the debris bed. The dominant processes i.e. the drag between the vapour and the unquenched bed are modelled by correlations but these models do not treat explicitly the processes at the single particle scale and therefore do not make use of heat transfer correlations. One-dimensional models have been derived which utilise the conservation equations for mass, momentum and energy for the flow in a porous medium.

11.4.1 Lumped Parameter Models P Hall and C Hall [1 1 . 1 ] developed a simple model for analysis of their experimental data. This work was focused at understanding the phenomena of hot debris quench and was largely motivated by the need to understand questions raised on debris coolability in the Three Mile Island accident. The basic premise of the Hall and Hall model was to assume that the two phase region was small compared with the heated region. Simple lumped mass, energy and momentum balance equations can then be derived: closure of the equation set was achieved by using standard correlations for the frictional pressure drop for flows through packed beds. This model reproduces the experimental data reasonably well including the steam binding effect. It would however be less applicable to shallow beds where the two phase region will not be short in comparison with the bed height

11.4.2 Quenching Models A simple quenching model was fonnulated by Tung and Dhir [ 1 1 .5] primarily for the purpose of analysing experimental data. The model made a number of assumptions which are not generally true and therefore its scope is limited. In particular it was assumed that the quenching process was one-dimensional, which limits its use to low flow rates. The model also assumed that the stored energy in the hot particles is released instantaneously and is entirely used in vapourising the liquid coolant. This will only apply to small particles. Other assumptions in the model are that the height of the two-phaselheat transfer region can be prescribed: also that thermal conduction in the bed can be neglected. Given these assumptions, an equation set to describe the quenching process (including the speed of the quench front) can then be derived: when the model was applied to the experimental data it was found that calculations including the steam cooling of the hot unquenched region gave good argument with the data over the full range of experimental conditions. Tung and Dhir also applied their model to experiments with stratified beds [1 1 .6] but in addition to the assumptions above also assumed that the two-phase region was of negligable thickness. This assumption is valid in the limit of flow rates.

Debris Coolability Models

1 32

1 1 .4.3 Models for Fluidisation The fluidisation of debris beds during bottom quenching at large coolant flow rates has also been investigated by Tung and Dhir [ 1 1 .22] . The pressure pulse produced by the quenching was compared with a simple criterion for the onset of fluidisation in the upper bed. The model predicted that the pressure pulse had a weaker dependence on the driving pressure than that observed. This was attributed to the use of a constant heat transfer coefficient, which therefore did not reflect the time heat transfer dependence on quality system pressure and coolant flow rate.

11.4.4 More Sophisticated Models Tutu et al developed a range of models for bottom-reflood [ 1 1 .3] , [ 1 1 04] . They considered the unsteady flow of a liquid/vapour mixture through a uniform cross­ sectional area debris bed. The model was based on a conservation of mass and energy momentum equation. The momentum equation was not based on D' Arcy' s Law or its variants. The energy conservation was based on the assumption that the liquid and vapour phases interchange energy with the solid directly but only indirectly with each other. Also it was assumed that vapour was produced if the temperature of the liquid was greater than the saturation temperature but otherwise evaporation did not occur. One of the problems of the model is that there is little quantative information on heat transfer and drag in two phase mixtures which is needed for closure of the equation system. A simplification of the above model was used in order to analyse experimental data. This was still a transient model but made simplifying assumptions such as the liquid entering the bed is at saturation, solid-vapour heat transfer and heat conduction through the solid can be neglected, the absolute liquid velocity is constant and vapour is produced at a temperature which is the mean of the solid and saturation temperature. The authors also developed a quasi-steady version of the model. This could be used for deep debris beds where the two-phase region height is small compared with the height of the bed and inlet velocities are small. This model assumed the existence of a quench front which divided the bed into essentially three regions, dry, two-phase and wet; the front travelled through the bed at uniform speed. Expressions for the quench front and heat flux can then be derived, again without the need for details of the heat transfer mechanisms. This particular simplification was recommended by Mueller and Sozer [ 1 1 .23].

1 1 .5

Top Reflood Models

The work on the development of top reflood models has taken account of the well established steady-state dryout modelling work. The physical phenomena at the quench point in top reflood are different from those encountered in dryout but it is possible to apply similar treatments.

Debris Coolability Models

•

0. 3

1 20

133

•

o. 1

t

•

435

•

o. 0 200

400

600

800

TEMPERATURE ·C

1 000

(a) PR E D I CTI O N

E

U

30

�

G -

"'"' %:

20 10 o

200

400

600

800

1 000

TEMPERATURE · C

(b) EXP E RI M ENT

FIGURE 1 1.2 DCC-2 EXPERIMENT: TEMPERATURE PROFILES

1 34

Debris Coolability Models

An important consideration in developing models for quench is to establish the spatial location of the physical process that limits the quenching rate - various locations have been assumed by the modellers: above the bed, at the top of the bed and at the quench front. Traditional dryout models derive their limits to coolability based on processes above or at the top of the bed. However, from analyses of data involving small quench rates in small particle beds it has proved necessary to consider the physical processes in the quench front region. Turland et al have investigated film boiling at the quench front. Barleon et al has examined the effects of using different capilliary pressure functions in the dry and quench regions.

1 1.5.1 Critical Heat Flux Models In these models the flow of coolant into the bed is limited by Rayleigh-Taylor instability in the vapour flow out of the bed surface. Those models where the processes above the bed surface control the influx of coolant are referred to as Critical Heat Flux models [ 1 1 .24]. It is commonly assumed that the bed surface can be modelled as a flat plate. Zuber [ 1 1 .25] has produced a correlation for pool boiling to bound the maximum vapour flux out of the bed. This model has fairly limited application: it is claimed that it is only valid for large diameter particles with the coolant/particle drag small. Even in this case however experiments have shown that the Zuber correlation for flat plate critical heat flux may underpredict the dryout heat flux [ 1 1 .24].

1 1.5.2 Upper Bed Quenching Rate Limitation Models Several models have been developed assuming that the limiting processes occur at the top of the bed. The Ostensen-Lipinski model assumes a flooding limit correlation to determine when dryout occurs. This model [ 1 1 .26] is based on the Wallis [ 1 1 .27] correlation developed from earlier work of Sherwood and Lobo [ 1 1 .28] . The model has been applied by Ginsberg et al to the quenching of hot stainless steel beds [ 1 1 . 1 3 ]. A variant of this model has been recommended by Marshall and Dhir [ 1 1 .29] . Several versions of the Lipinski model have been proposed. These models make no assumptions about the limiting mechanism. They allow the bed resistance and the competition between the liquid and the vapour flows to set their own limit. The Osten sen-Lipinski and the Lipinski models therefore differ in their treatment of coun ter-current vapour flow. The Quasi-Steady Lipinski Model for steady state debris bed dry-out [ 1 1 .30] is a separated flow model for two-phase flow in packed beds. It is based on mass, momentum and energy conservation equations for both liquid and vapour phases in a porous medium. The momentum equation is approximated by D' Arcy' s law, or the Ergun extension to it for large particles and turbulent flow. Dryout occurs by a process analogous to flooding. This is despite the fact that the standard Lipinski model does not model interphase drag explicitly.

Debris Coolability Models

1 35

The competition between liquid and vapour flow at the top of the bed is adequately represented using only simple functions for the relative penneabilities of the two phases. The Lipinski model perfonns well with prediction of quench heat fluxes for debris beds for particles with diameters greater than about 0.5mm [7] . The model tends to overpredict the quench heat flux for smaller particles. The Transient Lipinski Models [ 1 1.3 1 ] , [ 1 1 .32] , [1 1 . 33] contain transient equations for mass, momentum and energy conservation. Momentum conservation is based on D' Arcy 's law, or its Ergun extension, for flows in a porous medium. Models have been developed by Brearley and Ruel. The Brearley model takes the one-dimensional version of the transient equations and solves them using standard finite difference techniques. The debris and coolant are assumed to be in thermal equilibrium . Heat generated in the bed is deposited instantaneously into the coolant. The model has been used to simulate the DCC experiments. On quench the model produced quench times much shorter than those observed in the experiment by as much as a factor 2-3 times shorter. This has not been explained. The experiment exhibits two dimensional effects, such as separation of liquid and vapour flows, but these would be expected to increase the quench rate.

11.5.3 Qu en c h Front Limitation Models In order to improve predictions for certain experiments, in particular, the quenching in the DCC experiments [ 1 1 .20] and the KfK small particle experiments [ 1 1 . 16] , attention has turned to the processes occurring at the quench front. Barleon, Thomauske and Werle [ 1 1 . 1 6] developed a model, which introduces imbibition and drainage capillary pressure functions for the dry and wet regions of the bed. This reproduced the behaviour observed in the experiment by introducing new physical mechanisms into the models in the region of the front. Brealey suggested that the limiting factor may be boiling heat transfer close to the front, due to the large thennal disequilibrium between the bed and coolant in this region. Barleon, Thomauske and Werle [ 1 1 . 1 6] used a modified version of the DEBRIS code to model their experiments. The DEBRIS code is based on the standard Lipinski model and as such is a steady state code. Rewetting is predicted to occur at the same power as dryout. This was not consistent with experimental results and this provided the impetus for a new rewetting model. The main assumption in the new model was that the difference in dryout and rewetting heat fluxes was caused by different capillary pressures for imbibition and drainage and so the standard Leverett drainage function in DEBRIS was replaced by an imbibition Leverett function in the dry region. In the wet region the nonnal drainage function was retained. For a given saturation the imbibition Leverett function differs from the

Debris Coolability Models

1 36

drainage function by a factor of about 0.7. The use of two capillary pressure functions in this way has the same effect as a stratification boundary and forces a sharp change in saturation at the front. The results of this model show good qualitative agreement with the experiments. The quench heat flux is much lower than that predicted on the basis of drainage capillary pressure alone. The dry zone in the transient rewetting experiments persists without change of size after the bed power is reduced below the dryout power in the transient rewetting tests.

A Thermal Disequilibrium Model has been produced as an alternative to the use of two capillary pressure functions. This model assumes that the limiting process is due to thermal disequilibrium and its effect on heat transfer between debris and coolant To take account of this, Brearley postulated a boiling limit to heat transfer at the front which would determine the quench rate. Turland devised a model based on this premise. Turland started from the basic equations for mass, momentum and energy conservation in a porous medium and derived equations for the quasi-steady one-dimensional quenching of superheated core debris. The heat transfer from debris to liquid coolant was described using a modified version of the flat plate pool boiling curve. The solution of these model equations turned out to be complicated. However also, the results give a maximum quenching rate much lower than that predicted on the basis of flooding arguments.

11.6

Status of Modelling

11.6.1 Bottom Reflood In modelling the effects of bottom reflood, it has been necessary to start with the basic mass momentum and energy equations with closure relations for frictional drag and steam cooling in the hot unquenched region. It has not been possible to utilise classical steady-state dryout models. The simpler models are restricted to debris beds with small particles and low flow rates. In these circumstances it may be assumed that the heat from the particles can be transferred to the coolant simultaneously. None of the simplified models adequately predict the extent of the two phase region and in order to make satisfactory predictions either the thickness must be prescribed, or it must be assumed to be zero and the use of the model restricted to physical situations where this is the case. The more sophisticated models generally depend on a correlation for the heat transfer for the transfer of heat from the particles to the coolant. These cannot be measured easily and these models therefore need to be used only for scoping studies.

Debris Coolability Models

137

11.6.2 Top Reflood For top reflood the steady state Lipinski model can be used for particle sizes down to about I mm. For beds with very small particles these models are less adequate: quench rates may be under-estimated by factors of 2 to 3. New models are being considered to take account of the physical phenomena at the quench front. As for the bottom reflood case though. these models also rely on empirical relations e.g. the two phase region heat transfer coefficient which is difficult to obtain. At the present time these models will also need to be used only for seoping studies.

REFE R EN C ES 1 1 .1

P C Hall. C M Hall. Quenching o f Heated Particulate B ed s b y Bottom Flooding. Preliminary Results and Analysis. presented at the European Two-Phase Flow Group Meeting. Eindhoven. Holland ( 1 981). 1 1 .2 N K Tutu. T Ginsberg. J Klages. J Klein. C E Schwartz. Debris Bed Quenching Under B o ttom Flood Conditions (In- Vessel Degraded Core Cool ing Phenomenology). Brookhaven National Laboratory BNL-NUREG-5 1 7 88 . NUREG/CR-3850 (July 1984). 1 1 .3 N K Tutu. T Ginsberg. J Klein. C E Schwartz and J Klages. Transient Quenching of Superheated Core Debris During Bottom Reflood. 1 3- 1 . Proceedings of the Sixth International Exchange Meeting on Debris Coolability. UCLA. Los Angeles. CA (November 1 984). 1 1 .4 N K Tutu. T Ginsberg. Debris Bed Quench Characteristics Under Bottom-Flood Conditions. Topical Meeting on Thermal Reactor Safety. ANSIENS . San Diego. CA (2-6 February 1986). 1 1 .5 V X Tung. V K Dhir. Quenching of a Hot Particulate Bed by Bottom Flooding. 1 4 - 1 . Proceedings ASME-J S ME Thermal Engineering J oint Conference Honolulu. Hawaii. ( 1 983). 1 1 .6 V X Tung. V K Dhir. Quenching of Debris Beds Having Variable Permeability in the Axial and Radial Directions. Nucl. Eng. and Des. 99. 275-284 ( 1 987). 1 1 .7 D H Cho. D R Armstrong. L Bova, S H Chan. G R Thomas. Experiments on Quenching a Hot Debris Bed. Post Accident Debris Cooling. Proceedings of the Fifth Post Accident Heat Removal Information Exchange Meeting. Nuclear Research Centre. Karlsruhe (July 28-30 1982). 1 1 .8 D H Cho. D R Armstrong. L Bova. S H Chan. G R Thomas. Debris Bed Quenching Studies. Proceedings of the International Meeting on Thermal Nuclear Reactor Safety. Chicago IL (August 29-September 2 1 982). 1 1 .9 D H Cho. D R Armstrong. S H Chan. On the Pattern of Water Penetration into Hot Particle Beds. Nuclear Technology 65.23-3 1 ( 1 984). 1 1 . 1 0 V X Tung. V K Dhir. D Squarer. Quenching by Top Flooding of a Heat Generating Particulate Bed with Gas Injection at the Bottom. 12- 1 . Proceedings of the Sixth International Exchange Meeting on Debris Coolability. UCLA. Los Angeles. CA (November 1 984).

138

Debris Coolability Models

1 1 . 1 1 T Ginsberg, J Klein, J Klages, C E Schwartz, J C Chan. LWR Steam Spike Phenomenology: Debris Bed Quenching Experiments, Brookhaven National Laboratory BNL-NUREG-5 1 57 1 , NUREG/CR-2857 ( 1 982). 1 1 . 1 2 T Ginsberg, J Klein, J Klages, C E Schwartz, J C Chan. Phenomenology of Transient Debris Bed Heat Removal. Post Accident Debris Cooling, Proceedings of the Fifth Post Accident Heat Removal Infonnation Exchange Meeting, Nuclear Research Centre, Karlsruhe (July 28-30 1 982). 1 1 . 1 3 T Ginsberg, N K Tutu, J Klages, J Klein, C E Schwartz, Y Sanborn. Transient Core Debris Heat Removal Experiments and Analysis. Proceedings of the International Meeting on Thennal Nuclear Reactor Safety, Chicago IL (August 29-September 2 1982). 1 1 . 14 T Ginsberg, N K Tutu, J Klages, J Klein, C E Schwartz, Y Sanborn. Core Debris Quenching Heat Transfer Rates Under Top- and Bottom-Reflood Conditions, Paper TS 1 8.7. Proceedings of the International Meeting on Light Water Reactor Severe Accident Evaluation, Cambridge, MA (August 28-September 1 1 983). 1 1 . 1 5 L Barleon, K Thomauske, H Werle. Extended Dryout and Rewetting of Small­ Particle Debris, 1 7- 1 . Proceedings of the Sixth International Exchange Meeting on Debris Coolability, UCLA, Los Angeles, CA (November 1984). 1 1 . 1 6 L Barleon, K Thomauske, H Werle. Extended Dryout and Rewetting of Small­ Particle Core Debris. Nucl. Eng. and Des. 102, 59-69 ( 1 987). 1 1 . 1 7 G Hofmann, L Bar1eon. Reduced Coolability of Particle Beds as a Result of Capillary Effects at Horizontal Phase Boundaries. Topical Meeting on Thennal Reactor Safety, ANSIENS, San Diego, CA (2-6 February 1 986). 1 1 . 1 8 A W Reed, E D Bergeron, K R Boldt and T R Schmidt Coolability of UOz Debris Beds in Pressurised Water Pools: DCC I and DCC2 Experiment Results . Proceedings o f th e Sixth International Exchange Meeting o n Debris Coolability, UCLA, Los Angeles, CA (1 984). 1 1 . 1 9 A W Reed, K R Boldt, T R Schmidt. Coolability of LWR Debris: a Summary of the DCC Experiments, Paper VI.5. Proceedings of the International ANS/ENS Topical Meeting on Thennal Reactor Safety, San Diego ( 1 986). 1 1 .20 A W Reed, K R Boldt, E D Gorham-Bergeron, R J Lipinski, T R Schmidt. DCC­ I /DCC-2 Degraded Core Coolability Analysis, Sandia National Laboratories Report SAND85- 1967, NUREG/CR-4390 ( 1 985). 1 1 .21 K R Boldt, A W Reed, T R Schmidt, DCC-3 Degraded Core Coolability: Experiment and Analysis. Sandia National Laboratories Report SAND86- 1033, NUREG/CR-4606 (1986). 1 1 .22 V X Tung, V K Dhir. On Fluidization of a Particulate Bed During Quench by Flooding from Bottom, 14- 1 , Proceedings of the Sixth International Exchange Meeting on Debris Coolability, UCLA, Los Angeles, CA, (November 1984). 1 1 .23 G E Mueller and A Sozer. Thennal-Hydraulic and Characteristic Models for Packed Beds, Oak Ridge National Laboratory ORNL{fM- I 0 1 1 7 , NUREG/CR4689. 1 1 .24 D Squarer, A T Pieczyski, L Bova. Effect of Debris Bed Pressure, Particle Size and Distribution on Degraded Reactor Core Coolability, Nucl Sci and Eng 80.2- 1 3 ( 1 982). 1 1 .25 N Zuber. On the Stability of Boiling Heat Transfer, Trans. ASME, 80, 7 1 1 -720 ( 1 958).

Debris Coolability Models

139

1 1 .26 R W Ostensen, R J Lipinski, A Particle Bed Dryout Model Based on Flooding, Nucl. Sci and Eng. 79, 1 1 0- 1 1 3 (1981). 1 1 .27 G B Wallis, One-Dimensional Two-Phase Flow, McGraw-Hill, New York, 1 969. 1 1 .28 T K Sherwood, G H Shipley, F A L Hollaway, Ind. Eng. Chern. 30, 765 (1938). W E Lobo, L Friend. F H Ashmall. F Zenz. Trans A I Ch. 693-7 1 0 ( 1938). 1 1 .29 J Marshall. V K Dhir. On the Countercurrent Flow Limitations in Porous Media. Paper TS I 8.5, Proceedings of the International Meeting on Light Water Reactor Severe Accident Evaluation. Cambridge. MA (August 28-September 1 . 1983 . 1 1 .30 R J Lipinski. A Model for Boiling and Oryout in Particle Beds. Sandia Laboratories Report SAND 82-0765 [NUREG/CR-2646] ( 1982). 1 1 .3 1 T C Chawla. D R Pedersen. W J Minkowycz. Governing Equations for Heat and Mass Transfer in Heat-Generating Porous Beds - 1. Coolant Boiling and Transient Void Propagation, Int. J. Heat and Mass Transfer 28. 2 1 29-2 1 36 (198 1). 1 1 .32 E Gorham-Bergeron. A One-Dimensional Time-Dependent Debris Bed Model. Proc. ASME/JSME Thermal Engineering Joint Conference. Honolulu. Volume 2. p23 ( 1 983). 1 1 .33 E Gorham-Bergeron. An Analytical Model for Predicting Dryout and Quench Behaviour in a Volumetrically Heated Bed. Paper TS 1 5 .4 . Proceedings of the International Meeting on Light Water Reactor Severe Accident Evaluation. Cambridge MA (August 28-September 1. 1983). 1 1 .34 C H Wang. V K Ohir. An Experimental Investigation of Multidimensional Quenching of a Simulated Core Debris Bed. Nucl. Eng. and Des. 1 10. 6 1 -72 (1 988). 1 1 .35 V K Ohir. On the Coolability of Degraded LWR Cores, Nuclear Safety 24, 3 19337 (May-June. 1 983).

14 1

Chapter 12 DEB RIS INTERACTIONS WITHIN THE VESSEL

12.1

Introduction

This section addresses the mechanisms that occur during the later phases of in-vessel melt relocation including interactions with the vessel itself, and the release of debris from the vessel. Interaction of molten debris with water is likely to be of key importance concerning the progression of the accident. Steam explosions are treated as a separate topic in Chapter 10. The quenchability of debris has been treated in Chapter 1 1 . The core is supported by various steel structures, while the vessel itself presents a continuous steel barrier. For most sequences the lower head region will contain water and if debris falls into it, then steam will be generated and also possibly hydrogen if the debris contains unoxidised metallic components. These place an immediate additional load on the containment in the case of an already breached primary circuit or at vessel failure in the press urised sequences. The interactions in the lower head can have a major impact on later events in the accident sequence. The containment loading will depend on: (i)

the quantity of debris released at vessel failure

(ii)

its temperature

(iii)

its chemical composition; in particular its metallic content

(iv)

the mode of vessel failure e.g. if there is a localised failure or complete lower head failure, and

(v)

the timing of vessel failure after debris release to the lower head.

The short term threat to the containment is likely to arise immediately after vessel failure. The energy stored in the hot debris may be released to the atmosphere by steam explosions in the cavity, or by direct heating of the atmosphere. The latter may be increased by oxidation of metallic fragments. The quantity of hydrogen produced either through quench of the debris or from core concrete interactions will determine whether flammability conditions are reached in the containment The potential for fission product release or retention will be dependent of the distribution of the hot particles after vessel failure. S team or hydrogen flow through the breach at vessel failure will affect aerosol production which in tum impacts directly on fission product release and transport.

1 42

12.2

Debris Interactions within the Vessel Release of Material from the Core Region

The processes of melt progression have been considered in the earlier Chapter 9. There is little doubt that the thinking behind model development has been dominated by events in TMI-2 in which it appears that the majority of the melt relocation took place through the bypass region. Generally it is thought that melt can get to the lower head by various routes: it is not known whether flow through the bypass is likely to be a generic phenomenon. The computer codes for modelling these processes mechanistically are not validated. They also tend to suffer from the inherent weakness of coarse nodalisation which essentially implies a large pour, once the criterion for a pour has been satisfied.

12.3

Debris Interactions with the Lower Vessel Internals

Morgan and Turland [ 1 2. 1 ] reported calculations of two aspects concerning core slump in a PWR:

(1)

flow through the lower fuel nozzles

(2)

the potential for blocking or ablating the lower core plate.

Melts of control rod material, metallic Zircaloy and ceramic Zircaloy/U02 eutectic were considered and a simplified conduction freezing model was used for flow freezing in the lower nozzle. The indications are that the ceramic material would tend to plug the nozzles on a larger timescale than the metallic: it is not expected that the ceramic melt would ablate the nozzle plates significantly. Calculations have been performed exploring several extreme cases of melt flowing through holes in the lower core plate. The extent of plate ablation was found to be largely dependent on whether any crusts formed were stable or not: unstable crusts give greater ablation. The likely behaviour is expected to be closer to that of a stable crust, there the most likely situation is that the lower core plate will remain largely intact and that only small amounts of steel will be ablated. Kuan [ 1 2.2] has provided some supporting analysis to this thesis. Underneath the lower core plate lies the thick lower core support plate. This structure contains numerous large hole penetrations, Figure 1 2. 1 , and in principle the debris could interact with this structure in a similar way as described for the lower core plate. However (i)

the large holes militate against complete blockage of the plate by freezing debris, and

(ii)

the possibility of an interaction with water lying above the lower core support plate must be considered.

Some heat-up and ablation of the lower core support plate is therefore possible after the debris particulate remelts or due to melt jet attack.

Debris Interactions within the Vessel

143

However unless debris accumulates i n a vulnerable region e.g. a t the join with the core barrel, relatively little damage is expected from the initial pour. Failure at a later time could be possible due to radiation shine from any molten material in the lower head, but if the vessel fails quickly the in-vessel structure is likely to remain relatively intact.

-

0

0

0

0

0

0

0

(f

0

0

0

0

0

0

0

0

/

S MALL H O L ES

TH I N PLATE

(a) LOWE R CO R E PLATE

o 00

LAR G E H O L E S

TH I C K PLATE

(b) LOW E R CO R E SU PPORT P LATE

FIGURE 12.1 LOWER VESSEL INTERNAL PLATES

144 12.4

Debris Interactions within the Vessel Debris Interactions with Water

In a large modem PWR the region below the core could be filled with sufficient water to quench completely up to 50% of the core or provide significant cooling to a larger fraction of the whole core. The interaction with water has been investigated in the context of the likelihood of possible steam explosions. Studies carried out using the CHYMES Code [12.3] in which the melt is assumed to fall downwards in small jets have indicated initially rapid steam generation followed by expulsion of water up the downcomer or up into the core region (if unblocked), allowing the melt to collect in a pool on the lower head. Water could then cover this debris resulting in a stratified configuration. It is believed by some researchers (e.g. Bankoff [12.4]) that al low pressure the triggering of small steam explosions is inevitable. 1be consequences for the vessel and containment may be relatively benign although the course of the accident progression may be altered e.g. by quenching the melt or by encouraging fwther collapse of the core and lower structures. At higher pressures it is generally accepted that triggering is less likely to occur but if it does occur the consequences are likely to be more severe. Various scenarios have been proposed concerning the way the melt enters the lower head. These are: (a)

a coherent pour of melt depositing itself on the lower head

(b)

a melt layer under a water pool

(c)

a partially quenched melt with possible additions of core and structural debris.

There is considerable uncertainty on the geometrical configuration: in particular on the area of the interface between the debris and water. The area limits the steam generation associated with film boiling or an efficient mixing process leading to a steam explosion. The amount of quenching of the debris that can occur is dependent on whether the mixing process is self sustaining. If the mixture collapses in a short time relatively little quenching may have occurred, after that only a small amount of heat is released by film boiling. During the initial interactions with water there is interest in the amount of hydrogen that might be produced. Morgan and Turland [ 12. 1 ] have presented results obtained with the MARCH model [ 12.5] , a single drop model that assumes water availability. The oxidation of both Zircaloy and stainless steel produce hydrogen and the flow of steam/ hydrogen could have a significant effect on the material still left in the core region and also on the fission product release. A review of hydrogen production during melt - water interaction has been carried out recently. This noted that the limiting processes of oxygen diffusion though the oxide

Debris Interactions within the Vessel

145

layer and hydrogen diffusion through the vapour layer surrounding the melt [ 1 2.6] have to be treated. Ceramic melts are not expected to significantly ablate steel structures due to the formation of protective crusts. Overall the initial reactions in the lower head are uncertain. Quench of debris may occur at low pressures but the extent of oxidation during the debris/water interaction is uncertain.

12.5

Debris Interactions with the Vessel

The debris that arrives on the lower head may be a mixture of a number of different materials. These include: (i)

low melting point, high conductivity materials e.g. Ag-I in the case of a PWR

(ii)

intermediate metallic materials involving Zircaloy/stainless steel eutectics, and

(iii)

high melting point, low conductivity ceramic materials.

The extent to which these material separate into different phases is largely unresolved. Low melting point materials at the base of the vessel could result in an instrumentation tube failure. Calculations for lower head failure have been carried out by Morgan using the MELTPV2 code in which it was assumed that the metallic layer overlaid an oxidic layer. In the majority of cases there was no ablation of the vessel by the oxidic layer. This model is indicative of what might occur if the materials separate out but this assumption may not be justified. With this model the calculations indicate rapid local w hole head failures predicted in times of less than a few minutes. For the pressurised sequences instrument penetration due to heat transfer from the oxide layer is thought to occur on a similar timescale. If the residence time of the molten debris in the lower head is sufficiently long then steel would be incorporated into the debris either by melting, the mechanism being heat conduction between the debris and steel or by a radiation shine from the debris to a lower core structure. Calculations suggest that the barrel is likely to survive beyond core slump (e.g. in some designs there are baffle plates between the core and the barrel). Significant debris on the lower core support plate seem s unlikely so the lower core support plate is not expected to be threatened. The extent of attack on the lower head is dependent on whether the melt segregates into metallic and oxidic layers. However it is known that at high temperatures Zircaloy can dissolve oxide material and therefore the formation of a single metallic layer is not thought to be likely.

Debris Interactions within the Vessel

146

12.6

Vessel Failure Potential

Lower head failure is anticipated due to the pressure load imposed on the lower head and weakening of the steel as its temperature increases. This load is scenario dependent. For sequences in which the primary circuit is breached e.g. if the initiating event were a LOCA, the pressure load is only due to the weight of the debris and is therefore relatively modest. For sequences in which the primary circuit is not breached the pressure (in the case of a PWR) could be as high as l7MPa. Calculations by BoyIe and Lillington [ 1 2.7] and others have shown that in the case of a PWR, natural circulation heating may cause the primary circuit to breach before the lower head is threatened. Clearly there are a whole range of scenarios with the primary circuit at intennediate pressure where the degree of threat is much less clear. The pressure could also be affected by accident management strategies to depressurise the circuit at an earlier stage.

12.7

Vessel Response at Elevated Temperatures

The lower head and the cylindrical parts of the vessel will undergo membrane stresses due to the pressure loading. Estimates for the temperature gradient across the bottom of the vessel have been given to be at least 350K. The implication is that when the inside of the vessel is at high temperatures the outer part of the vessel will be much cooler. The stresses will only be unifonnly distributed if the temperature is unifonn and therefore the stresses in the outer half will be enhanced. Various mechanisms for failure of the upper head have been considered. These include: (a)

plastic failure

(b)

creep failure.

Plastic failure could occur as a consequence of the ultimate tensile strength of the material being reached. Creep failure would be a more likely failure mode at higher temperatures. The potential for plastic failure can be estimated relatively easily, provided the assumption is made that geometry does not change significantly. In these circumstances the temperature dependent ultimate tensile strength may be integrated across the vessel and can then be compared with membrane forces arising from the pressure difference. The data base for the ultimate tensile strength of steels is relatively sparse near the melting point of steel. This is because steels cannot be used as structural material at temperatures near their melting point because of their creep characteristics. Idaho National Engineering Laboratory in the USA have provided measurements for the ultimate tensile stress for A-508 carbon steel at temperatures up to about lOOK. These

Debris Interactions within the Vessel

147

data have been extrapolated to higher temperatures when failure is more likely to occur. The MARCH code [1 2.5] for example uses an extrapolation in which the ultimate tensile stress falls as the fourth power of temperature. These high temperature extrapolations are however suspect since they do not take account of phase transition that occurs at about 1 050K. Creep is strongly dependent on temperature via an Arrhenius relationship and the applied stress. This mode of failure can occur at stress levels lower than the ultimate tensile stress. Time to failure data for A-50S steel is reported in [ 12.S] but again there are limited data at higher temperatures. The data in [ 1 2.S] have been correlated in two different models. Larson-Miller and the Dom model. The latter fits the data well for stainless steel. The data were obtained and correlated at relatively high stress levels and at temperatures below the phase transition. At lower stress the Larson-Miller and Dom correlations give different behaviour. The Larson-Miller correlation implies failure at shorter times and at lower temperatures. For high pressure sequences there is a reasonable data-base at least for one pressure vessel steel. Failure would be expected to occur at about l l00K by plastic failure or creep. There are no reliable data for failure at lower pressure: the failure criteria available are based on unvalidated extrapolations.

12.8

Behaviour of Penetrations

For PWRs. the possible failure of penetrations through the lower head have also been considered. These penetrations carry in-vessel instrumentation and are potential failure sites. Several failure mechanisms have been considered: (i)

Local melting of the penetrations.

(ii)

Tube ejection following failure of weld(s).

Models for penetration have been developed following examination of the TMI-2 lower head, Figure 1 2.2. Cronenberg et al [ 1 2.9] attempted to assess the extent of melt penetration through the TMI-2 vessel penetrations by applying a bulk freezing model. As stated in Chapter 9 these would provide a lower estimate of penetration because no insulating crust is assumed on the steel wall. Cronenberg assessed the degree of penetration assuming various melt combinations of molten control rod material. Inconel and uranium dioxide. Not surprisingly deepest penetrations were found with the lowest melting point material i.e. control rod material. but all penetration distances were under-estimated compared with measurements made with wire probes. Order of magnitude estimates for the freezing times have been obtained from the conduction freezing model. by considering the timescale for a crust growth to block the penetration tube.

Debris Interactions within the Vessel

148

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:L:� /, v

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m

ill /1"/ /(/" I;;;' �

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Imm

EB

EB

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EB

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EB EB EB

I N C O R E I N STRU MENT i.. OCATION , I I G T R EMOVED WITH E F D P LATE

I N C O R E INSTRU M E NT LOCATION , I IGT NOT

REMOVED

V I D E O I N S PECTION SHOWS MAJOR DAMAGE TO IIGT

V I D E O I N S P ECTION SHOWS MINOR DAMAGE TO I I G T

V I D E O I N S P E CTION SHOWS NO DAMAGE TO I IGT

REDUCTION IN TH ERMOCOUPlE

erC) LENGTH TO B E LOW LOW ER GRID

FIGURE 12.2 TMI-2 LOWER PLENUM AND INDICA nONS OF GUIDE TUBE DAMAGE The degree of penetration into guide tubes will depend on the heat-up of guide tubes. Cronenberg in his analysis of TMI-2 assumed that the outside of the tubes was cooled, the mechanism being radiation. Cooling of the tubes via radiation and convection need to be modelled in order to provide a realistic estimate of the guide tube temperatures. There is little driving head to force material into the tubes, only that due to gravity and this may be opposed by the presence of water and possible steam generation. This factor may provide a saving mechanism for the guide tubes.

Debris Interactions within the Vessel

149

A possible failure mode for pressurised sequences i n a PWR with lower head penetrations is the failure of the weld retaining the instrument penetrations [ 1 2. 10] . Three ways in which the weld may be threatened include: (a)

thermal attack from a jet of melt entering the lower head,

(b)

heat transfer from particulate material in the weld vicinity,

(c)

heat transfer from a molten pool .

Mode (a) has been examined by Morgan and Turland [ 1 2 . 1 ] using a heat transfer correlation from the Zion Probabilistic Safety Study and the MELTPV Code. Calculations showed that even for a large jet, the degree of thermal ablation would be insufficient to fail the weld. Models have also been formulated to investigate failure of the weld by conduction of heat down the stubs of the nozzles and deposited in the weld area. However much of heat transferred to the welds would be transferred to the vessel surrounding the weld. Heat transfer to the welds from a molten pool could occur if debris in the vessel is uncooled. Heat transfer to peripheral welds away from the axis of the vessel could provide a potential weld failure mechanism since significant sideways heat flux could occur away from the base of the vessel. Meyer [ 1 2. 1 1 ] has investigated whether failure of a weld leads to melt ejection. Several studies have been carried out for various plants. A finite element heat-up and mechanical response model has been used at INEL which tends to imply penetration failure is more likely than vessel wall failure. However under some conditions, e.g. for the Grand Gulf BWR it has been predicted that 'binding' could occur. This phenomenon is possible when the vessel head and instrument penetrations are made of different steels with different thermal expansion coefficients. The situation is very complex under reactor conditions. The issue as to whether tube and melt ejection can occur must therefore be regarded as open at the present time.

12.9

Debris Behaviour at Vessel Failure

Failure of the lower head immediately following core slump is not considered lightly. In high pressure sequences failure would probably occur minutes later probably by tube ejection. For lower pressure sequences larger amounts of debris are expected to accumulate and the timescale for failure would be much longer.

1 50

Debris Interactions within the Vessel

The debris conditions at vessel failure are expected to be similar to those at core slump either because of minimal interactions in the lower head or because the vessel penetrations will not fail until the debris has re-heated. Most of the models have aimed to take account of information learned on melt progression from the TMI-2 accident and other experimental data. The predictive capabilities of the codes are very uncertain. Simpler in-vessel melt progression codes e.g. MARCH or MAAP can be used to provide scoping calculations provided uncertainties are taken properly into account. More detailed codes have also been developed: these claim to be mechanistic but are only validated against a small data-base largely consisting of data from the TMI-2 accident.

REFERENCES 1 2. 1

A T D Butland, et al Report on Phase 1 of the PWR Severe Accident Containment Study. AEE Winfrith Report AEEW-R I 842 ( 1 984). 1 2.2 P Kuan. Core Relocation in the TMI-2 Accident Idaho National Engineering Laboratory Informal Report EGG-TMI-7402 ( 1 986). 1 2.3 D F Fletcher, and A Thyagaraja. A Mathematical Model of Premixing in Proc. 25th National Heat Transfer Conference, Houston, Texas, July 1 988, ANS , HTC-3 , 1 84 - 1 94 ( 1988). 1 2.4 S G Bankoff and Y Yang. Studies Relevant to In-Vessel Steam Explosions. Proceedings of the Fourth International Conference on Nuclear Reactor Thermal­ Hydraulics [NURETH-4] , Karlsruhe, October 1 989 ( 1 989). 1 2.5 R O Wooton, P Cybulskis and S F Quayle. MARCH2 (Meltdown AccidentResponse Characteristics) Code Description and User Manual . B attelle Columbus Laboratory Report BMI-2 1 1 5 [NUREG/CR-3988] (1984). 1 2.6 M F Young, M Berman and L T Pong. Hydrogen Generation During FueV Coolant Interactions. Nucl. Sci. Eng 2..8. 1 - 1 5 ( 1988). 1 2.7 C B Boyle and J N Lillington. Natural Circulation Heating of the Sizewell Primary Circuit under Postulated Severe Accident Conditions. AEA Thermal Reactor Services Report AEA-1RS-5005 ( 1 990). 1 2.8 J Horhorst (ed). SCDAP/RELAP5/MOD2 Code Manual, Volume 4: MATPRO - A Library of Material Properties for Light Water Reactor Accident Analysis. EG&G Idaho Report EGG-2555, Volume 4 [NUREG/CR-5273] ( 1 990). 1 2.9 A W Cronenberg, S R Behling and J M Broughton. Assessment of Damage Potential to the TMI-2 Lower Head due to Thermal Attack by Core Debris. Idaho National Engineering Laboratory Informal Report EGG-TMI-7222 ( 1986). 12.10 Zion Probabilistic Safety Study. Commonwealth Edison Co. Docket No. 50-295 ( 1 98 1 ). 1 2. 1 1 R 0 Meyer. Bottom Head Failure Program Plan. Proceedings of the USNRC Seventeenth Water Reactor Safety Information Meeting, Rockville, October 1989, NUREG/CP-O I05 Vol 2, 2 1 7-222 ( 1 990).

Debris Interactions within the Vessel

151

12.12 S S Dosanjh (ed). rdELPROG-PWR/MOD I : A Two-dimensional, Mechanistic Code for Analysis of Reactor Core Melt Progression and Vessel Attack under Severe Accident Conditions. Sandia National Laboratories Report SAND88- 1 824 [NUREG/CR-5 1 93] (1 989). 12. 1 3 B D Turland and J Morgan (eds). Compendium of Post Accident Heat Removal Models for Liquid Metal Cooled Fast Breeder Reactors. European Applied Research Reports, Nuclear Science and Technology Section, Q (5)) 1 003 - 14 1 8 ( 1 985). 12. 14 D A Petti, Silver-Indium-Cadmium Control Rod Behaviour in Severe Reactor Accidents. Nuclear Technology Hi 1 28 - 1 5 1 (1989). 12. 15 J E Kelly, R J Henniger and J F Dearing, MELPROG-PWR-MOD I Analysis of a TMLB Accident Sequence. Sandia National Laboratories Report SAND86-2 1 75 [NUREG/CR-4742] 9 1 987). 12. 16 R W Wright, Current Understanding of In-Vessel Core Melt Progression. Paper lAEA-S M-296/95, Proceedings of International Symposium on Severe Accidents in Nuclear Power Plants, Sorrento, March 1 988. 12. 17 A Skokan. High Temperatures Phase Relations of the u-Zr-o System, Proceedings of the Fifth International Meeting on Thennal Nuclear Reactor Safety, Karlsruhe, September 1984 ( 1984). 12. 18 T J Theofanous et al. An Assessment of S team Explosion Induced Containment Failure. Nucl. Sci & Eng. , 97, 259 - 325 (4 parts) ( 1987). 12. 19 B D Turland and J Morgan. Thennal Attack of Core Debris on a PWR Reactor Vessel. Proc. International Meeting on LWR Severe Accident Evaluation, Cambridge (Mass) (1 983).

153

Chapter 13 FISSION PRODUCTS

13.1

Introduction

Under conditions of normal operation gaseous fission products collect in the pores of the fuel. As burn-up increases e.g. above 20MWd/kg U, the pellets swell due to the gas pressure. Release of gas from pores is fairly small at lower temperatures below about 1 50(tC but increases at high temperatures. Consequently fission product release from the fuel is of particular interest under severe accident conditions and there has been major model development activity in this field. The consequences of fission product release during a severe accident depend on the quantity, form and timing of the release of fission products from the containment. These entities are usually called collectively the 'Source Term' In order to release fission products into the environment various barriers have to be breached. These include: (a)

the fuel matrix

(b)

the fuel rods

(c)

the vessel and primary circuit

(d)

the containment.

Estimates of the Source Term from simplified analyses e.g. The Reactor Safety Study [ 1 3 . 1] were found to be conservati ve compared with the Source Term from Three Mile Island [ 1 3 .2] . One assumption resulting in this conservatism was the neglect of fission product retention in the reactor coolant system NUREG-0772 [ 1 3.3]. The quest to reduce this conservatism has led to studies of the detailed physical and chemical behavior of the fission products in the reactor coolant system. Many of the key processes are described by Silberberg et al 1986 [ 1 3 .4] . This chapter catalogues some of the models that have been developed for predicting the key mechanisms at various stages in the release process. A capability to calculate the initial inventory is the first basic requirement. Detailed mechanistic codes e.g. VICTORIA [ 1 3 .5] have been developed to predict fission product release from degrading fuel and its subsequent behavior in the reactor coolant system. Various versions of the code have been produced: later versions include detailed modelling for control rod behaviour, cladding failure, effects

1 54

Fission Products

of fuel oxidation, effects of structures, deposition and resuspension of aerosols and also the effects of chemistry. One of the problems in modelling the source tenn is that there are potentially hundreds of fission products species that may be released under the range of conditions that can exist under degraded core conditions. The geometry may change from the original fuel rod lattice geometry to a highly molten state involving a range of metallic and ceramic melts. Temperatures may be as high as the fuel melting point - 3 100K; pressures may be as high as 17MPa. The oxidation potential of the atmosphere may also change. It will be oxidising initially when a considerable amount of steam is present. Later it will tend to become reducing as the steam is converted to hydrogen via oxidation of the in-core metals, particularly zirconium. The timescales for fission product release are scenario dependent but typically involve the fIrst few tens of minutes up to the first few hours after core cooling is lost

13.2

Sta ges of Severe Accident Fission Product Release

Fission Product behaviour is typically analysed separately within: 1.

The vessel and reactor coolant circuit, and

2.

The containment, auxiliary buildings etc.

The fonner is mainly the phase of the accident up to the time of vessel failure and includes the time of ejection, quiescent or otherwise, of core and structural materials into the containment. The latter is concerned with the behavior of materials in the containment subsequent to the time of their release from the RCS . Prior to vessel failure, the main sources to the containment include steam, hydrogen, noble gases (Xe and Kr), low melting point (volatile) species involving (Cs, I and Te) and lesser amounts of other less volatile species. These sources pass to the containment through breaks (either the initiating event or as a consequence of primary circuit over-heating, or through leakages through valves as a consequence of system over-pressurisation). Much of the work carried out to date has concentrated on release within the vessel and transport within the RCS during this time. After vessel failure additional mechanisms take place in-vessel which could also impact on fIssion product release and the subsequent source tenns. The core materials not ejected will continue to heat up, possibly now through oxidation by air as well as steam. Also there may be revaporisation [ 1 3.6] or re-entrainment [ 1 3 .7] of fIssion products previously plated out on the primary circuit pipework. Anal yses and experiments have shown that many of the fIssion products released are retained in the vessel and the reactor coolant system either because they condense on to relativel y cold

Fission Products

1 55

structures or they condense on to aerosols and settle and deposit within the vessel and the primary circuit. The main exceptions to this rule include: (a)

the non-condensable noble gases Xe and Kr,

(b)

the highly volatile species that are released through breaks or relief valves as vapours,

(c)

those fission products that are carried out of the vessel and the reactor coolant circuit as aerosols.

There are various considerations in making quantitative assessments of fission product release and behaviour in the vessel and RCS. These include: 1.

Assessment o f release to the containment prior to and after vessel failure.

2.

Assessment of fractions of fission products released from the fuel but retained in the vessel and reactor coolant system .

3.

Detennination of the quantity of fission products that remain in the core debris.

4.

Determination of the behavior of fission products that remain in the core.

5.

Assessment of the potential for re-vaporisation and re-entrainment release from the reactor coolant circuit late in the accident sequence.

13.3

Fiss ion Product Behavior in the Fuel

Mechanistic fission product behaviour models have been developed and incorporated into advanced computer codes e.g. VICTORIA [ 1 3 .5] . This code treats the fuel as two separate regions, Figure 13. 1 : 1.

Intra-granular region which includes the fuel grains and their immediate surfaces.

2.

An extra-granular region which includes the pores, the gap and the cladding.

In the interior of the fuel grains, i.e. in region 1 above, the important phenomena are detennined by the equilibrium chemistry, intra-granular solid state diffusion and fission product gas bubble behavior. In region 2, the corresponding phenomena of interest include equilibrium chemistry, molecular diffusion and convective transport. Several models have been developed for FP release from within the grains. The Booth [ 1 3 .8] model assumes diffusion by a Fick's law relationship to provide a source for a transport equation in the extra-granular region. A more mechanistic diffusive release flow model has been developed by Rest and Zowadzki [ 1 3.9] .

Fission Products

1 56

The extra-granular regions can be considered as various sub-regions; the pores, the gap, the cladding and the fuel surface. Transport in the extra-granular region can be due to both convection and diffusion. This model receives its sources from the intra-granular region.

POROSITY CLAD D I N G G RAI N S

SU RFAC E S

B R EACH

G AP

FIGURE 13.1 INTACT FUEL FP RELEASE

The transport processes in both the grains and the porous region can be modelled by a similar transport equation, the output being the species concentration. The VICTORIA model considers various transport paths within the fuel grains. The fission processes create an initial grain concentration which can then diffuse to an open pore or be transported to the grain surfaces. A general transport equation can be written for these processes, acr(Jt

=

-V(-DVc) - ac jat

to provide c, the gas atom concentrations in the grains, given diffusion coefficients, D, and cb ' the gas bubble concentration. The solution of the general system produces a solution for all the concentrations. If it is assumed that there is a dominant path from the gas atoms to the open pore then the model is similar to the Booth model. In the Booth model the general transport equation is simplified.

Fission Products

1 57

For transport within the open pores both convection and diffusion processes are modelled via a similar equation to the Booth model.

13.3.1 Intra-granular Processes Mechanistic models e.g. in [ 1 3.5] have been developed to detennine the molar concentrations of atoms in the intra-granular gas bubbles, grain face bubbles and grain edge bubbles, by fonnulating rate equations and then solving the resulting matrix. Key quantitites that are considered include the bubble radii, the intra and inter-granular concentrations of fission products, the grain size and the probability of grain edge tunnel inter linkage. The concentration of noble gas atoms has been derived from models taking account of work by Olander 1 976 [ 1 3 . 1 0] and Matthews and Wood 1 980 [ 1 3 . 1 1 ] . Concentrations of the intra-granular bubbles in the VICTORIA code have been modelled via an analogous approach. An equation for the concentration of gas bubbles on the grain faces is also given in [ 1 3 .5] as are models for the grain edge bubble concentration including the loss of gas resulting from release through interconnection of grain edge pores to open pores. The model in VICTORIA makes the assumption that the number of gas atoms in a given bubble change by examining the bubble growth and shrinkage fluxes that influence an average size bubble. Contributions to fission gas release come from the venting of grain face gas into interconnected grain edge channels, from the venting of previously trapped grain edge gas and from the long range migration of fission gas bubbles up the temperature gradient. Models have been developed to take account of the difference in geometry depending on whether the bubbles are intra-granular or inter-granular. The diffusional growth of non-equilibrium intra-granular bubbles has been considered by Gruber 1 978 [ 1 3 . 1 2]. The diffusional growth for non-equilibrium inter-granular bubbles has been given by Speight and Beere 1975 [ 1 3 . 1 3] . Grain boundary sweeping o f gas bubbles accounts for the interaction between moving grain boundaries and bubbles on the grain faces. One theory, due to Speight and Greenwood 1967 [ 1 3. 14] , makes the assumption that small bubbles are swept along with the moving boundary but that larger bubbles, since they exert larger drag are absorbed by the grain . Grain growth kinetics are affected by fuel stoichiometry. Increased levels of oxygen in V02 lead to observed increases in the diffusivity of Xe and Kr; Belle 1961 [ 1 3 . 1 5] . Activation energies for hyperstiometric i.e. oxidised fuel, have been given by Turnbull 1972 [ 1 3 . 1 6] . These are significantly smaller than those for stoichiometric fuel see e.g. [ 1 3 . 17].

Fission Products

1 58

Fuel dissolution is typically characterised by two limiting conditions: 1.

Grain boundary dissolution occurs where there i s limited attack of molten a-Zr(O) on the fuel microstructure which results in a U-rich phase at the grain boundaries. This acts as an escape route for the fission products into the pellet surface.

2.

Complete dissolution of the fuel matrix occurs at higher temperatures where there is dissolution over the entire grain structure and fission product release is controlled by microbubble and atomic diffusion in the liquified fuel. For grain boundary dissolution, fission product release occurs mainly by fission product migration through the liquified film to the fuel surface. In fuel matrix dissolution, release occurs via fission product migration through the bulk melt to the fuel surface. Gas bubble mobility in the melt can occur via bubble rise in a viscous liquid, evaporation or condensation and volume diffusion. The dominant mechanism is determined by bubble size, Rest and Cronenberg [ 1 3 . 1 8] .

13.3.2 Extra-granular Processes Extra-granular processes occur once the fission product species arrive at open porosity surfaces. The fission products may be in condensed or gaseous form. In the mechanistic models, e.g. [ 1 3.5] , the gaseous species are transported by convection and diffusion. Permeable flow conditions may be assumed. Condensed species may be modelled by surface diffusion. Otherwise in other geometric configurations, transport may be by solid state or Knudsen diffusion. The physical state of the species can be determined assuming chemical equilibrium conditions. Concerning the condensed phase species diffusion, coefficients have been given for surface diffusion of U02 by Maiya 1 97 1 [ 1 3. 19]. In some models e.g. [ 1 3 .5] this coefficient has been assumed to apply to other condensed species as well. For the gas phase, diffusion coefficients have been modelled according to the Chapman Enskog model, see Bird et al 1960 [ 1 3 .20]. In reality each species diffuses at a different rate and there will be a number of diffusion boundary layers. Average diffusion coefficients corresponding to a single boundary layers are used in codes such as VICTORIA. The coefficients are calculated from the Chapman-Enskog kinetic theory for a binary system. The uncertainty in the pressure and temperature have an important effect on the calculated diffusion coefficent with this model. The correct functional form of the diffusion coefficient for the high temperature and pressure regime is not known. Under some circumstances, convection in the open porosity of the fuel can dominate over diffusion. Convection of the condensed phase is likely to be negligible but convection of the gaseous phase may be treated with a permeable flow model. In this model the pressure gradient in the gaseous phase is the driving force. This may be evaluated assuming an ideal gas law and summing the molar concentrations of each

Fission Products

1 59

gaseous species. Viscosity can be calculated using the Wilke semi-empirical fonnula for a mixed gas. Once fission product species reach the outside of the surface of the fuel pellet stack or an exposed debris bed surface then transport through a different medium is now the issue. In intact fuel the gap between the fuel and the cladding will provide a mechanism for axial gas phase diffusion. If the fuel-cladding gap was closed then there would be a suitable surface diffusion rate for the condensed phase. In degraded or non-lattice fuel geometries there may be obstructions or crusts and these may present changes in transport properties which must be treated. In any new geometry the intercharge between fuel, gap, cladding and the bulk gas region may need to be considered. Fission product transfer across the gap may be treated by models as described for the open pores region. Gas phase species can be modelled in a similar way to the open porosity region: diffusion coefficients can be specified in a similar manner. In [ 1 3 .5] diffusion of condensed species across the gap is not allowed. Surface diffusion models for condensed species will be introduced into the VICTORIA code in the future for the case when there is fuel/cladding contact In this particular model, the boundary layer film on the control rod surfaces is assumed to be in thennodynamic equilibrium and analogous to the film on fuel rods and structures. In other respects the physics modelling is similar to that used for the control rods. Just prior to bursting the internal pressure is calculated using the work of Powers [ 1 3 .22] . The partial pressure of the vapour species above the molten alloy is calculated from the temperature and the com position of the alloy and from thennodynamic properties such as activity coefficients and changes in Gibbs free energies. Once the failure temperature has been reached and the internal rod pressure exceeds the external pressure, release of the molten alloy takes place. The model assumes that a certain prescribed fraction of the material is released. Size distributions for aerosols are assumed to be log-nonnal based on experiments carried out at Winfrith, Bowsher et al 1 986 [ 1 3 .23] . Vapour species are added directly to the bulk gas.

13.4

Convective Transport

In this section the conservation of the gaseous species in the bulk gas flow is considered for flows within the reactor vessel. A transient two-dimensional model has been developed, see e.g. [ 1 3 .5] . In this model the geometry is concentric rings. Transport is by convection between adjacent radial rings on a given axial level and between different axial levels for a given radial ring.

1 60

Fission Products

Fuel surface films are located on vertical structures, Figure 1 3.2. The film geometry is calculated from a film thickness which is prescribed. The thickness is used to calculate species concentrations in the chemistry package and is distinguished from the diffusion length. These parameters fix the Sherwood number for diffusive transport and limit the size of the diffusion boundary layer. Transport between the fuel surface film and the bulk channel flow is by diffusion. The diffusion coefficient for each species is calculated in the gas mixture. The diffusion coefficient is calculated by averaging the species diffusion coefficients by their molar fractions in the bulk gas. A correction is also made for the resulting small area between the average length of diffusion and the species dependent diffusion coefficient. Structural surface films may be on either vertical (wall) or horizontal (floor or ceiling) surfaces. Film thickness and diffusion lengths are calculated as for the fuel surface film above. Structural and fuel film thickness vary with height, Nusselt number, and species in the bulk gas or film. The transport equations model diffusive transport of the gaseous species between the bulk (gas) flow and each of the three types of surface area described above.

CON D E N S E D S P E CI E S

STR U CT U R E LAYER

STR U CT U R E

VAPO U R

B U LK GAS AND AE ROSOL TRAN SPORT

FIGURE 13.2 SURFACE FILM LAYERS

Fission Products

161

Limi tations on the maximum concentration film thickness are placed in order to prevent axial flow being prevented entirely. Transport of species in the coolant channel is detennined by the flux of species between the fuel film , the bulk gas and the structure films. Chemistry sources are added in [ 1 3 .5] by casting the sources in tenns of a delta-function adjustment to the species initial condition at the beginning of each timestep.

13.5

Chem istry

The timing, chemical fonn and the quantity of fission products released from the vessel and/ or the reactor coolant system may be significantly affected by chemical interactions amongst themsel ves and the other core materials. Iodine is released at a relatively early stage under severe accident conditions. A particular issue of interest is the chemical fonn in which it exists, because this affects its volatility. Another issue of particular interest is the importance of resuspension by re-entrainment and by revapourisation of deposited aerosols, particularly if the reactor coolant system is suddenly depressurised, e.g. through failure of the lower head or a failure elsewhere in the reactor coolant system. The chemistry of fission products within the fuel has been investigated by x-ray diffraction techniques, Kleykamp, 1988 [ 1 3 .24] . These studies showed that fission product elements either combine with the V02 matrix or fonn separate phases at grain interstices. This chemical behavior can affect the concentration of fission products in the grains and therefore diffusion to the grain boundaries and into the open pores. Some fission-products e.g. telluriwn (Te) react with the Zircaloy cladding. Te reacts with both zirconium (Zr) and tin (Sn) a constituent of Zircaloy, Collins et al 1987 [13.25]. Sn is volatile and can react with fission products in the vapour phase (Grimley and Maudlin, 1986) [ 1 3 .26] . The cladding reaction with steam also affects the chemical behavior of fission products, through the increase in temperature cladding and reduction of the oxygen concentration in the steam. The partition between condensed aerosol species and gaseous species is influenced by the gas composition since fission products react with steam , hydrogen and other vaporised metallic components. Structural materials present include stainless steel, Inconel and other Zircaloy components. Other materials present include AgIIn/Cd in PWR control rods and B4C in BWR control rod blades. Many species and elements can be treated in the mechanistic codes if chemical rate limiting processes are not treated. In codes such as VICTORIA the assumption is made that chemical interactions are transport limited with chemical equilibrium prevailing. The assumption is often made that the chemical phases are ideal. This means the quantity of any species is detennined by the total mass of the individual elements. This is a

Fission Products

1 62

reasonable asswnption in many circumstances although it may be less true for condensates and low temperature gases.

13.5.1 Eq u ilibrium Chemistry For a system at temperature T and pressure P, the condition for chemical equilibriwn is given by l:J.ldn = 0 where J.l is the chemical potential of the individual species and n is the number of moles. This equation is often derived for a system held at constant temperature and pressure, by minimizing the Gibbs free energy of the system dG = 0. Advanced codes may perform equilibrium chemistry calculations for different regions of the reactor vessel. A large number of chemical species (-288) made up of about 26 elements are allowed for in the model [ 1 3 .5] . Chemical equilibrium is calculated for ideal solutions. Models employing these assumptions have been developed to apply to the bulk gas region between structures and fuel, to structure surface regions, to the fuel cladding! gap region, to the fuel open porosity region and to the fuel grains.

13.5.2 Non-Equilibrium Chemistry In some accident scenarios it is necessary to consider the effects of some specific non­ equilibrium chemistry. Models have been developed to treat the volatilisation of U02 [ 1 3 .27] , fuel oxidation effects on release rates [ 1 3 .28] , Zircaloy oxidation, the chemisorpton of tellurium by cladding based on data from [ 1 3 .29] and the interaction of CsOH with structural material [ 1 3.30].

13.6

Fission Product Effects on Decay Heating

Under reactor accident conditions vapours and aerosols released from the core region contain various fission products. Therefore if these are deposited on structural materials, these may heat up the structures, possibly inducing revaporisation of fission products already deposited, or perhaps even posing a threat to the integrity of the structures. Revaporisation is a concern during the later stages of an accident, particularly if the containment has already failed. The removal of fission products from the core region reduces the decay heat levels in the core region at the expense of increased heating of the primary circuit. This induces a temperature feedback mechanism between the core and the remainder of the primary circuit. The transport of elements away from the core region has been discussed earlier.

Fission Products

1 63

Basic data are needed for the decay heat per mass of isotope. The ANSI 1 979 [ 1 3 . 3 1 ] decay curve is commonly used. Data on the relative abundance of isotopes are also available Bennett, 1 979 [ 1 3 .32] and these data combined allow the amount of decay heat per mass of each element to be determined as a function of time after reactor shut down. The VICTORIA code makes the additional assumption that all isotopes of an element have the same diffusive and transport processes. This allows the local decay heating to be determined everywhere in the RCS as a function of mass of each element.

13.7

Aerosols

A recent tool developed by Wheatley 1988 [ 1 3.33] is the CHARM model for aerosol species and transport. This model assumes that aerosol particles have a single constant composition, can agglomerate and deposite on surfaces. Time-varying external conditions are assumed to be available as data to the model. Agglomeration and deposition models for the advanced VICTORIA code have been derived by Gelbard 1 982 [ 1 3 .34] for the MAEROS code and by Jorgan and Kohlman 1 985 [ 1 3 .35] for the 1RAPMELT 2 code. Aerosol number concentration may be modelled in a general equation of a form as given in [1 3.5] . The aerosol number concentration distribution depends on the agglomeration kernel, the removal rate for particles and the number concentration source rate for the particles. The detailed terms in the above equation are discussed by Drake 1 972 [ 1 3.36] . A n equation o f the form above contains various assumptions. These are essentially that the aerosol is well mixed, that particles are characterised by their size only, that boundary layers are not included in detail and that particles do not break up into smaller particles. Agglomeration and deposition rates depend on various gas properties including the viscosity, thermal conductivity, diffusion, mean free path and the mass fractions of the various species. To a first approximation the bulk gas may be assumed to be composed of steam and hydrogen and the properties cited may be determined accordingly. Codes such as VICTORIA allow for a more general gas composition. Schmidt and Knudson numbers as well as particle mobility are needed to determine aerosol behaviour. Turbulence parameters including viscous and diffusion boundary layer thicknesses are also required.

13.7.1 Deposition Mechanisms There are various aerosol deposition mechanisms, Figure 1 3.3: gravitational settling turbulent flow of particles

Fission Products

1 64 laminar induced deposition thermophoretic deposition

inertial deposition due to flow irregularities.

COLLI S I O N WITH OBSTR U CTI O N S VAP O U R D E POSITI O N TURB ULENCE

I N E RTIA B ROWN I AN M OTI O N TH E R M O P H O R ETIC

E-E NTRAI N M E NT

FIGURE 13.3 AEROSOL DEPOSITION MECHANISMS

Models have been produced for super-micron particle deposition Sehmel, 1 970 [ 1 3 .37] and sub-micron particle deposition, Davies 1966 [ 1 3 .38] . Deposition from laminar flow is a well understood phenomena, Gormley and Kennedy 1 949 [ 1 3 .39] . Thermophoretic deposition is the motion of particles due to a thermal gradient. A formula used in many aerosol physics codes is due to Brock 1 962 [ 1 3 .40] . Models for aerosol deposition due to geometric irregularities including aerosol deposition in bends have been produced by Pui 1987 [ 1 3.4 1 ]: for aerosol deposition due to sudden contraction in a pipe models are given by Ye and Pui 1 990 [ 1 3 .42]. Models have also been developed to calculate aerosol deposition in specific components e.g. in an up comer cyclone .. type steam separator, 1m et aI, 1 987 [ 1 3 .43].

13.7.2 Agglomeration Mechanisms Agglomeration processes are characterised by mechanisms that result in the collision and subsequent coalescence of particles. The typical processes include:

Fission Products

1 65

Brownian motion Gravity Turb ulence. Brownian motion agglomeration has been studied by Fuchs 1 964 [ 1 3.44] and Sitarski and Seinfeld 1 977 [ 1 3.45 ] . Gravitational agglomeration rates have been developed based on the work of Prupacker and Klett, a fonnation recommended by Dunbar [13.46] . Turbulent shear and inertial agglomeration rate have been given by Saffman and Turner 1 956 [ 1 3 .47] . Turbulent resuspension of deposited aerosol particles has recently been characterised by Wright et al 1992 [ 1 3 .48] and a model has been developed by Reeks - Reed - Hall [ 1 3 .49] .

13.8

Fission Product Release from Debris

Under severe accident conditions the core will heat up, melt and relocate to lower portions of the core, possibly fonning blockages, melt poo ls etc. If the core is quenched whilst in this state rubblised beds of debris may fonn. Fission product release models for both these circumstances have been discussed. For rubble beds, fission product species transport can be modelled via a Darcy like transport equation [ 1 3 .5] . For molten pool s, release models involving volume diffusion, bubble coalescence and interaction, and viscous bubble rise are available.

REFERENCES 13. 1

1 3 .2 1 3 .3

1 3 .4

1 3 .5

WASH- 1400, 1975, "Reactor Safety Study - An Assessment of Accident Risks in US Commercial Nuclear Power Plants," NUREG-75/104, US Nuclear Regulatory Commission. M Rogovin, 1979, "Three Mile Island: A Report to the Commission and to the Public," NUREG/CR- 1 250. US Nuclear Regulatory Commission, 1 98 1 , Technical Bases for Estimating Fission Product Behaviour During LWR Accidents, NUREG-0772, Office of Nuclear Regulatory Research, US Nuclear Regulatory Commission, Washington DC . M Silberberg, J A Mitchel, R 0 Meyer and C P Ryder, 1986, "Reassessment of the Technical Bases for Estimating Source Tenns," NUREG-0956, Section 6, US Nuclear Regulatory Commission, Washington DC. T J Heames et ai, VICTORIA: A Mechanistic Model of Radionuclide Behaviour in the Reactor Coolant System under Severe Accident Conditions.

1 66 1 3.6

13.7 1 3 .8

1 3 .9

1 3. 1 0 1 3. 1 1

13.12

13.13 1 3 . 14 13.15 13.16 13.17 13.18

13.19

1 3 .20 1 3 .2 1

Fission Products B R Sehgel and R L Ritzman, 1 987, "Integrated Analysis of Fission Product Transport in Reactor Coolant Systems," in Proceedings of the Symposium on Chemical Phenomena Associated with Radioactivity Releases During Severe Nuclear Plant Accidents, NUREG/CP-0078, US Nuclear Regulatory Commission. T J Heames and R C Smith, 1 99 1 , "Integrated MELPROG{fRAC Analyses of a PWR Station Blackout, "Nuclear Engineering and Design: 1 25, 175- 1 88. A Booth, 1 957, "A Method of Calculating Fission Gas Diffusion from UOz Fuel and its Application to the X-2-f Loop Test," Chalk River Report CRDC-72 1 , AECL, Chalk River, Ontario, Canada. J Rest and S A Zawadzki, 1 992, "FASTGRASS: A Mechanistic Model for the Prediction of Xe, I, Cs, Te, Ba, and Sr Release from Nuclear Fuel Under Nonnal­ and Severe-Accident Conditions: User's Guide for Maintenance, Workstation, and Personal Computer Applications," NUREG/CR-5840, ANL-92/3, Argonne National Laboratory, Argonne, IL. D R Olander, 1 976, "Fundamental Aspects of Nuclear Reactor Fuel Elements" TID-267 1 1-PI US Technical Infonnation Centre, Springfield, VA P320. J R Matthews and M H Wood, 1 980b, "An Efficient Method for Calculating Diffusive Flow to a Spherical Boundary," Nuclear Engineering and Design, 56 (No 2 Feb): 439-44. E E Gruber, 1 978, "The Role of Bubble-Size Equilibration in the Transient Behaviour of Fission Gas." ANL-78-36, Argonne National Laboratory, Argonne, IL. M V Speight and W Beere, 1975, "Vacancy Potential and Void Growth on Grain Boundaries," Metal Science, 9 (No 4, April): 1 90- 1 9 1 . M V Speight an d G W Greenwood, 1 964, "Grain Boundary Mobility and its Effects in Materials Containing Inert Gases," Philosphical Magazine 9, 683-689. J Belle, 1 96 1 , Uranium Dioxide, US Government Printing Office, Washington DC, p5 12. J A Turnbull 1972 "A Review of Rare Gas Diffusion in Uranium Dioxide," RD/ B/N2405, Berkeley Nuclear Laboratories, Berkeley, GLOS , United Kingdom. J R MacEwan and J Hayashi 1 967, "Nuclear and Engineering Ceram ics," Proceedings of the British Ceramic Society, 7:245. J Rest and A W Cronenberg, 1987, "Modelling the Behaviour of Xe, I, Cs, Te, Ba and Sr in Solid and Liquefied Fuel During Severe Accidents, "Journal of Nuclear Materials, 1 50 (No 2, Oct): 203-225. P S Maiya, 1 97 1 , "Surface Diffusion, Surface Free Energy, and Grain-Boundary Free Energy of Uranium Dioxide," Journal of Nuclear Materials, 50 (No I , July): 57-65 . R B Bird, W E Stewart and N E Lightfoot, 1 960, Transport Phenomena, John Wiley and Sons, New York, NY. N M Chown and D A Williams, 1989, "A Limited Investigation of the Sensitivity of the Containment Source Tenn to Certain Primary Circuit Parameters Under PWR Severe Accident Conditions," AEEW-R2429, UKAEA, Winfrith, United Kingdom .

Fission Products

1 67

13 .22 D A Powers, 1985, "Behaviour of Control Rods During Core Degradation: Pressurization of Silver-Indium-Cadmium Control Rods," NUREG/CR-440 1/ SAND85-0469, Sandia National Laboratories, Albuquerque, NM . 13 .23 B R Bowsher, R A Jenkins, A L Nichols, N A Rowe and J A Simpson, 1986, "Silvier­ Indium-Cadmium Control Rod Behaviour During a Severe Accident," AEEW­ R 1 99 1 , UKAEA, Winfrith, United Kingdom. 1 3 .24 H Kleykamp, 1988 "The Chemical State of Fission Products in Oxide Fuels at Different Stages of the Nuclear Fuel Cycle," Nuclear Technology, 80: 4 1 2-422. 13.25 J L Collins, M F Osborne, and R A Lorenz, 1 987, "Fission Product Telluriwn Release Behaviour Under Severe Light Water Reactor Accident Conditions," Nuclear Technology, 77 (No 1 , April): 1 8-3 1 . 13 .26 A J Grimley and P J Maudlin, 1 986 "Aerosol Fonnation o f Non-Radioactive Species: Implications for the Source Tenn ," Proceedings of the International ANS/ENS Topical Meeting on Thennal Reactor Safety, held in San Diego, CA (February 2-6) Vol 2, American Nuclear Society. 13.27 C A Alexander and J S Ogden, 1 990, "Vaporization of U02 at High Temperatures and High Pressures: A Generic Relation for Volatilization," High Temperatures/ High Pressures, 2 1 : 149. 13.28 B J Lewis, F C Iglesias, D S Cox, and E Gheorghiu, 1990 "A Model for Fission Gas Release and Fuel Oxidation Behaviour for Defective U02 Fuel Elements," Nuclear Technology 92: 353-362. 1 3 .29 B R Bowsher and R A Jenkins 1 99 1 , "High Temperature Studies of Simulant Fission Products: Part V, Temperature-Dependent Interaction of Tellurium Vapour with Zircaloy," AEEN-R I 9 1 5, UKAEA, Winfrith, United Kingdom. 13.30 M D Vine and P N Clough, 199 1 , "A Modelling Approach for Heterogeneous Chemistry in the VICTORIA Source Tenn Code," AEA TRS 5082, AEA Safety and Reliability, Culcheth, United Kingdom. 1 3 .3 1 ANS I, 1979, "American National Standard for Decay Heat Power in Light Water Reactors," ANSl/ANS-5. 1 - 1 979. August 29, 1 979. 13.32 D E Bennett, 1979, "Sandia-ORIGEN User's Manual," SAND79-0299, NUREG/ CR-0987, Sandia National Laboratories, Albuquerque, NM. 1 3 .33 C J Wheatley, 1986 "CHARM: A Model for Aerosol Behaviour in Time-Varying Thermal Hydraulic Conditions," S AND88-074 5, NUREG/CR-5 1 62, Sandia National Laboratories Albuquerque, NM. 1 3 .34 F Gelbard, 1982, "MAEROS User Manual" Sand80-0822, NUREG/CR 1 39 1 , Sandia National Laboratories, Albuquerque, NM. 1 3 .35 H Jordan and M R Kohlman, 1985, "TRAP-MELT2 User's Manual," NUREG/ CR-4205, BMI-262 1 , Battelle Columbus Laboratory, Columbus, OH. 1 3 .36 R L Drake, 1972, "A General Mathematical Survey of the Coagulation Equation, "G M Hidy and J R Brock, eds, International Reviews in Aerosol Physics and Chemistry, Vol 3, Pergamon Press, Elmsted, NY. 1 3 .37 G A Sehmel, 1970, "Particle Deposition from Turbulent Air Flow," Journal of Geophysical Research (No 9, March): 1 766- 1 78 1 . 1 3 .38 C N Davies, 1966, Aerosol Science, Academic Press, London, United Kingdom. 1 3 .39 P G Gormley and M Kennedy, 1949, Di ffusion from a Stream flowing Through a Cylindrical Tub, "Proceedings of the Royal Irish Academics" A52: 1 63- 1 69.

1 68

Fission Products

1 3 .40 J R Brock 1 962, "On the Theory of Thermal Forces Acting on Aerosol Particles, "Journal of Colloid Science, 1 7:768-780. 1 3 .4 1 D Y Pui, F Romay-Novas, and B Y Liu, 1987, "Experimental Study of Particle Deposition in Bends of Circular Cross Section," Aerosol Science and Technology, 7 (No 3):30 1 -3 1 5. 1 3 .42 Y Ye and D Y Pui, 1 990, Particle Degradation in a Tube with an Abrupt Contraction" Journal Aerosol Science, 2 1 , 1 :29-40. 1 3 .43 K H Im, R K Ahluwalia, H C Lin, 1 987, "The RAFT Computer Code for Calculating Aerosol Formation and Transport in Severe LWR Accidents," NP-5287-CCM, Argonne National Laboratory, Argonne, IL. 1 3 .44 N A Fuchs, 1964, The Mechanics of Aerosols, Pergamon Press, Elmsted, NY . 1 3 .45 M Sitarski and J H Seinfeld, 1977 '.'Brownian Coagulation in the Transition Regime, "Journal of Colloid and Interface Science, 61 (No 2, Sept): 26 1 -27 1 . 1 3 .46 I H Dunbar et al 1984 " Comparison of Sodium Aerosol Codes," EUR 9 1 72 en, Commission of the European Communities, Ispra, Italy. 1 3 .47 P G Saffman and J S Turner, 1956, "Corrigendum to "On the Collision of Drops in Turbulent Clouds," Journal of Fluid Mechanics, 1 : 1 6-30. 1 3 .48 A L Wright, W L Pattison and J Y King, 1 992 SERIES-2 Aerosol Resuspension Test Data Summary Report, Oak Ridge National Laboratory, Oak Ridge. To be published. 1 3 .49 M W Reeks, J Reed and D Hall, 1 988, "On the Resuspension of Small Particles by a Turbulent Flow, "Journal of Applied Physics, 2 1 :574-589.

1 69

Chapter 14 CAVITY PHENOMENA

14.1

Introduction

Previous chapters have been concerned primarily with phenomena that occur in-vessel and the corresponding models that have been developed to simulate these phenomena. This chapter is concerned with debris behaviour after vessel failure. This behaviour can impact significantly on subsequent events in the accident sequence. Many of the issues are discussed in [ 14 . 1 ] and [ 1 4.2] . Some of the models that have already been described for simulating in­ vessel events e.g. debris bed cooling are relevant to modelling some ex-vessel events: these models are briefly referred to here again but in this new context It is important to identify whether there are processes whereby the high heat content of the debris (possibly compounded by oxidation heating of metallic components) could be released to the containment possibly leading to an over-pressurisation failure. Two processes have been identified: dispersal of debris to containment causing direct heating of the atmosphere rapid quenching of the debris by water possibly accompanied by a steam explosion. An important issue is whether the debris is quenched and remains coolable or whether it is uncoolable. To determine which is the case it is necessary to develop models for cavity interactions to predict the extent of debris fragmentation and its dispersal. It should be noted that heat may also be dissipated through debris interaction within the concrete basement: the extent of this reaction will also impact on containment pressurisation. Various metals in the debris may be oxidised. leading to hydrogen and CO production. If a sufficiently high concentration is generated then there is a possibility of a hydrogen bum as a potential contributor to containment pressurisation.

14.2

Debris Formation and Mixing

Important factors concerning the formation of debris are the effects on the debris of expulsion from the vessel and the extent of mixing with water in the cavity. Mixing in the cavity is discussed in this Section. The role of possible steam explosions is discussed in Section 14.4.

Cavity Phenomena

170

For there to be significant cooling of the debris it is expected that there would need to be a large surface area between the debris and the water at all times. There are competing mechanisms however, concerning the likely effectiveness of mixing. Vigorous steam generation during the boiling process will tend to encourage mixing. However, against this, there is a large density difference between the two materials (corium and water) and the effects of steam blanketing could lead to a segregation of the water and the debris. Parametric models for debris fonnation, quench and debris bed fonnation have been developed in the MARCH code [14.3] and applied to various accident scenarios. In this model particles are assumed to be spherical: their sizes are required to be specified as user input. There are various parametric inputs into this model e.g. whether the debris is quenchable or not, whether a debris bed may fonn, the criteria for bed fonnation and whether the bed (if fonned) is fixed or may fluidize. Models have been developed for bed fluidisation which depend on flow rate [ 14.4] or particle steam generation rate. In the MARCH model this is based on the assumption of an isolated sphere. The effects of surrounding particles are ignored. Compared with reality, the boiling model has limitations in not taking account of these surrounding particles which also cause the absorption of thennal radiation by the water to be less complete. These effects tend to reduce (compared with the predictions of the MARCH model) the actual steam generation rate available for bed fluidisation. The critical flow rate and force to support an isolated sphere are given in the literature [ 1 4 .5] . The minimum fluidisation velocity for spherical particles and for non-spherical debris has been given by Wallis [14.6] . He suggests that the minimum fluidisation velocity is likely to be less than that predicted in the above model [ 14.4 ] , particularly for the case of non -spherical debris.

14.3

Hydrogen Production

Models have been produced for the oxidation of debris in the cavity. These require infonnation on the particle characteristics in order to define a geometric surface available for oxidation. Different models allow different composition of the debris but usually assume the particles are spheres. The MARCH parametric model [ 14 .3] assumes that particles of debris are spheres and composed of one or two compositions: (i)

fuel material surrounded by shells of zirconium and zirconium oxide,

(ii)

iron or structural material surrounded by an iron oxide shell.

The composition of debris upon vessel failure is subject to uncertainty. Additional uncertainties that need to be considered in modelling, concern the physical state of the particles whether these consist of: (iii)

a homogeneous mix of debris,

Cavity Phenomena

(iv)

oxidic debris,

(v)

metallic debris.

171

Where the debris has been segregated into different components one practice e.g. [ 14 .3] is toassume that the zirconium oxidation reaction takes place preferentially, i.e. available steam preferentially oxidises any zirconium: any residual steam is available to oxidise steel or structural material. There are various oxidation reactions of steel that have been considered. In a steam rich environment Fe)04 is formed [ 14 .7] , [ 14.8] and this is perhaps the most likely reaction. However Fe)04 oxidises to Fe20) on heating with air [ 14.8] . The FeO oxide is unstable: it can separate to a mixture of Fe and Fe)°4 upon standing [ 14.7] or can burn in air to gi ve Fez0) [14.8] . The stability or otherwise of these steel oxides depends on the nature of the oxidising (or reducing environment). The geometry of debris particles will be required and models have made various assumptions on whether the particles have: slab, thin shell or spherical geometries in order to determine the reaction rates. Different models make different assumptions concerning the heat distribution between the particles, steam and water. In [ 14.3] the hydrogen is released at the debris temperature. Various possibilities are then considered which include: (a)

the hydrogen mixes with steam at saturation temperature

(b)

as for (a) except that the gaseous mixture is cooled by water

(c)

the bed is assumed to heat the unreacted steam, which then mixes with the hydrogen.

In all cases the resultant gaseous mixture is eventually released to the containment.

14.4

Steam Explosions in the Cavity

Important considerations concerning this phenomenon are the amount of molten corium that might be released into the cavity and the amount of water that would mix effectively with the corium. Whether or not the cavity is dry is likely to be scenario dependent. This is discussed later. If the cavity is completely flooded then a large pressure pulse could result due to inertial effects of all the water. However, steam is produced as the corium interacts with the water and this mechanism introduces compressibility into the system. This has been demonstrated in experiments at Berkeley Laboratories [ 14.9] . The effectiveness of this mitigating

1 72

Cavity Phenomena

mechanism i.e. the generation of sufficient steam will be somewhat impaired if the cavity water is subcooled. The implications of the case with maximum water inertia have been studied from the point of view of the impact loads on the vessel. In the case of a damped explosion energy is dissipated as a pressure pulse through the cold water. This energy is typically a small fraction of the energy available. Experimental results have been obtained from the SPIT and IDPS experiments from Sandia National Laboratories, USA. These have been mentioned earlier. The SPIT series were at 1/20 scale of the cavity geometry of the Zion plant: HIPS was similar at 1/10 scale. One of the SPIT experiments showed a high pressure pulse in a fully flooded simulation where significant water inertia was present. Codes have been developed which replicate some of these results. One such code [ 14 . 1 ] , [ 14.2] assumes that the molten material fragments to particles o f different sizes based o n a log-normal distribution as suggested by experimental results from Winfrith [ 14 . 1 0] . Heat is released to the cavity by thermal diffusion from the particles. The water steam mixture is assumed to expand isentropically such that the work in expansion is balanced by the inertial energy transferred to the water in the cavity. Other circumstances where the cavity is neither dry nor completely flooded have been less thoroughly analysed because of the problems in modelling a freely vented surface. The experimental evidence is that if the surface is freely vented then there are low energy yields, particularly if subcooling is present. Some workers have investigated whether efficiencies close to those predicted by the Hicks and Menzies model [ 1 4 . 1 ] can be realised. The heat released from finely fragmented debris will generate steam at elevated pressures which then has the potential to do work on its surroundings. Hicks and Menzies estimated the maximum efficiency of this conversion likely in a steam explosion by considering the interaction as an adiabatic constant volume mixing phase during which the debris and water reach thermal equilibrium, followed by an isentropic expansion. There are various reasons why this upper bound is unlikely to be attained. For example the Hicks-Menzies description can only be considered as a useful approximation to the physical events if the timescale for heat transfer from the particles is less than the time to eject water from the cavity. The Hicks and Menzies assumption of initial mixing at constant volume cannot be realised since free volume will be generated in the mixing region by the steam ejecting some water. The efficiency also depends on the relative amounts of debris and coolant. The efficiency may be low with, for example, an excess of sub-cooled water. Tests have been carried out at Sandia National Laboratories involving models of flooded cavities. The models were destroyed by thermite melt water interactions. One of the SPIT experiments (SPIT1 7 [ 14 . 1 2]), Figure 14. 1 , showed very little free volume generated before

1 73

Cavity Phenomena

N

0

X C'l 'Vi C. W

c:::: ::::>

\n \n W

c:::: �

n""""'"'

' 3 2 00

BU RSTS

26 0 - 0 0

22400 '80

00

136 00

9 2 00

J - 40 - 00

\_---- -----------,/� .......

13',Ll 2-=] 4 8 0----38 0 ---'' 4-L 00----13' 4-L 2 0-----"4 4'4 '6'4 '-80 --'-'00 -0 --'-'0 0 ----"5 00 0 --

TI M E I N

ms

FIGURE 14.1 SANDIA EXPERIMENT - SPIT 17 - THERMITE INTERACTION PRESSURE PULSE

the equipment burst However, the HIPS tests indicated that there are a range of possibilities that can occur should debris enter a flooded cavity. The cavity models in two of the HIPS tests (HIPS-4W and HIPS-6W [ 14 . 1 3]) failed but only after much of the water and some of the debris had been injected by the blowdown gas stream. A fully resisted expansion did not occur in these HIPS experiments.

14.5

Debris Transport within the Cavity

The behaviour of debris within the cavity will be scenario dependent and in particular will depend on the mode of vessel failure. For some pressurised sequences debris could be ejected into the cavity with some considerable force. This could lead to debris being swept along in the cavity and fmally up into the containment. Following water and gas discharge could further enhance the sweep-out of debris. The CaROE code or Core Debris Ejection Model has been developed at Winfrith to model this situation. The code is based on a lumped param eter, one-dimensional flow through the cavity together with phenomenological models and correlations for cavity inputs and special phenomena. The main purpose of the code is to provide sources of particulate and gaseous material for containment analysis codes and in particular to be used in the assessment of risk from Direct Containment Heating or OCH. The phenomena modelled, Figure 14.2, have been identified from various investigations into high pressure melt ejection [ 14.4] , [ 14.5] , [14.6] , [14.7] and [ 14.8] .

Cavity Phenomena

1 74

The initial conditions for the CORDE code are based on the assumption that a pool of molten core debris exists in the bottom head of the reactor pressure vessel. Molten debris are then discharged through a user prescribed orifice e.g. an instrument guide tube penetration. The model allows the hole size to expand through thermal ablation and as the hole grows the over gas (steam and hydrogen) will tend to penetrate the pool surface and the two-phase mixture of gas and corium will be discharged from the vessel. Correlations for this phenomenon can be found in [ 14. 19]. Once the corium is completely exhausted the discharge will be steam and hydrogen only. Different models have been considered for thermal ablation. Annular flow models assume that the area for heat transfer between the debris jet and the vessel wall is unaffected by gas blowthrough. Heat transfer correlations for short tubes [8] are used in the CORDE code. However, some modellers e.g. Pilch and Tarbell [14.20] allow for the break-up of the liquid film. The fragmentation of the jet of molten debris will be caused by dissolved gas and by atomisation. These processes are not modelled mechanistically in the CORDE code, the size distribution of the debris particles is user specified. The basic assumption is that debris will accumulate as a liquid pool at the bottom of the cavity prior to the gas blowdown phase. Once the gas blowdown begins the molten debris in the cavity will be pushed away and if the gas velocities are sufficiently high then debris particles will be entrained in the gas flow. Figure 14.3 shows a detailed Computational Fluids Dynamics (CFD) code simulation of possible flow directions within a reactor cavity during the blowdown phase. The extent by which debris could be ejected from the cavity is likely to depend on various processes. These include the inertial deposition in bends as the flow changes directions, the splashing, re-entrainment of debris from surfaces, the crusting of debris and the levitation of debris particles in upward flows. In the CORDE code drag coefficients are based on [ 14.21 ] . The chemical reactions between the zirconium and iron and other metallic components in the melt also require modelling since these materials will be present in particles exposed to the air. These particles will be at high temperature and have relatively large surface area.

14.6

Debris Coolability

To achieve long term coolability, it is necessary for the molten material to be quenched and for water to penetrate all the debris. Heat can then be removed by boiling and convection. Otherwise the bed will dry out. Large Break LOCA sequences [ 14.22] are thought to provide the greatest challenge from the point of view of debris coolability. In these circumstances melt is not expected to be energetically ejected from the vessel and therefore there is likely to be less dispersion of debris in the cavity. For vessel failures at higher pressure where there is significant dispersion of debris out of the cavity, the debris beds may be expected to be more spread out and shallower, with increased potential or likelihood of contacting water and getting quenched.

Cavity Phenomena

175

There are considerable uncertainties concerning the nature of the debris beds that may form. The key parameters affecting long term coolability are: (i)

debris size: for debris involved with steam explosions particle diameters can be expected to be in the range O . l mm to 1 .0mm; otherwise particulate debris diameters are likely to be in excess of several millimetres,

(ii)

the structure of the bed: in particular whether stratification exists,

(iii)

the location of debris: the extent of debris dispersion from the cavity and the depths of any debris beds formed.

For the experiments performed, it is generally recognised [ 14.23] , [ 14.24] , [ 14.25] , [ 14.26] that porous media based models provide reasonable agreement with the data for the range of particle diameters, bed structures and coolants studied. Porous media based models are based on mass and energy conservation but with momentum equations based on Darcy's law. Some modellers include the Ergon modification applied separately to each phase [ 14.27] . These models have been found to be largely successful. However workers at Sandia [ 14.28] and Winfrith [ 14.29] have shown some discrepancies in the details. There is considerably less dependence of dryout heat flux on pressure that is predicted by the porous media based models. Barleon [ 14.30] noted that the bed power had to be substantially reduced, below the power needed to create the dry zone, before re-wetting can start and this observation has also been supported in Winfrith work.

14.7

Uncoolable Debris

If molten material enters the cavity it may or may not contain water. Without water, and particularly for the case of vessel failure at high pressure, the debris particles could transfer their heat preferentially to the vapour phase in the containment There is then little chance that water would be effective in dissipating heat Even if water is present in the cavity unquenchable debris may form a molten poo l on the floor, under any water and with the potential to attack the concrete. These mechanisms are considered in Chapter 1 5. A mitigating factor concerning the issue of unquenchable debris is that if this concrete reaction occurs then evolved gas may assist in promoting better mixing of water and the debris and thereby effect cooling.

Even if the debris is initially quenchable, it is possible that all the water may be evaporated, after which time the bed will dry out and heat up. Until this time the debris bed physics may be modelled by a porous medium based approach. Benocci et al [ 14.3 1 ] however modelled the dry debris bed using a porous medium model, assuming that the bed is saturated with vapour. Convective flow effects were modelled

176

Cavity Phenomena

G ", s /p ... F l ow

r�

.

. ' . ( l 1t� ... l '!'r ... n s [ e r • Ox l d a ti o " & • • l I y d r oq e n P r()( I I ) ( : 1 i nn )

•

• •

• •

• •

•

•

� . Tr app i n " by St ruc t ures •

FIGURE 14.2 SCHEMATIC REPRESENTATION OF CAVITY PHENOMENA

by a Boussinesq approximation. The approach was originally developed for Liquid Metal Fast Breeder Reactors (LMFBRs). A limitation of this model is that the important effect of convection is not included. This could provide a mechanism for introducing further steam into the debris bed, resulting in more debris heat-up through oxidation. The debris may be quenched but the debris bed fonned may be uncoolable and the comments and approach above on dry debris beds appl y if the supply of water is exhausted. Even if there is an overlying pool of water it is not clear how much heat may be lost from the debris to the water. Gases released from the core concrete interaction may reduce cooling in the bed by expelling water from the bed but sufficiently increased gas flow may in contrast produce greater cooling.

14.8

Uncertainties in Debris Behaviour in the Cavity

There are uncertainties in phenomenology concerning debris behaviour in the cavity, particularly if the primary circuit is still at pressure. (It should be remembered that in the case of the PWR there is a potential for failure elsewhere in the primary circuit due to natural circulation phenomena, prior to failure of the lower vessel head). For accident scenarios with the primary circuit close to the containment pressure, the accumulators will have already discharged. In these circumstances in the case of a severe

1 77

Cavity Phenomena

accident the cavity could be dry unless for example the break is directly to the cavity or perhaps other engineered safeguards e.g. containment sprays are in operation. If the cavity is dry then little fragmentation of debris and transfer of heat to the containment atmosphere is expected for a low pressure scenario. If the debris had already been in contact with water in the vessel and the vessel failed, this would be indicative of high temperature debris. The water would follow the debris to the cavity. Little further interaction between the debris and water in the cavity might be expected and the debris might be expected to attack the concrete.

I I I

I

I

I I

I

I I

I

I I I I I

I I I

)

FIGURE 14.3 DETAILED FLOW PATTERNS IN THE CAVITY

For a high pressure sequence, there is some evidence that if the vessel fails initially through a penetration then some debris may be ejected from the cavity. Experimental evidence suggests debris in sub-millimetre scale particles but there are uncertainties in scaling to the reactor. However, whatever fonn the debris may take it is likely that there will be disentrainment of the debris from the gas flow due to the tortuous routes to the containment. These routes will be design dependent. A discussion of uncertainties of this phase of the accident have been given in the SAUNA report [ 1 4.32]. This is a generic study with the emphasis on the reactions with water rather than whether the cavity is wet or dry. These have also been other studies. High pressure melt ejection phenomena were recognised in the APS review [ 14.33] . An ANS report on Source Tenns [ 14.34] considered the high pressure melt ejection phenomena in high pressure sequences. The IDCOR report [ 14.35] also considered the potential for direct heating in high pressure sequences.

1 78 14.9

Cavity Phenomena Summary

Early models e.g. as in the MARCH code assumed that debris always remained in the cavity but more recent data and CORDE analysis indicates that debris may reach other parts of the containment building. The CORDE code has been specifically developed to address sweep out from the cavity in high pressure sequences. The mode of vessel or primary circuit failure which could give rise to a high pressure melt ejection in one of a nom ber of possibilities. It is not clear whether a significant amount of material will reach the containment and this could be design dependent. Models have been developed and studies have been carried out to assess the likelihood of steam explosions in a fully flooded cavity. The tamper produced by the water could allow high pressures to be generated and lead to an efficient conversion of the slug thermal energy to kinetic energy.

REFERENCES

14. 1 14.2 14.3 14.4 14.5 14.6 14.7 14.8 14.9

14. 10 14. 1 1 14.12 14. 1 3

A T D BuLland et at. Report on Phase 1 of the PWR Severe Accident Containment StUdy. Winfrith Report AEEW-R I842 ( 1984). A T D Butland et at. Report on Phase 2 of the PWR Severe Accident Containment StUdy. Winfrith Report AEEW-R I964 (1985). R 0 Wooton, P Cybulskis and S F Quayle. MARCH2 Code Description and User's Manual. NUREG/CR-3988 ( 1 984). J F Davidson and D Harrison (eds) 'Fluidization' Academic Press (197 1 ). G K Batchelor ' An Introduction to Fluid Dynamics' Cambridge University Press (P34 1) ( 1967). G B Wallis 'One-Dimensional Two-Phase Flow' McGraw-Hill (p1 82 Equation 8.35) ( 1969). E Hutchinson 'Chemistry: The Elements and Their Reactions 2nd ed, W B Saunders Co. ( 1 964). F Sherwood Taylor 'Inorganic and Theoretical Chemistry' 10th ed, Heinmann (1960). M Baines (May 1983). Preliminary Measurement of Steam Explosion Work Yields in a Constrained System. First UK National Heat Transfer Conference, Leeds, 3-5 July 1 984. D Fletcher ( 1984). The Particle Size Distribution of Fragmented Melt Debris from Molten Fuel Coolant Interactions. AEEW M 2 1 03. E P Hicks and D C Menzies. Theoretical Studies on the Fast Reactor Maximum Accident. ANL 7 1 20, 654-670 (1965). W W Tarbell. Initial Test Results - SPIT 17 - Sandia Laboratories Minute (revised), 5 December 1 984. W W Tarbell, M Pilch and J E Brockman. Behaviour of Core Debris Ejected from a Pressurised Vessel into Scaled Reactor Cavities. Sandia Laboratories report (1 984).

Cavity Phenomena

179

14. 14 B W Spencer, D Kilsdonk, J J Sienicki and G R Thomas. "Phenomenological investigations of Cavity Interactions Following Postulated Vessel Melt Through", Proc. Int. Meeting on Thennal Reactor Safety, Chicago ( 1 983), NUREG/CP-0027, Vol. 2. 14. 15 B W Spencer, L M McUmber and J J Sienicki. "Results and Analysis of Reactor Material Experiments on Ex -Vessel Corium Quench and Dispersal" Proc. Int. Meeting on Thennal Nuclear Reactor Safety, Karlsruhe (1984). 14. 16 W W Tarbell et al "Pressurised Melt Ejection into Scaled Reactor Cavities", NUREG/ CR-45 1 2 ( 1986). 14.17 R V Macbeth and R Trenberth "Experimental Modelling of Core Debris Dispersion from the Vault under a PWR Pressure Vessel, Part 1 : Preliminary Experimental Results", AEEW-R I 888 (1987). 14.18 P W Rose "Experimental Modelling of Core Debris Dispersion from the Vault under a PWR Pressure Vessel, Part 2: Results Including the Intsrument Cable Support Structure", AEEW-R2 143 ( 1987). 14. 19 J Reimann and M Khan "Flow Through a Small Break at the Bottom of a Large Pipe with Stratified Flow", Int J. Multiphase Flow 12, 4 ( 1986). 14.20 M Pilch and M Tarbell "High Pressure Ejection of Melt from a Reactor Pressure Vessel - The Discharge Phase", NUREG/CR-4383 ( 1985). 14.21 R Clift, J R Grace and M E Weber "Bubbles, Drops and Particles" Academic Press ( 1978). 14.22 Zion Probabilistic Safety Study (1981). Commonwealth Edison Co., Docket No. 50-295. 14.23 R J Lipinski (1982). A Model for Boiling and Dryout in Particulate Beds. Sandia Laboratories Report SAND82-0765 (NUREG/CR-2646). 14.24 R J Lipinski ( 1983). A Review of Debris Coolability Models. Proc. ANS International Meeting on LWR Severe Accident Evaluation (Cambridge, Mass.), Paper 1 8.2. 14.25 T Ginsberg, J Klein, J Klages, C E Schwarz and J C Chen (1983). Phenomenology of Transient Debris Bed Heat Removal, in "Post Accident Debris Cooling" (U Muller and C Gunther eds) pp 1 5 1 - 1 58, G Braun, Karlsruhe ( 1 983). 14.26 D Squarer, L E Hochreiter and A T Pieczynski (1983). Some aspects of Decay Heat Removal from a Debris Bed and its Implication to Degraded Core Coolability. ibid 14.25. pp 1 59- 164. 14.27 K More and B D Turland ( 1 983). Developments of Models for Boiling and Dryout in LWR debris beds. ibid 14.25. 14.28 K R Boldt, P A Kuenstler Jr and T R Schmidt ( 1984). Results of the In-Pile Degraded Core Coolability Experiments DCCI and DCC2. Presented at the 5th International Conference on Thennal Reactor Safety, Karlsruhe. 14.29 G F Stevens (1984). Experimental Studies of Dryout During Boiling in Particle Beds at AEE Winfrith (UKAEA). Presented at the Information Exchange Meeting on Debris Coolability, UCLA. 14.30 L Barleon, K Thomauske and H Werle ( 1984). Extended Dryout and Rewetting of Particulate Core Debris Beds. ibid 14.28. 14.3 1 C Benocci, J-M Buchlin, C Joly and A Siebertz (1 983). A Two-Dimensional Finite Difference Modelling of the Thennohydraulic Behaviour of the PARR Debris Bed up to Extended Dryout, ibid. 14.25. pp. 265-27 1 .

1 80

Cavity Phenomena

14.32 J B Rivard et al. Identification of Severe Accident Uncertainties. NUREG/CR-3440 (SAND83- 1 689) ( 1984). 14.33 R Wilson et al. Radionuclide Release from Severe Accidents at Nuclear Power Plants. Report of a Study Group of the American Physical Society. Rev. Mod. Phys., 57 (3), Part IT (1985). 14.34 American Nuclear Society. Report of the Special Committee on Source Terms (1984). 14.35 Nuclear Power Plant Response to Severe Accidents. IDCOR ( 1984).

181

Chapter 15 CORE DEBRIS INTERACTIONS WITH CONC RETE

15.1

Phenomenology

If debris is released from the bottom of the reactor vessel then the interaction between the debris and the underlying cavity needs to be considered. From earlier discussion there is a potential for this phenomenon to occur for some hypothesised severe accident conditions. For a coherent mass of material the debris is maintained at high temperature by the decay heat associated with the less volatile fission products. The temperatures and heat fluxes are sufficiently high that any concrete underlying the debris may be ablated and decomposed. This offers a potential route for release of radioactive material to the soil and thence into the environment. The phenomenon is known as basemat penetration. In addition the interaction process results in a copious production of water, carbon dioxide and other gases as the concrete decomposes. These gases are reduced on contact with any metallic melt to hydrogen and carbon monoxide. These gases contribute to the potential of a possible over-pressurisation threat to the containment, particularly if they are ignited. In addition the gases and aerosol produced by the core concrete interaction provide carriers for the transport of condensed fission products from the melt into the containment. An additional risk is the ablation of and failure of internal structures in the containment e.g.

the reactor pedestal in the BWR. A mechanical failure of this component could lead to failure of the component and then the containment. There are therefore fundamental safety issues associated with molten core concrete interactions.

15.2

Relevant System Components

Codes such as CORCON [ 1 5 . 1 ] have been developed which include models for three basic system components or physical volumes, the debris pool , the concrete and the atmosphere above the pool . The composition of each component consists of four basic material groups: oxidic compounds, metals and other elements, gases and other various compounds. The debris poo l is assumed to consist of a number oflayers. These could be from the bottom of the concrete cavity upwards: heavy oxide phase heterogeneous mixture of oxides and metals metallic phase

1 82

Core Debris Interactions with Concrete

coolant atmosphere. Such a model enables pool s of various layers to be modelled. The concrete cavity in which the core debris resides is a chamber: it is convenient for modelling purposes to assume that it is axisymmetric. Concrete composition must also be specified since different concretes exhibit different interactions. Examples of such concretes include: basaltic aggregate concrete limestone aggregate common sand concrete and others. Data required to model the ablation process include liquidus and solidus temperatures and ablation temperatures of the materials. The principal constituents of concrete include CaCOJ or Ca(OH)2. Computer codes, such as CORCON, also have provision for the user to define non-standard concrete. The capability to include steel reinforcing bar is also included. The atmosphere above the pool provides a heat sink for the evolved gas and energy emitted via convection and radiation from the pool surface. Convection heat losses may be calculated once sink temperature or heat flux conditions have been prescribed. Radiation heat transfer is based on the surroundings temperature and the opacity of the atmosphere. The aerosol concentration is modelled in some codes to allow the optical thickness of the cavity atmosphere to be estimated and therefore modelled. There are a range of physical processes that need to be considered in modelling core/concrete interactions. Examples include: internal heat generation heat and mass transfer chemistry concrete response bubble mechanisms. Some of the fundamental models for these various phenomena are considered in the sections below. The mechanisms are tightly coupled and this must be accounted for in the computational modelling,

Core Debris Interactions with Concrete 15.3

1 83

Heat Generation

Thermal mechanisms dominate the attack of molten corium on concrete in a LWR. Decay heat and some oxidation heating are generated in the IXlOI and are dissipated either through the top surface or into the concrete. The distribution of heat between these regions is determined by the thermal resistance of the various heat flow paths. Heat transferconstituti ve relations are therefore extremely important. Since the more volatile fission products will be released prior to formation of the pool , standard decay curves e.g. ANS are not correct (without modification). Decay heating can be calculated using detailed decay chains [15.2] but codes such as CORCON have applied a simpler approach. It has been shown that the decay power between one hour and ten days is reasonably proportional to operating power and less sensitive to burn up. Codes such as CORCON­ MOD2 have therefore employed an ORIGEN [ 1 5.3] calculation for a reference PWR core at equilibrium burn up to identify a selection of elements which essentially account for all the heat production. Assuming the initial core inventory is proportional to the core operating power, the initial pool inventory can then be determined by the mass of the core in the melt multiplied by a retention factor for each element The decay power is then calculated from the specific decay power (W/g-atom) generated from each element (taken from ORIGEN) and the inventory of the element Decay powers associated with each element are given in [ 15.3]: these are calculated using simple correlations of the form: P(t) = mc exp (-A.t) where P(t) is the decay power (w) t is time (days) m is the elemental mass (gram-atoms) c is a coefficient (w/gram-atom) A. is a coefficient (day·I). Other losses may occur from the melt due to vaporisation, mechanical sparging and aerosols driven by the concrete decomposition gas. Models such as CORCON consider vaporisation of alkali metals and halogens. According to Powers, aerosol generation is not a significant factor in material loss or reduction in decay power.

15.4

Pool Heat Transfer

From experimental studies, it has been established that if the heat source is sufficiently large

1 84

Core Debris Interactions with Concrete

then the internal temperature of the pool adjusts itself so that the heat losses are balanced by the internal heat generation. This observation has led to the development of quasi - steady heat transfer models which therefore have the advantage that the history of the pool can be neglected. In codes such as CORCON-MOD2 where a multi-layered poo l is modelled, the interfacial temperatures are adjusted subject to the criterion that the heat fluxes are continuous across the interfaces. The layers can be in various physical states e.g. completely molten a solid crust (with or without one or more surfaces). Considering fIrst a liquid layer (the other case is considered later) heat transfer coeffIcients are required from its interior to its surfaces. Boundaries may be between layers or between the pool and the concrete. Various models have been put forward to take account of the effects of bubbles e.g. the Blottner [ 15.4] modifIcation of the Konsetov model [ 15.5] for vertical surfaces to take account of tmbulent material convection associated with bubble buoyancy. Models for horizontal liquid/liquid interfaces have been developed by Greene [ 1 5.6] based on his data and the data of Werle [ 15.7] [ 15.8]. Another similar model has been developed by Szekely [ 1 5.9]. For the coolant (water) layer, boiling heat transfer models are required. Standard pool boiling correlations that have been used include those given by Bergles [15. 10] . Nucleate boiling may be treated using the Rohsenow [ 1 5. 1 1 ] correlation for the temperature rise and the Zuber [ 1 5. 1 2] [ 1 5. 1 3] correlation for critical heat flux (NB Rohsenow recommended some modification of the coefficients). The effect of subcooling on nucleate boiling critical heat flux has been addressed by Ivey [ 1 5 . 14]. The actual heat flux can then be calculated using the methods of Rohsenow. For film boiling the Berenson correlation [ 1 5. 1 5] is available. Above the Leidenfrost point the total heat flux depends on both convection and also a radiation component Transition boiling regimes can be treated in the standard way by appropriate interpolation, between the nucleate and ftlm boiling regimes. At low velocities heat transfer in the coolant and the molten debris is dominated by natural convection and standard Nusselt Rayleigh correlations [ 1 5. 16] are available. These cover both the laminar and turbulent regimes (low and high Rayleigh number respectively). For pool heat transfer applications natural convection limits can be imposed by choosing the greater of the Nusselt numbers calculated for bubble-enhanced and for natural convection. For very thin or viscous layers, the natural circulation correlations can be replaced by a simple conduction model.

Core Debris Interactions with Concrete 15.5

185

Surface Heat Transfer

The pool surface provides an interface with the containment. Models may need to be interfaced with those in a containment code. The energy exchange problem can therefore be split into two halves i.e. the determination of heat fluxes to and from the surface as a function of surface temperature. The problem then reduces to finding the surface temperature which provides the desired continuity of heat flux. Heat loss from the pool surface includes convective heat transfer to the atmosphere and radiative heat transfer to the surroundings. Simple models have been formulated. Experience indicates that convection and radiation are strongly coupled: radiation tends to increase thermal stability and reduce convection.

15.6

Heat Transfer Between the Melt and the Concrete

Various models have been formulated for prediction of heat through this interface. The interface is a thin region and various gases and molten oxides associated with concrete decomposition are present in it Examples of models are given in [ 1 5. 1 7] , [ 15.4], [ 1 5. 1 8] , [15. 1 9] and [15.20]. The model in CORCON-MOD2 makes the assumption that the boundary region contains a gas film. This film is modelled differently depending on whether the boundary is horizontal or vertical.

C R UST � x :::> -J u. I­ « w

I

NO C R U S T �

SOLI D I F ICATI O N TE M PE RATU R E

I

w

o Vi -J

a a Q..

I N TE RFACE TE M P E RATU RE BETWE E N TH E M E LT AN D TH E GAS F I L M

FIGURE 15.1 DEPENDENCE OF POOLSIDE HEAT FLUX ON SURFACE TEMPERATURE

Core Debris Interactions with Concrete

1 86

For horizontal surfaces the bubbles enter the melt via the Taylor instability mechanism. Thus for the horizontal surface a heat transfer coefficient can be derived based on a momentum balance in a bubbling cell [ 1 5 . 1 8] . For more inclined surfaces the gas makes a flowing film. For this case a mechanistic model may be used for both laminar and turbulent films, based on momentum balances. For the transition between laminar and turbulent flow, models have been formulated which preserve film thickness and heat transfer continuity e.g. the model by Persh [ 1 5. 1 ] . For the transition between the horizontal bubbling model c ase and the vertical fum-flow model, a mechanistic model has been formulated based on a momentum balance between the fraction of injected gas entering into bubbles and the remaining fraction in establishing the film. As for the cases described earlier, radiation across the fum must also be modelled. A form for a transparent gas between infinite parallel grey walls has been used in CORCON. The contributions to the total heat flux between convection and radiation are approximately the same in typical applications. Heat transfer through the gas film is dependent on the temperatures of its surfaces. The dependency of pool side heat flux on surface temperature is shown in Figure 1 5 . 1 . If this surface temperature is greater than the solidification temperature then no crust will exist and small changes in surface temperature will result in large changes in heat transfer. Otherwise a crust will be present and small change in surface temperature will merely affect the crust thickness with little change in the heat flux.

IS.7

Crust Behaviour

After interacting for some period the temperatures of the poo l will drop to a level where the pool will begin to solidify. There are various possibilities: at early times the crusts may form at an interface with the centre of the layer remaining as liquid, the crusts may be unstable, the whole melt could form a slurry. At later times considerable freezing may occur. Once a substantial part of the melt becomes frozen heat conduction remains the only mechanism for heat removal and this process is much less effective than convection. The tendency is that only thin crusts tend to form. Substantial freezing of the metallic layer may occur but oxidic layers containing fuel have a higher internal heating source and thermal conductivities are much lower.

Core Debris Interactions with Concrete

1 87

It is possible. for some accident scenarios. that core debris may initially be solid or partly solid. If there is sufficient internal heating and this cannot be removed then material will melt until a balance between heat production and heat losses is achieved. This means that melting tends usually to be from the centre outwards. A rigorous solution to the problem would require a transient two (or even three dimensional) heat transfer model including conduction. convection and allowing for change of phase. It would need to include sufficient spatial resolution to model crusts of several orders of magnitude smaller dimensions compared with the whole. Codes such as CORCON MOD2 employ a simpler approach i.e. quasi-steady state models[IS.2 1 ] . Convection losses are limited because the boundary temperature of the liquid interface cannot fall below the solidification temperature and crusts provide an effective thennal resistance. These effects reduce heat losses and slow cooling rates and therefore a steady state approach is acceptable. The CORCON model reduces the calculational domain to an idealised cylinder whose temperatures match those of the actual layer. The problem is further simplified by reducing it to two independent one-dimensional problems (axial and radial). For the one-dimensional calculation the layer may be solid. liquid or liquid with a solid crust. In all liquid regions heat transfer is by natural or bubble enhanced convection with a conduction limit as described earlier. In solid regions it is by conduction only. For a liquid with crusts a composition method is used in which the liquid and solid layers are treated separately subject to the assumption of continuity of heat flux at the liquid/solid interface.

15.8

Concrete Ablation

As concrete is heated. it undergoes various changes in composition. These include:

TABLE 15.1 CONCRETE CHANGES ON HEATING Event

Vaporisation of absorbed water Decomposition of calcium hydroxide Decomposition of calcium carbonate Melting of remaining oxide

Temperature (K)

400 700 1 0S0 I S00- 1 900

Carbon dioxide and water vapour are released through the pores of the concrete as pressure gradients build up. Since the oxide matrix is a mixture of compounds. melting takes place over a range of temperatures. During the molten core concrete interaction molten or semi­ molten materials are absorbed into the pool . Early codes providing a detailed treatment of the concrete are given in reference [ I S. 1 ] .

1 88

Core Debris Interactions with Concrete

Codes such as CORCON-MOD2 employ a simplified model based on a steady state one dimensional heat balance. Basically a simple heat balance at the concrete surface is linked with the melt/concrete heat transfer model referred to earlier, to detennine whether the temperature of the pool inside is high enough to promote ablation. Studies using a more detailed ablation code show that ablation processes reach a quasi-steady state fairly quickly i.e. within a minute. The generation of gases during the decomposition process can be calculated with a steady­ state model. In CORCON gas released in front of the ablation front is ignored. This assumption is less valid at early times due to the initial burst of gases that is released. The assumption is also less valid at late times since ablation ceases. One input required for modelling is the ablation energy for concrete. This consists of both sensible and chemical energy. Sensible energy includes the energy necessary to raise the gaseous decomposition products up to the concrete ablation temperature. The chemical energy for various reactions has been detennined experimentally. Important interactions include [ 1 5 . 1 ] : TABLE 15.2 CONCRETE RELEASES O N HEATING Energy (kcaVmole)

Event

Evaporation of fill water Release of chemically bound water from hydroxides Release of carbdn dioxide from carbonates

11 25 40

There is no well defined ablation temperature for concrete since the material may not be completely molten. Melting ranges are defined by the concrete solidus and liquidus temperatures. The ablation temperature is usually chosen to be bounded by these temperatures. The exact ablation temperature affects the calculated heat of ablation but provided all enthalpies are computed from the same database the choice of ablation temperature should not affect energy conservation. Concrete decomposition products enter the gas film of the pool at the ablation temperature, with a corresponding enthalpy for that temperature.

15.9

Chemical Interactions

The most important chern ical interactions concern the oxidation of metals by the decomposition gases from the concrete namely carbon dioxide and water vapour. A general calculational package has been developed by Powers using an original method of Van Zeggeren and Storey [ 15.22]. The approach is to perform a Gibbs function minimisation for some 38 chemical species comprising 1 1 elements. All the relevant condensed phases were included (metals, metal oxides, carbon) the main gaseous phases (water vapour, hydrogen. carbon dioxide and monoxide) and some additional light hydrocarbons.

Core Debris Interactions with Concrete

1 89

General packages based on this kind of approach have the advantage that specific reactions need not be specified. If the entropy of mixing and heating of solutions of the reactions are neglected then reaction precedence can be determined e.g. zirconium metal is oxidised to depletion in preference to stainless steel.

15.10 Mass and Energy Transfer

Mass transfer processes are important in the modelling of molten core/concrete interactions. These processes are closely coupled with chemical processes and this fact must be accounted for in the modelling. Mass transfer processes control the injection of concrete decomposition products up into the pool and the addition into the pool of other core or structural materials or water falling from above. Codes such as CORCON model the rising gases and condensed-phase materials from the concrete decomposition or other interactions. Relative densities determine the direction of

ATM OSP H E R E

COOLANT LIG HT OX I D E LIG HT M I XTU R E M ETAL H EAVY M I XTU R E H EAVY OX I DE G AS F I LM

FIGURE 15.2 POOL LAYERS

Core Debris Interactions with Concrete

1 90

motion. As chemical reactions take place the composition and enthalpies of these materials are modified. Materials are assumed to be thermally equilibriated with any layers through which they pass. A typical layer structure is shown in Figure 1 5.2. The energy of these materials is finally added to the layer in which they remain: any heats of reaction are deposited in the layer in which the interaction occurred. Final layers are assumed to be the first oxide layer encountered for oxides, the frrst metal layer for metals and the atmosphere for the gases. The reverse process for downward flow is modelled similarly.

15.1 1 Energy Conservation

Energy balances may be formulated for each layer of the pool , including energy sources entering and leaving each layer, energy sources associated with decay heat, chemical reactions and the ablation of concrete.

15.12 B ubble Behaviour

As gas rises up through the pool the gas bubbles cause the volume of the pool to increase, a phenomenon often referred to as level swell. Heat transfer is also affected. Void fractions are typically -0.4 [15.4]. The bubble velocity is sometimes linked to the terminal velocity see e.g. [15.23]. Codes such as CORCON consider three separate situations: small bubbles in Stokes flow medium sized bubbles large spherically capped bubbles. Bubble size as the gas enters the pool can be determined using the Taylor instability bubbling model, based on data from [ 1 5.24] . Other gas velocity/bubble-size regimes have been investigated in earlier models/codes e.g. the analogue of nucleate bubbling. However, this approach has not been continued in CORCON since the bubbles sizes were not in agreement with prototypical melt/concrete experiments. Bubble sizes are recalculated for each layer surface. Predictions of the CORCON model of the terminal velocities of single bubbles have been compared with experimental data [ 15.25] and void fractions with simulate data [ 1 5.26] .

REFERENCES

15.1

R K Cole, DP Kelly, M A Ellis, "CORCON MOD2: A Computer Program for Analysis of Molten Core Concrete Interactions" NUREG/CR-3920 ( 1 984).

Core Debris Interactions with Concrete

15.2

K K Murataetal "User's Manual forCONTAIN, a Computer Code for Severe Nuclear Reactor Accident Containment Analysis" SAND2039, NUREG/CR-5026 Sandia National Laboratories, Albuquerque, NM 1 989. 0 E Bennett "SANDIA-ORIGEN User's Manual" SAND79-0299 (NUREG/CR0987), Sandia National Laboratories, Albuquerque, NM, October 1 979. F G Blottner "Hydrodynamics and Heat Transfer Characteristics of Liquid Pools with B ubble Agitation" SAND79- 1 132 (NUREG/CR-0944), Sandia National Laboratories, Albuquerque, NM November 1 979. V V Konsetov "Heat Transfer During Bubbling of Gas Through Liquid" Int J Heat Mass Transfer, Vol 9 pp 1 103- 1 108, 1 966. T Ginsberg and G A Green "BNL Program in Support of LWR Degraded Core Accident Analysis" in Proceedings of the US Nuclear Regulatory Commission Tenth Water Reactor Safety Research Information Meeting, NUREG/CP-004 1 , Vol 2, pp 364-395, 1 983. H Werle "Modelexperiments zum Kernschmelzen" Halbjahresbericht 1 97811 . PNS 4332. 1 978. H Werle "Enhancement of Heat Transfer between Two Horizontal Liquid Layers by Gas Injection at the Bottom, KfK 3223, Kernforschungszentrum Karlsruhe, FRG, 1 98 1 . J Szekely "Mathematical Model for Heat or Mass Transfer at the Bubble-Stirred Interface of Two Immiscible Fluids" International Journal of Heat and Mass Transfer Volume 6. A E Bergles et al "Two-Phase Flow and Heat Transfer in the Power and Process Industries" Chapter 7, Hermisphere Publishing Corporation, Washington, McGraw Hill, New York, 1 98 1 . W M Rohsenow " A Method of Correlating Heat Transfer for Surface Boiling of Liquids" Trans ASME Volume 74, pp 969-976, 1952. N Zuber "On Stability of Boiling Heat Transfer" Trans ASME Volwne 80, pp 7 1 1 720 1 958. N Zuber et al "The Hydrodynamic Crisis in Pool Boiling of Saturated and Subcooled Liquids" International Developments in Heat Transfer, Part II, pp 230-235, ASME New York, 1 96 1 . H J Ivey "Acceleration and the Critical Heat Flux in Pool Boiling" Chartered Mechanical Engineering. Volume 9, pp 4 13-427. 1 962. P J Berenson "Transition Boiling Heat Transfer from a Horizontal Surface" Journal of Heat Transfer. Volwne 83, pp 3 5 1 -358, 1 96 1 . W H McAdams, Heat Transmission, McGraw-Hill Book Co, New York, NY , 1 954. V K Dhir, I Catton, J Castle "Role ofTaylor Instability on Sublimation of a Horizontal Slab of Dry Ice" Joumal of Heat Transfer Vol 99, No 3 , August 1 977, P 4 1 1 . H Alsmeyer and M Reinmann "On the Heat and Mass Transport Processes of a Horizontal Melting or Decomposing Layer under a Molten Pool" Nuclear Reactor Safety Heat Transfer, Winter Annual Meeting ASME. Atlanta, GA pp 47-53, 1977. A S Benjamin "Core-Concrete Molten Pool Dynamics and Interfacial Heat Transfer" Proceedings of the ANS/ASME/NRC International Topical Meeting on Nuclear Reactor Thermal-Hydraulics, NUREG/CP-OOI4, Vol 2, p 1437, October 1980. ,

15.3 15.4

,

15.5 15.6

15.7 15.8

15.9

15.10

15. 1 1 15. 1 2 15.13

15.14 15. 1 5 15.16 15.17 15.18

15.19

191

1 92

Core Debris Interactions with Concrete

1 5.20 M Lee, M S Kazimi and G Brown "A Heat Transfer Model for the Corium/Concrete Interface" Paper 12.6 in Proceedings, International Meeting of Light Water Reactor Severe Accident Evaluation, Cambridge, MA, 1983. 1 5.21 R K Cole Jr "A Crust Formation and Refreezing Model for Molten-FueVConcrete Interactions Codes" Paper 1 2.5 in Proceedings, International Meeting on Light Water Reactor Severe Accident Evaluation, Cambridge, MA 1 983. 1 5.22 F H Van Zeggeren and S H Storey ''The Computation of Chemical Equilibrium" Cambridge University Press, Cambridge, MA 1 970. 1 5.23 G B Wallis "One-Dimensional, Two-Phase Flow, McGraw-Hill Inc, New York, NY , 1 969. 1 5.24 E R Hosler and J W Westwater "Film Boiling on a Horizontal Plate" ARS Journal pp 553-558, April 1 962. 1 5.25 R J Andreini, J S Foster and R W Callen "Characterization of Gas Bubbles Injected into Molten Metals Under Laminar Flow Conditions" Metallurgical Transactions, Volume 8B, pp 625-63 1 , December 1977. 1 5.26 G A Greene and T Ginsberg "BNL Program in S upport of LWR Degraded Core Accident Analysis" in Proceedings of the US Nuclear Regulatory Commission Ninth Water Reactor Safety Research Information Meeting, NUREG/CP-0024 Volume 3, 1 982. ,

,

193

Chapter 16 AEROSOL PRODUCTION FROM CORE-CONCRETE INTERACTIONS 16.1

Introduction

Aerosol production from core concrete interaction forms a sizeable contribution to the total aerosol and fission product source to the containment, even for large LOCA severe accident conditions. In general high volatility fission products will be released in-vessel during the core degradation phase. Aerosols produced by core concrete interaction produce a long term aerosol and fission product source, long after aerosols produced from the LOCA phase have plated out, Table 16. 1 . The latter could however still contribute to the environmental source term in the event of a late containment failure, together with any additional aerosol production associated with gases coming out of the melt, via resuspension of already deposited aerosols. This would particularly be the case in the event of a significant depressurisation.

16.2

Important Mechanisms

A number of observations have emerged from experimental programmes at Sandia National Laboratories, VSA and elsewhere [ 16. 1]. Melts of 1 2-200 kg consisting of stainless steel, V02, Zr02 and fission product simulants have been heated up to about 3000°C and dropped into concrete crucibles. Various points have emerged. Four important mechanisms of aerosol generation have been identified [ 16.2] : bursting of bubbles, entrainment of melt in gas flow and subsequent break-up, vapour condensation following vapour release from bubbles, condensation of vapours released from the surface of the melt. Data on particle size distribution have been produced. The distribution appears to be bimodal in which each component is of log - normal profile with means typically of 1 .4J.UD and 4.5J.UD [ 1 6.3] , [ 16.4] . Aerosol concentrations in the gas emitted are found to be proportional to the gas evolution rate. The melt temperature also has an important influence on the quantity of aerosol generated [ 16.3].

194

Aerosol Production/rom Core-Concrete Interactions TABLE 16.1 TIMESCALES OF FISSION PRODUCT RELEASE

Location

Released Fission Products

Mechanism

Timescale

In-vessel

High Volatility

Core degradation. Release from the fuel matrix

Early before 1-4 hours (scenario dependent)

Ex-vessel

Low Volatility

Aerosol production in core/concrete interaction

Late after 1-4 hours (scenario dependent)

16.3

Aerosol Production Models

An early model produced by Murfin and Powers [ 16.6] was derived for the Zion/lndian Point Study in 1980. This was based on a simple correlation for the aerosol concenttation in the evolved gas at STP. This correlation required as input, the superficial gas velocity or the volumetric flow rate per unit area of melt in order that the aerosol production rate could be calculated for a melt of given temperature. Mechanistic models have been developed for the VANESA code [16.2] developed at Sandia National Laboratories, USA. Four key mechanisms for aerosol generation were identified in Section 16.2. Models for these mechanisms are discussed in the succeeding sections. These models are available in the V ANESA code.

16.4

Material Entrainment

Powers [ 16.7] gave a criterion for material entrainment if the gas velocity is sufficiently large. Typical recommended parameters for material entrainment result in velocities of ....30m/s. Velocities of an order of a magnitude higher are required for aerodynamic break up of entrained material (based on a minimum Weber (We) number criterion, We >We crit .... 12 [ 16.9] , [ 16. 10]. The bubble rise velocity is typically around O.3m/s [ 16.2] and therefore far too low to produce fragmentation and entrainment, at least after the initial stages of core/concrete interaction. There is the possibility of mechanical aerosol production if gases pass up the cavity walls, between the melt and the concrete [ 16.3], [16.4], [ 16.5] .

Aerosol Production/rom Core-Concrete Interactions 16.5

195

Aerosol Production via Bubble Collapse

If the effects of chemical reactions between vapours and the melt are neglected, then it would be consistent to assume that the size of a spherical bubble would increase as it rises to the surface since the internal gas pressure decreases as it moves upwards. Codes such as CORCON assume that the size of bubbles remain constant [ 16. 10] as the bubble rises to the surface. Bubbles rising in a plume e.g. from nucleation centres on the cavity floor do tend to interact and in such a way that their size remains constanl The ttansittimes for bubble rise in plumes are less (by about a factor of3 [ 16.6]) than for single bubbles. Smaller transit times and smaller bubble size tends to reduce the aerosol production rate from reactive vaporisation. A study of pool scrubbing is given [ 1 6. 1 1] which is relevant to the physics here. The CORCON model for aerosol production from bubble bursting is taken from data in [16. 12] which applies to single bubbles in aqueous solutions. The model assumes that a fixed number of aerosol particles are produced from each burst.

16.6

Vaporisation

Volatile fission product species will vaporise from the melt surface. This is an important effect during the early stages of the core/concrete interaction. It is considered in more detail later. 16.7

Condensation

Gases produced by the core/concrete interaction react with the melt as the gas bubbles rise up to the surface. On reaching the surface the gas bubbles burst releasing vapours which then condense and form aerosol particles. Models have been produced to take account of the gas/melt interactions and allow for ttansient effects. The transient times are too fast to assume equilibrium and rate-limiting processes need to be modelled. The V ANESA code takes account of mechanisms such as mass transport in the melt, vaporisation and reaction at the bubble surface, and mass transport away from the bubble surface into the gas. These mechanisms were originally identified in the metallurgy and iron and steel industries [ 16. 10] , [ 16. 1 3] , [ 1 6. 14] , [16. 1 5] . Mass ttansport i n the melt can be treated using mass transfer coefficients see e.g. [ 16. 10] which utilise diffusivity coefficients in [ 16. 13].

Models typically assume that the reactive vaporisation at the bubble surface for a particular species is proportional to the pressure difference between the equilibrium and actual partial pressure exerted by that species.

1 96

Aerosol Production/rom Core-Concrete Interactions

For mass transport in the gas phase, models have been produced assuming steady state diffusion with linear concentration gradients, with the boundary layer thickness length scale equal to that of the bubble diameter. Models for diffusivities are given in [16.6] but mass transport of gaseous reactants to the bubble's surface is not considered in some models: gas phase mass transport is found to be very fast and not rate controlling.

16.8

Aerosol Particle Size

Models have bee n developed assuming a log-normal particle size distribution with standard deviation derived from experiments at Sandia. An empirical correlation [ 16.2] has been given for the mass medium diameter. The VANESA code for example, assumes that the number density of particles for the gas evolving, is constant at STP.

16.9

Gas Composition at Equilibrium Over the Melt

A dominant mechanism for aerosol formation at early times is associated with bubble burst as bubbles reach the cool atmosphere above the melt, where they burst causing the species in the bubble gas to condense and form aerosol particles. Estimates have been made of the gas composition at equilibrium over the melt due to vaporisation of chemical species from the surface of the corium melt as the gases from the core-concrete attack pass through the melt. Codes such as SOLGASMIX [ 16. 1 7] can be used to determine the equilibrium composition by a direct minimisation of the Gibbs energy of the system. It is important that all the dominant species (elemental, oxide, hydroxides etc) are included, Table 1 6.2. The [mal equilibrium depends upon the species included in the thermodynamic database and the thermodynamic data used for each species, Table 1 6.3.

TABLE 16.2 DOMINANT SPECIES

fuel species (V02 Pu02, ) ' fission products (Cs, Te, Sr + oxides, . . . . . ) cladding components (Zr, Zr02' ) control rod materials (Ag, In, Cd, . . . . . . ) structural components (Fe, Cr, FeO, . . . . . . ) gases (CO, CO2, 1\, 1\0, . . . . . ). •

•

•

•

•

•

•

•

•

•

In modelling the melt various layer configurations have been considered. Oxide and metallic components may be assumed to form a homogenous liquid (although this is not thought to be physically realistic) or more likely to exist as two immiscible liquids, with either the metal

Aerosol Production/rom Core-Concrete Interactions

197

phase above the oxide phase or vice-versa depending on the relative densities of the two liquids. The gas and the oxide and metal layers may or may not be in equilibrium depending on the effects of sparging and the system can be modelled accordingly. The quality ofresults or predictions is clearly dependent on the adequacy of the thermodynamic database. The data for some of these species are reasonably well known for the gas oxides. However, current predictions are sensitive to uncertainties in the data. For two sources of estimated data for the actinides, Krikorian [16. 1 8] and Jackson [16. 19] , a difference of an order of magnitude has been found in the release of plutonium in one particular application. The indications from a number of studies are that volatile fission products (Cs, Te) are released rapidly during the core concrete interaction. For the less volatile components, the non-volatile release fraction depends on the molten debris structure. As the melt becomes more reacted with concrete components, releases are decreased. Reduction in melt temperature also produces this effect. Temperature has an important effect on the equilibrium composition, Figure 16. 1 . The more volatile components e.g. Cs are relatively unaffected but other groups e.g. La, Pu, U are substantially affected. Release ofthese components is also dependent on the layer configuration, being higher if the oxide phase is above the metal phase. The change in the heat content of the melt as a result of the removal of radionuclides generating a large amount of decay heat and due to the enthalpy change during vaporisation needs to be considered. The latent heat of vaporised species needs to be accounted for in thermal hydraulic calculations since it is comparable with the decay heat in the early stages.

TABLE 16.3 CHEMISTRY MODELLING Input

Method

Thermodynamic data

Total Gibbs free energy minimisation for a set of chemical species

Stoichiometric matrix

Output

Equilibrium composition of multi component, multiphase systems

16.10 Bubble Rise Phenomena

The closeness to equilibrium composition in the gas is determined by the mass transport in the melt and gas and by the reactive vaporisation at the interface. The area of the interface between the melt and the gas phases and the contact time are therefore important parameters. Models for bubble formation and ascent through the melt have been developed. A criterion for gas evolved at the floor of the cavity to rise as bubbly flow is established in Reference [16.20].

198

Aerosol Production/rom Core-Concrete Interactions

A mechanism for bubble release has been proposed which assumes that a gas film exists on the cavity floor and that bubble release is then dependent on the wavelength of the fastest growing mode of the Rayleigh-Taylor instability. Experimental results of Dhir et al [ 1 6.21] and Powers [16.22] appear to support this thesis. Lee et al [ 16.23] have proposed a variation to the Dhir et al model which makes the additional assumption that the melt periodically contacts the concrete, after bubble release. The next bubble is then formed and released from the concrete when its buoyancy force overcomes the surface tension face holding it to the concrete. Such models make the asswnption that the gas bubbles rise up through the melt It is possible that gas evolved from cavity walls could flow as a film up the walls ( 16.24] or could fonn additional bubbles [16.22]. Bubbles will remain spherical provided that the swface tension forces are sttonger than the 1lOIl­ spherically symmetric drag force and this indeed appears to be the case [16.25]. A criterion for bubble stability has been given by Levich [16.26] and isolated bubbles are expected to be stable. However for streams of bubbles there are other mechanisms which tend to keep the bubble size down and may result in diameters less than the minimum stable diameter for single bubbles. A correlation for the velocity of interacting bubbles has been proposed by Woodford and Scriven [16.27] . It is generally found that larger velocities are predicted for interacting bubbles than for isolated bubbles. Taking these velocities enables residence times for the layers to be determined. Bartolomei and Alkutov [ 16.28] state that the wall effects can only be neglected if the vessel diameter exceeds a minimum value.

16.1 1 Mass Transport

As the bubble rises, the chemical species must be transported through the liquid melt to the surface of the bubble, where they may then vaporise and then the vapour will be transported into the interior of the bubble. Models have been fonnulated to treat three processes: transport in the melt, surface vaporisation in the gas, transport in the gas. Other mechanisms that may occur include transport of the gases to the bubble surface and chemical reactions at the surface. Mass transport coefficients in the melt are given by Calderbank and Moo-Young [16. 19] for the vapour species in the melt These are in terms of a Schmidt number which can be estimated from the Scheibel modification of the Wilke-Change correlation [ 1 6.30].

Aerosol Production/rom Core-Concrete Interactions

199

Comparisons have been made ofpredictions with the Calderbankand Moo-Young correlation against measured values [16.3 1 ] . Calculated values are within a factor 2 of the measured values. Surface vaporisation mass transfer coefficients for the individual species have been given by Tabor [ 1 6.32]: these are incorporated into the V ANESA code. This correlation is an estimate of the rate at which molecules in the surface melt layer will gain sufficient kinetic energy to exceed the latent heat of vaporisation and escape into the gas phase. Mass transport coefficients in the gas have been given by the Koenig and Brink equation for spherical bubbles [16.33] . These incorporate gas phase diffusivities calculated from the Wilke-Lee modification of the Hirschfelder-Bird-Spatz method [16.34] . Other correlations for the diffusivity have been given by Gilliland [ 1 6.35].

:::-=====::- (s ��: K

Te Si Ag

10 -5

2733

2633

2333

TEMPERATURE

K

2233

2133

2033

FIGURE 16.1 THE EFFECT OF TEMPERATURE UPON THE EQUILIBRIUM RELEASE

Aerosol Production/rom Core-Concrete Interactions

200

16.12 Departure from Equilibrium Conditions

The fractional departures from equilibrium for the mass of each species vaporised into the bubbles can be calculated once the mass transfer coefficients are known. Significant departures have been observed for Cs but uncertainties are compounded by reductions in surface area, and shortening of bubble rise times due to convection in the melt. The departures are dependent on the layer configuration. Conservative calculations show departures for Cs and Te from their equilibriwn conditions, but most of the other species are predicted to achieve their equilibrium concentrations in the gas bubbles. However, best estimate calculations show that both Cs or Te (particularly Te) also come close to reaching equilibrium conditions.

16.13 Mechanical Aerosol Production

Ginsberg [16.36] has reviewed the literature on mechanical production of aerosol particles by gas sparging through a liquid. For the first hour of interaction, releases of some species e.g. Mo, U and II are dependent on the layer configuration, since mechanical aerosol production mechanisms are important. However, for many species this is less so since vaporisation processes dominate the release in the first hour. The estimated release rate due to mechanical aerosol formation remains fairly constant over the first few hours. The relative importance of mechanical release increases at later times. Models have been formulated e.g. VANESA in which there is an asswnption that each gas bubble that bursts at a surface releases a fixed number of identical aerosol particles. The justification for this is uncertain. Aerosol production from bubbly flow is uncertain and is an area that requires further investigation since mechanical aerosol production becomes important at late times.

REFERENCES

16. 1 16.2

16.3

D A Dahlgren et aI, "Molten LWR Core Material Interactions with Water and with Concrete" SAND-77- 12 16C. D A Powers and ] E Brockmann, Chapter 6 of "Review of the Status of Validation of the Computer Codes used in the NRC Accident Source Term Reassessment Study (BMI-2 104)" ORNL{fM-8842. D A Powers and ] F Muir "Melt/Concrete Interactions: The Sandia Experimental Program, Model Development and Code Comparison Test", Proceedings of the Seventh Water Reactor Safety Research Information Meeting, Gaithersburg, 1979. SAND79- 19 18C.

Aerosol Production/rom Core-Concrete Interactions

16.4 16.5 16.6 16.7 16.8 16.9 16. 10 16. 1 1

16. 12 16. 1 3 16. 14 16. 1 5 16. 16 16. 1 7 16. 1 8 16. 1 9 16.20 16.21 16.22

16.23

1 6.24 16.25 16.26 16.27 1 6.28 1 6.29 16.30

201

D A Powers "Sustained Molten SteeVConcrete lnteractions Tests". SAND77- 1479C. D A Powers "Influence of Gas Generation on Melt/Concrete Interactions" SAND78-0939C. W B Murfm "Report of the Zion/lndian Point Study: Volume I" NUREG/CR-2247. D A Powers et aI, "Ex-Vessel Core Debris Interactions". NUREG/CR-2679 91 of 40, SAND82-0904 (1 of 4). M Pilch et al "Acceleration Induced Fragmentation of Liquid Drops" NUREG/CR2247. G M Hidy and J R Brook "The Dynamics of Aerocoloidal Systems" Volume 1 , Pergamon Press, 1 970. F D Richardson "Physical Chemistry of Melts in Metallurgy" Volume 2, Academic Press, 1 974. C N Amos ( 1984) "Evaluation of Contact Time for Pool Scrubbing". Paper to the ANS Topical Meeting on Fission Product Behaviour and Source Term Research, Snowbird Utah, 1 5 - 1 9 July, 1 984. M Tomaides and K T Whitby in "Fine Particles", Ed B. Y H Liu, Academic Press, 1 976. R D Fehlke, ed "BOF Steelmaking Theory", Volume 2, AIME Iron and Steel Society, 1 975. A F Ellis and J Glover, J Iron and Steel Inst., 197 1 . J Szekely an d N J Thermelis, "Rate Phenomena in Process Metallurgy", Wiley Interscience, 1 97 1. L Anclrussow, Z Elecktrochem, 54, 567, 1950. G Eriksson, Chern. Scr 8 ( 1 975) 100. 0 H Krikorian, High Temp - High Press., 14 ( 1982) 4, p387. D D Jackson, UCRL-5 1 1 37 ( 1 97 1). I Kataoka, M Ishii "Mechanistic Modelling of Pool Entrainment Phenomenon", Int. J. Heat Mass Transfer 27 1 1 ( 1984). V K Dhir, J N Castle and I J Catton, Heat Transfer 99 3 ( 1977) 4 1 1 . D A Powers, Appendix E ofT S Kress et al "Review of the Status of Validation of the Computer Codes used in the NRC Accident Source Term Reassessment Study (BMI 2104)" ORNL/fM-8842 ( 1985). M Lee, M S Kazimi and G Brown "A Heat Transfer Model for the Corium/Concrete Interface" Proc. Int. Meet. LWR Severe Accident Evaluation., Cambridge, Mass ( 1983). R K Cole, D P Kelly, M A Ellis "CORCON MOD2: A Computer Program for Analysis of Molten-Core Concrete Interactions" NUREG/CR-3920 (1984). G K Batchelor "An Introduction to Fluid Dynamics" CUP ( 1970). Levich "Physiochemical Hydrodynamics" Prentice-Hall (1962). D J Woodford, A H Scriven ''The Rise Velocities of Bubbles and the Distribution of Vapour in a Liquid Pool during Depressurisation" TPRD/L/241 3/N82. G G Bartolomei, M S Alkhutov, Thermal Engineering 14 ( 12) ( 1967) 1 12- 1 14. P H Calderbank, M B Moo-Young Chern. Eng. Sci. 16 39 (196 1 ). E G Scheibel, Ind. Eng. Chern., 46, 2007 ( 1954).

202

Aerosol Production/rom Core-Concrete Interactions

16.3 1 D A Powers, J E Brockmann, A W Shiver "V ANESA: A mechanistic Model of Radionuclide Release and Aerosol Generation During Core Debris Interactions with Concrete" NUREG/CR-4308, SAND85- 1 370. 16.32 D Tabor "Gases, Liquids and Solids" Penguin (1969). 1 6.33 R Koenig, J C Brink, Appl. Sci. Res. Sect. A, 8 142 (1 950). 1 6.34 C R Wilke, C Y Lee , Ind. Eng. Chern., 47 1253 ( 1955). 16.35 E R Gilliland, Ind. Eng. Chern., 26 68 1 ( 1934). 1 6.36 T Ginsberg "Aerosol Generation from Sparging of Molten Pools of Corium by Gases Released from Core-Concrete Interactions", Proceedings of the International Meeting on Light Water Reactor Severe Accident Evaluation, Cambridge, Massachusetts, August 28 to September 1 , 1 983.

203

Chapter 17 CONTAINMENT THERMAL-HYDRAULICS

17.1

Introduction

The aim of this chapter is to describe some of the main models that have been developed for predicting thennal-hydraulics related phenomena in the containment. The containment is a critical safety feature of modern LWRs and in the event of a severe accident serves as a final barrier for fission product release to the environment. The potential forrelease depends on various factors, the containment temperature and pressure loading, the fission product release to the containment from the reactor coolant system and the subsequent behaviour of fission products within the containment. Fission product related issues are discussed in Chapter 1 8 . The CONTAIN [ 1 7 . 1 ] code developed b y th e USNRC is o ne o f the most important and significant mechanistic containment analytical tools currently in use. It is a best estimate code and aims to predict the complete thennal, physical, chemical and radiological response of the containment and its surroundings under both design basis and severe accident conditions. The scope of the code includes a capability to calculate the load on the containment in accidents involving significant loss of primary circuit/water inventory to the containment e.g. large leak LOCA and also the wide ranging and complex phenomena existing in the event of melt release to the containment The major processes (excluding fission product processes) that need to be modelled include: Fundamental Thennal-Hydraulics BWR Specific Processes Gas Burning Processes Energy and Mass Transfer Heat Conduction in Structures and Engineered Safety Features (ESFs). There are various feed back mechanisms between these processes as indicated in Table 1 7. 1 .

204 17.2

Containment Thermal-Hydraulics Major Phenomena Affecting Thermal Hydraulics

It is clear from the earlier discussion that thermal hydraulics models are crucial in predicting many facets of the containment function. The major phenomena include: gas flow within and between containment compartments atmospheric temperature and pressure response coolant pool thermal response and boiling phenomena relating to injection of sources from the primary circuit convection and radiation heat transfer mass transfer relating to condensation and evaporation structure heat conduction burning of inflammable gases.

TABLE 17. 1 FEEDBACK MECHANISMS IN THE CONTAINMENT

Thermal Hydraulics • • • •

gas and liquid flow heat transfer thermodynamics engineered safety features

Deposition/Agglomeration

Aerosols • • •

particle size distribution material composition deposition

Fission Products • • •

radioisotope inventory decay and heating release and acceptance

Distribution of Fission Products

Intercell Transport Evaporation/Coolant Inventory Heat to gas, walls, pool Transport of gas/fission products

17.3

Flow Between Compartments

Containments are composed of various compartments, see Figure 17. 1 . Lumped parameter computer models take advantage of this feature by employing computational volumes or cells which correspond to some fraction of a compartment or to one or more compartments.

Containment Thermal-Hydraulics

205

Codes such as CONTAIN have the capability to utilise arbitrary arrangements of cells and flow path connections. In the CONTAIN code for example two types of flow path connections are available: single flow paths between specified cells engineered vent flow paths (these are described in more detail below). Single flow paths connecting separated cells represent physical pipes, ducts or possibly a cable room. Flow can occur though the common connecting area of two adjacent cells. The code also includes the capability of allowing the flow areas to be time or pressure dependent. The engineered vent flow path model allows two or more parallel flow paths to occur between the containment volumes. The latter might represent containment room s with multiple connecting cable riser or horizontal ducts. The latter option allows a more detailed flow configuration to be modelled.

17.4

Fundamental Thermal-Hydraulics

Lumped parameter system codes suffer from the restriction that momentum conservation is not treated within a control volume. The basic assumption in such codes is that the control volumes are joined by simple flow paths. The flow takes account of inertial effects, the frictional flow resistance in the path and gravity forces. The basic inertial flow model in the CONTAIN code takes the fonn: dw/dt where &p Cf� P

A L

= pressure difference, including the gravitational head = irreversible loss coefficient = gas flow density = flow path area = inertial length of path.

The irreversible loss coefficient above is assumed to account for both discharge and flow loss coefficients. Discharge coefficients used by containment codes are given by [ 1 7.2] and loss coefficients are given in [ 17.3] and [17.4]. The CONTAIN code includes an option to neglect flowing gas inertia or acceleration in order to improve calculational speed. Laminar flow is also allowed for with this option. Models available in the code are shown in Table 17.2. For high mass flow rates when choked or critical gaseous flow occurs, the mass flow rate becomes independent of the pressure gradient across the cell connecting flow paths or through the orifice. Critical flow under these circumstances has been modelled as a vena contracta [17.6].

206

Containment Thermal-Hydraulics

U PP E R DOME FLOWPATH TH RO U G H COM MON AREA

NODALISATI ON FOR CONTAI N - -- - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - -

S I M U LATI ON

..... - •. - - 1 :: : 1

12

FLOWPATH TH ROU G H DUCT/PI PE 10

16

I:::::J Reinforced concrete

CAVITY

FIGURE 17.1 PWR CONTAINMENT BUILDING REDUCED TO 18 INTERCONNECTED CELLS/COMPARTMENTS

There are a number of processes involving mass transfer between the phases that need to be modelled. Sources and sinks to the basic mass conservation equation include the amounts of combustible gases consumed or burnt, the amount of mass removed or added by various ESF systems, the amount of mass condensed on or evaporated from the structures, the amount of mass removed or added to from the aerosol field and the amount of mass added through pool boiling. These processes also give rise to corresponding energy transfer. Conservation of energy is achieved by performing an energy balance on the amount of heat entering and leaving a cell taking account of intra cell sources such as chemical or decay heating and sinks such as condensation or the ESFs.

Containment Thermal-Hydraulics 17.5

207

BWR Specific Processes

Special models have been developed for the clearing ofB WR suppression pool vents. In such models, liquid levels in the wet well and dry well are computed if the poo l inventory or density changes. Quasi-steady gas flow models have been developed to detennine the suppression pool gas flow once the suppression pool vent has cleared. The energy and mass exchanges between the vented gas and the poo l can be calculated assuming that the gas and pool temperatures are at equilibrium and the gas is saturated. Models for the vent clearing time have been developed based on the time required for the liquid level on either side of the suppression pool to drop to a level at which the vent begins to clear. The model in the CONTAIN code is based on a Bernoulli energy balance to determine the liquid level velocity and applies to all the Mark I, II and III BWR containments. It should be noted that to model the Mark III containment correctly, multiple vents need to be represented since the drywell contains a weir wall. The pressure drop in the flow from the drywell to the wetwell can be modelled by an expression of the form: �p = 1(2 KpV2 where K is a turbulent loss factor [ 1 7.3] . K is well tabulated for a sudden expansion. In some models, additional losses have been introduced which are dependent on whether the flow is from the wetwell to the drywell or vice versa. Thus the vent clearing time depends on the containment pressures and the initial water level on the dry well side of the suppression pool . It can also be affected by the overflow from the suppression pool in a Mark III containment when the wetwell pressure is higher than the drywell pressure. Suppression pool gas flow can be calculated using a quasi -steady flow assumption introduced earlier. The effective flow area can be taken to open linearly with pressure. To account for the mass and energy transferred during vent flow, it should be noted that there are three main stages of transfer for the suppression poo l: the enthalpy added to the pool due to the flow of non-condensable gas and condensation of vapour flashing because of vent flow vapour taken up by non-condensable gas of the vent flow.

208

Containment Thermal-Hydraulics TABLE 17.2 INTERCOMPARTMENTAL FLOW MODELS

Cell 1 --------------------- - ----------------------------- Cell 2

17.6

(a)

Inertial Flow

(b)

Quasi-Steady Flow

(c)

Critical Flow

Material Properties

For the computation or modelling of containment systems, physical properties for the order of fifty standard materials are required. These include temperature-dependent specific enthalpy, specific heal, density, viscosity and thermal conducti vity for a variety of materials found in containments under severe accident conditions. A table of typical materials is given in [ 1 7.7] . Ice/water phase changes are required for modelling ice melting in ice condenser Engineered Safety Feature (ESF) models. Steam/water phase changes are needed for the condensation of steam on to structures and aerosols, poo l boiling and evaporation of films of water. These can be obtained from standard references e.g. the enthalpy function for water vapour can be obtained in [1 7.7] . For vapour saturation conditions various thermodynamic quantities e.g. vapour saturation pressure can also be obtained from [ 1 7.7] . Bulk thermodynamic properties are required if more than one material components are present. The weighted average property is defined in the usual way but the weights may be different depending on the property in question. In the CONTAIN code for example the bulk gas thermal conductivity is weighted by the number of moles of each component Other weighting conventions are also given in Reference [ 1 7. 1 ] .

17.7

Gas Burning Models

If hydrogen or carbon monoxide gases accumulate in sufficient proportions then they may cause deflagrations to occur in the containment. The conditions for deflagration have been measured experimentally. Combustion models were developed for the HECTR code [ 1 7.8] and the basis of this model has been incorporated into the CONTAIN code. The chemical reactions that occur during a hydrogen or carbon monoxide burn are: 2� + 02 = 2H20 + Qm 2CO + 02 = 2C02 + Qco where if the Q ' s are positive then the reactions are exothermic. Typical release energies are given in [ 17. 1].

Containment Thermal-Hydraulics

209

Models have been developed based on a minimum threshold principle i.e. when a minimum concentration of the combustible gas and oxygen are reached. Burns in codes like CONTAIN are allowed to propagate from neighbouring cells. The burn typically continues for a time based on an internally calculated flame speed described later or this may be user prescribed. It should be noted that many of the models are derived from experiments which have been perfonned with ignition sources. They may not be entirely typical since in a reactor accident higher concentrations may exist prior to the ignition. For mixtures containing more than one combustible gas, the critical concentration can be related to the concentration required for a mixture with a single combustible gas via Le Chatelier's principle. For typical accident conditions a burn is initiated spontaneously provided that the mixed mean combustible mole fraction is >7%, the oxygen mole fraction is >5% and the sum of carbon dioxide and water vapour mole fractions is <55%. For the front to propagate these concentration criteria must apply to the neighbouring cell into which the deflagration front propagates. Deflagration models have been developed which assume that the flames travel a characteristic distance in the room along a flame wave front travelling at some speed. The burn time can then be calculated in the obvious way. Flame speed has been calculated from experimentally derived correlations. Both CO and � are accounted for using Le Chatilier's fonn ula in the flame speed correlation [ 1 7 . 1]. The basic approach in detennining a burn rate is to estimate the number of moles remaining to bum over the rest of the bum time.

17.8

Energy and Mass Transfer

A key requirement of containment modelling is an ability to calculate heat and mass transfer between the atmosphere and the structures. Material layers affecting heat transfer are shown in Table 1 7.3. Specific additional requirements in LWRs may be a need to calculate heat transfer to ice condensers, from coolers and to poo ls of water. The convective heat flux can be modelled by a standard heat transfer coefficient or Nusselt number correlation. A number of standard correlations are available in codes such as CONTAIN depending on whether conditions are forced or natural convection and/or laminar or turbulent. For radiation heat transfer, various models have been used including: net enclosure radiation model, simple radiation models (of the kind described below).

210

Containment Thermal-Hydraulics

The net model allows for simultaneous radiation transfer between all the heat structures and the atmosphere. A formulation for grey surfaces is given in [ 1 7.9] . The basic assumption is that the atmosphere is completely surrounded by heat transfer surfaces. These may be at different temperatures and the model accounts for secondary surface to surface reflections. Simpler models have been introduced in codes such as CONTAIN which only treat the direct radiative heat exchange between the atmosphere and the heat structures. All radiation models require optical properties for the atmosphere/gas mixture (1\0 vapour, CO2, CO and also aerosols). The Modak model [ 17. 10] calculates the total gas mixture em issivity and absorptivity taking into account the opaque gases above. The emissivity of individual opaque gases can be obtained by summing over spectral absorption bands [ 17. 1 1], although containment analysis codes such as CONTAIN have used an approximate method. The gas mixture emissivity is computed by summing over the spectral emissivities together with an overlap correction [ 1 7. 12] but modified to include temperature variation [ 1 7. 1 1 ] . The emissivity of particles has been calculated b y Felske an d Tier [ 1 7. 1 3] with the assumption that the particles are so small as to produce negligible scattering. The aerosol cloud absorption has been derived by Pilat and Ensor [ 1 7. 14] . A second radiation model available in the CONTAIN code is the Cess-Lian model [17. 1 5] . This model is based on an analytic approach and is applicable to calculating steam/air emissivity and absorptivity. In this model data from Hottel emittance charts [17.16] are fitted to a single band exponential form which requires less computing time than a wide band model. TABLE 17 .3 MATERIAL LAYERS AFFECTING HEAT TRANSFER

Bulk Atmosphere (including non-condensables) Gas Film Boundary Layer* Water Condensate Film Layer Paint Layer Structure *Heat Transfer Models

(a) Condensation (b) Convection (c) Radiation

Containment Thermal-Hydraulics

21 1

Mass transfer governing condensation and evaporation of vapour on to heat structure surfaces depends mainly on the partial pressure difference of the condensable vapour. If the condensate layer exists on a structural surface, the partial pressure of the non-condensable gas will be larger at this film layer than in the bulk gas region. As vapour moves towards the surface where it accumulates it drags the non-condensable gas vapour with it. This ultimately leads to a force to drive the non-condensable gas away from the surface and oppose the motion of the bulk mixture to the surface. For condensation the vapour partial pressure is lower at the surface than in the bulk mixture for evaporation the vapour pressure is higher than in the bulk mixture. Molar fluxes for non-condensable gases and condensable vapour have been described in [17. 1 7] in terms of the drift flux velocity of the mixture toward the interface. Interface conditions need to be known but for containment applications the condensate film resistance is expected to be quite small relative to other resistances in series with it. In the early mechanistic codes this film resistance was neglected. However, in the later more advanced models the temperature gradient across the film is modelled. The interface temperature can be determined by equating the heat flux via conduction through the film with the mass and convection heat transfer on the gas side of the interface. 17.9

Heat Conduction in the Structures

The pressure and temperature loadings on the containment are dependent on the various heat sources to the containment and removal of heat to the structures. As described in the previous Section models have been developed to account for convection, radiation and condensation (and evaporation) between the atmosphere and the structures. Heat conduction within each structure is usually accounted for by treating the structures as idealised slabs, cylinders or spheres within an appropriate co-ordinate system and with appropriate boundary conditions. Methods have been described earlier in this book . 17.10 Engineered Safety Features

Specific models have been developed in containment codes for various engineered safety systems including containment sprays, ice condensers and fan coolers, Figure 1 7.2. Most LWRs include containment sprays as a safety feature. Typical sources of spray water include the refuelling water storage tank (RWST) in PWRs or a coolant storage tank in BWRs. Sprayed water is collected in a sump at the bottom of the containment and when water from the sources above is exhausted, water is pumped up from the sump through a heat exchanger. It then passes to the spray nozzles. Sprays are effective because heat transfer to and condensation of steam on to droplets provides an effective mechanism for the reduction of temperature and pressure and also for removing fission products from the atmosphere (this is addressed in the next section).

212

Containment Thermal-Hydraulics

H OTS I D E

� OLD L E G

TO � CONTAI N M E NT ATMO SPH E R E

__

�

C O L D SI D E COLD L E G (CO O L I N G WAT E R)

LOW D I R E CTION

�

H OT S I D E H OT LEG

FROM CO NTAI N M E NT ATM O SP H E R E DRIVEN BY E LECTRIC M OTO R

'---

O UTLET M AN I F O LD

COLD S I D E ( H OT L E G WAR M E D COO LI N G WAT E R)

(a) FAN COO L E R H O RIZONTAL STRAPS

ICE BASKETS H E LD IN P E R F O RATE D CYLI N D E RS

(b) I C E COM PARTM ENT

CONTAI N M E NT SPRAY / \/ \/ \/ \/ \

(c) CONTAI N M E NT S PRAYS FIGURE 17.2 ENGINEERED SAFETY FEATURES

Containment Thermal-Hydraulics

213

Models have been developed which enable heat transfer between the droplets and the atmosphere to be calculated and the condensation of steam on to or evaporation from the droplets. As water droplets fall through the atmosphere their diameters can change as a result of condensation or evaporation. Heat and mass transfer can be modelled as described before for heat structures. However for convective heat transfer around a spray droplet a Nusselt number correlation for a sphere can be used [ 1 7. 18]. Mass transfer is determined by the diffusion of water vapour through the gas boundary layer at the drop surface and is driven by the water vapour pressure difference between the droplet surface and the atmosphere. As spray droplets fall some evaporate but others reach the bottom and deposit their energy. When sprays are on, the containment conditions tend to approach a quasi-steady state where the balance is between the steam and heat being removed by sprays from the containment and the various sources of steam and heat to the containment Additionally the droplet itself tends to reach a quasi-steady state very early during its fall. The exceptions to this rule occur following major transients. Ice condensers were introduced by Westinghouse to suppress the temperature and pressure rise within a containment arising from a severe LOCA. Ice condenser melting models have been treated in a similar manner as for cylindrical wall structures. Models have been produced for both turbulent natural convection and forced convection. Fan coolers are normally provided in containments to provide cooling under other than emergency conditions and to augment steam removal from water sprays by providing enhanced mixing. The air is typically circulated horizontally by an electric motor across pipes carrying water throughout. Steam is condensed and falls downwards beneath the horizontal pipes. Simple models have been developed e.g. as in the MARCH code. In this model a correlation for the effective heat transfer coefficient is used as a function of steam vapour fraction and derived from saturated conditions considered in the Oconee safety analysis report. Inputs have also been taken from fan coolers in the Zion nuclear power plant [ 1 7. 1 9] . This model does not address condensation and therefore does not remove water from the system. More mechanistic models have been produced based upon condensation mass transfer and convective heat transfer. An available Nusselt number is a Reynolds-Prandtl correlation for flow over horizontal tubes [ 1 7.20] .

17.11 Summary

The chapter has briefly referenced some of the physical situations and associated models that have been developed in state of the art mechanistic containment codes. The models described include those available for modelling mass and energy transfer in the containment and

214

Containment Thermal-Hydraulics

between comparunents, the exchange of energy between the atmosphere and the heat structures, the deflagration of flammable gases and some of the approaches that are adopted for modelling the ESF functions. REFERENCES

17.1 17.2 1 7.3 17.4 1 7.5

1 7.6 17.7

1 7.8

KE Washington etal "Reference Manual for the CONTAIN 1.1 Code for Containment Severe Accident Analysis" NUREG/CR-57 1 5 , July 1 99 1 . "Flow of Fluids through Valves, Fittings and Pipes" Technical Paper No 4 10, Crane Co, New York, NY, 1 979. R B Bird, W E Stewart and E N Lightfoot "Transport Phenomena" John Wiley and Sons, New York, NY, 1 960. W M Kays and A L London "Compact Heat Exchangers" 2nd Edition, McGraw-Hill Book Co, Inc, New York , NY 1 964. K K Murata et al "User' s Manual for CONTAIN 1 . 1 , A Computer Code for Severe Nuclear Reactor Accident Containment Analysis" NUREG/CR-5026, SAND-2309, Sandia National Laboratories, Albuquerque, NM 1 989. H Lamb "Hydrodynamics" Dover Pub, New York, NY, 1 945. W C Reynolds "Thennodynamic Properties in SI, Graphs Tables and Computational Equations for 40 Substances" Deparunent of Mechanical Engineering, Stanford University, 1 979. S E Dingman et al "HECTR Version 1 .5 User ' s Manual" NUREG/CR-4507, SAND86-0 101 , Sandia National Laboratories, Albuquerque, NM 1 986. R Siegel and J R Howell ''Thennal Radiation Heat Transfer" 2nd Edition, McGraw­ Hill Co, Inc, New York, NY 198 1 . A T Modak "Radiation from Products of Combustion", Fire Research, Vol 1 , pp 339361 , 1 979. D K Edwards and R Matavosian "Scaling Rates for Total Absorptivity and Emissivity of Gases", Transactions, ASME Joumal ofHeat Transfer, Vol l 06, No 4, pp 684-689, 1984. E E Lewis "Nuclear Power Reactor Safety" John Wiley and Sons, Inc, NY 1 977. J D Felske and C L Tien "Calculation of the Emissivity of Luminous Flames" Combustion Science and Technology, Vol 7, pp 25-3 1 , 1 973. M J Pilat and D S Ensor "Plume Opacity and Particulate Mass Concentration" Atmospheric Environment, Vol 4, pp 163- 173, 1970. R D Cess and M S Lian "A Simple Parameterization for the Water Vapour Emissivity" Transaction, ASME Journal of Heat Transfer, Vol 98, No 4, pp 676-678, 1 976. H C HOllell and A F Sarofim "Radiative Transfer, McGraw-Hill Book Co, Inc, New York, NY, 1967. J G Collier "Convective Boiling and Condensation" 2nd ed, McGraw-Hill Book Co (UK) Limited, Great Britain, 198 1 . W E Ranz and W R Marshall "Evaporation from Drops" Chemical Engineering Progress, Vol 48, No 3, April 1952. ''Zion Station, Final Safety Analysis Report, Vol VI" Commonwealth Edison Company, US Atomic Energy Commission Docket 50295- 1 8. J P Holman "Heat Transfer" McGraw-Hill Book Co, New York, NY, 1 986. ,

1 7 .9

,

1 7 . 10 1 7. 1 1

17. 1 2 17. 1 3 1 7 . 14 17.15 17.16 17.17 17. 1 8 1 7. 19 1 7.20

215

Chapter 18 CONTAINMENT AEROSOL AND FIS SION PRODUCT MODELS 18.1

Introduction

The previous chapter was concerned with the prediction of pressures, temperatures and material distribution in the containment from the point of view of predicting loads and associated threats to containment integrity. This chapter is concerned with models for aerosol and fission product behaviour and transport which provide the input for calculating any fission product release to the environment in the event of containment failure. Various mechanistic models have been developed for the representation of the behaviour of aerosols, e.g. the aerosol model in the advanced containment code CONTAIN [ 1 8. 1 ] is derived from the MAEROS [18.2] code and allows for multi-sectional and multicomponent modelling of aerosols. In the former case the particle size distribution is allowed to have an arbitrary shape. In the latter case, particles consisting of different materials can be modelled. Aerosols may be generated from various events occurring at different times in an LWR accident. Fission products and core materials may be expelled from the reactor coolant circuit during the early phase. Later aerosols may be generated by core-concrete interactions, pool boiling and resuspension processes also leading to fission product release. Mechanisms affecting fission product behaviour are summarised in Table 1 8. 1 . In codes such as CONTAIN water condensation on to and evaporation from aerosols is modelled in a similar manner as the aunosphere thermal-hydraulics calculation for heat transfer to structures, i.e. the latent heat associated with the coolant mass transfer between the annosphere and the aerosol surfaces is incorporated in the total internal energy to and from the aunosphere. In addition evaporation and condensation of coolant mechanisms for the aerosols and heat structures can occur at the same time as condensation or evaporation from heat sinks and agglomeration and deposition of aerosols on surfaces. The main aerosol quantities of interest are the mass and chemical composition distributions of aerosol particles throughout the containment There are four general processes that determine the aerosol particle size distribution: agglomeration - two particles collide and form one larger particle, size change caused by condensation and evaporation of water on to or from the particle,

216

Containment Aerosol and Fission Product Models

aerosol sources, particle deposition on to surfaces. 18.2

Size Distribution

In advanced containment codes, the particle size distribution is discretised into particle size classes or sections. Within a particular section the distribution of aerosol may be assumed to be constant with respect to the logarithm of the particle mass. In general this constant will vary between sections and may also be time dependent The particle size distribution is unrestricted in some models e.g. in the CONTAIN code. Different particle size distributions may be specified for each aerosol component. Particle size distributions may be taken as log - nonnal. Aerosol concentrations for each section may be obtained by integrating individual concentrations for each component. Section boundaries may be defined by partitioning the particle size domain geometrically. Particles are also assumed to have spherical geometry. Provision has to be made to allow for the situation when aerosol particles either grow smaller through evaporation ofwater or larger due to condensation ofwater vapour or by condensation. Mass that leaves the computational domain is typically assumed to be deposited on floor structures. These aerosols are described as undersized or oversized in the models as in [ 1 8.3] . Particle agglomeration, deposition, condensation and evaporation are all processes which change the aerosol size distributions. Current models take account of sources from in-vessel processes, aerosols from core-concrete interactions, water vapour from pool boiling, aerosol sources from adjacent cells and the effects of changes in atmosphere temperature and pressure. A dynamic model is needed to calculate new aerosol mass and composition infonnation. Aerosol size distribution is governed by a complex integro-differential equation. Numerical fonns has been developed by Gelbard and Seinfield [ 1 8.4] . In their method the range of aerosol masses is divided into an arbitrary number of sections and the method enables the mass of a given component in a particular section to be calculated. Hence the evolution of the aerosol size and composition distributions can be calculated for each computational cell.

18.3

Agglomeration

Agglomeration or coagulation can occur if two aerosol particles collide. Agglomeration processes include: Brownian diffusion Differential Gravitational Settling

Containment Aerosol and Fission Product Models

217

Turbulent Shear Agglomeration Turbulent Inertial Agglomeration.

TABLE 18.1 MECHANISMS AFFECTING FISSION PRODUCT BEHAVIOUR

Fission Product Decay Chains Fission Product Decay Heating Transport between Hosts e.g. Gas, Aerosols, Heat Structures, Pools etc Liquid Transport

The dependence of the agglomeration coefficients upon aerosol properties is given in [ 1 8.5] . Some dependency on the pressure, temperature and viscosity of the containment atmosphere is known. These however are not the main source of uncertainty in aerosol calculations. In physical reality, aerosol particles may not be spherical and also the effective aerosol densities may be less than the bulk density of the materials of which the aerosols are composed. To accommodate this observation, agglomeration and dynamic shape factors are introduced. Dense aerosols of spherical shape will have factors close to unity. Given a description of aerosol shapes and densities, these factors could be calculated theoretically. However in practice empirical approaches tend to be used, fitting code calculations to appropriate experiments. Turbulence in the containment atmosphere can also lead to enhanced agglomeration. This is only recently being modelled. Collisions between falling aerosols have been studied both theoretically and experimentally. Collision coefficients of unity correspond to the case when the collision cross sections correspond to the geometric cross sections. However, hydrodynamic interactions can lead to collision efficiencies rather less than unity. This topic has been studied by Dunbar and Ramsdale [ 1 8.6]. Turbulent agglomeration collision efficiencies have been rather less studied than gravitational collision efficiencies. The effects of collision efficiencies, aerosol shape factors and turbulence are coupled and dependence upon the various parameters depends on the agglomeration mechanism in question. However in general all agglomeration processes are increased by increases in agglomeration shape factor: the dependence is greatest for turbulent shear agglomeration and smallest for Brownian agglomeration. Large values of dynamic shape factor have the effect of reducing the coefficients, apart from the turbulent shear case which is not affected.

218

Containment Aerosol a nd Fission Product Models

Models as included in the CONTAIN code have the level of modelling described above. The code takes the temperature, pressure, and mass flow rate information from its atmosphere model, to compute the agglomeration coefficients and input the information to the aerosol dynamics model.

18.4

Aerosol Condensation and Evaporation

This section is concerned with the condensation and evaporation of water to and from the surfaces of particles. Models for the rate of condensation on to a dense spherical particle, which take account of the diffusi vity of water vapour in air and the conduction of the heat of condensation have been developed. Rate equations for the diffusion of water vapour to and from the aerosol are given in reference [ 1 8.7] . The core of these spheres may be composed of solid materials. The condensation rate can be affected by the number of nucleation sites that are available. Rates can decrease if particles grow. Aerosol particles which act as seed nuclei are available in codes such as CONTAIN. Condensation and evaporation rates are affected by the degree of superheat in the atmosphere and are driven by the steam mass available in the atmosphere. Given the steam mass the

condensation and evaporation rates may be calculated using the Mason equation, see for example [ 18.7] .

18.5

Deposition

There are various mechanisms through which containment aerosols can settle or plate out on to surfaces or structures. These include: Gravitational settling Diffusion to surfaces Thermophoresis Diffusiophoresis Inertial compaction (relevant to ESF operation). In many cases gravitational settling is the most important mechanism. A correlation for the downward settling terminal velocity is given in [ 18 . 1 ] . Aerosol depletion can occur through the diffusion o f aerosols down through a concentration gradient. This mechanism can become important for smaller aerosol sizes [ 1 8 . 1 ] .

Containment Aerosol and Fission Product Models

219

In thennophoresis the force to produce the aerosol depletion i s generated b y the temperature gradient [ 1 8. 1 ] . Diffusiophoresis occurs i n circumstances where aerosols are present i n a gas mixture: a force results due to concentration gradients of the gas components. An important example of diffusiophoresis occurs during the condensation of vapour on to a surface or the evaporation of water from the surface. If the water vapour is denser than the air then the water vapour molecules will exert greater momentum on to the aerosols and force the aerosols in the direction of diffusion of the water vapour molecules. The aerosol deposition velocity for water vapour will depend on the diffusiophoresis forces and a competing aerodynamic (Stefan) flow near the condensing/evaporating surface. Given these tenninal velocities the aerosol renewal rate may be defined.

18.6

Aerosol Sources

Aerosols are produced from a number of sources e.g. core-concrete interactions: most of these were described earlier in Chapter 16. Aerosols are removed in other ways e.g. by ESFs: these mechanisms are described later in this chapter. These sources and sinks need to be added to the aerosol inventory at the time of production/removal and will correspond to different materials and size groups. In some circumstances e.g. in BWRs, the aerosols may pass through a suppression pool where they would be scrubbed, if for example the vents are under water. These processes are treated in codes such as CONTAIN and these codes in some cases allow aerosol sources/sinks to be explicitly stated by the user.

18.7

Scrubbing

Aerosol scrubbing models have been developed for passing aerosols through suppression pool vents. Examples are provided for in the SPARC code [ 18.8] and the VANESA code [ 1 8.9]. The aim of these models is to provide decontamination factors as the gas bubbles rise up in the pool . One such model takes account of steam condensation in the bubbles at the inlet, after assuming initially saturated bubbles. The model is based on Fuchs ' [18. 1 0] treatment of sedimentation, impaction and diffusion in spherical bubbles. In addition, the efficiency of scrubbing depends on the ratio of the gas circulation velocity to the bubble rise velocity. Elliptical bubbles providing enhanced scrubbing can be modelled by prescribing this ratio greater than unity. The effects of surface impurities which inhibit circulation can be modelled, by prescribing a value of this ratio less than unity.

220

Containment Aerosol and Fission Product Models

The SP ARC code have been developed to provide a more mechanistic approach. It extends Fuch ' s model to elliptic bubbles explicitly, treats deposition arising from initial steam condensation and includes sedimentation, inertial deposition and diffusion in rising bubbles. Bubble growth and deposition Iimi ting effects of vapour evolution are accounted for during bubble rise. The model in the SPARC code also includes particle growth due to condensation, interior heat transfer in the bubble and particle solubility effects. Scrubbing models require as input the depth for scrubbing. The above models apply to the scrubbing of aerosols and FPs hosted by these aerosols. The scrubbing ofFPs, not hosted by aerosols, is a separate issue and is not considered here.

TABLE 18.2 AEROSOL FISSION PRODUCT REMOV AL MECHANISMS IN ENGINEERED SAFETY FEATURES

18.8

Uti

Fan Coolers Diffusiophoresis

ill

Ice Condensers Settling Impaction/lnterception Brownian diffusion Diffusiophoresis Thennophoresis

�

Containment Sprays Impaction/lnterception Brownian Diffusion Diffusiophoresis Thennophoresis

Radionuclide Behavior Modelling

The behavior of radionclides or fission products in modelling depends on various factors including the decay chains, decay heating, the criteria for release and acceptance of fission products on hosts and the transport models. Given an initial inventory and location of FPs, codes such as CONTAIN can track FP decay and detennine the inventories of each isotope. There are clearly a large number of decay chains that could be specified but in practice a selected subset can be chosen. FPs are typically associated with various hosts, such as gas, aerosols, heat structures or liquid layers or pool s. Once attached to mobile hosts, the transport of the FPs then takes place: examples of available models are given below.

Containment Aerosol and Fission Product Models

221

These hosts fall into two categories, mobile and stationary. Mobile hosts include gases and aerosols and mechanistic models have been developed. Other mechanistic models are available for the removal of gaseous iodine from the aunosphere and also for condensate films draining from heat transfer structures and for pool to pool liquid transport. The philosophy in most of the models is that the FPs are assumed to have no mass i.e. the dynamics of the hosts are not affected by the presence of FPs. Heating effects though have been considered where appropriate. Other models, usually not mechanistic, assume that FPs can transfer between different hosts. Transfer rates tend to be temperature dependent. 18.9

Decay Chains

To simplify decay chains, techniques have been developed to define linear decay chains [ 18 . 1 1 ] . The basic approach is to decouple the differential equations for decay by breaking the branched decay chains into a system of linear chains. The basic requirements for the model are the decay constants and initial masses. Where a fission product appears in more than one chain, it is necessary to define a branching ratio or the proportion of the disintegration that corresponds to each decay chain. Branching ratios and decay constants are tabulated in [ 1 8. 12]. There results a linear set of first order partial differential equations and therefore as sources of FPs arise, these can be added into the system as sources. Solutions of the system of equations can be obtained using standard Laplace transport theory. 18.10 Decay Heating Models

FP decay heating is a mechanism through which the behavior of fission products is coupled to the thermal-hydraulics. If the decay of each radioisotope is modelled as a decay chain, as described in the previous subsection, then the total heat produced can be calculated given the half-life and decay power of each daughter. Daughter products may be a different element, which may be released or accepted among various hosts. It is therefore reasonable e.g. as in codes such as the CONTAIN code that the heat of decay is deposited on the particular host. Thus decay heat associated with FPs hosted by the gas or aerosols is deposited into the atmosphere: for FPs hosted by structures heat is deposited on the structures. In practice it is not necessary to consider in detail all radioisotopes, and only those of interest for radiological reasons need be treated individually or possibly in small groups. The total power may be modelled using the ANSI/ANS-S . 1 1 979 Standard [ 1 8. 1 3] and any remaining power can be deposited at locations prescribed by the modeller/code user.

222

Containment Aerosol and Fission Product Models

The ANS-5. 1 model takes account of all alpha, beta and gamma emissions as a function of time after the reactor is shut down. For LWR applications the key elements include the thennal fission of235U and 239Pu and the fast fission of238U. This model requires as initial input the time and power before the accident initiation. For application to containment heating modelling, the time the containment modelling is initiated will also be a required input. 18.11 Transfer Rates in the Atmosphere

Models have been developed to enable fission products to transfer between different hosts e.g. FPs may be assumed to transfer at mass rates proportional to the amount of FP mass present For simple coupled equations e.g. transfer between a single pair of hosts, mass redistributions can be calculated from an analytic solution. More complicated coupled equations can be solved using the methods of Lee [ 1 8. 14] , [ 1 8. 1 5] . Typically it is assumed that all chain elements having the same name will transfer at the same rate. In the model in the CONTAIN code source and destination FPs, release rates and host target relations must be prescribed. FP initial locations to cells and aerosol behavior models are then specified. 18.12 Transfer in the Liquid

FPs can transfer across liquid pathways by several mechanisms. Condensate film may drain from structural surfaces, the liquid having been fonned as a result of vapour or aerosol condensation. FPs condensed or deposited into coolant pool s can clearly also be transported with the coolant if there is flow between the pool s. For empirical modelling purposes various fractions of condensate may be assumed to transfer during condensate drain or pool to pool fluid transfer. 18.13 Engineered Safety Features

Diffusiophoresis is an important mechanism for diffusiophoresis across a fan cooler. A model for the deposition rate, which is driven by the differential vapour pressure, is given in [ 1 8 . 1 ] . Condensed materials can then be routed to the destination of the bulk condensate. A number of aerosol removal mechanisms can occur in ice condensers. A model has been developed by Winegardner et al [ 1 8. 15] , [ 18. 16] which includes settling, impaction and interception, Brownian diffusion, diffusiophoresis and thennophoresis. Turbulent deposition is thought to have less important effect.

Models have been developed based on the perfect sink assumption i.e. aerosol particles adhere irreversibly to surfaces if they came into contact with the surface, i.e.

223

Containment Aerosol and Fission Product Models 1 00

To�1

1 0- '

/

.

/

./

/ / /

1 0-2

Collection Efficiencies

1 0- )

1 ()-4

.

I

"g; .::�"'(_J ')'\ /h ": ,:",.r.:.� � '/

/•

/

.

.

•

\

- _ ... · ·

· · ·

...

Thermophoresis

4. ,--. ,

· ·

· · . · · · · · · · · · · · · · · · ·

�

,

1 0- 5 1 0-1

' "

... ....

1 ()-6 Particle Diameter · Microns

FIGURE 18.1 QUALITATIVE COMPARISON OF AEROSOL COLLECTION EFFICIENCIES WITH CONTAINMENT SPRAYS R = dc/dt

=

-kc

where k = removal constant and c = gas concentration. The removal rate for gravitational settling is the product of the settling velocity and the surface area. The former can be related to the particle sizes and the gas properties by means of Stokes law [ 1 8. 1 5] .

224

Containment Aerosol and Fission Product Models

Aerosol and fission product removal mechanisms in Engineered Safety Features are summarised in Table 1 8.2. Impaction and interception of aerosols occurs on the horizontal strips in ice condensers. Removal rates can be modelled via an efficiency parameter. Impaction efficiency has been related to particle and flow parameters via correlations developed for cylinders [ 1 8. 1 5 ] . Interception efficiencies of particles has been related to particle size by Fuchs [18. 15]. Brownian diffusion occurs as a result of momentum exchanges with surrounding gas molecules: it occurs on surfaces because particle capture at the surface reduces the gas phase concentration of the particles on the surface. The total Brownian diffusion removal rate can result from various flow patterns, flows: around ice baskets parallel to surfaces around the ice [ 1 8 . 1 ] . Thermophoretic deposition in ice condensers is modelled in the CONTAIN code [ 1 8. 1] but this is not considered to be an effective particle trapping mechanism. Diffusiophoretic deposition arises due to the net flow of gas towards the ice which then convects the particles towards the surface. 18.14 Containment Sprays

For the purposes of determining the depletion of airbourne fission products by containment sprays, fission products have been categorised into a number of groups [ 1 8 . 17]: noble gases molecular iodine organic iodine compounds aerosols. Spray systems have no effect on the noble gases. The efficiency of removal of iodine depends on the form of iodine: there are several gaseous forms molecular � which is very reactive or methyl iodide C�I which is less reactive. The initial removal of iodine by caesium iodide has been observed by experiment to be a relatively rapid process, until equilibrium is reached between absorption and deposition [ 1 8 . 18].

The rate at which iodine is removed depends on the absorption efficiency and the extent the

Containment Aerosol and Fission Product Models

225

compartment is sprayed. Efficiency has been modelled as the diffusion rate of iodine through the gas and liquid side boundary layer of the droplet's surface. Additives have been added to spray systems e.g. NaOH to improve the extent of iodine removal. The molecular absorption depends on the type of additive. The influence on absorption of a wide variety of spray additives is given in [ 1 8. 19]. Models for the depletion rate for airborne aerosols have been developed based on collection efficiency and the fraction of all the volume swept out by the spray per unit time. Various collection mechanisms for aerosol removal by spray droplets have been modelled including interception, inertial impaction, Brownian diffusion, diffusiophoresis and thennophoresis. A qualitative comparison of aerosol collection efficiencies with containment sprays [ 1 8. 1 ] is shown in Figure 1 8. 1 . The first three effects depend on the particle size, the phoretic effects depend primarily on the temperature and humidity of the atmosphere and on the drop temperature. The collection efficiencies have been tabulated as a function of particle diameter in [ 1 8.20].

REFERENCES

18.1

K E Washington et al. Reference Manual for the CONTAIN 1 . 1 Code for Containment Severe Accident Analysis, NUREG/CR-57 15, July 1 99 1 . 18.2 F Gelbard "MAEROS Users Manual" NUREG/CR- 1 39 1 , SAND80-0822, Sandia National Laboratories, Albuquerque, NM 1982. 18.3 K K Murata et al, "User's Manual for CONTAIN 1 . 1 , A Computer Code for Severe Nuclear Reactor Accident Containment Analysis" NUREG/CR-5026, SAND872309, Sandia National Laboratories, Albuquerque, NM 1 989. 18.4 F Gelbard and J H Seinfield, "Simulation of Multicomponent Aerosol Dynamics" Journal of Colloid and Interface Science, Vol 78, December 1 980. 1 8.5 DC Williams, K D Bergeron, P E Rexroth and J L Tills "Integrated Phenomenological Analysis of Containment Response to Severe Core Damage Accidents" Progress in Nuclear Energy, Vol 19, 1 987. 1 8.6 I H Dunbar and S N Ramsdale "Improvements in the Modelling of Sedimentation and Gravitational Agglomeration" CSNI Specialists' Meeting on Nuclear Aerosols in Reactor Safety, Karlsruhe, Gennany, September 4-6, 1984 as cited in Reference [ 1 8.5]. 1 8.7 H R Byer "Elements of Cloud Physics" University of Chicago Press, Chicago, IL, 1 965. 1 8.8 PC Owczarski and R I Schrenk "Technical Bases and User's Manual for the Prototype of a Suppression Pool Aerosol Removal Code"NUREG/CR-33 1 7, Pacific Northwest Laboratory, Richland, WA, 1985. 1 8.9 D A Powers, J E Brockman, and A W Shiver, "V ANESA: A Mechanistic Model of Radionuclide Release and Aerosol Generation During Core Debris Interactions with Concrete" NUREG/CR-4308, SAND8 5 - 1 370, Sandia National Laboratories, Albuquerque, NM 1 986. 18.10 N A Fuchs "The Mechanics of Aerosols" , Pergamon Press, 1 964. ,

226

Containment Aerosol and Fission Product Models

1 8. 1 1 T R England "CINDER, A One Point Depletion and Fission Program, UW Version" WAPD-TM-334 Bettis Atomic Power Laboratory, Pittsburg, PA, 1 962, revised 1968. 1 8. 1 2 David C Kocher "Radioactive Decay Data Tables" DOEmC- l l026, Technical Information Centre, US Department of Energy, Washington DC, 1 98 1 . 1 8. 1 3 "American National Standard for Decay Heat Power in Light Water Reactors" ANSI/ ANS-5. 1-1979, available from American National Standards Institute, 1430 Broadway, New York, NY 1 00 18, Copyrighted 1 979. 1 8 . 1 4 C E Lee "ASH: An Isotope Transmutation Program Using Matrix Operators" Proceedings, International Conference on Nuclear Waste Transm utation, Austin, TX, 1 980. 1 8 . 1 5 W K Winegardner, A K Postma and M W Jankowski "Studies of Fission Product Scrubbing Within Ice Compartments "NUREG/CR-3248, PNL-469 1 , Pacific Northwest Laboratory, Richland, WA, 1983. 1 8. 16 P C Owczarski et aI, "ICEDF: A Code for Aerosol Particle Capture in Ice Compartments" NUREG/CR-4 1 30, PNL-5379, Pacific Northwest Laboratory, Richland, WA, 1 985. 1 8. 1 7 E E Lewis, Nuclear Power Reactor Safety, John Wiley & Sons, Inc, NY, 1 977. 1 8 . 1 8 R K Hilliard et al "Removal of Iodine and Particles by Sprays in the Containment Systems Experiment" Nuclear Technology, Vol 10, 197 1 . 1 8 . 1 9 D R Grist, "Spray Removal of Fission Products in PWR Containments" SRD-R267, UKAEA Safety and Reliability Directorate, 1 982. 1 8 .20 K D Bergeron, D C Williams, P E Rescroth and J L Tills "Integrated Severe Accident Containment Analysis with the CONTAIN Computer Code" NUREG/CR-4343, SAND85- 1639, Sandia National Laboratories, Albuquerque, NM, December 1985.

227

Chapter 19 THERMOPHYSICAL PROPERTIES

19.1

Introduction

In many accident conditions, the temperature of the fuel may rise, particularly if the core is uncovered, e.g. in the design basis large break LOCA. To describe the core response thennophysical properties of the principal core materials and two-phase water steam mixtures are required. Within the design basis class of accidents, there may be some defonnation of the original geometry, e.g. due to clad ballooning but the extent of core damage will be limited. To sustain severe fuel damage, it is necessary to uncover the core for a prolonged period. In this case, as discussed earlier, it is possible for a layer of molten material mixtures to form. This is therefore a requirement for additional data/models for the prediction of severe accidents. This chapter gives a brief survey of the thermophysical properties data that are available and that are needed for the purposes ofaccident analysis including severe accidents. Thermophysical data for fuel, fuel pin cladding, fuel/clad eutectic material, control rod material, control rod cladding, control rodIclad eutectic material, struc tural materials and core/concrete interaction materials are described.

19.2

Fuel Requirements

To detennine the thermal performance of a fuel rod for a complete range of accident conditions, material properties from operating temperature up to temperatures of the order of the fuel melting point are required. Densities, thennal conductivities, specific heats depend on the temperature and on the physical nature of the fuel in particular the extent of voidage, with some weak dependence on the composition. Depending on the degree of bum up, fission products accumulate in the grains in the fuel but these will only have marginal effects on the thermophysical properties. Heat transfer from the fuel depends on the thennal conduction in the fuel, gap gas and cladding and also on the emission and absorption of radiation in the gas and from the cladding surface. Radiative heat transfer data e.g. emissivities are therefore also required. Mechanical properties of the fuel are needed. Fission gas release from the fuel depends on the chemical state of the fuel e.g. the extent of fuel oxidation but also significantly on the

228

Thermophysical Properties

physical state of the fuel i.e. the degree of cracking. Fuel fracture properties are important in determining the integrity of a fuel bundle following a quench, or following loss of a considerable amount of clad through liquefaction. As fuel heats up it may contact the Zircaloy clad due to thermal expansion of the fuel coupled with possible clad collapse. This can give rise to complex eutectic formation and the lowering of effective fueVclad melting temperatures. However in certain blockage conditions temperatures may approach fuel melting point and therefore data for the melting temperature and heat of fusion for fuel are required. To determine the extent of bulk fuel motion, viscosity data are also required for molten fuel. Melting Point

The principal LWR fuels consist of V02 and Pu02 in varying composition depending on enrichment and burnup. In general, fuel properties may vary somewhat as a function of composition. The melting point of unirradiated fuel has been considered by a number of investigators e.g. Lyons [19. 1 ] . Lyons and Bail y [1 9.2] have determined a phase diagram for stoichiometric (U,Pu)02 mixed oxide fuels and shown the solidus and liquidus temperatures depending on the Pu02 concentration. There is some disagreement concerning the effects of burn-up on the melting point. Krankota [ 19.3] shows a decrease in the melting point of irradiated fuel but Reavis and Green have found no significant reduction in the melting point of V02 due to irradiation. Thermal Conductivity

Investigations have assessed the dependence of thermal conductivity data on the temperature, density, oxygen-to-metal ratio, and plutonium content of the fuel. Two types of experiment have been carried out, those involving a radial heat flow in which the conductivity is directly measured, and transient pulsed methods in which the measured quantity is the thermal diffusivity. V02 thermal conductivity has been estimated by Goldsmith and Douglas [ 19.5] over the temperature range 670-1270K, by Hobson et al (547-2500K) [ 1 9.6] , by Weilbacher (9743027K) [19.7] and others for theoretical V02 densities in the range 95-99%. The results show reasonable agreement. Data also exist for various (U,Pu)02 compositions, see for example Gibby [ 1 9.8] and V02+1I see for example Hobson et al [ 19.6] . Kim et al [19.9] have measured the thermal diffusivity for thin layers o f molten V02 i n the temperature range 3 1 87-33 1 5K. There is a significant data-base for solid fuel: some data exist for the liquid phase. Thermal Capacity

The dependence of the specific heat capacity of fuel on temperature, composition, molten fraction and oxygen-to-metal ratio has been investigated.

Thermophysical Properties

229

Kerrisk and Clifton [19. 10] have reviewed data for mainly stoichiometric V02 over a wide range of temperature 483-3 107K. For mixed oxide fuel, data are given by Gibby [ 19. 1 1 ] and by Leibowitz et al (2350-2730K) [ 1 9. 12] . These data are in good agreement. Leibowitz et al data [ 1 9.3] exist for liquid V02• This reference also gives heats of fusion for unirradiated V02• Leibowitz [19.4] has also detennined the heat of fusion for mixed oxides, obtaining values about 1 0% lowerthan the V02 values. This is reasonable given the similarities in crystal structure between V02 and the mixed oxides. As for thennal conductivity data are sparser in the liquid phase than in the solid phase. Thermal Expansion and Density

Fuel thennal expansion data have been found to depend on the temperature, composition, oxygen-to-metal ratio and the fraction of molten fuel. Data for V02 exist from a number of sources in the temperature range 300-3400K, e.g. Christensen [ 1 9.5] including the phase change. There is some increase in the scatter of the data which increases with temperature. Data from other oxide fuels are only available over a more limited temperature range e.g. for Pu02 in the range 300- 1 700K, Tokar and Nutt [ 1 9. 16] . Comparison between V02, Pu02 and mixed oxide data show very similar values up to 2000K . The density of fuel may be calculated from the theoretical density using reference temperature data and the estimated thennal expansion strains as obtained from data described above. Data are sparse for temperatures near and above the melting point. Emissivity

Data exist for V02 from a variety of sources up to 2400K. Analysis of the data shows that the only significant dependence is on temperature. Data are given by Held and Wilder [ 19. 1 7] for temperatures in the range 450-2600K . Mechanical Properties

There is a substantial amount of data available for fuel properties concerning elasticity, plasticity, potential for swelling during irradiation, morphology and brittleness. Many of the data apply to fuel under nonnal operation and so such data are only relevant in derming the initial conditions existing at the start of the accident Only a brief coverage of the available literature is therefore given here. Concerning elasticity, data exist from a number of sources for Young's modulus over the temperature range 200- 1 700K mainly for stoichiometric V02, e.g. Padel and de Novion [19. 18]. Limited data exist for (U.Pu)02 mixtures also Padel and de Novion [ 19. 1 8] , the mixed oxide values showing slightly larger modulus than the stoichiometric V02 values. Fuel deforms through a number of creep mechanisms depending on the stress, density, temperature, oxygen-to-metal ratio, burnup and grain size. Data have been given by Dienst

230

Thermophysical Properties

[ 19.19]. Data also exist to calculate fuel dimensional changes due to irradiation of U02 and (U'pU)02 fuels during initial and subsequent operation, changes due to exposure to high hydrostatic pressures, and changes due to high temperatures etc. Fuel swells as a result of the production of both solid and gaseous fission products. There are two types of fission gas swelling data. These are isothennal and unrestrained. Typical results are summarised by Collins and Hargreaves [ 19.20]. The morphology of fuel changes as a function of time, temperature, bumup and energy density. This is relevant in tenns of estimating the fission product release from the fuel during the high temperature excursion. There is a plentiful supply of unirradiated data, but few data exist for irradiated fuel especially at high bumup. Data exist for grain growth at high temperature, Ronchi and Sari (22003000K) [ 19.2 1 ] and many others. Ainscough et al [ 19.22] have detennined the effect of bumup on grain growth rates although over a more limited temperature range. The degree of brittleness will influence the extent to which the bundle is likely to remain intact should a substantial quantity of the Zircaloy clad relocate downwards exposing the fuel. U02 begins to change from a brittle to a ductile material at temperatures above l 000 K and this is accompanied by a change in fracture strength. Cannon et al [ 19.23] gives an out-of-pile transition temperature in the range 1 373- 1 723K. Experimental data for the fracture strength in the brittle region are given by Roberts and Ueda [ 19.24] . Cannon et al [ 19.23] gives data for the elastic-plastic region above the transition temperature. Fuel Oxidation

At high temperatures, U02 fuel may act as a large reservoir of oxygen which may be absorbed or released depending on the oxygen potential external to the fuel. Rotwell [ 19.25] reports oxygen release at temperatures greater than 2073K in a reducing environment, whilst analysis of the PBF Severe Fuel Damage Scoping tests by Knipe et al [19.26] suggests possible hydrogen generation by further oxidation of the fuel to hyperstoichimetric conditions. Viscosity

The viscosity of fuel is expected to depend on the temperature, the composition, the melting temperature, the liquid fraction and the oxygen-to-metal ratio. Data applicable to completely molten U02 are given by Woodley [ 19.27] for temperatures in the range 3 1 00-3350K. Fuel Vapour Pressure

There is a substantial data-base for actinide and actinide oxide vapour pressures for the fuel. Data for urania exist over a wide temperature range 2000-7200K. There are many sources of data in the lower range e.g. Benezech et al [ 19.28] and others which show reasonable agreement. Urania is the dominant phase contributing to the fuel vapour pressure. Much of the material referred to in this section is available in the excellent MATPRO data computer library of Hagrman et al [ 19.29].

Thermophysical Properties 19.3

23 1

Fuel Pin Cladding

Requirements

Clad behaviour can be important in determining the consequences of LWR accidents. The clad provides both a physical and chemical barrier to fission product release from the pin. Under severe accident conditions Zircaloy reacts with steam to produce hydrogen and very significant heat The clad and fuel react to form relatively low melting point eutectics. The pin cladding thermophysical properties will therefore change during the course of a severe accident due to these phenomena. At temperatures below 1200K, the clad will consist of mainly Zircaloy with a very thin oxide coating. The heat-up characteristics will depend on the density , thermal conductivity and the specific heat ofZircaloy metal. However. radiative surface emission and absorption data will also be required for the oxide phase due to the thin surface oxide fum. At temperatures in the range 1 ()()()- 12ooK the clad will tend to balloon outwards, e.g. in PWR depressurising accident sequences, due to high internal gas pressure compared with the system pressure. Mechanical properties such as elasticity and plastic deformation creep data will be required to determine the rate and extent of strain. The extent of strain will influence the thermal-hydraulics and flow and also the degree of fueVclad interaction. In the press uri sed sequences, the latter may also be influenced by clad collapse. The material strength properties of the clad depend on the extent of oxidation. Embrittlement is important for determining the extent of clad failure during quench. Even without quench the strength of the oxide layer is important in determining how long any liquid u-o-Zr eutectic may be confined before breach and flow. As the temperatures rise above 13ooK, the Zircaloy/steam reaction becomes very significant and large quantities of hydrogen are produced. Thus data for all stages of clad oxidation are required over a temperature range up to the oxide melting point. Zircaloy melts at about 21 OOK and thermophysical data are also required for Zircaloy in the liquid phase. Similar data are also required for stoichiometric zirconium dioxide with melting point of approximately 2900K and for u-o-Zr phases of intermediate oxygen concentration. Zirconium/Oxygen Phases

The material used for the fuel pin cladding is zirconium alloy e.g. Zircaloy-4, which (e.g. for PWR typically) consists of zirconium alloyed with approximately 1 .9% tin. 0. 1 2% iron, 0. 1 % chromium and 0.05% nickel. Under normal reactor conditions it has a hexagonal close­ packed structure known as alpha-Zircaloy. At higher temperatures up to melting, the stable form is beta-Zircaloy, with a body centred cubic structure. At these high temperatures, Zircaloy will oxidise in steam to zirconia (Zr02), and an oxygen-stabilised alpha-Zircaloy phase may also be formed. At low temperature, the oxide phase has a monoclinic structure, whilst at higher temperatures a tetragonal phase becomes the more stable. Further temperature increase favours the formation of a cubic phase.

232

Thermophysical Properties

A phase diagram for the (Zr-O) system is given by Lustman and Kerze [ 1 9.30]. Solidus and liquidus temperatures have been published by Ruh and Garrett [ 19.3 1 ] , who also examined the Zr02 cubic/tetragonal transition temperature. The effect of the minor alloying components is also known to lower the melting temperature of Zircaloy from that of pure zirconium . Thermal Conductivity

The thennal conductivity of metallic Zircaloy has been investigated by Feith [19.32] up to a temperature of 1 770K. For molten Zircaloy Nazare et al [ 1 9.33] have reported data for a number of Zircaloy-like metals. For Zr02 Gilchrist [ 19.34] has published data for black oxide films fonned on Zircaloy wbing up to 1450K but reports considerable uncertainty due to difficulty measuring the oxide ftIm thickness. No experimental data for molten ZI02 or (Zr-O) are available. Thermal Capacity

The specific heat of alpha-Zircaloy has been measured up to 1 320K by Eldridge and Deem [ 19.35] whose data cover the phase transition to beta-Zircaloy. Experimental data are sparse at higher temperatures. Zirconium oxide specific heat data up to melting at 2973K and for molten Zr02 have been reviewed by Hammer [19.36] . Density and Thermal Expansivity

Data are given by Bunnell et al [ 1 9.37] for the alpha phase up to 1244K. There are few data available for the higher temperature beta phase and for the density change on melting. The density and thennal strain ofZI02 have also been given by Hammer [ 19.36] , who predicts a density increase on transition from monoclinic to tetragonal, and also on melting, due to the removal of porosity. Emissivity

The cladding emissi vity is affected by the presence of oxide layers on the surface. Thin layers, with a thickness of several wavelengths of near infrared radiation, are semi-transparent, and the emissivity is strongly dependent on layer thickness. Thicker opaque layers lead to emissivities dependent on temperature and chemical environment The changes in emissivity during oxidation in steam have been measured by luenke and Sjodahl [ 19.38] in the temperature range 1 125K to 1 979K with oxide thickness up to 129 mm. Mechanical Properties

The mechanical properties of Zircaloy have been extensively investigated in order to

Thermophysical Properties

233

detennine the cladding perfonnance under nonnal operating conditions. Much of this work, for example on annealing and cyclic fatigue is of little relevance to accident analysis. Ballooning or clad collapse is governed by the elastic and plastic properties of the cladding. Bunnell et al [ 19.37] provides Young ' s measurements for Zircaloy tubing up to 1 500K in the axial and circumferential direction for various concentrations of oxygen in the metal. Tube burst tests have provided the most relevant biaxial and multiaxial data on plastic defonnation. Hardy [ 19.39] gives diametrical expansions up to 1400K. Yield points and ultimate strength data were obtained by Chapman et al [ 19.40] from the USNRC sponsored Multigrid Test Programme. These tests were perfonned in steam covering a temperature range from 1 030K to 1444K. The diffusion of oxygen into high temperature beta-Zircaloy may cause embrittlement leading to failure on cooling. The criterion for embrittlement leading to failure depends on whether slow cooling or fast cooling (quench) is being considered. Pawel [ 19.4 1 ] suggests a criterion for fast cooling. The effects of iodine, caesium (present as fission products) and cadmium (present in control rods) on clad embrittlement are considered by Haddad and Cox [ 19.42]. Oxygen Uptake

The highly exothennic oxidation of Zircaloy in steam is one of the more significant phenomena in a severe accident, and much effort has been devoted to a determination of the reaction kinetics and rate constants worldwide. These were described in Chapter 8. Hydrogen Uptake

Under normal operating conditions, the presence of dissolved hydrogen in the Zircaloy cladding may affect its mechanical properties. Hydrogen uptake under these conditions is considered by Millner [ 1 9.43]. The retention at high temperatures has been considered by Futura and Kawasaki [ 19.44] . Viscosity

Correlations for the temperature dependent viscosity of liquid zirconium and zr02 are given by Nazare, Ondracek and Schulz [ 19.33]. Vapour Pressure

The vapour pressure of zirconium metal is given in the range of 2450K to 3850K (the boiling point) by Quill [ 19.45] . As for the fuels data much of this material is contained in the MATPRO library of Hagrman [ 19.29].

234 19.4

Thermophysical Properties Control Rod Material

Requirements

Control rod material in most PWR designs consists of an alloy of sil ver. indiwn and cadmiwn. Under accident conditions as the temperature increases to about l l00K the alloy will melt. and the more volatile components will vaporise. e.g. cadmium. Recent data from Winfrith also show significant indium release. Basic thermophysical data associated with these materials. melting points. vapour pressures etc are therefore required to define the solidi liquid transition. together with data on density. thermal conductivity. specific heats etc for the liquid phase. .•

.•

In contrast for BWR. the melting temperature for B4C is very high 2700K but similar thermophysical data are required. Melting Temperatures

For Ag-In-Cd alloy. the consensus from in-pile and out-of-pile experiments at Oak Ridge [ 19.46]. Winfrith [ 19.47] and elsewhere is that the alloy melts in the range 1073-1 123K. The phase change temperatures appear well characterised. For B 4C. solidus and liquidus. temperatures are given in [19.48] . Thermal Conductivity

Ag-In-Cd thermal conductivities are given by Cohen et al [ 19.49] for the temperature range 323-873K. The data for the liquid phase are sparse though Nazare et aI [ 1 9.50] recommend a 50% reduction on melting by arguing that this is the reduction expected when a face centred cubic solid melts. B 4C thermal conductivities are given by Goldsmith et al [ 19.29] . Thermal Capacity

Molar heat capacities up to beginning of melting at approximately 1 050K are given for the individual elements Ag. In. and Cd by Lynch [ 1 9 .50] . Values for the alloy may be obtained by interpolation. Heat capacities for B4C are given in [ 19.29] . Elemental heats for fusion are given by Lynch [19.5 1 ] for Ag. In and Cd. As for the specific heat, values for the alloy may be obtained by interpolation. Thermal Expansion and Density

Data for the range 300- 1050K are given by Cohen [1 9.49] . Smithells and Brandes [ 19.52]

Therrrwphysical Properties

235

quote a value for the change in volume of Ag. the major component of the alloy. on melting but otherwise data for the molten phase are sparse. A reference density at 300K for Ag-In-Cd is given in [ 19.48] . Densities may be evaluated from the reference density and the thermal expansion data which are available over the required range. A reference density for B4C at 300K can be found in [ 19.29] . Viscosity

Viscosities for the Ag. In and Cd components individually are given by Nazare. Ondraek and Schulz [ 19.50] for temperatures greater than 1 100K. Surface Tension

Engineering estimates for the interfacial surface tension of absorbed material on stainless steel cladding are given in [ 19.29] . Vapour Pressure

Powers [ 1 9.53] gives theoretical data for the vapour pressures over the pure elements. silver (1600- 3000K). indium ( 1500-2900K) and cadmiwn (650- 1 750K). In addition vapour pressures over binary (Ag-In). (Ag-Cd). (In-Cd) and ternary alloys are considered. These data are used to formulate a model for the vaporisation of representative control rod alloys during reactor accidents. A considerable number of experiments have been performed by Bowsher et al [ 19.54] investigating vapour pressures for a variety of alloys. Bowsher's work indicates that steam may have an important impact. for example in the formation of 1,\ which is far more volatile than the metal.

19.5

Control Rod Cladding

Requirements

PWR control rod materials are sheathed in 304 stainless steel clad and hence there is a need for basic thermophysical data for steel and its oxides up to at least the melting point of steel. The timing and nature of control rod rupture are important because of radionuclide material transport on control rod aerosols and also because any formation of eutectics of control rod materials with Zircaloy clad could result in an early removal ofZircaloy from the core region. Eutectic properties are considered below. Material strength properties of stainless steel are required. together with the effects of steam atmosphere on these properties. The stainless steel in the control rod cladding will contribute to the hydrogen produced via oxidation by steam. Oxidation kinetics for stainless steel were considered in Chapter 8.

236

Thermophysical Properties

Melting Temperatures

The solidus and liquidus temperatures of 304 stainless steel may be obtained from Peckner et al [ 19.55] . Few data on the stability and melting temperatures of the various iron oxide phases seem to exist, suggesting that the (Fe-O) phase diagram has not been extensively investigated at high temperatures. Thermal Conductivity

Thermal conductivities for 304 stainless steel are also given by Peckner et al [19.54] for temperatures between 374K and 965K. To determine the thermal conductivity in the liquid phase, typical models assume that there is a 50% reduction in thermal conducti vity on melting [ 19.55]. Thermal Capacity

These data for 304 stainless steel are also given by Peckner et al [ 19.54] over the range 263K to 1 1 1 9K. Heats of fusion of the component elements Fe, Cr, Ni may be used to interpolate an overall value for stainless steel based on the atomic fraction of each element. Typical elemental values are given by Brassfield et al [ 19.56]. Thermal Expansion and Density

Data on thermal expansivity are gi ven by Peckner et al [ 19.55] at 455K and 959K. As for the other materials the density can be calculated from the thermal expansivity together with a reference density. Room temperature values of density are given by Brassfield et al [ 19.56]. A theoretical volume increase of 3% is typically assumed on melting. Emissivity

The emissivity of 304 stainless steel fuel cladding is given by Spore et al [ 19.57] . Mechanical Properties

The mechanical properties of 304 stainless steel have been extensively documented up to its melting poin� for example, by Boyer and Gall [ 19.58] . The data are thought to be adequate to enable the calculations of the temperature dependent failure stress in stainless steel cladding. Oxygen Uptake

As forZircaloy the oxidation of304 stainless steel by steam can generate hydrogen in a severe accident. Oxidation kinetics are given in Chapter 8.

Thermophysical Properties

237

Viscosity

Choony [ 19.59] has detennined a theoretical value for the viscosity of molten steel based on measured values for iron. Joshikiyo [ 19.60] has found that the viscosity decreases with increasing oxygen content Vapour Pressure

The vapour pressure of iron for temperatures from 1 564K to 2755K is given by Boyer and Gall [ 1 9.58]. 19.6

Fuel Cladding Eutectic

Requirements

If as temperatures increase fuel cladding contact occurs, eutectic material resulting from dissolution of fuel may form at temperatures below the melting point of Zircaloy and well below the fuel melting point Phase diagrams for the v-o-Zr system are therefore required, as are any changes in the thermophysical properties e.g. density, thermal conductivity, specific heat etc resulting from the mixture composition. An important phenomenon concerning the potential for flow of liquefied material is the volume expansion associated with the dissolution process. Once the material has breached the clad, its flow will depend on viscosity and possibly surface tension phenomena. If this material is exposed to steam, further oxidation may also ensue.

As the material flows downward it will cool and eventually solidify. It follows that the thermophysical properties are required for both the solid and liquid phase. Because the material is likely to be close to or above its melting point, mechanical properties for such eutectics are not required. Phase Diagrams

v-o-Zr ternary system data have been published by Hofmann and Politis [ 1 9.61 ] and Skokan

[ 19.62] for the temperature range 1400 2273K. Data for models for temperatures up to 3 1 OOK have to be interpolated from binary phase diagrams. -

Thermal Properties

Deem [19.63] has given thennal conductivity data for a number ofV02-Zr02 compounds in the solid phase for temperatures in a maximum range 423-2373K. The compositions varied from O.2 V02 -O.8 Zr02 to 0.94 V02 - 0.06 Zr02 weight fractions. Current models are obtained via interpolation (frequently atomic fraction weighted averages) from data for the individual phases V02', Zr02 and Zr.

238

Thermophysical Properties

As for thennal conductivity very limited specific heat data exist Deem [ 19.63] gives enthalpies for several V02 - zr02 compounds in a maximum temperature range 273-2480K. As for the thennal conductivity, models are produced by interpolation from data for the individual components. Thennal expansion strains were also measured by Deem [19.63] in the temperature range 293-2273K for the range of V02 - zr02 compositions previously described for the thennal conductivity and specific heats. To derive densities from the thennal expansion coefficient, Deem also measured densities at a reference temperature 293K for the range of V02 - zr02 compositions. Viscosity

Data are available for the ternary V-O-Zr system, Prater and Courtright [ 19.64] for a range of V02 mole fractions in a Zr + V02 compound and over a temperature range 2 1 73-2423K. As long as the liquidus temperature for the particular composition was exceeded, minimal dependence on composition was found. 19.7

Control Rod Eutectic

Requirements

For PWR there are a range of possible control rod material eutectics for which data are required. At low system pressure failure of the Zircaloy guide tube is expected to occur at about 1 500K because of thennal expansion and physical contact between the stainless steel clad and Zircaloy guide tube. At high system pressure without Zircaloy/c1ad contact, the Zircaloy guide tubes would be expected to fail near 1700K as a result of stainless steel melting enabling contact between the molten control rod alloy and the Zircaloy. In either case, the control rod alloy will travel down the Zircaloy guide tube and chemical reactions are likely to occur between the Zircaloy, silver and indium. Low melting point Zr­ Ag and Zr-In eutectics are known to fonn. There may be other eutectics involving these materials and steel which require attention. Phase Diagrams

An Ag-Zr phase diagram can be derived from data by Elliott [19.65] . Data exist for the Zr­ rich part of the diagram above 0.65 atomic fraction Zr and for the Ag-rich part below 0.5 atomic fraction Zr. For temperatures below 1230K the data indicates no liquids present. For composites up to 0.65 atomic fraction Zr the liquidus temperature is below 1 500K rising to 2 1 OOK for pure Zircaloy. Work by Bowsher [19.66] indicates that molten sil ver-indium alloy fonns a low-melting solidus (l 500K) with Zircaloy; this behaviour is not inhibited by limited

Thermophysical Properties

239

pre-oxidation of the Zircaloy. Quantitative data for the solubility of Zircaloy or zirconium­ oxide in Ag-In-Cd absorber material are available from Hagen [ 19.67] , sufficient for approximate modelling of the solubility of the Zircaloy cladding in Ag-In-Cd solvents. 19.8

Structural Material

Requirements

There are a large number of major structural stainless steel components in the vessel. The extent to which some of these components reach high temperature under e.g. high temperature severe accident conditions is difficult to determine. Other components in the core, e.g. grids provide an additional potential for hydrogen production and therefore require suitable data. Data are required concerning the reaction of nickel-based steels (Inconel etc) with Zircaloy. The formation of a low melting point Inconel-Zircaloy eutectic at the grid/clad contact could induce rod failure at the point of contact. The mechanical performance and loading on core support structures must be known in order to estimate the likelihood of major core collapse. The major structures in the pressure vessel are made of stainless steel, for example, the core barrel, upper and lower grid plates and support plates, and the support columns. The thermophysical properties data for stainless steel have already been discussed for control rod cladding. The remaining material of significance is Inconel, of which the grid spacers are made. Melting Temperatures

Solidus and liquidus temperatures of Inconel-600 are given by Weast, Astle and Beyer [ 19.68]. A low melting temperature phase has been observed to form when Inconel contacts Zircaloy by Daniel, Nichols and Simpson [ 19.69] . Thermal Properties

Thermal conductivities are given in Tipton [ 19.70] from room temperature up to 1 144K. The room temperature specific heat of Inconel is also given in Tipton [19.70] . The density and thermal expansion of Inconel is given in Beyer and Gall [ 19.7 1]. Mechanical Properties

The mechanical properties of Inconel have been extensively documented up to its melting temperature, and are available from manufacturers' specifications. They are summarised in Tipton [19.70] and in Beyer and Gall [ 1 9.7 1 ] .

240

Thermophysical Properties

Vapour Pressure

The vapour pressures of nickel, chromium and iron, the major constituents of Inconel, are given in Beyer and Gall [ 19.71] up to 3000K . 19.9

Core-Concrete Material

Requirements

In an unrecovered accident in which the pressure vessel fails, molten core and structural materials will be deposited on to the reactor floor in the cavity beneath the vessel. The decay heat (and possibly chemical heat of reaction) may decompose the concrete and this process results in additional fission product release to the containment atmosphere. The attack by the debris on the concrete is largely a thermal process. As the material is ablated, heat is lost through the top surface of the pool or into the concrete. The progress of the front will depend on the relative thermal resistances of the melt and the concrete. As the concrete is dissolved water vapour, carbon dioxide and other gases are released. The gases are reduced to hydrogen and carbon monoxide etc on contact with metals in the melt and various oxide compounds are formed. There are therefore three main groups of chemical species that need to be considered, oxidic compounds, metals and gases. In general in the oxidic phase there is a wide range of materials. The principal materials include Si02, CaO, A12 03 from the concrete, V02, Zr02 ' from the fuel melt and various iron oxides from structural materials. The metals include any Zr which has remained unoxidised and various components from the steel structures. Phase diagrams for the liquidus and solidus temperatures for the metallic and oxidic mixtures will be required. For the internal thermal transport, thermal conductivities, sjJecific heats, densi ties are necessary. Emissivities are required to determine heat losses to the containment via radiation. The transport processes require viscosities, surface tensions etc. Phases

In order to predict the melting and freezing behaviour it is necessary to have the concrete melting point and more generally the liquidus and solidus temperatures for the various metallic and oxidic mixtures formed during the core-concrete interaction. The melting ranges for typical concretes have been determined experimentally Cole [ 1 9.72] . For the metallic phase a ternary phase diagram is given for iron-chromium metal in Speich [ 1 9.73].

Thermophysical Properties

24 1

Concerning the oxidic behaviour the major constituents of the concrete, Si02 , CaO and A 1 2OJ known to form a complicated ternary system, see Sandia Review [ 19.74] .

are

Consideration of the complete concrete oxide and fuel oxide phase diagrams is very complicated. Skokan [ 19.75] has measured the composition solidification temperature for mixtures of corium at different oxidation levels with different types of concrete over a temperature range 1300-2800K. Thermal Properties

Values of thermal conductivity are given for miscellaneous condensed phase species materials in a Sandia Review [ 19.74]. Further data relate to specific materials, rather than mixtures. Specific heat data exist up to temperatures in the range 2()()()-2500K but for specific materials rather than mixtures. Densities are available for individual condensed phases in Sandia Review [19.74]. The temperature range of data available varies from species to species. For many of the oxides the data lie in the range 1 500-2100K although some data cover the eutectic range from melting to boiling. Emissivity data have been given for uranium and zirconium oxides solids in earlier sections but not for the liquid phases. The extent to which such material can be extrapolated to melts is very uncertain. Viscosity

For the metallic phase, viscosity data exists for the major constituent, iron. Data for the viscosity of iron are given in Sandia Review [ 19.74]. For the oxidic phase, the viscosity of the molten oxides increase with increasing silica content. For low-silica mixtures the viscosity can be computed as in the Sandia Review [19.74]. For higher-silica mixtures, data are given by Bottinga and Weill [ 19.76]. A viscosity model based on these data has been proposed by Shaw [ 19.77] . Surface Tension

Values of surface tension for condensed phase species are given in Sandia Review [ 19.74].

19.10 Water/Steam Properties

A fluid such as water exists as liquid or as a vapour. The state of the fluid is described by its thermodynamic properties such as pressure, temperature and specific volume. Two such

242

Thermophysical Properties

properties are needed to describe the state of the fluid, as any thennodynamic property can be expressed as a function of any other two: e.g.

p = fl (v,T)

T = f3 (p,v) where p = pressure v = specific volume T = temperature. Under saturation conditions, each phase can be fully described by only one thennodynamic property, and any such property can be expressed as a function of any other, for example:

This equation relating the saturation temperature and pressure defines the saturation line. At the critical temperature and pressure (Tc and PC> the thennodynamic properties of saturated liquid and vapour become identical, and the saturation line ends at this point (the critical point). The thermodynamic and transport properties are typically based on polynomial fits to steam table data for water, and ideal gas behaviour for any noncondensable gas component. Non­ condensable gases are considered in the next section. Functions required in a six-equation model such as in the TRAC code include the saturation temperature corresponding to total pressure; the saturation temperature corresponding to the partial pressure of steam; the specific internal energies ofliquid, gas phase, and noncondensable; the saturated liquid and steam enthalpies corresponding to the partial pressure of steam; the liquid, gas-phase and noncondensable densities; the derivatives of saturation temperatures and saturation enthalpies with respect to pressure; the partial derivatives ofliquid, steam, and noncondensable internal energies and densities with respect to pressure (at constant temperature) and with respect to temperature (at constant pressure). Other codes employ different combinations of thennodynamic variables. The range of validity for the thennodynamic properties has to be large to cover the complete range of accident conditions, 0. 1 -20 MPa, 273K - 3000K . It is important that the models give adequate values for the thennophysical properties over a whole range of temperatures and pressures. Polynomials to fit the experimental data for a particular property are unlikely to be valid out of the range of the original data. These decrepancies usually occur near the critical

Thermophysical Properties

243

temperature, the critical pressure and also at low pressure, when it is difficult for experiments to be carried out. 19.1 1 Non-condensable Gases

Various non-condensable gases can be present depending on the accident conditions. Non­ condensable gases are included as internal gases within LWR fuel rods. Properties for ten gases are given in MATPRO [ 19.29] , including helium, argon, krypton, xenon, hydrogen, nitrogen, oxygen, carbon monoxide, carbon dioxide and water mixtures. Thermal conductivites are modelled (with any combination of mixtures), gas viscosities are similarly modelled. Also included are specific heat capacity, emissivity and mean free path. In many cases the assumption is made that the gases are ideal. 19.12 Other Gases

In order to calculate the rate of evaporation or condensation of various species on to fixed surfaces or aerosols, equilibrium vapour concentrations and vapour pressures must be known. In MATPRO [ 19.29] these are given fora range of additional gases. The gases include iodine, caesium iodide, caesium hydroxide, tellurium, �Te, HI, tin, tin telluride, zirconium dioxide, uranium dioxide, carbon and silver iodide. REFERENCES

19. 1

M F Lyons et aI, "V02 Properties Affecting Performance", Nuclear Engineering and Design, 21 ( 1972) p. 167. 19.2 W F Lyon and W E Baily, "The Solid-Liquid Phase Diagram for the V02 - Pu02 System" Journal of Nuclear Materials, 22, 332 ( 1 967). 19.3 J L Krankota and C NCraig, "Melting PointofHigh Burnup Pu02 - V02, "Transactions of the American Nuclear Society, 1 1 , 1 32 (1968). 19.4 J G Reavis and J L Green, "Transformation Temperatures of Irradiated V02 - Pu02 Fast Reactor Fuels", Transactions of the American Nuclear Society, 14, 595 ( 1 97 1). 19.5 L A Goldsmith and J A M Douglas, "Measurements of the Thermal Conductivity of Vranium Dioxide at 670- 1270 K", Journal of Nuclear Materials, 47 (1 973) pp. 3 1-42. 19.6 I C Hobson, R Taylor, and J B Ainscough, "Effect of Porosity and Stoichiometry on the Thermal Conductivity of Vranium Dioxide", Journal of Physics Section D: Applied Physics, 7 ( 1974) pp. 1003- 1 0 1 5. 19.7 J C Weilbacher, "Diffusivite Thermique de 1'0xyde d'Vranium et de 1'0xyde de Thorium a Haute Temperature, "High Temperatures - High Pressure, 4 ( 1972) pp. 43 1-438.

19.8 19.9

R L Gibby, "The Effect of Plutonium Content on the Thermal Conductivity of (D, Pu) 02 Solid Solutions", Journal of Nuclear Materials, 38 ( 1 97 1 ) pp. 163- 177. C S Kim et al, "Measurement of Thermal Diffusivity of Molten V02, "Proceedings of the Seventh Symposium on Thermophysical Properties at the National Bureau of Standards, Gaithersburg, Maryland, May 10- 12, 1977, pp. 338-343 Published by the American Society of Mechanical Engineers, CONF 770537-3.

244

Thermophysical Properties

19. 10 J F Kerrisk and D G Clifton, "Smoothed Values of the Enthalpy and Heat Capacity of UOz", Nuclear Technology, 16 (December 1 972) pp. 5 3 1 -535. 1 9. 1 1 R L Gibby et al, "Analytical Expressions for Enthalpy and Heat Capacity for Uranium - Plutonium Oxide", Journal of Nuclear Materials, 50 ( 1974) pp. 1 5 5- 1 6 1 . 1 9 . 1 2 L Leibowitz, D F Fishcer, M G Chasanov, "Enthalpy of Uranium-Plutoniwn Oxides (UO.8' PUO) 01 .07 from 2350 to 3OOOK", Journal of Nuclear Materials, 42 (1972) pp. 1 131 16. 19. 1 3 L Leibowitz et aI, "Enthalpy of Liquid Uranium Dioxide to 3500K", Journal of Nuclear Materials, 39 ( 197 1) pp. 1 1 5- 1 16. 19. 14 L Leibowitz, D F Fischer, M G Chasanov, Enthalpy of Molten Uranium-Plutonium Oxide, ANL-8082 (February 1975). 19. 1 5 J A Christensen, "Thermal Expansion and Change in Volume ofUraniwn Dioxide on Melting", Journal of the American Ceramic Society, 46 ( 1 963) pp. 607-608. 1 9 . 1 6 M Tokar and A W Nutt, "Thermal Expansion ofPu02 from 25 to 1420·C", Transactions of the American Nuclear Society, 1 5 (June 1 972) pp. 2 10-21 1 . 19. 1 7 P C Held and D R Wilder, "High Temperature Hemisperical Spectral Emittance of Uranium Oxides at 0.65 and 0.70 Jl.Ill " , Journal of the American Ceramic Society, 52 ( 1 969). 19. 1 8 A Padel and C de Novion, "Constants Elastique des Carbures, Nitures et Oxydes d ' Uranium et de Plutonium", Journal of Nuclear Materials, 33 (1 969) pp. 40-5 1 . 19. 1 9 W Dienst, "Irradiation Induced Creep of Ceramic Nuclear Fuels", Journal of Nuclear Materials, 65 ( 1977). 1 9.20 D A Collins and R Hargreaves, "Performance-Limiting Phenomena in Irradiated UOz", Paper No. 50, Proceedings of the BNES International Conference on Nuclear Fuel Performance, October 1 5- 19, 1 973, London (CONF-73 1004). 1 9.21 C Ronchi and C Sari, "Properties of Lenticular pores in UOz (U, Pu)Oz and Pu0z", ' Journal of Nuclear Materials, 50 ( 1 974) pp. 9 1 -97. 1 9.22 J B Ainscough, B W Oldfield and J ° Ware "Isothermal Grain Growth Kinetics in Sintered UOz Pellets", Journal of Nuclear Materials, 49 (1973f74) pp. 1 17- 128. 19.23 R F Cannon, J T A Roberts, R J Beals, "Deformation of UOz at High Temperatures", Journal of the American Ceramic Society, 54 ( 197 1) pp. 105- 1 12. 19.24 J T A Roberts and Y Ueda, "Infl uence of Porosity on Deformation and Fracture of UOz", Journal of the American Ceramic Society, 55, 3 ( 1972) pp. 1 1 7- 1 24. 19.25 E Rotwell, "High Temperature Substoichiometry in Uraniwn Dioxide", Journal of Nuclear Materials 6 ( 1 972) pp. 229-236. 19.26 A D Knipe, S A Ploger, D J Osetek. PBF Severe Fuel Damage Scoping Test - Test Results Report EGG-24 1 3 , NUREG/CR-4683, August 1 986. 19.27 R E Woodley, "The Viscosity of Molten Uranium Dioxide", Journal of Nuclear Materials, 50 ( 1974) pp. 1 03-106. 1 9.28 G Benezech, J P Coutures and M Fox, Transition Study of Uranium Dioxide Vaporization Processes between 200K and 2600K , ANL-TRANS-972 ( 1974). 1 9.29 A Handbook of Materials Properties for Use in the Analysis of Light water Reactor Accident Analysis, NUREG/CR-5273, 1 990. 1 9.30 B Lustman and F Kerze, "The Metallurgy of Zirconium", New York: McGraw-Hill Book Company, Inc. (1955).

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245

19.3 1 R Ruh and H J Garret� "Nonstoichiometry of Zr02 and its Relation to Tetragonal­ Cubic Inversion in zr02", Journal of the American Ceramic Society, 50 (1966) pp. 257-26 1 . 19.32 A D Feith, ''Thermal Conductivity and Electrical Resistivity ofZircaloy-4", GEMP669 (October 1 966). 19.33 S Nazare, G Ondracek, and B Schulz, "Properties of Light Water Reactor Core Melts", Nuclear Technology, 32 ( 1 977) pp. 239-246. 19.34 K E Gilchrist, "Thermal Property Measurements on Ziracloy-2 and Associated Oxide Layers", Journal of Nuclear Materials, 62, (1 976) pp. 257-264. 19.35 E A Eldridge and H W Deem, "Specific Heats and Heats of Transfonnation of Zircaloy-2 and Low Nickel Zircaloy-2", USAEC BMI- 1 803 (May 3 1 , 1 967). 19.36 R R Hammer, "Zircaloy-4, Uranium Dioxide and Materials formed by their Interaction. A Literature Review with Extrapolation of Physical Properties to High Temperatures", IN- I093, (September 1967). 19.37 L R Bunnell et al, "High Temperature Properties of Zircaloy-Oxygen Alloys", EPRI NP-524 (March 1 977). 19.38 E F Juenke and L H Sjodahl, "Physical and Mechanical Properties: Emittance Measurements", AEC Fuels and Materials Development Program, GEMP 1 008 ( 1 968) pp. 239-242. 19.39 D G Hardy, "High Temperature Expansion and Rupture Behaviour of Zircaloy Tubing", Topical Meeting on Water Reactor Safety, Salt Lake City, American Nuclear Society, March 26-28, 1 973, CONF-730304. 19.40 R H Chapman, J L Crowley, A W Longest and E G Sewell, "Effect of Creep Time and Heating Rate on Deformation of Zircaloy-4 Tubes tested in Steam with Internal Heaters, ORNL/NUREG/fM-245 and NUREG/CR-0345 (October 1 978). 19.4 1 R E Pawel, "Oxygen Diffusion in Beta Zircaloy During Steam Oxidation", Journal of Nuclear Materials, 50 (1974) pp. 247-258. 19.42 R Haddad and B Cox, "On the Initiation of Cracks in Zircaloy Tubes by 12 and Cs/Cd Vapours", Journal of Nuclear Materials 1 38 ( 1 986) pp. 8 1 -88. 19.43 E Millner, "Hydrogen Absorption in Zircaloy During Aqueous Corrosion, Effect of Environment", WAPD-TM-4 1 1 (November 1 964). 19.44 T Futura and S Kawasaki, "Reaction Behaviour of Zircaloy-4 in Steam-Hydrogen Mixtures at High Temperatures", Journal of Nuclear Materials 105 ( 1982), pp. 1 19131. 1 9.45 L L Quill, "The Chemistry and Metallurgy of Miscellaneous Materials", New York ( 1950) pp. 144- 1 5 1 . 19.46 G W Parker, G E Creek and A L Sutton, "Influence of Variable Physical Process Assumptions on Core Melt Aerosol Release", Proceedings of the International Meeting on Thermal Nuclear Reactor Safety, Chicago, IL, August 29-September 2, 1 982, NUREG/CP-0027, Vol 2. 19.47 D A Petti, Silver-Indium-Cadmium Control Rod Behaviour and Aerosol Formation in Severe Reactor Accidents, NUREG/CR-4876, EGG-2501 , April 1 987. 19.48 Chase et al, JANEF Thennochemical Tables, ( 1986), pp. 54 1 -543. 19.49 I Cohen, E F Losco and J D Eichenberg, "Metallurgical Design and Properties of Silver-Indium-Cadmium Alloys for PWR Control Rods", Bettis Technical Review, ( 1958), WAPD-BT-6.

246

Thermophysical Properties

19.50 S Nazare, G Ondracek and B Schulz, "Properties of Light Water Reactor Core Melts", Nuclear Technology, 32, ( 1977), pp. 239-246. 19.5 1 C T Lynch, ed, "Handbook of Materials Science, Volume II: Metals, Composites and Refractory Materials", Cleveland, Ohio: CRe Press, Inc. (TA403.4 L94). 19.52 C J Smithells and E A Brandes (eds) "Metals Reference Book", London and Boston: Butterworths (TN 67 1 S55 1 956). 1 9.53 D A Powers, "Behaviour of Control Rods during Core Degradation: Pressurisation of Silver-Indiwn-Cadmium control rods", NUREG/CR-4401 , (September 1 985). 1 9.54 B R Bowsher, R A Jenkins, A L Nichols, N A Rowe and J A H Simpson, "Silver­ Indium-Cadmium Control Rod Behaviour During a Severe Reactor Accident", AEEW-M 1 99 1 , 1986. 1 9.55 D Peckner and I M Bernstein, (cds) "Handbook of Stainless Steel , New York: McGraw-Hill Book Company, ( 1 977). 19.56 H C Brassfield, J F White, L Sjodahl and J T Bittel, "Recommended Property and Reaction Kinetics Data for Use in Evaluating a Light-Water-Cooled Reactor Loss-of­ Coolant Incident Involving Zircaloy-4 304 SS Clad U02", GEMP 482, ( 1 968). 1 9.57 J W Spore et al., "TRAC-BDI An Advanced Best Estimate Computer Program for Boiling Water Reactor Loss of Coolant Accident Analysis, Volume 1 Model Description", Appendix B, NUREG/CR-2/78, EGG-2 109. 1 9.58 HE Boyer and L Gall (ed), "Metals Handbook , Desk Edition", American Society for Metals, (1 984). 1 9.59 K S Choony, "Thermophysical Properties of Stainless Steel", ANL-75-55. 1 9.60 Joshikiyo, "On the Viscosity of Liquid Iron", Journal of the Japan Institute of Metals 37 (1 973) pp. 1230- 1 235. 19.61 P Hofman and C Politis, "The Kinetics of the Uranium Dioxide-Zircaloy Reactions at High Temperatures", Journal of Nuclear Materials, 87 ( 1 975), pp. 375-397. 19.62 A Skokan, "High Temperature Phase Relations in the U-O-Zr System", Fifth International Meeting on Thermal Nuclear Reactor Safety, Karlsruhe, German Federal Republic, (September 9- 1 3), ( 1984), KfK 38801 1 , (December 1 984), pp. 1035- 1042. 1 9.63 H W Deem, "Fabrication Characterisation, and Thermal-Property Measurements of Zr02-Base Fuels", BMI- 1 775, (June 1966). 1 9.64 J T Prater and E L Courtright, "Zr Oxidation and Zr-U02 Viscosity Studies", Severe Fuel Damage and Source Term Research Program Review Meeting, Idaho Falls, Idaho, April 16- 19, 1985. 1 9.65 R P Elliott, "Constitution of Binary Alloys, First Supplement", New York: McGraw­ Hill Book Company, Inc, ( 1 965). 19.66 B R Bowsher, R A Jenkins, A L Nichols, N A Rowe and J A H Simpson, "Silver­ Indium-Cadmium Control Rod Behaviour during a Severe Reactor Accident", AEEW-R 1 99 1 , 1 986. 19.67 S Hagen, "Absorber Rod Tests in the NIELS Facility", Severe Fuel Damage and Source Term Research Program Review Meeting, Idaho Falls, Idaho, April 1 6- 19, ( 1985). 19.68 R C Weast, M J Astle, W H Beyer, "CRC Handbook ofChemistry and Physics", CRC Press Inc., ( 1983/84).

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247

19.69 B Daniel, A L Nichols, J A H Simpson, "The Liquefaction of Alloys Identified with Structural Materials in a PWR Core: Zircaloy, Inconel-600 and 304 Stainless Steel", AEEW-M 2250, ( 1 985). 19.70 C R Tipton (ed), "Reactor Handbook Volume 1: Materials" 19.7 1 H E Beyer, T L Gall (eds), "Metals Handbook (Desk Edition)" ASME, ( 1984). 19.72 R K Cole, In., D P Kellu, M A Ellis, "CORCON-MOD2: A Computer Program for Analysis of Molten Core-Concrete Interactions", NUREG/CR 3920, SAND84-1246. (August 1984). 9.73 G R Speich, "Cr-Fe-Ni (Chromium-Iron-Nickel)", Metals Handbook , American Society for Metals, Metals Park, Ohio, Volume 8, p 424, ( 1973). 19.74 "Core Meltdown Experimental Review", SAND74-0382, (NUREG-0205), Sandia National Laboratories, Albuquerque, NW, (March 1977). 19.75 A Skokan, H Hollek and M Peehs, "Chemical Reactions Between Light Water Reactor Core Melt and Concrete", Nucl. Techn. 46 2, p 255, (1979). 19.76 Y Bottinga and D F Weill, "The Viscosity of Magmatic Silicate Liquids: A Model for Calculation", American Journal of Science, Volume 272, pp 438-475, (1 972). 19.77 H R Shaw, "Viscosities of Magmatic Silicate Liquids: An Empirical Method of Prediction", American Journal of Science, Volume 272, pp 870-893, ( 1972). 19.78 W A Coffman and L L Lynn, "WATER: A Large Range Thermodynamic and Transport Water Property FORTRAN-IV Computer Program", Bettis Atomic Power Laboratory Report W APD-TM-568 (December 1966) .

249

C hapter 20 COM PUTER C ODES

20.1

Introduction

The purpose of this chapter is to summarise the main features of the major computer codes that have been developed for reactor safety research. The original emphasis of research was to verify LWR designs and that the protection systems would prevent overheating of the core. This gave rise to an active period of code development to provide thermal-hydraulic system codes for predicting Loss of Cooling Accidents (LOCAs). There were many allied experimental programmes to provide insights into model requirements and code validation. After the core melt accident at Three Mile Island, there was a change in emphasis towards the development of codes for predicting core melt, fission product transport and the ultimate threat to the containment -the source tenn referred to earlier. Here again the codes were supported by ambitious experimental programmes. The chapter falls into two halves: Heat Transfer and Hydraulics Codes are summarised first. Severe Accident codes are discussed later in somewhat more detail since these codes are less well known and severe accident issues constitute the more active area of current research.

20.2

Heat Transfer and H y draulics

20.2.1 Thermal-Hy draulics System Codes

The USNRC Codes RELAP5 and 1RAC are among the most commonly used thennal­ hydraulic system codes. Versions of these codes have been developed for wide ranging accident scenarios and for both pressurised and boiling water reactors. Other European codes at advanced stages of development include CATHARE (France) and ATHLET (Germany). The main features of 1RAC and RELAP are summarised below as representative examples of thermal-hydraulic system codes. 20.2.2 Thermal-Hydraulic Reactor Coolant System Codes RELAP5

RELAP5 is a light water reactor (LWR) transient analysis code, under development at the Idaho National Engineering Laboratory (INEL) for the U.S . Nuclear Regulatory Commission

250

Computer Codes

(NRC). The main code applications include analysis required to support rulemaking, licensing, evaluation of accident management strategies and operator guidelines, and experiment planning and analysis. Other applications include simulations of transients in LWR systems that lead to severe accidents, such as LOCAs, anticipated transients without scram (ATWS), operational transients, loss of offsite power, station blackout, and turbine trip. The code is also used as a driver for a nuclear plant anal yzer. RELAP5 is a highly generic thermal-hydraulic code that can be used for simulation of a wide variety of hydraulic and thermal transients in both nuclear and nonnuclear systems. The code includes many generic component models which enable general systems to be simulated. The component models include pumps, valves, pipes, heat structures, reactor point kinetics, electric heaters, turbines, separators. accumulators and control system components. There are also various special process models that have been included to enable effects such as form loss. flow at an abrupt area change, branching, choke flow, boron tracking, and noncondensable gas transport to be modelled. The development of the models and codes that are included in RELAP5 has spanned approximately 1 2 years from the early stages of RELAP5 numerical scheme development to the present. RELAP5 represents the accumulated experience that has been gained in modelling core behaviour during severe accidents, two-phase flow process and LWR systems. The code development and validation have been extensive via application and comparison with experimental data in the LOFf, PBF, Semiscale, ACRR. NRU. BETHSY and other experimental programmes. These programmes have been discussed earlier. The most recent developments with the code are to extend its capability to address thennodynamic issues associated with advanced LWRs and Eastern European VVERs. TRAC

The Transient Reactor Analysis Code TRAC is a family of specialised computer codes for modelling transients in Light Water Reactors and associated experimental rigs. The codes are designed to model Loss-of-Coolant Accidents (LOCA), and Pressurised Transients, and were originally developed at Los Alamos National Laboratory on behalf of the USNRC. There are PWR [20.3] and BWR [20.4] versions of the codes. The main features of the TRAC suite of codes are summarised in Table 20. 1 . Later versions of the code can model the whole of the reactor coolant system, the Emergency Core Cooling System (ECCS), and trip control functions. The central feature of the code is its hydrodynamic model. In addition, the code contains models of fuel pin heat generation and conduction, pump behaviour, valve operation etc. The modelling of solid structures is restricted to fixed geometry. The TRAC hydrodynamics is based on a two-fluid, six-equation model. Thus mechanical and thermal dis-equilibrium between the phases is allowed. The reactor pressure vessel can be modelled by three-dimensional (r, 9, z) or one-dimensional equations. The remainder of the coolant circuit is modelled by one-dimensional equations. The degree of detail in the modelling is determined by the user, and is dependent on computer running time considerations.

Computer Codes

25 1

TABLE 20.1 FEATURES OF TRAC

Multi-Dimensional Fluid Dynamics Nonhomogeneous, Disequilibrium Hydrodynamic model Flow Regime Dependent Constitutive Relations Comprehensive Heat-Transfer Capability General Accident Analysis Capability Modular Construction

The thennal-hydraulic equations describe the transfer of mass, energy, and momentum between the steam-water phases and the interaction of these phases with the heat flow from the system structures. A wide range of two phase flows are possible and a flow-regime­ dependent constitutive equation package has therefore been incorporated into the code. The code accommodates detailed heat-transfer anal yses of the vessel and the loop components. Included is a two-dimensional (r,z) treatment of fuel-rod heat conduction with flow-regime­ dependent heat-transfer coefficients, dynamic fine-mesh rezoning is also incorporated in later versions of the code to resolve both bottom-flood and falling-film quench fronts. Entire accident sequences can be computed enabling a consistent and continuous calculation. For example, the code models the complete blowdown, refill, and reflood phases of a LOCA. This capability eliminates the need to perfonn calculations using different codes to analyze a given accident and the possibility of inconsistency, mass or energy loss etc. TRAC has been developed and assessed via many separate effects and integral scale experimental programmes including LOFT, LOBI, SEMISCALE and more recently UPTF. Other Major System Codes

A .code for Analysis of Ihennal-Hydraulics during an Accident of Reactor and Safety Evaluation (CATHARE) [20.5] is under development by the French Atomic Energy Commission (C.E.A.), the French utility (E.D.F.) and the French vendor (FRAMATOME). CATHARE may be used to simulate wide ranging accidents in PWR-type installations during which little severe damage occurs to fuel rods including transients, intennediate and large break LOCAs.

252

Computer Codes

CA THARE models take into account mechanical and thennal dis-equilibrium that can occur in situations such as blowdown. refill and reflood. A large assessment program is being carried out, including both separate effects and integral tests. e.g. Bethsy covering the whole domain of PWR accidents. Another European thennal-hydraulics code ATHLET (Analysis of Thennal-hydraulics of Leaks and Transients) [20.6] is under development in Gennany. The field of application covers the complete spectrum of operational and fault transients and LOCAs for PWRs and BWRs. 20.2.3 Fuel and Cladding Behaviour

In order to satisfy licensing criteria various limits on peak clad temperature and the degree of oxidation are typically imposed. This has led to the development of various codes to predict fuel and cladding behaviour under LOCA and transient conditions. Fuel perfonnance is affected by a number of factors including the fuel operating history. and the thennal-hydraulic boundary conditions. The objective is to calculate fission product release to the coolant. the extent of coolant blockage due to changes in the fuel lattice geometry and various critical factors for licensing. e.g. peak clad temperature and extent of cladding oxidation. In the U.K. various fuel perfonnance codes have been developed including SLEUTH [20.7] . [20.8]. MINIPAT [20.8] and HOTROD [20.9]. As described earlier in LOCAs the clad will creep once the temperatures have risen sufficiently. The extent of clad defonnation will be driven by the pressure in the gap between the fuel and the cladding and the cladding temperature which is in turn affected by the gap conductance. The GAPCON code developed in the USA [20. 10] provides the gap conductance under steady state conditions and these provide the initial conditions for accident analysis. Cladding defonnation is calculated by the FRAPT6 [20. 1 1] and CANSWEL [20. 1 2] codes. given the thennal-hydraulic boundary conditions. To provide this necessary feedback during ballooning under LOCA conditions the MABEL code [20. 1 3] was developed. The MABEL code calculated the coupled effects of subchannel heat transfer. fuel and cladding temperature and the rod internal pressure of the cladding deformation. by coupling to CANSWEL. The code took fuel perfonnance input from the codes SLEUTH. MINIPAT. and HOTROD referred to above. 20.2.4 Containment

Under large-break LOCA conditions large quantities of water and steam would exit the break and produce a substantial pressure spike. Western style type PWRs and BWRs have containments which are designed to withstand these pressure loads. Computer codes have been developed to predict the containment pressure and temperature under design basis LOCA and also severe accident conditions.

Computer Codes

253

The CONTAIN code [20. 14] has been developed at Sandia National Laboratories and sponsored by the USNRC to provide a model for the thennal response of the containment and the material (water, steam, and possibly non-condensables) mass flow. Critical flow conditions arise during a severe LOCA and a key issue is to compute the critical massflow . A series of experiments have been carried out at Marviken and these programmes have been extensively used in containment code development and validation.

20.3

Severe Accident Codes

The accident at Unit 2 of the Three Mile Island Plant (TMI-2) spawned considerable research into severe accident phenomena. Large computer codes were developed in parallel with expensive and wide ranging experimental programmes. Much of this work was carried out in the USA and a so-called two-tier code development approach was adopted. The codes fell into two categories: 1.

Integrated Codes.

These aimed to calculate complete sequences up to the release of fission products from the containment. The codes employed engineering based models, were parametric and fast running in order that parametric variations could be explored. 2.

Mechanistic Codes.

These codes were more detailed and usually only considered a particular part or phase of an accident sequence. These codes were in general more expensive in computer time than the parametric codes. 20.3.1 Integrated Codes

It is now recognised that two different classes of integrated codes have been developed. The frrst class were developed before there was much understanding of the main phenomenology. The intention was that specific components or modules ofthese codes would be benchmarked against an appropriate mechanistic code. The first code or more accurately suite of codes to be developed in the USA was the Source Tenn Code Package (STCP) [20. 1 5] . Following a better understanding of the phenomenology, second generation codes are now being developed. These aim to take advantage of the knowledge gained from the mechanistic codes and to synthesise this knowledge into a faster running model. The most significant code in this category is MELCOR [20. 1 6] , which is arguably on the fringe between integrated and mechanistic codes. Another example is the US industry code MAAP[20. 1 7] . STCP

The Source Term Code Package has been developed by Battelle Columbus for the USNRC. The package is comprised of various individual codes, MARCH2 [20. 1 8] , CORCON MOD2

Computer Codes

254

[20. 1 9] , CORSOR-M [20.20], V ANESA [20.2 1 ] , MERGE [20.22] , TRAP-MELT2 [20.23] , NAUA-4 [20.24], SPARC-B [20.25] and ICEDF [20.26] . The aim of the STCP i s to retain the basic mechanisms in these codes but to couple the codes together to produce a practical engineering tool. Generally the STCP has taken the individual codes as written but in some cases the codes have been combined to provide integrated calculations. There are now four major categories represented by the codes MARCH3, TRAP-MELT3, V ANESA, and NAUNSPARC/ICEDF. The MARCH3 code is a combination of the MARCH2, CORSOR-M, and CORCON-MOD2 codes. The TRAP-MELT3 code is a combination of the TRAP-MELT2 and MERGE codes. V ANESA is an individual code. The NAUA 4/SPARC and ICEDF codes are currently treated separately but are considered within a specific category. MARCH3 treats the thermal-hydraulic conditions of the reactor coolant system together with core-concrete thermal interactions and fission product release. TRAPMELT3 provides an additional RCS thermal-hydraulic capability and fission product transport in the RCS. VANESA is discussed later in view of its role as a general fission product code. NAUA4/ SPARC and ICEDF provide the containment capability. MELCOR

MELCOR is a fully 2nd generation integrated computer code that models the progression of severe accidents in light water reactor (LWR) nuclear power plants. MELCOR is under development at Sandia National Laboratories for the U.S. Nuclear Regulatory Commission (USNRC) as a probabilistic assessment (PRA) tool. The code models the whole spectrum of severe accident phenomena and may be applied to both boiling water reactors and pressurized water reactors. To overcome certain limitations in the STCP it was determined that two new codes should be written, one to analyze the system response through release of radionuclides to the environment, and one to calculate the consequences of that release. The resulting accident progression and source term code was designated MELCOR, the companion consequences code was designated MACCS (MELCOR Accident Consequences Code System). MELCOR is capable of modelling the following severe accident phenomonology: the thermal-hydraulics of the reactor coolant system; the thermal-hydraulic response of the containment and auxiliary buildings; thermal response of structures; core heatup and degradation; reactor cavity interactions (including molten core-concrete interactions);

Computer Codes

255

hydrogen production, transport and combustion; fission product release, transport, and deposition and various engineered safety features. Many of the MELCOR models are mechanistic, with capabilities similar to those in the detailed codes, however, many of the mechanistic models have been coded with optional adjustable parameters. Parametric models are available in areas where mechanisms are poorly understood. This enables the use of MELCOR to include uncertainty analyses and sensitivity studies. Other Integrated Codes

Other codes include ESCADRE (France) [20.28] , MAAP (USA), THALES [20.29] and the European code ESTER [20.30]. 20.3.2 Mechanistic System Codes

Mechanistic codes have grown from mechanistic models that have been developed for individual phenomena. They provide the coupling effects between the various phenomena. In severe accident analysis coupling and feedback effects can be extremely important In the development of mechanistic codes advantage has been taken of the many advanced thennal­ hydraulic system codes already in existence to provide the thennal-hydraulic feedbacks and boundary conditions. This approach has been followed for all the following thennal­ hydraulic codes referred to earlier.

TABLE 20.2 MECHANISTIC SEVERE ACCIDENT CODES Thermal Hydraulic Module

Fuel Damage Module

RELAP5

SCDAP

CATHARE

ICARE

ATHLET

KESS

Attempts have also been made to couple fission product release and traflSJX>rt codes such as VICTORIA to an appropriate reactor coolant system/fuel damage code. This allows for the thennal­ hydraulic feedbacks on fission products deposition, retention and resUSjXmsion to be investigated. Attempts are also being made now to couple thennal-hydraulic system codes to containment codes. One way coupling is necessary for conventional LWR accident analysis but for many advanced reactor safety issues a two-way coupling is necessary.

2 56

Computer Codes

SCDAPIRELAP5

SCDAP/RELAP5 [20.3 1 ] has been developed at the Idaho National Engineering Laboratory (INEL) for the US Nuclear Regulatory Commission (NRC) to provide an advanced best­ estimate predictive capability for use in severe accident applications. As its name suggests it has the thermal-hydraulic capabilities ofRELAP5 but these are extended and other models are included to provide the severe accident analysis capability. SCDAP/RELAP5 was developed by integrating two separate codes, RELAP5/MOD2 and SCDAP. These codes were combined to model the coupled interactions that occur between the core and the RCS during a severe accident. For example, changes in core geometry caused by fuel rod ballooning and meltdown, can have a significant effect on RCS flows. SCDAPj RELAP5 has undergone extensive validation against various experimental programmes. These include the PBF SFD tests, LOFT and selected tests in the CORA melt progression programme. The SCDAP component models all aspects of core behaviour during a severe accident. Treatment of the core includes fuel rod heatup, ballooning and rupture, rapid oxidation, Zircaloy melting, U02 dissolution, zr02 breach, flow and freezing of molten fuel and cladding, and debris formation and behaviour. The code also models control rod and shroud behaviour (for the purposes of experiment design, analysis and model validation). Early versions of the code were linked with the TRAP-MELT code [20.23] . The TRAP­ MELT code models the behaviour of fission products and aerosols release and transport within the RCS. This treatment includes aerosol agglomeration (including Brownian motion, gravitational settling, and turbulent eddy effects, aerosol deposition (including gravitational settling, thermophoresis, and diffusion from laminar or turbulent flow), fission product evaporation and condensation, and chemisorption of vapours by stainless steel. CATHAREIICARE

This code, developed by the Institute for Protection and Nuclear Safety (IPSN) of CEA France, have been obtained via merging the thermal-hydraulic code CATHARE with the core degradation code ICARE. ICARE models the progression of reactor core damage through core heat up, melting and material (cladding) embrittlement, material relocation and blockage formation. For fission product release and transport this code system also includes TRAPF, a derivative of TRAP-MELT. The code has been validated against the PHEB US SFD series, PBF, selected CORA tests and the TMI-2 accident. ATHLET-SA

This code, under development at GRS, results from the coupling of the ATHLET and KESS codes [20.32] , [20.33]. KESS developed by IKE Stuttgart addresses the physical and chemical behaviour associated with core heat up and degradation and fission product release.

Computer Codes

257

The code is validated against most of the available fuel damage tests data, including PHEBUS, CORA and TMI-2. VICTORIA

The behaviour of the radionuclides within the reactor vessel and coolant system will have a significant effect in determining the quantity, nature and timing of release of radionuclides to the containment during an accident. VICTORIA [20.34] is a computer code that mechanistically predicts radionuclide release, transport, and deposition within the reactor vessel and coolant system during a severe reactor accident. The code development is funded by the USNRC and carried out at Sandia Laboratories USA. VICTORIA follows the destination of many different materials, including fission products and the major vessel materials, as they interact with each other. The interactions may be sensitive to changing pressures, temperatures, material motions, and oxidation environment during a severe reactor accident given an initial element distribution, which can be provided by an inventory code and thermal-hydraulic conditions. VICTORIA determines the transport of the fission products from fuel grains through the open porosity in the fuel and into the coolant channels provided fuel cladding failure has occurred. The fission products are allowed to interact chemically and form gaseous or condensed species that will either enhance or retard the transport processes. Aerosols can form once the species reach the coolant channel. These aerosols are allowed to deposit on structural surfaces where they can reheat, revaporise, resuspend, and chemically interact with each other and with the surface itself. VICTORIA provides the fission product fuel release fractions, the relative quantities by vapour and aerosol, and the particular radionuclides that have been interacted chemically with, and possibly condensed onto, a surface. This output provides the input for source term calculations to the containment and thence to the environment. CONTAIN

The CONTAIN models provide the capability to calculate mechanistically the containment internal thermal-hydraulic conditions and the amount of radioactive matter that would be released to the environment if there were a leak from the containment. The code applies to design basis thermal-hydraulics and has undergone extensive development to attain a capability for determining containment threat under severe accident conditions. The code has been used for modelling PWR and BWR containments. The models available include those to predict the flows of mass and energy between containment compartments, the exchange of energy between the atmosphere and heat structures, the thermodynamic conditions, the distributions of aerosols, the decay and transport of fission products, the deflagration of hydrogen and carbon monoxide, boiling water reactor suppression pool behaviour, and engineered safety features, including sprays, fan coolers, and ice condensers.

258

Computer Codes

CONTAIN has been assessed against a range of experiments in the Battelle Model Containment, experiments in the LACE programme and HDR. CONTAIN includes other codes as modules which provide more detailed modelling capability for certain processes. For example the consequences of molten core-concrete interactions are included in various separate effects codes. CORCON provides thermal­ hydraulics and heat transfer. Fission product release is provided for by V ANESA and SOLGASMIX. Another example is direct containment heating (OCH). A code CORDE has been written specifically to model this phenomenon. 20.3.3 Separate Effects Codes

Many of the integrated and mechanistic systems discussed above have been built up from separate effects codes. They have been benchmarked against independent, specific separate effects codes. In some cases few experimental data exist e.g. a plausible comparison with a separate effects code provides a measure of confidence in the more general code. CFD Codes

Computational Fluid Dynamics (CFD) codes have been used to provide finer detail in flow modelling than is possible in the integral and mechanistic system codes. Two codes which have been used in this manner are the COMMIX code [20.35] developed by Argonne National Laboratory, USA and FLOW3D [20.36] developed by AEA Technology in the UK . These codes include a generall y rigorous fonnulation of the three-dimensional fluid dynamics equations for laminar and turbulent flow. The usual closure assumptions are included to allow for sub-grid scale and turbulence modelling. The codes incorporate a porous medium description for sub-grid scale structure (necessary for modelling the fuel pins in a reactor vessel). There is considerable flexibility in the prescription of friction factor and heat transfer coefficient correlations. Turbulence is modelled using the k-epsilon turbulence model with standard closure assumptions. It is believed that the physical models are adequate for most single phase flow applications: the codes are currently under development for two-phase flows. Cladding Oxidation and Melt Progression

The kinetics processes for cladding oxidation are modelled in the systems codes via simple kinetics correlations. The PECLOX code has been written in Gennany to provide a rigorous diffusion calculation for the oxygen profiles through an oxidising cladding. Such treatments are necessary to provide the detailed modelling of cladding in an oxidising or reducing environment. For melt progression, the PLUGM code allows much greater mesh refinement than would be possible in a full melt progression code (results are known to be sensitive to coarse noding) and it allows a variety of initial conditions to be considered economically.

Computer Codes

259

The PLUGM code is essentially a coupled, thennal-hydraulic computer model for the freezing of melt flow in a channel. That is, it models the time dependent flow of a molten material through a user-specified flow channel, taking into account heat transfer between the molten material and the channel wall, and possible constriction of the channel due to surface crust fonnation [20.37]. The bulk hydrodynamics is provided by a discretised momentum equation appropriate for quasi one-dimensional pipe flow. This is linked with an equation for the conservation of mass which neglects fluid velocities nonnal to the pipe axis, but includes the gradient in axial velocity along the channel. The momentum equation assumes a flat velocity profile and, as such, is applicable to turbulence flow problems. The code also models the process of liquid film deposition using empirically based correlations. In modelling the heat transfer, the PLUG M code allows the user to specify one of two models, on a mutually exclusive basis. The first of these is a conductive freezing model, where a crust is allowed to build up on the cool surfaces of the flow channel. Heat conduction through the crust and channel wall is calculated using a one-dimensional finite difference scheme to solve the heat conduction equation in the crust and channel wall, and crust growth/remelt is calculated via a modified Stefan condition. The second option is a bulk freezing model, where a crust is assumed not to fonn on cold surfaces. When the molten material cools to its liquidus temperature a solid particle fraction is calculated, allowing for the release oflatent heat which is assumed to vary linearly with the bulk melt temperature between its solidus and liquidus values. The channel is assumed to plug when this fraction reaches some prescribed critical value. Fuel Coolant Interactions

The IFCI computer code (Integrated Fuel-Coolant Interactions) [20.38]has been developed as a tool to provide researchers with a best estimate tool to study FCIs in reactor geometries. It has been developed based on known physical laws and the results of available experiments. An assessment of the code's perfonnance against available experimental data and parametric analyses at reactor scale have been carried out, e.g. the boiling and fragmentation models in IFCI have been assessed against the EJET series of boiling jet experiments perfonned at Sandia. Other codes developed for modelling aspects of steam explosions include CHYMES and CUL DE SAC . CHYMES was developed to model the pre-mixing phase. It has a 2-D 3-phase transient capability. The three-phases are melt, steam and water which are all assumed to be at saturation. CHYMES has been validated against tests in Oxford, France, Gennany and Italy. CUL DE SAC is a code for modelling the detonation phase. This code is limited to 1-D planes or spherical geometry. The phases include melt droplets, melt fragments and water/steam.

260

Computer Codes

Core-Concrete Interactions Heat Transfer and Thermal-Hydraulics CORCON

Under LWR severe accident conditions, the interaction between molten core materials and concrete provides a significant source term contribution to the containment. CORCON [20.39] is a computer code funded by the USNRC and developed at Sandia National Laboratories, USA, for modelling these interactions. The code predicts the behaviour of the system, including heat transfer, concrete ablation, cavity shape change, and gas generation. Later versions of the code include models for solidification of the melt and for its (non-explosive) interactions with coolant water. CORCON has received extensive validation against a set of experiments (the metallic melt dry tests in SURC and BETA programmes; metallic melt wet tests in the SWISS series; and several oxide melt dry tests in the ACE series. Chemistry

VANESA [20.2 1] calculates the thermodynamic equilibrium of a melt by considering a few key reactions. The important reactions involving fission products are then calculated by a perturbation method. This approach allows the releases to be calculated provided that the data defining the major reactions (the thermodynamic database) are well defined. SOLGASMIX [20.40] is a code that calculates the equilibrium composition of multicomponent, multiphase systems by the direct minimisation of the total Gibbs free energy. The system is defined by a thermodynamic database comprised of a set of chemical species, a stoichiometric matrix and thermodynamic data supplied by the user. In order to perform efficient calculations of the equilibrium chemistry during MCCIs, the CORSOL code was developed to provide a simple coupling between CORCON and SOLGASMIX. High Pressure Melt Ejection Direct Containment Heating

COROE [20.4 1] has been developed by the AEA at the Winfrith Technology Centre as a High Pressure Melt Ejection (HPME) module. The COROE module can be used in stand alone mode to predict the ejection of molten debris from a hole in the bottom of the pressure vessel and the subsequent transport of this material into the containment, Stand-alone COROE also has a simple, single-cell, Oirect Containment Heating (OCR) model; however, in order to make more realistic estimates of the containment loads due to OCH in a multicell containment, COROE has been interfaced to CONTAIN.

Computer Codes

261

The code has been validated against various experiments: Surtsey; HIPS/SPIT; CWTI; Winfrith Dispersal - Experiments; Berkeley Dispersal Experiments and the Berkeley "Cavity FCr' Experiment. REFERENCES

20. 1 20.2 20.3

20.4 20.5 20.6 20.7 20.8 20.9 20. 10 20. 1 1 20. 1 2 20. 1 3 20. 14 20. 1 5 20. 16

V H Ransom et al, RELAP5/MOD2 Code Manual, Volumes I and II, NUREG/CR43 12, EGG-2396, (August and December 1985, revised April 1987). R A Dimenna et al, RELAP5/MOD2 Models and Correlations, NUREG/CR-5 194, EGG-253 1 , (August 1988). Safety Code Development Group, "TRAC-PF1/MOD1 : An Advanced Best-Estimate Computer Program for Pressurized Water Reactor Thermal-Hydraulic Analysis", Los Alamos National Laboratory Report LA- I01 57-MS (NUREG/CR-3858, (July 1986). D Taylor, "TRAC-BDl/MOD1 : An Advanced Best-Estimate Computer Program for Boiling Water Reactor Transient Analysis", EG&G Idaho inc. report NUREG/CR36233 (April 1984). A Forge et al, "Comparison of Thermal-Hydraulic Safety Codes for PWR Systems", Graham and Trotman/CEC. J Burwell et aI, "The Thermal-Hydraulics Code ATHLET for Analysis of PWR and BWR Systems", Fourth Inter. Topical Meeting on Nuclear Reactor Ther. Hydrau., Karlsruhe FRG, (October 1989). H Hughes, R F Cameron, JE Sinclair and T J Haste, "Fuel Rod Behaviour during Transients: Part 1 : Description of Codes", ND-R -702 (S) ( 1982). R Hargreaves and D A Collins, "A Quantitative Model for Fission Gas Release and Swelling in Irradiated Uranium Dioxide" Journal British Nuclear Engineering Society Vol. 1 5, pages 3 1 1 -3 1 8, (1976). L F A Raven, "Comparison of HOTROD Code Predictions with PIE Data". ANS Meeting on Water Reactor Fuel Performance, Chicago, (May 1977). J Gamier and S Begij, "Ex Reactor Determination of Thermal Gap Conductance between Uranium Dioxide and Zircaloy-4", NUREG/CR-00330, (1980). L J Siefken, et al, "FRAPT6: A Computer Code for the Transient Analysis of Oxide Fuel Rods", EGG-CDAP-54 10, April 1977. T J Haste, "CANSWEL-2: A Computer Model of the Creep Deformation ofZircaloy Cladding under Loss-of -Coolant Accident Conditions", Part I - Model Description. UKAEA Report No. ND-R-8 14(S), (1982). R W Bowring and C A Cooper, "MABEL-2: A Code to Analyse Cladding Deformation in a Loss of Coolant Accident", AEEW-R I 530, (1983). K E Washington et aI, "CONTAIN 1 . 1 , Code for Containment Severe Accident Analysis, NUREG/CR-57 15, (July 199 1). "Source Term Code Package: A User ' s Guide (MODI )", NUREG/CR-4587, BMI2 1 38, Battelle Columbus Division, Ohio, USA, (July 21, 1986). R M Summers et aI, "MELCOR 1 .8.0: A Computer Code for Nuclear Reactor Severe Accident Source Term and Risk AssessmentAnalyses", NUREG/CR-553 I , SAND900364, R3, (October 1990).

262

Computer Codes

20. 17 MAAP 3.0B Users Manual Vol. 1 and 2, Fauske & Associates Inc., (April 199 1). 20. 1 8 R O Wooton, P Cybulskis and S F Quayle, "MARCH-2 Code Description and User's Manual, NUREG/CR-3988, report to U.S. Nuclear Regulatory Commission (September 1984). 20. 19 R K Cole Jr., D P Kelly and M A Ellis, "CORCON-MOD2: A Computer Program for Analysis of Molten Core-Concrete Interactions", NUREG/CR-3920 (August 1984). 20.20 M R Kuhlman, D J Lehmicke and R 0 Meyer, "CORSOR User's Manual", NUREG/ CR-4 173, Report to U.S. Nuclear Regulatory Commission (March 1985). 20.2 1 D A Powers, J E Brockman and A W Shiver, "VANESA: A Mechanistic Model of Radionuclide Release and Aerosol Generation During Core-Debris Interactions with Concrete (Draft)" , NUREG/CR-4308, SAND85- 1370, (June 1985). 20.22 R G Freeman-Kelly and R G Jung, "A User's Guide forMERGE", NUREG/CR-4 1 72, report to U.S. Nuclear Regulatory Commission, (March 1985). 20.23 H Jordan and M R Kuhlman, "lRAP-MELT2 User's Manual" , NUREG/CR-4205, report to U.S. Nuclear Regulatory Commission (May 1985). 20.24 H Bunz, M Koyro and W Schock, "A Code for Calculating Aerosol Behaviour in LWR Core Melt Accidents - Code Description and User's Manual" presented at a Workshop at EPRI, Palo Alto, California, (March 29-30 1982). 20.25 P C Owczaraki, A K Postma and R I Schreck, "Technical Bases and User's Manual for SPARC - Suppression Pool Aerosol Removal Code", NUREG/CR-33 1 7, Report to U.S. Nuclear Regulatory Commission (May 1983). 20.26 W K Winegardner, A K Postman and M W Jankowski, "Studies of Fission Product Scrubbing Within Ice Compartments", NUREG/CR -3248, PNL-469 1 , Report to U.S. Nuclear Regulatory Commission (May 1983). 20.27 D I Chanin et aI, MELCOR Accident Consequences Code System (MACCS), Vols. I, II and III, NUREG/CR-469 1 , SAND86- 1 562, Sandia National Laboratories, Albuquerque, NM (1990). 20.28 J Dufresne et aI, "Presentation of the ESCADRE System, Together with a Practical Application", International Symposium on Severe Accidents in Nuclear Power Plants, Sorrento, Italy (March 1988). 20.29 K J Abe of Atomic Energy Society, Japan, Vol. 27, No. 1 1 ( 1985). 20.30 "ESTER code- Contratde Recherche a Frais PartagesEntre la Communaute Europeenne de I 'Energie Atomique" - CISI Ingenierie - IKE Stuttgart I2 3870-89- 12 EL ISP F­ , (December 1989). 20.3 1 C Allison et aI, "SCDAPjRELAP5/MOD2 Code Manual, Volume: I, II and III, NUREG/CR-5273, EGG-2555, (June 1989). 20.32 K D Hocke et aI, "KES S-III-Advanced Modelling of Reactor Core Behaviour under LWR Severe Accident" 20.33 A Schatz et aI, "KESS-A Modular System for the Simulation of Severe LWR Accidents", EUfO. Simulation Multiconference, Nurnberg, FRG, (1990). 20.34 T J Heames et al, "VICTORIA: A Mechanistic Model of Radionuclide Behaviour in the Reactor Coolant System under Severe Accident Conditions, NUREG/CR-5545, (December 1992). 20.35 W T Sha et al, "COMMIX-1B: A Three Dimensional Transient Single-Phase Computer Program for Thermal-Hydraulic Analysis of Single and Multi-component Systems Vol 1 and 2, NUREG/CR-4348, ANL 85-42 (January 3 1 , 1986).

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20.36 A D Bums, D Ingram, I P Jones, J R Kightley, S Lo, N S Wilkes, "FLOW3D, The Development and Applications of Release 2, (June 1987). 20.37 M Pilch and P K Mast, "PLUGM: A Coupled Thermal-Hydraulic Computer Model for Freezing Melt Flow in a Channel, SAND82- 1 580, NUREG/CR-3 1 90, (1984). 20.38 M F Young, (1987), "IFCI: An Integrated Code for Calculation of all Phases of Fuel­ CoolantInteraction, NUREG/CR-5084, SAND87- 1048, SandiaNational Laboratories, Albuquerque, NM, September. 20.39 R K Cole et aI, "CORCON-MOD2: A Computer Program for Analysis of Molten Core-Concrete Interactions, NUREG/CR-3920, 1984. 20.40 G Eriksson, "Thermodynamic Studies of High Temperature Equilibria: XII SOLGASMIX, A Computer Program for Calculation of Equilibriwn Compositions in Multiphase Systems' , Chemica Scripta 8 (1975) 100. 20.4 1 B W Morris and G J Roberts, "User's Manual for CORDE, AEA-TRS-5033, (November 1990).

265

Chapter 21 CODE VALIDATION

21.1

Introduction

The aim of this chapter is to provide examples of analyses that have been carried out with some of the codes cited in the previous chapter for the purposes of code validation. Two main classes of experiments are included: thennal-hydraulic experiments severe accident experiments. It is not possible in the broad survey here to provide details of the systematic validation of

all the models and their interactions within even a single version of one code. Some specific

examples have been chosen from selected key areas. 21.2

Thermal-Hydraulics Experiments

The areas covered for thennal-hydraulics include: press uri sed faults, small breaks, large breaks. Two computer codes RELAP5 and TRAC have been developed under US NRC sponsorship to provide simulation of the reactor coolant system under transient accident conditions. In the UK, RELAP5 is used predominantly for pressurised faults and small breaks. TRAC is used for large breaks. The UK has had access to these codes under various agreements in exchange to providing assessments and making agreed improvements to the codes. Important examples of integral tests include the LOFT and LOBI experiments and specific examples ofRELAP5 and TRAC assessments against these experiments are described below. There are also numerous separate effects tests, the most important of these from the point of view of model validation have been documented recently in [2 1 . 1]. An important series was the ACHll.L . ES series of tests conducted at Winfrith. The purpose of these tests was to provide a demonstration of the reflood phase of a large break LOCA: in particular that a partial blockage would not lead to excessive local fuel temperatures.

266

Code Validation

2 1.2.1 Intact Circuit Faults

Intact circuit faults fall into various categories: Decrease in SG heat removal, Increase in SG heat removal, Decrease in primary coolant system flow, •

Reactivity abnormalities, Anticipated transients without Scram (ATWS).

Both LOFT and LOBI experiments have been used for the purpose of code validation of the thennal-hydraulic codes for intact circuit faults. Examples of LOFT tests that have been used for code validation of RELAP5 include: loss of feedwater (LOFW) with steam generator (SG) refill, LOFW with ATWS, loss of offsite and onsite power (LOOP) with ATWS. Examples of LOBI tests include: SG overfill, natural circulation, LOOP ATWS, small steam line breaks (SLB), asymmetric cooldown, large SLB, LOFW with SG refill. Steam Line Break

The steam line break poses one of the more extreme intact circuit faults. The main phenomena and causes of concern are: 1 . The initial SG inventory,

Code Validation

267

2. The changes in SG heat transfer as the SG blows down via the steam line, 3. Separator performance, 4. Reverse heat transfer from SGs, 5. Pressuriser response during the outs urge as the primary coolant cools down, 6. Boron and thermal mixing in the reactor vessel, 7. Reactivity feedback. By way of example, results are shown, Figure 2 1 . 1 , of a typical RELAP5 MOD2 prediction for one of the LOBI steam line break simulation tests (BT- 12). LOBI was a two loop (single and triple) full height electrically heated facility, see Chapter 4. The principal phenomena addressed in this test [21 .2], [2 1 .3] and [2 1 .9] include: •

•

•

•

The changes in SG heat transfer as the SG blows down via the steam line, Separator performance, Reverse heat transfer from SGs, and Pressuriser response during the outsurge as the primary coolant cools down.

The experiment was initiated from simulated hot shutdown conditions. The B T- 1 2 experiment was designed to investigate SG and primary system response to a large steam line break, in which liquid carryover occurs. The break was scaled to represent a 100 percent break in a steam line downstream of the flow restrictor. The depressurisation rate was therefore controlled by the flow restrictor rather than by the size of break. The faulty SG blew down over a period of about 200 s. The cooldown induced by the SG depressurisation resulted in minimum temperature of about 520 K in the cold leg of the affected loop, just before the SG blew dry and heat transfer was lost The primary depressurisation fell to about 13.0 MPa at the end of the test, 600 s. The pressuriser did not empty and the HPI's set point was not reached. There was evidence of significant carry-over of liquid, which was responsible for the smaller cooldown of the primary fluid. RELAP5/MOD2 calculations [2 1 .5] reproduced the course of the transient, the main features, and the timings quite well. The rate of depressurisation of the faulty SG and the volumetric break flow rate were calculated, although the shape of the calculated depressurisation curve differed from the data. The comparison with data suggests some overprediction ofliquid flow through the break, but this might be influenced by experimental uncertainty. The period of high steam flow persisted longer in the calculation.

268

Code Validation

7.0 I NTAG LOOP

6.0 5.0 to a..

�

'-'

w

a:: => I.I'l I.I'l W

--

R E LAP 5 CALC U LAT I O N EXP E R I M E NT

4.0 3 .0

a:: a..

2.0 1 .0

B R O K E N LOOP

o �--�--��==�==� o 1 00 200 300 400 500 600 700 TI M E, S E C O N D S

F1GURE 21.1 LOBI LARGE STEAM LINE BREAK TEST BT-12

This test therefore demonstrates the effect of SG heat transfer during blowdown on primary cooldown in an integral system transient It also allows the effect of the blow down on separator performance, and hence the effect of liquid carryover on the cooldown to be investigated. 21.2.2 Loss of Coolant Accidents

LOCAs are characterised by a decrease in primary coolant systems inventory. The LOFT programme was significant since it included integral LOCA experiments conducted in a reactor with real fuel. The complete sequences of loss of coolant accidents could be studied. Experiments performed included: small breaks in the hot legs, small breaks in the cold legs, -

large breaks in the cold legs.

Code Validation

269

The cold leg small breaks were carried out in the US NRC programme. The hot leg small breaks carried out later were carried out in the follow-on OECD NEA programme. The latter highlighted some inadequacies in early code versions and led to various code improvements. The large break in the cold leg represents the design basis accident for the PWR. Small Break LOCA Experiments

The principle issues of concern and phenomena include: 1 . The effects of reactor coolant pump operation, 2. The primary coolant inventory and the effectiveness of ECCS, 3. Plant recovery using secondary feed and bleed, 4. The effectiveness of natural circulation heat removal, 5. Pressurised thermal shock, 6. Accident diagnostic techniques, 7. Core recovery and dryout. To illustrate typical code prediction capabilities for small break LOCAs, analysis is shown below, of one of the LOFf small break tests (LP-SB-3) [21 .6] , [2 1 .7] carried out using the RELAP/MOD2 code. The particular issues addressed by this experiment included: The effectiveness of steam generator feed and bleed as an effective cooling procedure, •

Accumulator injection when only a low pressure differential to the primary circuit pressure exists.

This experiment was designed to provide data on various phenomena but did not aim to represent a particular accident sequence. The effectiveness of the steam generator operation was confirmed and accumulator injection was also demonstrated. RELAP/MOD2 analysis is described in [21 .8] . The predictions showed pump trip times to be correctly predicted and the primary coolant inventory and initiation of fuel cladding heat­ up was also calculated correctly. Accumulator injection was adequately predicted. The core heat-up rate during core uncovery was somewhat higher than measured. The predictions of the codes in simulating this experiment were in general good and the calculated parameters (break isolation time, primary coolant pump operation, auxiliary feed criteria, initiation of secondary bleed) were within the bounds of the experimental data.

270

Code Validation

Large Break LOCA Experiments

The main issues in large break LOCA concern: I . ECCS performance, 2. Core Thermal Response. Large break analysis code capabilities are illustrated by a lRAC-PFI/MODI analysis of a LOFT test (LP-LB-OI) [2 1 . 1 0] . The objectives of this test were: •

•

To determine the core response with minimum ECCS conditions, To include a rapid pump coast down and to maximise the fraction of the core that has not rewetted by the end of the blowdown.

The main features of the LOFT facility for modelling a large break LOCA in a four loop PWR were discussed in Chapter 4. The loop containing the break was modelled as one. The three intact loops were lumped together. The break was simulated by operating valves. The experiment selected LP-LB- 1 [21 .6], [21 .9] was the last of the large break - 200 % double­ ended cold leg break - transients to be performed in the LOFT facility and simulated the range of phenomena expected to occur in a large break loss-of-coolant accident with minimum safeguards emergency core cooling system assumptions. The experiment was initiated by opening the quick acting blowdown valves, so that the reactor quickly depressurised. The primary coolant pumps were then tripped. The reactor core was quickly shutdown, first by the core voiding and then by the insertion of the control rods which are tripped on low system pressure. After about 5 seconds, the loss of coolant reduced as the fluid approached saturation but the flow into the downcomer from the intact loop was not sufficient to produce a "bottom-up" flow of liquid into the core. Peak clad temperature attained was 1263K. The calculated values for the loop and vessel flows during the blowdown period (.... 1 7 seconds) were in good agreement with the experiment. At 1 7.5 seconds, the accumulator discharged and between 1 7.5 to 45 seconds there were marked flow and pressure oscillations around the intact loop. The accumulator liquid flowing into the downcomer reached the lower plenum while a saturated twcrphase flow from the lower plenum travelled up the downcomer and out of the break. This flow pattern was deduced from a comparison of the lRAC and experimental downcomer and broken loop fluid temperatures. After 40 seconds, when the accumulator tank emptied, nitrogen from the accumulator flowed along the ECCS injection line and into the intact loop cold leg. This forced extra water into the downcomer and the resultant surge of liquid into the core produced a rapid cooling of the fuel rods of .... lOOK at the highest power elevation, and a quenching of some of the cooler

27 1

Code Validation

peripheral parts of the core. Following this the liquid level in the core showed an oscillatory behaviour. The core was completely quenched by 70 seconds. lRAC calculations [2 1 .90] did not reproduce the intact loop flow and pressure oscillations seen in the experiment, but correctly predicted no direct downcomer bypass of the ECCS liquid. The refill of the vessel was therefore in good agreement with the experimental data and the initial reflood of the core began at about 32 seconds in line with the experiment. The calculated fuel rod cladding temperatures for the central fuel assembly - which did not quench during blowdown - were in good agreement with the experiment up to -45 seconds. The final quench was prolonged in the calculation, compared with the experiment. to about 100 seconds. Analyses of large break tests provided a better understanding of early rewet phenomena and the importance of three-dimensional flow patterns in the vessel. A separate effects test for reflood was conducted in the ACIDLLES test facility at Winfrith [21 . 1 1]. The purpose was to simulate the end of the accumulator discharge period in a postulated large break loss of coolant accident when the nitrogen, which is used to pre-pressurise the accumulator, enters the primary circuit. The resultant decrease in pressure drop between the accumulators and the pressure vessel causes an increase in the pressure at the top of the downcomer which in tum produces a surge of water into the core with subsequent oscillatory flow occurring between the core and downcomer. This test was the subject of a CSNI Internal Standard Problem ISP 25 [2 1 . 12] to assess advanced thennal-hydraulic code perfonnance for large LOCA.

* 1 04 30.0 .......

co

e::.

25.0

w

20.0

V"I V"I w

1 5.0

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1 0.0 5.0 0.0

400.0

800.0

TI M E (SEC)

FIGURE 21.2 "DR T31.S EXPERIMENT

1 200.0

272

Code Validation

This test provides a severe challenge to codes. The modelling of the non-condensable field presented problems which were overcome by a variety of means. Heat transfer during the initial surge of water into the core was generally well predicted although none of the participants were able to predict the correct timing for this period. Heat transfer at later times was generally over-predicted. The test provided some evidence for the adequacy of the current advanced codes in predicting the nitrogen injection phase of a large break LOCA. The main problems were considered to be associated with inadequate heat transfer packages in the codes causing vapour temperatures to be generally under-predicted. This coupled with problems with the core hydraulics resulted in an under-prediction of rod temperatures. Various assessments of the CONTAIN code as a design basis LOCA containment code, have been carried out via experiments in the HDR facility. Several International Standard Problem Code Comparison exercises have involved experiments in this facility. The purpose of the T3 1 .5 experiment [2 1 . 1 3] was to provide an assessment of the CONTAIN code for the design basis large break LOCA. CONTAIN calculations were carried out by Tills [21 . 14], Figure 2 1 .2. This experiment is mainly relevant to design basis but also serves to provide more general validation of the CONTAIN thermal-hydraulic model for a compartmentalised containment The HDR facility is the containment building of a disused reactor: it was about the correct height for a large modem PWR but was about one third in diameter. The facility had a large open dome representative of a large dry PWR but the lower level was more compartmentalised. The CONTAIN calculations gave a good prediction of the transient and long-term pressurisation, Figure 21 .2. Forced convection heat transfer was the important mechanism during the blowdown. The airborne water transport modelling was important for good prediction of the dome temperature during this phase. HDR Experiment 1 1 .2 represented a small break LOCA followed by light gas injection. The objectives were to simulate a LOCA in the upper part of the containment together with hydrogen production resulting from zirconium oxidation. CONTAIN calculations were carried out for this test also as part of an international standard problem exercise. These are described in [21 . 1 5]. The main features [2 1 . 16]. [2 1. 1 7] were a distinct stratification with almost no steam penetration in the lower containment rooms with most of the light gas contained in the dome. The early part of this test was relevant to design basis small break LOCA. To provide a better insight of the phenomena, CFD calculations were also carried out using the FLOW3D code.

Code Validation 21.3

273

Severe Accidents

The main characteristic of many severe accidents is that the core becomes uncovered for an extended period. The main aim of experiments has been to provide understanding of the phenomena occurring in a degraded core and the implications for reactor coolant system integrity and on the behaviour in the containment if the RCS boundary is breached. A wide range of phenomena are involved and the interactions are very complicated. Most severe accidents start with a thennal-hydraulic phase, and, as discussed in the previous chapter the principal predictive computer codes have been developed. Codes were originally developed for purely the thennal-hydraulic phase. The new models allow the fuel rods to degrade at high temperature. 21.3.1 Early Phase Core Degradation

The key outputs required from assessments of the thennal hydraulic response and core damage up to vessel failure include: fuel and primary circuit temperature, hydrogen production rate, timing of debris fonnation, debris quantity and composition. For early phase the main issues and phenomenology relate to: •

Pre
•

Coolant boil-off,

•

Natural circulation,

•

Quench,

•

Component heat transfer,

•

Oxidation kinetics,

•

Clad ballooning.

•

Control rod behaviour.

•

Effects of grids,

•

Fuel/cladding dissolution,

274 •

Code Validation

Zircaloy/U02 melt relocation, Blockage formation, Ceramic melting, Debris formation.

The early phase melt progression integral test data arose from various USNRC founded programmes i.e: PBF, ACRR, NRU and the last experiment in the LOFf series LP-FP-02. The data were used for various assessments of the SCDAP stand-alone code above and the SCDAP/RELAP5 coupled code as the code matured. The main objectives of the LP-FP-02 experiment [21 .6] , [2 1. 1 8] were: 1 . To provide data on melt progression from an in-pile test that was driven by decay heat. (The other in-pile tests were fission power driven), 2. To determine the fission product release from fuel during a severe fuel damage scenario and the transport of these fission products in a vapour/aerosol environment. The thermal-hydraulic conditions for the experiment were aimed to simulate those in a reactor if a break occurred in the low pressure injection system outside the containment with failure to isolate the system. ECCS injection is delayed until after the system has depressurised and fuel damage has occurred. The SCDAP/RELAP code predictions reproduced quantitatively the principal thermal­ hydraulic processes in the experiment, Figure 2 1 .3. However quantitatively the results were very sensitive to the relative distribution of fuel between the centre fuel module and the peripheral channels. Concerning the melt progression the calculations showed a lower blockage consisting of control rod material, that had relocated to the bottom of the bundle and a higher blockage of Zr-u-o eutectic material. The presence of two blockages is consistent with the experimental data. The hydrogen calculated during the transient prior to the quench was consistent with the estimated experimental figure, however the overall hydrogen prediction tended to be under­ estimated due to the difficulties in dealing with hydrogen production in the reflood phase. More recently priority has been given to PHEBUS and CORA amongst the melt progression experiments used for code validation. Collectively these experiments include most of the phenomena associated with early phase melt progression, in particular hydrogen production, melt relocation, ballooning, control rod material relocation and grid related phenomena etc.

275

Code Validation is

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The PHEBUS SFD B9+ experiment has recently been the subject of a CSNI standard problem. The PHEBUS SFD overall programme is described in [2 1 . 19]. The B9+ experiment was a nuclear heated small bundle melt progression test aimed at studying materials interactions in

276

Code Validation

the early phase of a severe accident. The test rig consisted of a bundle of 2 1 PWR -type fuel rods surrounded by a Zircaloy liner and zirconia shroud. There were three main phases in the high temperature transient: a heat-up phase in flowing steam to obtain an oxide thickness axial profile. reaching bundle temperatures up to 2 1 OOK. a helium flow phase where the power was increased to raise peak temperatures to about 2700K. causing substantial bundle degradation. a helium flow cooling down phase. With careful modelling SCDAP/RELAP5 gives good agreement with the observed temperature history. particularly regarding the timing of the oxidation excursion and also the hydrogen production [21 .20] . Limited amounts of uranium dioxide/Zircaloy interaction. fuel melting and slumping were subsequently calculated during the second phase of the test 21.3.2 Natural Circulation

Buoyancy driven flows are an important phenomenon in nuclear reactor high pressure accidents. Natural circulation can provide a mechanism for transferring heat away from the core region up to the cooler parts of the primary circuit in PWRs. The main issues are : the threat to primary circuit integrity due to overheating of the pipework. the effects of melt progression which could influence the threat to the lower head. In the first instance experiments carried out by Westinghouse have shown that complex natural circulation patterns involving counter-current flow in the hot leg can develop. Some details are given in [2 1 .2 1 ] . These data have been used for validation of the COMMIX code for buoyancy driven flow prediction under LWR severe accident conditions. To understand better the characteristics of in-vessel natural circulation. experiments have been conducted at the University of Stuttgart [21 .22] which consider the buoyancy driven flow in rod bundles. The main purpose of these experiments is: to provide data for in-vessel code validation. to provide flow visualisation. The data have been used to validate the 2D thermal-hydraulics code FRECON which will be used as a sub-module in KESS. The codes referred to above are computational fluid dynamics codes which provide a fine numerical grid for good resolution of the flow physics. The strategy for validation of the

Code Validation

277

system codes for natural circulation has been to benchmark these system codes against the CFD codes, once these CFD codes have been validated against experiments. 21.3.3 Fission Product Transport in the Reactor Coolant System

The main requirement is to calculate the release and transport of fission products from the fuel to the remainder of the RCS and provide the source tenn to the containment. The principal phenomena/safety issues fall into three main categories: Fission product release from the fuel. Release rates, timing and magnitude of release. Aerosol behaviour. Particular phenomena include deposition, agglomeration, nucleation effects of bends, resuspension, release rates, timing and location. Chemistry interactions. Vapour/structure, vapour/aerosol, vapour/vapour. Chemical species, interaction rates, location etc. Nucleation. Effects of chemical species. Data are available from various integral and separate effects test programmes including Oak Ridge National Laboratory (ORNL) [21 .23], Marviken [21 .24],LACEandrecently FALCON [21 .25]. VICTORIA code assessments have been made against some of these data. Predictions of release from one of the ORNL experiments have been made using the VICTORIA code. The ORNL experiments were a series of tests for investigation of fission product release from intact fuel at high temperature i.e. greater than 2000K in chemically reactive atmospheres of steam and hydrogen. The experimental apparatus was an induction furnace: optical pyrometers were used for the temperature measurement. Fission products were sampled in thennal gradient tubes. The main purpose of the tests was to quantify the release of fission products from representative LWR fuel at high temperatures under conditions relevant to severe accidents. Results[21 .27] show caesium and iodine fractional releases predicted by the code and compared with the experimental data. Agreement between calculation and experiment during the transient release was reasonable. The Marviken tests were designed to investigate the transport of fission products and structural materials in the primary circuit many of the experiments which were carried out used simulant core materials and fission products but useful data were obtained for code assessment. The experiments were also close to full scale. The experimental facility is shown in [2 1 .27] . The simulant fission products and other core materials were injected in various sites together with the steam and carrier gases. The objective of the tests was to generate a large data base on the transport and attenuation of

Code Validation

278

aerosols and volatile fission products in representative LWR primary coolant circuits under severe accident conditions. Figure 2 1 .4 shows some results for Test 7 [2 1 .28], [2 1 .29] where the injection site was at the bottom of the reactor vessel. The simulant fission product species were caesium hydroxide, caesium iodide and tellurium. Overall the agreement is reasonable but some discrepancies for individual components did arise (not shown here). 21.3.4 Core-Concrete Interaction

The key phenomena can be characterised as: concrete ablation (vertical and horizontal), composition of the off-gas, state of chemical equilibrium of melt, layer structure of the debris/melt, profile of melt temperature, heat fluxes particularly upward heat flux from radiative cooling, quenchability of the melt by water. Assessments of the CORCON code have been made via comparisons of tests in the SURC, BETA and ACE programmes. A representative calculation is discussed below for the BETA 5. 1 experiment. The BETA 5. 1 test [2 1 .30] used siliceous concrete with a stainless steel melt and Zircaloy addition. Zircaloy is highly reactive and there are large chemical and energy releases associated with zirconium oxidation, both of which subsequently affect the rates of concrete ablation and gas release. The objectives of Test 5.1 were to produce data for code validation for 2D behaviour with Zircaloy addition. CORCON predictions for crucible erosion are shown in [21 .3 1 ] . Although there are differences through the transient, the final values after the cooling period are in reasonable agreement. Production of hydrogen, steam, carbon dioxide and methane also occur: all were consistent with experimental data except the methane prediction. The SURC-4 test [2 1 .32] involved the interaction of a melt of stainless steel containing fission product simulants with a basaltic concrete. The experiment was the subject of a CSNI standard problem. A comparison of the experimental results with predicted releases from single layer CORSOL calculations and stand alone SOLGASMIX calculations has been made. Apart from Mo, the predicted releases for single layer CORSOL and stand alone SOLGASMIX are similar and in reasonable agreement with the values measured over the sampling period.

Code Validation

279

6 5

Pressuriser

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--

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___

volumes

Injection Point

I'

1

(a) SCHEMATIC REPRESENTATION 1 . 0 �-+----�----��--r---+---�-

0.9

--

Experiment

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(b) TOTAL I DEPOS ITI ON

FIGURE 21.4 MARVIKEN-V TEST 7

Code Validation

280

Discrepancies between CORSOL and the experimental data which occurred cannot be explained in tenns of the different concrete ablation rates, since over the sampling period they were in good agreement The predicted release of Mo could have been affected by the overpredicted release of the alkali metals Na and K. 21.3.5 Containment Phenomena

The main areas concern thennal-hydraulics, aerosol behaviour and fission product transport and decay. The important key phenomena include: heat conduction is structures, heat and mass transfer to structures, heat transfer in water pool s, humidity/airborne condensation, aerosol agglomeration, condensation (insoluble) on to aerosols (soluble), aerosol deposition, sprays, fan coolers, fission product decay heating, chemistry, inter-cell flows, light gas distribution, combustion. Two CEC code comparison exercises have been based on the DEMONA B3 experiment in the Battelle Model Containment. This is a 640 m3 concrete model containment with an open space dome and a lower containment consisting of a number of room s. The facility is equipped with aerosol gases which can supply large quantities of particles of different compositions. The main purpose of the B3 test was to address condensation of steam on to airborne aerosols [21 .33] followed by a period during which aerosol depletion occurred via diffusiophoresis to the walls.

Code Validation

28 1

Two CEC comparison exercises was carried out for this experiment. CONTAIN calculations have been perfonned for the thennal-hydraulic conditions [21 .34] and the aerosol behaviour [21 .35]. REFERENCES

21.1 21.2 2 1 .3 2 1 .4 2 1 .5 2 1 .6 2 1 .7 2 1 .8 2 1 .9 2 1 . 10 21.1 1 2 1 . 12 21.13 2 1 . 14 21.15 2 1 . 16 2 1 . 17 21.18

"Separate Effects Test Matrix for Thennal-hydraulic Code Validation", NEA/CSNI R (93) 14, Volumes 1 and 2 (December 1993). "Experimental Data Report", LOBI-MOD2 Test BT- 12 (Steam Line Break), JRC Ispra Communication No. 42 1 7. C Addabbo and G Leva,"Quick Look Report", LOBI-MOD2 Test BT- 12, JRC Ispra Communication No. 4340, (December 199 1). C Addabbo. LOBI Seminar, 3 1 March - 2 April 1 992, ECSC-EEC-EAEC Brussels, Luxembourg 1992. A J Smethurst, "Post Test Analysis of LOBI BT-12 Test using RELAP5/MOD2", AEEW R 2645. J Fell and S M Madro, "An Account of the OECD LOFT Project", OECD LOFT-T3907, (May 1 990). A Alemberti, "Comparison Report on OECD LOFT Experiment LP-SB-3", OECD LOFT-T-3606, (February 1987). A Alemberti, "Experimental Analysis and Summary Report on OECD LOFT Experiment LP-SB-3", OECD LOFT-T-36OS. J P Adams and J C Birchley, "Quick Look Report on OECD LOFT Experiment LP­ LB- l ", OECD LOFT-T-3504, (February 1984). P Coddington, "Analysis of LOFT Experiment LP-LB- l using the TRAC-PFI­ MOD I Code", AEEW R2039, Winfrith, (January 1986). M K Denham et aI, "Achilles Unballooned Cluster Experiment, Part 1 , Description of The Achilles Rig. Test Section and Experimental Procedures", AEEW R2336, (1988). B J Holmes, "ISP 25 Comparison Report", AEA-TRS-I043, NEA/CSNI/R(9 1 ) 1 1 , (February 199 1). H Karwat, "ISP-23: Rupture of a Large Diameter Pipe in the HDR Containment", CSNI Report No. 160, Volumes 1 and 2 ( 1989). J Tills, "Selected Post-Test CONTAIN Calculations for the International Standard Problem 23", Viewgraphs from Presentation at the ISP-23 Workshop, Garching, (45 April 1989). L Valencia, "Comparison of Submissions to the PHDR Blind Calculational Exercise on HDR E l 1 .2", PHDR Report to Participants (1990). P N Smith and P Ellicott, "The UK ' s CONTAIN Submissions to the PHDR Blind Calculational Exercise on the Light Gas Distribution Test E 1 1 .2 in the HDR Facility" , AEA TRS 5059 (1990). PN Smith and P Ellicott, "A UK Analysis of Light Gas Distribution Experiment E 1 1 .2 in the HDR Facility", Proc. Seminar on Containment of Nuclear Reactors, Shanghai, 14- 15 August 199 1 , SMIRT 1 1 ( 199 1). M L Carboneau et aI, "Experiment Analysis and Summary Report for OECD LOFT Project Fission Product Experiment LP-FP-2", OECD LOFT-T-3806, (June 1 989).

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2 1 . 19 B Adroguer, A Commande and C Rongier, "ISP-28: Phebus-SFD B 9+ Experiment on Degradation of a PWR Core Type; Preliminary Comparison Report" , CEA Cadarache, Note Technique DRS/SEMAR 16/9 1 , Note Phebus CSD 122/9 1 , (May 199 1). 21 .20 T J Haste, J N Lillington and A J Lyons, "International Standard Problem 28: Summary ofUK Semi-Blind Calculations of the PHEBUS SFD B9+ Melt Progression Experiment using the SCDAPjRELAP5 Code", AEA-TRS-5058, (November 1990). 2 1 .21 W A Stewart et al, "Experiments on Natural Circulation Flows in Steam Generators during Severe Accidents", ANS/ENS Topical Meeting, San Diego, California, (February 2-6 1986). 2 1 .22 R. Kulenovic, S RosIer, M Groll,J Un fried and M Biirger, "Experimental Investigation of Natural Convection in a Vessel with Rod Bundle Configuration and Comparison with Results from the 2D- Thermal-hydraulics Computer Code FRECON" 2 1 .23 M F Osborne, J L Collins, R A Lorenze and R V S train, "Fission Product Behaviour in Tests ofL WR Fuel Under Accident Conditions", Proc. Int Symp. on Source Term Evaluation of Accident Conditions, Columbus, Ohio, IAEA-SM-28 1 (Oct 28 - Nov 1 1985). 21 .24 J Collen, H Unneberg and D Mecham, "Overview of Marviken Experimental Procedures", Proc. ANS Meeting on Fission Product Behaviour and Source Term Research, Snowbird, Utah, (July 1 5- 19 1984). 21 .25 A M Beard, C G Benson and B R Bowsher, "Fission Product Vapour-Aerosol Interactions in the Containment: Simulant Fuel Studies", AEEW-R 2449, (December 1988). 2 1 .26 M F Osborne, J L Collins, R A Lorenze, J R Travis, C S Webster and T Yamashita, "Data Summary Report for Fission Product Release Test VI- I ", ( 1989), NUREG/CR5339. 21 .27 D A Williams and H S Bond, "Analysis of the ORNL VI Experiments using VICTORIA" (June 199 1), AEA RS 5 1 39. 2 1 .28 MARVIKEN-V Aerosol Transport Tests, Test-7 Results, MXE-207. 2 1 .29 H S Bond and N A Johns, "Analysis of the Marviken-V ATI Experiments using VICTORIA", (June 199 1), AEA RS 5200. 2 1 .30 H Alsmeyer and M Firnhaber, "Specification of the International Standard Problem ISP-30: Beta V5. 1 Experiment on Melt-Concrete Interaction", (September 1990). 2 1 .3 1 L D Howe, R J Humphreys and P M Keeping, "Assessment of the CORCON Code Version 2.04 and Preliminary Version 2.05 against BETA V5. 1", PWR/SATRG/ P(9 1)E43 [RS(9 1)Nl l (Revised)] , (February 199 1). 2 1 .32 E R Copus, R E Blose, J E Brockman, R D Gomez and D A Lucero, "Core-Concrete Interactions using Molten Steel with Zirconium on a Basaltic Basemat The SURC4 Experiment", NUREG/CR-4994. 2 1 .33 H Bunz, M Koyro, W Schock, D Haschke, A Fromentin, R. Taubenberger, T. F. Kanzleiter, Th SchrOder, G. Weber, H. Ruhmann, M. Fischer, "DEMONA Experiment B3, 28. 1 1 . 1984, Short Information", Battelle Frankfurt Report 2 1 .34 J Gauvain, "Post-test Calculation of Thermal-hydraulic Behaviour in DEMONA Experiment B3 with Various Computer Codes used in EC Member States", EUR 12 197 EN (1989). 2 1 .35 W Schock, "Post Test Calculations of Aerosol Behaviour in DEMONA Experiment B3 with Various Computer Codes used in CEC Member States", EUR 1 1374 EN (1 988).

283

Chapter 22 THE ACCID ENT AT THREE MILE ISLAND UNIT 2 : IMPLICATIONS FOR MODEL DEVELOPM ENT 22.1

Introduction

The Three Mile Island - Unit 2 (TMI-2) pressurised water reactor accident occurred on March 28, 1 979. During the accident a large amount of molten core material (approximately 20 tonnes) relocated from the core into the lower head. The accident is unique (fortunately) in PWR operating history but has provided the only available data on plant-scale degraded core behaviour. The main purpose of this chapter is to summarise some lessons learned on severe accident phenomenology and the implications for model development. Much of the research over the last decade has been carried out within programmes initiated by the USOOE and OECD/ NEA community. The OECD/NEA and USOOE funded international programme for TMI-2 investigation [22. 1 ] was based on two parallel enterprises: (a)

an International Analysis Exercise and;

(b)

a Sample Examination Programme.

The International Analysis exercise was initially mainly concerned with thermal-hydraulic system code analysis of the earlier phases of the accident Later, attention was turned to the heat-up and recovery phases, with the application of more separate effects codes to provide greater understanding of individual events. The Sample Examination Programme was concerned with the identification of materials in various parts of the vessel together with quantifying thermophysical data (e.g. thermal diffusivity, specific heat, porosity, density etc) for the purposes of model development and verification. More recently the OECD/NEA have formed a consortium to examine samples from the lower head itself to learn more about the extent of damage and the margin to vessel failure. This Chapter includes a review of the final state of the plant and the accident scenario on the basis of findings from the defuelling programme, from analysis of core samples and from examination of the vessel itself. The end-state core contained a large coherent central blockage, together with a substantial quantity of debris in the lower head. Fission product

The Accident at Three Mile Island

284

release and transport analysis and infonnation are included in this summary: mass balance for most of the significant species has now been obtained. The paper considers the completeness of current understanding of TMI-2. The 'lessons learned' from TMI-2 are many and varied and this chapter in no way claims to be exhaustive. TMI-2 provides important data on debris attack of the lower head. Other examples of important data include the phenomenon of hydrogen blocking of the steam generators, the nature of the blockage, the constituency of the surrounding crust and mode of crust failure, and the lessons learned on any threat to primary circuit integrity arising from the severely degraded core conditions. The paper includes a summary of both integral systems analysis and some of the important separate effects analysis that has been carried out. Although the accident is not appropriate for code benchmarking per se, it has provided a useful vehicle for code assessment and for overall understanding. Melt progression data at plant scale are unique. The system codes were found to be in reasonable agreement for the early thennal-hydraulic phase but large areas of disagreement are seen in the melt progression phase predictions. Part of this disagreement results from uncertainties in the thennal-hydraulic boundary conditions, a point discussed further below. 22.2

Accident Scenario

The accident started [22.2] with a technician attempting to clear a resin blockage in the pipework of the condensate polishing plant. The condensate pump was turned off which supplied water to the condenser. The reactor and turbine were then tripped, shutting off the main feed pumps and initiating the auxiliary feed pumps. The power operated relief valve on the pressuriser had also stuck open, having opened to relieve the primary pressure following the turbine trip. The operators were not aware of this situation and made various errors in their management of the plant transient, by then a loss of coolant accident. The most serious error was their failure to realise that the pressuriser level was high due to a steam bubble that had fonned in the head of the reactor vessel. The ECCS was cut off despite the fact that the inventory was reducing and the core subsequently uncovered. Substantial core damage ensued before cooling was re-established by operation of the B loop pump (TMI-2 had two primary loops designated A' and 'B '). It took about 1 6 hours to get the plant to a stable cooling mode and about a month before a long tenn stable natural circulation state was achieved. I

Certain key events perfonned by the operators, or deduced from on-line instrumentation measurements enable the accident to be divided into various separate phases for the purpose of analysis. These events are shown on the pressure trace, see e.g. [22. 1]. Phase 1 : Small Break LOCA (0-100 min)

Accident initiation occurred due to a secondary cooling system problem. Primary coolant inventory then decreased due to PORV flow, and insufficient water make-up. Liquid was held-up in the pressuriser. By 100 min, neither reactor coolant pump was operating.

The Accident at Three Mile Island

28 5

Phase 2: Core Boil-off and Initial Damage (100-1 74 min)

Core uncovery took place during this period. The system continued to lose primary coolant and depressurise until - l40 min. when the pressuriser block valve was manually closed. The block valve was cycled subsequently to maintain system pressure. Ballooning rod failures occ urred during this time. inferred from the containment radiation monitor response. Cladding oxidation. liquefaction and melt relocation followed as the temperature rose. The response of incore instrumentation and source range monitors strongly imply that a region of partially-molten core materials formed by the time of the B loop pump transient -174 min. Phase 3: Pump Transient and Continued Damage (1 74-200 min)

The B loop pump was operational from about 174-1 80 min producing a surge of cold water into a highly degraded core. The effects of this surge are being studied in detail. It is postulated to have caused: (a)

fragmentation of rods to produce the upper core debris bed.

(b)

some damage to the upper plenum assembly. and

(c)

the sweep out of materials found in the primary system.

It is probable that continued heat-up. fission product release and substantial further oxidation occ urred during this period. High pressure injection started at 200 min.

Phase 4: HPI Injection and Relocation to Lower Plenum (200-300 min)

Indications from the primary system pressure monitor. the self-powered neutron detectors and the source range monitors indicate that the primary core relocation event occurred between 224 and 226 min. The postulated flow path will be discussed in more detail later. Interaction of core material with core support structures was found to have occurred. 22.3

Thermal-Hydraulic System Analysis

A number of code predictions were produced as contributions to the OECD/NEA and USOOE Analysis Exercise [22.2] . Mechanistic codes used included CATHARE/ICARE and SCDAP/RELAP5: engineering codes used included MAAP and MARCH3. Key parameters for comparison included: Reactor Coolant System Pressure Hydrogen Generation

The Accident at Three Mile Island

286

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Core Temperatures Pressuriser Level Core Liquid Level. SCDAP/RELAP5 predictions [22.2] for some of these parameters are shown in Figure 22. 1 . One problem was that there was considerable uncertainty in the thennal-hydraulic boundary conditions. Indeed several key boundary conditions concerning the water inventory were not recorded such as the high pressure injection make-up, let down and emergency feed water flows. These data are crucial for accurate prediction of the core heat up phase. Since there was this uncertainty in the boundary conditions, the core results tended to deviate from the recorded data. However, even with the same imposed boundary conditions, there were quantitatively different predictions for certain key parameters e.g. RCS pressure and core temperatures. This reflected a significant code 'user effect'. Concerning the thennal-hydraulics modelling during the core heat-up phase two principal points emerged: uncertainties in the water availability to a degraded core have a large effect on the temperature excursion. A key uncertainty in TMI-2 was the HPI rate. hydrogen blocking of the steam generators is significant during the Zircaloy oxidation excursion phase. The pressure increase observed dwing the heat up phase correlated well with the calculated time of rapid oxidation. 22.4

Upper Vessel Structural Temperatures.

An interesting question that has been posed following in-vessel investigations concerns the

limited extent of upper vessel damage and that the primary circuit remained intact despite severely degraded core conditions. Several mechanisms have been proposed to explain the lack of damage to the upper plenum. These include ballooning, hydrogen stratification in the upper head, pressuriser drainback, and timely operation of the B-loop pump which tenninated the initial heat-up phase. Calculations [22.3] imply that the upper grid plate would have been close to its melting point, just prior to the pump transient Ballooning reduces the rate of heat-up of the upper plenum, at the expense of increased core heat-up, but calculations suggest that this was not a significant mechanism in TMI-2 particularly as a bypass region would have existed once the core melt progression process started. The codes also do not show the existence of a stratified region of hydrogen (though it should be noted that the relevant models have not been assessed against experimental data) which would have resulted in lowering upper plenum temperatures.

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The Accident at Three Mile Island

It is further claimed that the pressuriser drainage was probably counter-current flow limited, though this issue has not been settled. The main reason that the upper grid plate above the core in the upper plenum did not completely melt and that the temperatures in the upper plenum remained low, would seem to be the timely operation of the pump, with some small contribution to temperature reduction from ballooning. It was demonstrated [22.4] that 3D buoyancy-driven flow probably existed during the heat­ up phase, induced by the upper plenum geometry. This flow is consistent with the observed limited bi-polar upper grid plate damage, Figure 22.2. Further damage could have been caused by the B-loop pump transient Concerning lessons learned for severe accidents in PWRs the following points emerge. With regard to the potential for damage to upper plenum structures, in-vessel steam and hydrogen flows may be complex, variable in composition and three dimensional. The nature of these flows and therefore the potential for damage may be dependent on details of the particular plant geometry. 3D buoyancy -driven flow patterns have been shown to exist in TMI -2 arising as a consequence of the upper plenum cylinder geometry. For the depressurising sequences, ballooning may have initially a reducing effect on natural circulation although its significance will depend on the extent of flow bypass. Core temperatures are hotter and plenum temperatures colder than in the unballooned case. However if degraded core conditions are reached there will be an increased potential for flow bypass, diminishing the impact of any previously ballooned rods. Blockages produced in the core may induce 3D effects in buoyancy-driven flows over the whole vessel volume. The 3D nature of the flows calculated in TMI-2 is substantially enhanced if a core blockage is taken into account The introduction of water into the TMI-2 degraded core had a number of important effects. These included limiting the temperature excursion, but not after increased hydrogen production and effect on core degradation.

22.5

Final State of the Plant

The major physical damage to the plant was restricted to the reactor vessel internals [22.5] , Figure 22.3. Some small amounts of fuel debris have been found elsewhere e.g. in the pressuriser (-10 kg), B loop steam generator (-50 kg) and the decay heat line (-50 kg). The upper plenum assembly received no major structural damage, apart from limited oxidation in two damage zones on the underside of the upper grid plate. The implications on the flow patterns and gas composition were discussed above. The end-state core region consisted of a voided region at the top, below which there was a loose debris bed resting on a resolidified mass of material supported by fuel rod stubs, Figure

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22.3. The debris was found to be principally oxidised U, Zr and 0, with small amounts of oxidised structural materials, and traces of control materials. The quench of the degraded core during the B-loop pump transient probably produced a bulk material relocation process in addition to existing candling processes. It would have produced fragmentation of highly oxidised standing fuel rods, caused the material to collapse and form a debris bed. The centre of this mass consisted of ceramic material, mainly U, Zr, 0 with small amounts of structural and control materials and metallic material, principally Fe-Ni and Ag-In phases and some mixed ceramic and metallic material.

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FIGURE 22.2 AZIMUTHAL LOCATIONS OF MAIN FLOW HOLES IN THE PLENUM CYLINDER IN RELATION TO PLENUM ASSEMBLY DAMAGE PATTERN

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The upper and peripheral crusts consisted of similar ceramic material. with some solid V02 particles. The metallic material in the upper crust consisted of principally Fe-Ni and Ag-In­ V phases. The metallic material in the peripheral crust was principally Zr-Fe-Ni-Cr, Zr-Ni­ In and Ag-In phases. Peak temperatures of ceramic material in the upper and lower plenum debris, the upper and peripheral crusts and the central, previously molten region were at least 2700K and may have reached 3 1 OOK. The lower crust was more metallic. The metallic material consisted of Zr-Fe-Ni-Cr, Zr-Ni­ In and Ag-In phases surrounding intact fuel pellets. The peak temperature of the lower crust was cooler, between 1400 and 2200K. On the east side of the core, there existed previously molten material in place of four assemblies. There was a large hole in the baffle plates and some material also solidified in the bypass. There was a large quantity of previously molten material in the lower head, particle sizes varying from large rocks to small particles. The nature of this material is considered later.

22.6

Melt Progression

After the core was uncovered. at about 100 mins from the start of the accident, the oxidation excursion escalated and core degradation ensued until ECCS injection filled the vessel at about 200 mins. Early phase melt relocation involving initial melt relocation from intact rods and followed by blockage formation was the first relocation event (or sequence of events) that occurred. A major second relocation event was deduced to take place between 224 and 226 mins, detected from instrument alarms. Some pressure increase was also observed a few minutes subsequently associated with melt interaction with water in the lower vessel. Relocation is thought to have occurred when the blockage crust failed near its top and on the east side of the core. Molten material flowed laterally and relocated through the peripheral subassemblies and the core bypass region. Access to the core bypass region was through holes in about three baffle plates. The melt is thought to have entered the lower head through the periphery of the core support assembly. Resolidified material was found on the circumferences of these structures. The damage to the core former plates in the bypass and the core support assembly was not extensive apart from a 1 5 cm diameter hole in the lower grid. The melt would appear to be a fully oxidised ceramic consisting of V02 and zr02 • Analysis of melt that remained in the core region has indicated that peak temperatures may have

The Accident at Three Mile Island

29 1

reached 3 1 OOK the melting point of U02. It is thought that the temperature of the melt that relocated was at a temperature in the region of 3000K . In summary therefore, melt progression in TMI-2 was a multi-stage process involving initial melt relocation from intact rods, blockage fonnation, blockage heat-up, crust failure and further melt relocation. These processes are simulated in the advanced mechanistic melt progression codes. Many earlier simpler models, e.g. MARCH, assumed a single core-wide stage process which is not supported by the evidence in the layered structure of the particles found in the lower plenum in TMI-2. This suggests a prolonged candling sequence and not a single core-wide evenl A large molten pool can exist encased in an oxidic crust. The existence of such a pool encased in a crust is unique to TMI-2. No comparable blockage has been found in the bundle experiments of limited radial scale, perfonned to date. Crust failure did not occur at its lowest point. The possibility of crust failure at the periphery of the surrounding crust containing a molten pool , as apparently occurred in TMI-2, and not at the centre of the lower crust, had not been considered previously. Certain degraded core configurations are uncoolable. TMI-2 provides evidence of this important observation. In the accident, the molten core probably continued to heat up following reflood of the core, and the major melt progression event took place under water. Few experiments to date have addressed quench issues and none at plant scale and therefore whether degraded core blockage configurations are likely to be coolable or uncoolable. 22.7

Debris in the Lower head

The top debris present in the lower head was found to be in a range of sizes from large rocks (up to O.2m diameter) to small particles (less than 1 mm). Beneath the upper layer was a hard pan with some small amounts of material in the instrument nozzles. Most of the debris was found to be (U,Zr)02 ceramic material. There is similarity between this material and the ceramic material in the central previously molten region indicating that the lower plenum material came from the central region. Small amounts of structural material were found. Brown et al [22.6] and Strain et al [22.7] demonstrated the presence of a spinel or ferrite phase in the lower head debris samples. The materials fonned a eutectic system. The damage to the steel structures was relativel y small. Codes such as PLUGM and data from experiments such as BLOKKER [22.8] imply that the ceramic melt has a self-insulating effect causing only minimal damage. Little damage is therefore consistent with a higher melting point melt, increased damage may imply the presence of a material (metallic component) that lowers the melting point of the debris. TMI-2 supports the thesis that steam explosions in pressurised sequences are unlikely. The primary pressure response in the accident implies that no energetic steam explosion took

292

The Accident at Three Mile Island

place. The nature of debris material in the lower plenum provides additional evidence that no steam explosion occurred.

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22.8

State of the Lower Head

Wedge-shaped samples were taken for analysis from the lower head of TMI-2. Analysis work was carried out at the Argonne National Laboratory and INEL in the USA and in the UK. Samples were also taken from another vessel and subjected to known heating cycles.

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Comparisons could then be made with the TMI-2 10wer head samples and by comparison with the control samples, the temperatures in the TMI-2 samples could be deduced. A few samples were estimated to have reached temperatures in excess of I 320K. However, most of the samples were estimated to not have exceeded lOOOK , even those close to the believed main relocation route. The maximum temperatures may have been sustained for about 30 minutes. An elliptical hot spot was found in the vicinity of the samples with the highest temperatures. There was an interface between the stainless steel clad (5mm depth) and the base metal (50mm depth).

The cladding was cracked in the hottest region but there was no significant extension of the cracks into the base metal. Temperatures at 45mm into the vessel wall were approximately lOOK lower than those at the interface with the clad on the hot spot vicinity. Of particular interest in the vessel examination is the extent of damage to and melt in the lower head instrument penetrations. The region of maximum damage to the nozzles was found to be the same as the region of maximum vessel temperatures. An interesting observation was that there was more limited damage in the region under the major relocation pathway that at some other regions where nozzles were also completely melted off to within 25mm of the vessel. The peripheral nozzles were all undamaged.

22.9

Lower Head Failure Analysis

Various analyses have been carried out by INEL to investigate the threat to the lower head in TMJ-2. Moore and Tolman [22.9] considered various configurations of debris and various assumptions concerning the quenching of the debris. They concluded that the potential for failure depended mainly on the debris configuration. The margin to failure of the vessel has also been investigated by INEL as part of the extended TMJ-2 Vessel Investigations Programme. This work took into account new ultimate strength data for the vessel steel. The main objective was to make predictions based on inferred temperatures and reconcile predictions against the fact that the vessel remained intact. Early calculations were carried out by Wooten [22. 1] using the MARCH Code. These concluded that the survivability or otherwise of the lower head also depended on the state of the debris. A molten debris pool would pose a greater threat than non-uniform particle beds. JAERI have considered the debris configurations of Moore and Tolman [22.9] . Their work has helped to explain the cracking observed at the interface of the clad and base metal. They also showed that temperatures beneath the nozzle were - l OOK hotter than elsewhere and that high tensile stresses would be produced in the nozzle vicinity under quench conditions.

294

The Accident at Three Mile Island

Calculations have also been performed at the Culham Laboratory using a two-dimensional thermal response model of the vessel, including the possibility of ablation. The debris is allowed to be solid, partially molten, on liquid. Simple models were used to predict vessel failure criteria. 22.10 Fission Product Release and Transport in TMI-2

During the course of the TMI-2 accident, fission products (noble gases, caesium and iodine) were released to both the containment and the auxiliary building. The transport paths and timings were deduced from signals from process and area radiation monitors though these were not designed to give quantitative information under these severe accident conditions. Shortly after the accident, samples from the containment (sumps, air, and surfaces) and from the auxiliary building (tanks, sumps, and air) indicated that some 30% to 50% of the total fuel inventory of the gaseous and volatile fission products had been released to the containment (via the pressuriser and reactor coolant drain tank). The release to the auxiliary building was about 7% caesium and 5% iodine and about 8% of the noble gases, (via the let-down system to the CVCS); most of these noble gases were then released to the environment, some via a hole in the waste gas system, and some by deliberate discharge of waste gas delay tanks. Retention of caesium and iodine in the auxiliary building was mainly in the primary coolant clean-up system. There was no release to atmosphere from the containment during the accident, since the pressure rose above atmospheric only for a very short period. Most of the volatiles in the containment were found in the sump by the time this was sampled. There was a small release of noble gases when the containment was vented some time after the accident. The distribution of fission products in the plant is given in [22.5]. The table does not include the noble gases released to atmosphere. Another total mass balance was published in Gaithersburg [22. 10] . (The noble gases released to atmosphere are included). The Wlcertainties due to extrapolation from small samples to large volumes should be noted. 22.11 Implications for Modelling and Conclusions arising from the TMI-2 Accident

TMI -2 is a poor vehicle for benchmarking codes due to the lack of data available, particularl y the lack of thermal-hydraulic data. However in the OECD-CSNI TMI-2 exercise attempts

were made to define a common set of data and boundary conditions, which provided a useful framework for assessing modelling differences between the codes. The primary circuit remained intact despite very high temperature material in the core region. The upper plenum structures, and ex-vessel hot components remained sufficiently cool in TMI-2 that the integrity of these structures was not threatened.

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Differences exist between different PWR plants in respect of the upper plenum geometry and steam generator designs which impact on the generalisation to other plants of TMI-2 conclusions concerning primary circuit integrity. A major difference is the V-tube steam generator design in Westinghouse plants compared with the once-through steam generator designs in Babcock and Wilcox plants. There was no evidence of an energetic steam explosion. The lack of any direct instrumentation to detennine the core conditions is an obvious defect which should be addressed in future plant operations. Thennocouple infonnation would be the most useful, since the state of a degrading core is well characterised if its temperature conditions are known. (Various different time/temperature windows are known in which ballooning, control rod failure and fuel rod failure take place respectively). The severe core damage system codes showed reasonable agreement [22. 1 1] for the pre­ oxidation phase of TMI-2 in the OECD-CSNI exercise. Differences at later times were due to differences between the mechanistic and parametric codes in respect of their treatment of melt progression. Melt progression appeared to occur in TMI-2 as a stage by stage process, at relatively slow flow rates, not from the lowest point of blockage, rather than as a coherent pouring process. Although damage was very extensive the accident was halted by a modest amount of water. However a delayed quench seems possible. The codes have poor quench models and do not take account of surface cooling effects properly. The degree of thennal contact between the vessel and the debris may be variable, good to poor. The debris interactions and pour had three dimensional features but there is little justification to develop higher dimensional models given other uncertainties in the melt progression. For the lower head, the vessel may get close to failure before there is any significant melting or ablation of the vessel. However, penetration failures still require more investigation. Substantial core melting and fission product release to the containment can occur, with only negligible release to the environment from the containment provided containment cooling functions operate. The bulk of the radionuclides released from the TMI-2 vessel were retained in the primary circuit or containment, with little transport to the auxiliary building. Despite significant releases to the containment the only release to the environment during the accident was via the auxiliary building. Mechanistic codes can only predict events that occur over length and time scales greater than certain minima. These limits are constrained by available resources and computing power.

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The Accident at Three Mile Island

Therefore from a practical viewpoint, it is not feasible to expect to be able to follow every detail of accident scenarios of the complexity of TMI-2, although the general trends should be predictable mechanistically. The wealth of data generated by the TMI-2 programmes is providing valuable input into severe accident understanding and model development. However it must also be remembered that TMI-2 was one accident, providing one set of data, and care must be taken in extrapolating the findings of TMI-2 to all classes of severe accident scenarios.

REFERENCES

22. 1

D W Golden (ed), TMI-2 Analysis ExerciseFinaI Report. OECD-NEA-PWG 2 Draft Report (199 1). 22.2 TMI-2: Materials Behaviour, Nuclear Technology. A Journal of the American Nuclear Society, Vol 87, No 1 NUTYBB 87(1) 1 -334 (1989), August 1989. 22.3 J N Lillington, A J Lyons, I M Lovely, "UK Thennal Hydraulics Calculations in TMI2 Accident Analysis", CSNI Meeting of Analysis Exercise, Paris, March 1 988. 22.4 J N Lillington, A J Lyons, "Detailed Flow Analysis for the Three Mile Island Unit 2 Reactor Accident, Nucl. Energy, 1990, 29, No 4 August, pp 235-24 1 . 22.5 G R Eidam, E L Tolman, J M Broughton, R K McCardell and W R Stratton, "TMI2 Defuelling Conditions and Summary of Research Findings", International lAEA Symposium on Severe Accidents in Nuclear Power Plants, Sorrento, Italy, 2 1 -25 March 1988. 22.6 A Brown, G J McIntyre, C Graslund, "Analysis of Crystalline Phases in Core Bore Materials from Three Mile Island Unit 2". Nuclear Technology, 87, 1 370 145 ( 1 989). 22.7 R V Strain, L A Neimark, J E Sanecki, "Fuel Relocation Mechanisms based on Microstructures of Debris" Nuclear Technology , 87, 1 87- 190 (1989). 22.8 H Hohmann, D Magallon, A Benuzzi, A V Jones, A Yerkess "Results of the FARO Programme in Reactor Safety Research - the CEC Contribution (W Krischer, ed). Elsevier Applied Science (1990). 22.9 R L Moore and E L Tolman, Estimated TMI-2 Vessel Response Bounds based on the Lower Plenum Debris Configuration in Proceedings of the 1988 National Heat Transfer Conference, Houston, July 24-27, 1988 ( 1988) 22. 10 Transactions of the Sixteenth Water Reactor Safety Infonnation Meeting, NUREGI CP-0096, Gaithersburg, Maryland, 24-27 October 1988. 22. 1 1 D W Golden, "Summary of the Results of the TMI-2 Analysis Exercise", Paper presented at the ANS/ENS TMI-2 International Conference, Washington DC, 30 October - 4 November 1988.

297

Chapter 23 PLANT STUDIES 23.1

Introduction

The purpose of this chapter is to describe the status of current code prediction capabilities for plant This will be achieved by providing some examples of different code predictions of various key accident scenarios at representative plant scale. The codes chosen have all been developed under US NRC sponsorship and are considered to be state-of-the-art and mechanistic. The scenarios are typical of those considered in fault analysis. For accidents within the design basis the codes below, Table 23. 1 , have been applied in the UK for different scenarios as follows:

TABLE 23.1 THERMAL HYDRAULICS CODES Code

Accident Scenario

RELAP5 [23 . 1 ]

- Small Break Loss o f Cooling Accidents (SBLOCAs) - Loss of Feed Water - Steam Line Break - Reactivity Transients - Anticipated Transients without Scram - Steam Generator Tube Rupture

TRAC [23.2]

- Large Break Loss of Cooling Accident (LBLOCAs)

RELAP5 has been widely used for intact circuit faults and small break loss of cooling accidents. TRAC is considered to be a more rigorous code for larger breaks because of its treatment of multi-dimensional effects in the vessel. Analysis of some selected design basis scenarios are described in the sequel. For severe accidents the codes shown in Table 23.2 have been used in the UK to predict various scenarios. For the station blackout two possible scenarios are considered assuming either (a) that the circuit remains intact (a high pressure case) or (b) the surge line fails (a low pressure case).SCDAP/RELAP5 is the leading USNRC severe accident primary circuit and core degradation code. The station blackout scenario has been a widely studied case. As an example of a containment bypass scenario an interfacing LOCA scenario is used which

298

Plant Studies

represents a medium size break through the RHR system. Accidents at shutdown are at low pressure with incondensables present, e.g. if the primary circuit is drained to the mid-loop level. VICTORIA is the leading USNRC primary circuit fission product transport and chemistry code. TABLE 23.2 SEVERE ACCIDENTS Code

Accident Scenario

Primarv Circuit Thennal Hydraulics and Core Degradation SCDAP/RELAP5 [23.3]

- Station Blackout - SBLOCAs - Containment Bypass - Shutdown Accidents

Fission Product Transport VICTORIA [23.4]

Station Blackout

Containment CONTAIN [23.5]

Station Blackout

CONTAIN is a leading US NRC mechanistic containment code. The station blackout scenario covers a range of conditions. The early phase will be high pressure since the circuit will be intact: however if the surge line fails due to natural circulation heating as is predicted by some current codes, the later phase will be at lower pressure, (containment pressure). The VICTORIA and CONTAIN codes have been used to provide predictions for either situations. It is impossible within the scope of this book to discuss the details of different predictions between different versions of the codes. These differences may be significant It may be assumed however that the calculations cited have been predicted with state-of-the-art frozen versions of the codes. The objective of this chapter is not to describe defmitive results but to give broad descriptions of current code capabilities. 23.2

Intact Circuit Faults

23.2.1 Loss of Feed Water

A loss-of-feed water transient is a particular example of an intact circuit fault involving a decrease in SG heat removal. The sequence described below for a large Westinghouse 4 loop PWR is one assuming complete loss-of-feed water to all SGs with the auxiliary feed and pressuriser spray failing to function.

299

Plant Studies

This is a very pessimistic scenario and would be beyond design basis for some plants. In this scenario no operator action is assumed for the first 30 min (1 800s). At this point the primary coolant pumps are tripped, two out of three of the Power Operatored Safety Relief Valves (POSRVs) are assumed latched open and the High Pressure Injection System (HPIS) is initiated. RELAP5 calculations predicted the following events. The reactor trips at low sa level, following the loss of feed. The pressure falls initially then rises again as heat renewal is limited by the sa relief valves and then as the sas boil dry. There is a sustained insurge to the pressuriser. The primary pressure reaches the POSRV set point at about 1270s and is controlled by valve cycling. Feed and bleed is initiated at 1 830s (assuming an operator action). m 0..

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The important phenomena in such a transient are the changes in sa inventory, the changes in sa heat transfer following loss of feed and during the boil down and the pressuriser response during the insurge as the primary coolant heats up. A potential problem with sa inventory may be associated with coarse noding and sensitivities may need to be considered in modelling. The assessment of plant transients of excessive heat transfer at low sa inventory is not likely to cause problems for progressive sa boildown because heat removal is primarily determined by the inventory The prediction of pressure rise and pressure peak will depend on the adequacy of modelling of the various heat transfer mechanisms in the pressuriser e.g. the heat transfer to the walls and the steam/water interface.

300

Plant Studies

23.2.2 Steam Line Break

This section discusses some typical results for a large steam line break in a 4 loop WestinghousePWR. The results were obtained with the RELAP5/MOD2 code and cover the thermal hydraulics and integral system response: core reactivity data were provided by an independent three dimensional kinetics calculation using a neutronics code. The sequence calculated makes various conservative assumptions which maximise the likelihood of return to power. These include maximum auxiliary feedwater flow, limited mixing in the vessel, any emergency b
Reactivity changes often concern core behaviour effects rather than system effects, and therefore rather than system thermal hydraulics. However, for a rod withdrawal the transient may be sufficiently slow that thermal hydraulics modelling is an issue. Representative RELAP5 calculations for a rod cluster control assembly withdrawal for a large 4 loop PWR are briefly described below. For a slow withdrawal, equivalent to a reactivity excursion of (-3mN/s) quasi-equilibrium conditions remain and the mechanism limiting the power is moderator temperature and density.

Plant Studies

30 1

For fast withdrawal (-48mN/s) the power changes significantly before the fluid conditions change significantly. The power limiting mechanism is then Doppler feedback. The safety related issues relate to margin to Departure from Nucleate Boiling (DNB) and fuel damage and the effectiveness of the reactor protection system in limiting the overall overpower and thermal-hydraulic transient. The main phenomena concerning the plant response are the nuclear feedback, the pressuriser response to insurge (prior to the trip) and the pressuriser response to outsurge (post reactor trip).

23.2.4 Anticipated Transient Without Scram (ATWS)

The potential seriousness of an ATWS depends on the particular failures suffered by the fluid system. In general both thermal-hydraulic and neutronic feedbacks are important. One severe ATWS is that associated with loss of off-site power together with flow run-down. In these circumstances core heat removal is reduced and any emergency boration systems (EBS) operate less effectively. Thermal-hydraulics calculations using system codes such as RELAP5 require both more detailed kinetics input from a neutronics code and also more detailed information on vessel mixing which needs to be supplied by a computational fluid dynamics code. The most important safety issues concern the effectiveness of any EBS , the potential for DNB when the power is high during the earl y part of the transient and the SG heat renewal function. The main modelling considerations require the response characteristics of most of the main primary system components including SG inventory and heat transfer, the pressuriser response to insurge, the performance of pressuriser sprays, the EBS system, the effectiveness or otherwise of boron and thermal mixing in the reactor vessel and the reactivity feedbacks.

23.3

Loss or Coolant Accidents

23.3.1 Steam Generation Tube Rupture (SGTR)

The SGlR scenarios belong to a category of small-break LOCA transients. Important phenomena include SG overfill by two-phase water, cool down with reverse heat transfer and pressuriser phenomena including outsurge and effectiveness of sprays. Calculations have been carried out for a real plant transient that occurred on the DoeI-2 plant [23.6]. This is a Westinghouse two loop PWR and the transient was considered as the subject of a CSNI International Standard Problem, ISP 20 [23.7] . Following the tube rupture, the faulty SG was isolated and heat was removed by depressurising the intact SG. The charging system and the HPIS was utilised to maintain primary side inventory. The primary coolant pump of the faulty loop was tripped to reduce heat input on

302

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Plant Studies

303

the intact side and later restarted to equilibriate primary and secondary pressures on the faulty loop. The HPIS and pressuriser sprays were operated to reduce primary pressure until the pressure level rose to too high a level. Spray operation was then stopped and the pressure increased, finally stabilising at the HPIS cut off pressure. Calculations with the system code RELAP5 showed that the main features could be calculated. Different calculations submitted for the ISP showed differences in detail and in the rates of change of certain important parameters. Early primary depressuration was well calculated but the pressure decrease was dependent on the assumed spray flow rate. Secondary pressures were less well calculated due to difficulty in modelling interphase heat transfer. In general timings and rates of change were dependent on uncertainties in the boundary conditions and the noding of the numerical models. 23.3.2 Large Break LOCA

This section describes the results of the TRAC-PFI/MODI calculation [23.8] for a 200% double-ended cold leg break loss-of-coolant accident (LOCA) in a Westinghouse 4-loop Pressurised Water Reactor. The break represents a complete severance of the cold leg pipe. TRAC-PFI/MODI code is a best-estimate code containing no built-in conservatisms. Operational and transient boundary conditions were specified consistent with the UK Sizewel1 PWR minimum safeguards requirements, i.e. the accident was assumed to occur at the worst time in life, with a design value of peak: power rating and with a loss of off-site power. The calculation provides an insight into current state-of-the-art modelling of the thermal-hydraulic response of a system during a double-ended cold-leg break LOCA, particularly during the vessel refill and reflood phases of the transient The nodalisation used for the TRAC calculation is shown in Figure 23.2. The important points to note from the calculation are as follows: A peak clad temperature of 1033K for the highest powered rod (peak: linear rating 12.9kW/ ft) occurs during blowdown at three seconds. For a core average rod (peak linear rating 8.4kW/ft) the peak temperature is 9 10K, Figure 23.3. The maximum clad temperature of the highest powered rod remains below 875K (and 750K for the average rod) during the refill and reflood phases of the transient. The core is cooled and partially quenched during the first 20s, as the primary circuit blows down, due to a surge of water from the bottom and from liquid draining down from the upper head. Initiation of the accumulator flow in the intact loops occurs at 14s. The blowdown ends at 26s. Emergency core cooling (ECC) liquid first enters the core just after 30 seconds. The accumulators empty of liquid at about 45s and at this time nitrogen begins to flow into the cold legs just before 45s producing an increase in cold leg pressure and a surge of liquid into the core, at about 50s. Following the end of accumulator flow the core fills slowly with liquid from the downcomer. The fuel rods continue to cool slowly with the final quench of the highest powered rods at approximately I lOs.

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304 23.4

Core Degradation and Thermal Hydraulics

This section describes calculations with the SCDAP/RELAP5 code of various postulated severe accident scenarios for a large Westinghouse-type PWR. Examples include Station Blackout, Small Break LOCA. Interfacing LOCA and a representative Shutdown Accident.

5.0 4.5 4.0 (0 0..

6 w 0::: =>

V"I

� 0::: 0..

3. 5 3 .0

C OL D L E G

2.5 2.0

ACC U M U LATO R

1 .5 1 .0 0.5 0.0 -+--�-""""";"""':-----'--r--,--f 20 40 60 o 1 00 1 20 80 TI M E (s) 950 �----. 900 850 800

Q

750

� 700 0::: => t-

�

w 0..

�

w t-

650 600 550 500 450 400 �-�--�----r--'-�� 20 40 1 00 1 20 60 o 80 TI M E (s)

FIGURE 23.3 TRAC CALCULATIONS FOR PWR LARGE BREAK LOCA

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23.4.1 Station Blackout Intact Circuit

The station blxkout accident sequence is firstly an intact circuit scenario, initiated by total loss of all elecnical JX>wer to the plant With all emergency cooling systems inoperable, the core eventually Wlcovers, high core temperatures result and without recovery of cooling systems, failwe of the v�l and primary circuit boWldary is the ultimate consequence. The threat to the containment in the Station Blackout accident depends on whether: (i)

the primary circuit remains intact and at high pressure up to lower head failure; or

(ii)

the primary circuit fails due to creep failure of a vulnerable primary circuit component (e.g. surge line or hot leg), resulting in a primary circuit depressurisation.

Calculations have been carried out using the SCDAP/RELAP5 code. The heat up phase is shown in Figure 23.4. Details of the primary circuit and vessel nodalisation can be found in [23.9] . The core was modelled as three radial and ten axial regions. The SCDAP/RELAP5 calculation firstly corresponds to nominal flow and power conditions before the reactor scram and pump run-down is initiated at 4OOs . After this time, the core continues to generate heat from fission product decay. With no feedwater supply, the secondary side water gradually boils off, with dryout occurring at approximately 4300s. After the loss of the secondary side heat sink, the primary circuit pressure and temperature begins to increase, with the POSRV opening and releasing primary circuit coolant to the containment. As a result of the continuing loss of coolant inventory, a general core heat-up begins at approximately 7200s. As core uncovery progresses, the uncovered region undergoes a rapid temperature increase, Figure 23.4. Zircaloy oxidation begins at 7800s, with accompanying hydrogen production. Control rod material relocates at about 8500s when the steel cladding melting temperature (1700K) is reached. At 8500s the oxide phase change temperature of 1850K is reached and there is a rapid temperature excursion up to the assumed cladding failure temperature of 2500K. The first fuel rod material relocation takes place at 8550s in the central core region, following clad failure. By 9500s, the core configuration predicted could be divided into three regions: (I)

the top third of the core consisting primarily of uraniwn dioxide and zirconiwn dioxide, having reached temperatures in the vicinity of 2500K;

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306 (2)

the central region of the core consisting of material at elevated temperatures with. in some cases. a large amount of material relocated from higher up in the core: temperatures were below the uranium dioxide - zirconium eutectic dissolution temperature;

(3)

the lower part of the core consisting of intact, relatively unoxidised rods. at temperatures between saturation and about 1200K; there was some additional relocated control rod material and uranium dioxide-zirconium eutectic from higher up in the core.

To establish the threat to the primary circuit due to overheating. more sophisticated SCDAP/ RELAP5 nodalisations can be used for the vessel and hot leg than were employed in the calculation described above to provide a better representation of the thennal hydraulics: in particular to include the effects of in-vessel natural circulation and hot leg counter-current flow. The mesh can be found in [23.9]. However such a representation requires qualification since it is based on a split hot leg nodalisation for modelling the counter-current flow. This prevents mechanistic modelling of drag and mixing between the forward and return flows. It should be noted that the nodal isation is similar to that used in analogous calculations carried out in the US . System codes such as RELAP5 do not at present have the capability to model single phase. counter-current flow within a component. With this representation. a reasonably steady natural circulation flow was established by about 7000s . Flow patterns were similar to those observed in the Westinghouse natural circulation experiments described in the previous chapter.

3 000

Q

2500

.......

w � ::> I« ca:: w 0...

2000 1 500

�

w I-

1 000 500 7000

7500

8000

8500 TI M E (s)

9000

9500

F U E L CLAD D I N G TE M P E RATU R E FIGURE 23.4 SCDAP/RELAP CALCULATIONS FOR A STATION BLACKOUT SEVERE ACCIDENT

The main features of calculations where natural circulation is modelled are an extended heat­ up period. Due to the long period of slow heat-up. the cladding is heavily oxidised by the time

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307

of the rapid temperature excursion. There is then less eutectic relocation than for the case when natural circulation is not modelled since the clad is more fully oxidised. The effect of enhanced energy transfer to the structures temperatures in the primary loops results in raised hot leg, surge line, and steam generator tube temperatures. The hot leg temperature decreases from the reactor vessel to the steam generator. The steam generator tube temperature also steadily decreases from the inlet plenum to the outlet plenum along the outflow tubes and from the outlet plenum to the inlet plenum along the cold flow return tubes. The surge line temperature increases more rapidly when the POSRV s are open because there is more flow, and hence more heat transfer. through the surge line. The average temperature of the surge line is higher than that of the other structures in the loop during the heat up phase when natural circulation is significant and is sufficiently high for creep rupture failure of the surge line to be possible before reactor vessel failure. The presence of the hot leg flow carries hot vapour to the steam generator tubes, causing them to heat up as well. However, compared with the surge line temperatures, the temperature of the SG tubes is calculated to still remain low enough that these would not be expected to fail. Surge Line Failure

RELAP5 calculations for a station blackout with assumed hot leg failure have also been perfonned. The surge line was the assumed breach site: failure was assumed at l 00s from the start of core heat up. Prior to the initiation of the hot leg breach the transient behaviour of the primary circuit was the same as for the station blackout without breach summarised above. The calculations showed that following the hot leg failure the primary circuit pressure falls rapidly to the accumulator set point. The coolant flow induced by the depressurisation is sufficient to cool the fuel and end the temperature and oxidation excursions before the core is reflooded by the accumulators. Accumulator injection occurs 50s after the hot leg failure and the subsequent condensation of steam on the subcooled liquid causes a further depressurisation until the core is recovered. At the time of accumulator injection the peak clad temperature falls to 835K. The core is recovered by 8280s. By 8300 seconds the molten absorber material at the top of each control rod component is refrozen in situ. Once the accumulator injection tenninates, the primary system inventory again starts to fall as decay heat in the core is removed by boiling. However, rapid vapour generation during this low pressure boil-down maintains adequate core cooling for a period of almost 1000 seconds and the second core heat-up is delayed until 93 15 seconds. Control rod material is the first to relocate, between 10300 and 10600s. Relocation in the inner fuel channels takes place at about 1 1 000s . The initial relocation is to the bottom of the channel but no fuel or cladding is lost from the core. During the second core uncovery and heat-up a further 62 kg of hydrogen is generated, compared with 7.8 kg during the frrst uncovery.

308

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23.4.2 Small Break Loss of Cooling Severe Accident

A SCDAP/RELAP5 calculation of a small break. two inch diameter. cold leg loss of coolant accident was performed in which the engineered safety features were assumed to fail. In this transient the breach was the initiating event The reactor is tripped at about 70s after initiation of the break after the pressure had fallen to 12 MPa. The calculation showed that after the pumps run down. decay heat from the core is transported to the boilers by natural circulation around the loops. The system pressure during the period from the start of the transient up to 3500s is set by the steam generator safety relief valves. The start of significant steam flow from the break occ urs at around 1400s with a transition to high quality flow at around 2500s. The primary systems pressure remains high until around 3400s as the core uncovers and the energy input to the coolant begins to fall. Fuel temperatures begin to rise from around 3 100s as the core uncoveres. with oxidation of the fuel beginning shortly afterwards at 3250s. The more rapid oxidation, associated with the change of phase of the zirconium oxide at 1 850K, begins at around 3900s. At4330s, when the primary system pressure falls t04.0 MPa, accumulator injection starts and approximately 23 tonnes of water are injected. Core recovery is prevented by a rise in pressure which stops the accumulator injection. This is caused by a combination of evaporation of part of the primary circuit liquid inventory and the decrease in the energy flow from the breach as the quality falls. Part of the injected liquid flows out of the breach. A second accumulator injection at about 4800s, when approximately 4.8 tonnes have injected, also fails to recover the core for the same reason. Relocation of the core materials begins at 3650s when the control rod material begins to melt. The control rod relocation begins in the central region of the core and progresses to the outer region. Fuel and clad relocation begin in the central region at 400s half way up the core. Two further relocations are followed by a total blockage of the central channel at 4220s and some further relocations occur. Significant amounts of the control rod material are lost from the bottom of the core by 5000s . The hydrogen production rate is significantly reduced by the relocation of the cladding materials at 4020s with subsequent small increases each time the accumulators inject Thermal-hydraulic, chemical and mechanical models are all required for prediction of these core melt scenarios. The most important mechanisms relate to boil-off, accumulator discharge and nitrogen injection, clad ballooning in the case of primary circuit depressurisation and natural circulation which affects the temperature distribution in the core and the timing of the temperature excursion, hydrogen production and melt progression in intact circuit accidents. Radiation heat transfer is the most important mechanism for distributing heat within the high temperature core and surrounding structures. Zircaloy oxidation kinetics determine the rate of the initial high temperature excursion and the hydrogen production throughout the accident (steel oxidation is of lesser importance compared with Zircaloy

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oxidation since the steel inventory in the core is low in comparison with the Zircaloy inventory). Control rod materials interactions, the dissolution of fuel by liquid Zircaloy and ZircaloyfU02 melt relocation strongly influence the peak temperatures attained and hydrogen production prior to initial melt relocation. The main interest in late phase melt relocation concerns the potential threat to lower head integrity. 23.4.3 Containment Bypass: Interfacing LOCA An interfacing LOCA could occur if a direct path is opened from the primary circuit of a PWR to the residual heat removal (RHR) system outside the primary containment while the reactor is in nonnal operation. It might be initiated by failure of the check valves separating the high pressure circuit from the RHR system, causing the RHR system to rupture and primary coolant to be expelled to the auxiliary building exterior to the containment

Since the containment is bypassed the extent of fission product release is a key issue and this depends on the ability to predict temperatures of the pipe work in the pathway. Calculations perfonned with the SCDAP/RELAP5 code show that substantial voiding of the coolant circuit follows blowdown. However, the core remains cold as long as the ECCS pumps operate. The ECCS water supply is eventually exhausted because ECCS water is lost outside the containment and therefore cannot be recirculated e.g. from the containment sump. Once the ECCS supply runs out after about three quarters of an hour, the water remaining in the vessel cannot cool the fuel for long and severe overheating of the fuel follows within about half an hour. The calculations show that the residual heat removal (RHR) system pipe, which provides a pathway to the auxiliary building, remains cold. Deposition of volatile fission products in the pipe would be important with the prospect of a significant reduction or delay in the source tenn. 23.4.4 Shutdown Accidents

SCDAP/RELAP5 calculations for accidents during shutdown have also been carried out. One particular issue was to examine gravity drain of the Residual Waste Storage Tank (RWST) following a loss ofRHR during mid-loop operation. The objective was to examine an accident management strategy for an accident involving loss of the RHR system while the reactor is shut down and operating with the primary coolant drained to mid-loop level. In the example considered below, the loss of RHR is assumed to occur while a steam generator man way is open in the primary circuit, and where flow from the manway is initially prevented by the presence of nozzle dams, but which then subsequently fail. Shutdown accidents fonn a significant contribution to overall risk [23 . 1 0] [23. 1 1 ] . The calculations examined the effect of aligning the Refuelling Water Storage Tank (RWST) with the cold legs and allowing water to flow into the primary circuit under gravity. The CONTAIN code was used to calculate the containment pressure. It was shown that gravity drain of the RWST is able to delay the onset of core damage: timescales are further extended if the containment atmosphere can be cooled. Calculations were able to provide insights on:

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the length of time before core damage occurs, in the case of no operator intervention, the effectiveness of aligning the RHR system to allow gravity drain of the RWST into the primary circuit, as a means of delaying core damage, how the gravity drain might be improved through action to keep the containment pressure low. Modelling issues include conditions of reflux condensation with an air bubble in the steam generator tubes and conditions of low pressure and low flow. The degree of validation of RELAP5 models with respect to the local condensation heat transfer correlation [23. 12] , the calculation of the distribution of air in the tubes, and the effects of natural circulation on the secondary side, is very limited. The noding scheme used is another issue: the noding scheme used may be too coarse to provide a "converged" result. 23.S

Fission Product Transport under Severe Accident Conditions

23.S.1 Station Blackout Intact Circuit

Fission product transport calculations have been perfonned with the VICTORIA code [23.4] using thennal hydraulic input taken from the SCDAP/RELAP5 results for the Station Blackout sequence above. Fission products exiting the core pass into the pressuriser and loops with a fixed proportion being passed to each. Around 80% of the released fission products pass through the pressuriser loop. Results have been produced for a high diffusivity for the fuel open porosity, a value chosen to give good agreement between the release fractions using VICTORIA and the empirical correlations based on low pressure experiments. Data are not available at high pressure. Release fractions were calculated of eleven fission product elements released from the reactor pressure vessel to the reactor coolant system from 8300s to 8900s. These compared reasonably with release fractions from the fuel using correlations from [23 . 1 3] and [23. 14] . This i s assuming modifications to the correlation for tellurium release in [23 . 1 2] i n line with suggestions in [23. 1 5] . Concerning the vapour and aerosol chemical species releases from the reactor pressure vessel to the reactor coolant it was found that caesium exits the reactorpressure vessel predominantl y as the vapour species CsOH. However, as the core continues to heat up, and release the fission product molybdenum, Cs2MoO4 aerosol increases in importance to dominate the aerosol species over CsOH aerosol. Iodine is mainly in the aerosol species CsI with a small fraction of the vapour species as HI and �. Barium and strontium molybdates exit the reactor pressure vessel predominantly as aerosol species. Molybdenum release is predominantly as aerosol species Cs2Mo04 and Mo02• Tellurium is released primarily as the aerosol species CdTe. The interaction of tellurium with unoxidised Zircaloy clad was not modelled. Boric acid plays a minor role at the exit from the reactor pressure vessel: caesium borate is only a small fraction of the total caesium.

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For fission product aerosols exiting the top of the reactor vessel it is instructive to consider where the aerosol species are predominantly fonned. From the middle of the fuelled core to the top of the fuelled core the fission products are present predominantly as vapour species. Aerosol fonnation at 8700s takes place predominantly above the fuelled core. For these calculations, the VICTORIA code was used incorporating a single component aerosol model and an ideal solution model for the chemistry. Fission products enter the pressuriser and intact loops as vapour and aerosol species and in the pressuriser loop can exit the reactor coolant system via the pressure operated relief valve (PORV) on top of the pressuriser. There is virtually no release through the PORV as vapour species, with the exception of the noble gases. At 8900s, essentially all the noble gases entering the pressuriser leg had been released through the POR V. All of the other fission product elements behave similarly to each other, with between 24 and 27% of the fission product elements entering the pressuriser leg released from the PORV. Release from the PORV is predominantly by aerosol species, except for the noble gases. 21 -22% of the fission products exiting the reactor pressure vessel enter the intact legs. The noble gases do not deposit but all the remaining fission products deposit on surfaces in the intact loops. Caesium and iodine deposit predominantly in the hot legs with little penetration of either fission product to the steam generators. CsOH deposition is on the floor, walls and ceiling of the intact leg. Deposition on the floor dominates completely and all fission products except for the noble gases behave similarly as aerosol species. There is more growth of aerosol particles in the intact loop hot legs than in the pressuriser loop, a lower flow velocity and consequently high residence time, allowing gravitational settling to dominate. Surge Line Failure

If surge line failure occurs then all of the gas flow from the Reactor Pressure Vessel (RPV) enters the broken leg (pressuriser leg) and all fission products follow the flow. Concerning the dominant physical fonns of fission products leaving the reactor pressure vessel it is predicted that aerosols dominate the physical fonns of all the fission products except the noble gases and the vapour fraction corresponds to the temperature variation in the top of the core. There is lower molybdenum release than in the intact circuit case and higher caesiwn release. For the important individual fission products therefore CsOH dominates the caesium speciation. CsI again is the dominating iodine chemical species. Barium and strontium speciation are dominated by the hydroxides, again because of the lower molybdenum release. Caesiwn molybdate dominates the molybdenum speciation and CdTe dominates tellurium speciation. In this case, there is very little retention of fission products in the broken hot leg. There is also little aerosol growth along the broken hot leg.

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3 12 23.6

Containment Behaviour Under Severe Accident Conditions

23.6.1 Station Blackout Intact Circuit

Some representative CONTAIN predictions for the thermal-hydraulic response of the containment are discussed finally in this Section. This transient can be divided into three main phases: prior to vessel failure; short term following vessel failure: the issue here is direct containment heating discussed earlier; long term following vessel failure. Figure 23.5 shows the calculated temperature and pressure of the atmosphere in one of the lower compartments. The long-term pressure response is similar between the fourcontainment cells (the computational model was based on four cells) because they are linked by large flow paths. More variation is seen during the transient following vessel failure, when the pressure in the cavity becomes very high for a few seconds. The transient pressure peak resulting from direct containment heating and a coincident hydrogen burn, cause the containment pressure to rise quickly to about O.7MPa. The long­ term pressure rises, as decay heat from the debris boils water in the cavity and the lower containment compartments. Containment failure is not modelled in this calculation, the pressure continues to rise. Prior to vessel failure the calculated pressure in the containment rises as steam and hydrogen are released from the primary circuit through the pressure relief valves. The hot source of gas might rise as a plume through the main body of the containment and form a stratified layer in the dome. At vessel failure estimates indicate that about 2/3 of the core material would be ejected at high pressure into the dry cavity when the lower head fails, at about 12300s after reactor shutdown. The calculated OCH pressure is significantly lower than would be indicated by calculations of the equilibrium conditions between the debris and the atmosphere. About half the debris remains in the lower containment rooms and is isolated from the bulk of the containment atmosphere. Also much of the free metal in the melt is oxidised by steam, rather than oxygen; the steam reaction is much less exothermic. Heat losses to containment structures are also a factor. Early peak pressures are reduced if rapid debris trapping is assumed or that a hydrogen burn does not occur, but increased if the containment is modelled as a single cell or it is assumed that most of the core debris participates in direct containment heating.

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1 600 1 400 1 200 -

Q

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800 -

w

-

600 -

0-

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w .....

400 -

)

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200 O �------�I--�I�--� 000 0. 5 1 .0 1 .5 2 . 0 x 1 03 TI M E

(s)

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1 .2 1 .0 ro- 0.8 0-

S w

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0.4

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2 . 0 x 1 03

(s)

(b) PRESS U R E R ES PO N S E I N LOW E R O UTE R CONTAI N M ENT FIGURE 23.5 CONTAIN CALCULATIONS FOR STATION BLACKOUT

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3 14

The long tenn thennal-hydraulic response is sensitive to the distribution of decay heat and whether or not containment safeguards are recovered. This has been investigated using CONTAIN and the CORCON code. If the containment safeguards spray system is recovered after 24 hours, the pressure falls and the threat to the containment is avoided. Higher containment pressures arise if allowance is made for the additional decay heat associated with fission products released from the core-melt in-vessel and which reach the containment. CONTAIN predictions also can provide indications of the long-tenn fission product behaviour in the containment. Sources can be included which allow for fission products released from the core-melt in-vessel and which reach the containment. Aerosols expected to be released during direct containment heating can also be modelled. Over 99% of the fission product aerosols from the core-melt and from direct containment heating are predicted to be deposited on surfaces, mainly by gravitational settling, before the long-tenn pressure rise approaches a realistic containment failure pressure. This analysis referred to here took no account of fission product aerosols which would be released during core-concrete interactions or of the re-release of volatile fission products in­ vessel, which would accompany the depressurisation following vessel failure. Surge Line Failure

In this sequence the containment response can be described by the same three major phases as for the intact circuit assumption. However, the direct containment issue now disappears and the only issue is long tenn over pressurisation. The sources that threaten the reactor containment integrity are: thennal hydraulic sources of water, steam and hydrogen; fission product decay heat sources, and; heat generated from the chemical reaction between the core debris and the concrete in the cavity.

REFERENCES

23. 1 23.2

23.3 23.4

RELAP5/MOD2 Code Manual, Volumes 1 and 2 , Victor H Ransom et al. NUREG/ CR-43 12. 'TRAC-PFI/MOD 1 . An Advanced Best-Estimate Computer Program for Pressurised Water Reactor Analysis" Los Alamos National Laboratory Report. LA- l 0 1 57 -MS , NUREG/CR-3858. SCDAPjRELAP5/MOD2 Code Manual, Volumes 1-3. C M Allison and EC Johnson. NUREG/CR-5273, September 1989. T J Heames, D A Williams, N E Bixler, A J Grimley, C J Wheatley, N A Johns and N M Chown, VICTORIA: A Mechanistic Model of Radionuclide Behaviour in the Reactor Coolant System Under Severe Accident Conditions, NUREG/CR-5545, 1990.

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23.5 23.6

23.7 23.8

23.9 23. 1 0 23. 1 1 23. 1 2 23. 1 3 23. 14 23. 1 5

315

K K Murata, et al. Users Manual for CONTAIN 1 . 1 . A Computer Code for Severe Nuclear Reactor Accident Containment Analysis. Sand 87-2309 ( 1 987). Report on the Incident at DoeI-2 Nuclear Power Plant: Severe Leakage in Steam Generator B on June 25, 1 979. Tractabel, Brussels, Belgium, 25 October 1 979. PDf VEF rev. A. ISP-20 International Standard Problem No 20; Steam Generator Tube Rupture; Nuclear Power Plant DoeI-2, Belgium, M de Feu et al. CSNI Report No 1 54. P Coddington, F Motley, TRAC-PFI/MODI . Analysis of a Minimum Safeguards Large-Break LOCA in a 4-loop PWR with 1 7 x 1 7 fuel, AEEW-RI772, February 1 985. J N Lillington et al, SCDAP/RELAP: Recent UK Experience, Presentation to International RELAP5 Users Seminar, 1992. Shutdown and Low Power Operation at Commercial Nuclear Power Plants in the United States. NUREG- I449, February 1 992. The Use of PSA Results in France by Safety Authorities. N Tellier. NUREG/CP0 1 1 5 pp 1 59- 164, 199 1 . RELAP5/MOD3 Code Manual Volume IV: Models and Correlations. K E Carlson et al, NUREG/CR-5535, June 1 990. US Nuclear Regulatory Commission, Technical Bases for Estimating Fission Product Behaviour During LWR Accidents, NUREG-0772, 198 1 . J L Kelley, A B Reynolds and M E McGowan. Temperature Dependence o f Fission Product Release Rates, Nucl. Sci. Eng., 88, 144, 1984. M Silberberg, J A Mitchell, R 0 Meyer and C P Ryder, Re-assessment of the Technical Bases for Estimating Source Terms, NUREG-0956, 1 986.

317

Chapter 24 ACCIDENT MANAGEMENT

24.1

Introduction

Accident Management (AM) procedures are being developed to help derme Station Operating Instructions whose purpose, in the event of an accident, is to prevent the plant reaching a serious damage state e.g. core degradation. Early research was carried out in Germany [24. 1] [24.2]. Now most of the OECD member states have programmes in place to investigate accident management, including accident management of severe accidents [24.3]. The term 'accident management' is used here to refer to all actions which prevent core damage, contain the core within the primary circuit boundary, maintain containment integrity and minimise the source term to the public. Accident management measures have been broken down into several categories. Preventative measures aim to avoid damage to the reactor core. These measures include the use of safety or non-safety systems to restore the plant in the time period between initiation of the accident and the onset of core degradation. For many existing designs and scenarios this time frame is relatively long e.g. hours. The objective in many future designs is to extend this period substantially to allow operators more time to carry out appropriate action. Mitigative measures are mainly concerned with minimising the consequences once a core melt has occurred. Without such action core melting may progressively worsen possibly culminating in a large and uncontrolled source term to the environment. Mitigation may be achieved in various ways either by maintaining containment integrity or by restricting release of fission products by other means. Best estimate thermal-hydraulic codes have been developed for traditional accident analysis. These are now being required to assess the effectiveness of AM procedures and therefore need to be validated and verified beyond their original scope of application. This is being achieved via various experimental programmes: a good example is the BETHSY programme currently underway in France. These programmes aim to support planning and assessment of AM procedures as well as providing data for code validation. For accident mitigation severe accident codes are also being developed and adapted from their original intended use as analysis tools for unrecovered sequences. Experiments are underway to understand phenomena, support planning and provide data for code validation. An example here is the CORA programme which includes melt progression tests that are quenched with water at high temperature.

318

Accident Management

The purpose of this chapter is to briefly describe for both preventative and mitigati ve accident management catagories, the current status of knowledge. Various representative scenarios and associated possible accident management measures are considered. Computer code requirements are described together with assessments that have been carried out for accident management code validation.

24.2

Preventative Accident Management

Various studies are showing that preventative accident management measures can have a significant impact on reducing the frequency of core melt. It has been found that effective procedures could reduce the overall frequency of core melt by an order of magnitude. Accident management measures may be effective in dealing with various transients and small break LOCAs where the development of the accident proceeds relatively slowly. In many circumstances the cause of problem can result from loss of feed to the steam generators. In this case 'feed and bleed' procedures on the secondary system are likely to be the most effective. In some circumstances primary circuit pressure relief may also be achieved via direct operator intervention on the primary side. Accident scenarios can therefore be categorised into those that can benefit from accident measures on the secondary side and those that can benefit from accident management measures on the primary side. Important scenarios for the PWR [24.4] are shown in Table 24. 1 :

TABLE 24. 1 PWR ACCIDENT SCENARIOS Scenario

Station Blackout Loss of Feedwater

AM on Secondary Side

AM on Primary Side

lie *

LOCAs with loss of secondary heat sink LOCAs with loss of HPIS

lie lie

Loss of secondary heat sink

*

SG tube ruptures with loss of pressuriser sprays

lie

Some particular examples are considered below:

Accident Management

3 19

24.2.1 Total Loss of Feed Water The scenario here makes the assumption that there is total loss of main feed water together with total loss of auxiliary feed water. This accident is an important contributor to the frequency of accident sequences that are not covered by the safety systems. A probable sequence of events might be as follows, Figure 24. 1. Initially the secondary water level will decrease which causes the reactor and turbine to scram. The steam valves will be closed causing the pressure of the secondary and also the primary side to rise. The pressure of the secondary side will level out at the secondary relief valve set point at about 7.5MPa. After about 20 mins the secondary side of the steam generator will have boiled dry. Pressure control on the primary will be limited by the pressuriser sprays but ultimately the pressuriser will fill with water and water will be discharged through the pressuriser relief valves.

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I N ITI ATIO N O F S E CON DARY B L E E D S E CO N DARY

\

5 F E E DWATER TAN K

o �--�----�--��- 1 000 0 1 000 2000 3 000 4000 5000 6000 7000 s TI M E -.

PRI MARY AN D S ECON DARY P R ESS U R E FIGURE 24.1 TOTAL LOSS OF FEEDWATER WITH SECONDARY FEED AND BLEED

If secondary bleed is initiated at this point then the secondary side pressure can be reduced in a short time down to the saturation pressure of the pre-heated feed water which will then flash to produce steam. This flashing delays further secondary side pressure decrease. However, despite this the primary pressure side can be lowered and the loss of inventory halted. Control and shut off valves between the feedwater system and the steam generator are opened if the pressure on the steam generator side falls too low. This produces a high fluid flow into the steam generator and a pressure peak.

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Accident Management

Eventually a constant emptying of the feedwater tank into the steam generator is achieved, causing a continuous depressurisation of the primary. Eventually the primary pressure will fall sufficiently low that the LPIS pumps can refill the primary side with heated water and the primary side can be cooled by the RHR system. For this procedure to be effective it is clearly necessary that the feed water does not exhaust before the LPIS pumps can operate.

24.3

Computer Code Requirements

This section considers some of the additional requirements for the codes for simulation of transients where some accident management action is taken. A best estimate analysis capability is required and the existing system analysis codes are being extended The RELAP5 and TRAC codes are at an advanced stage in the US [24.5]. In Europe the ATHLET code [24.6] and CATHARE code [24.7] , [24.8] are under active development S ince exact predictions are not possible some measure of the uncertainties in certain parameters is required. These might include peak clad temperatures, collapsed water levels or times in reaching certain setpoints. The best estimate method with uncertainty approach is now starting to be adopted as a licensing approach for certain accidents e.g. LOCAs. Calculations that simulate AM actions are often associated with complex thermal-hydraulic boundary conditions. All the safety and non-safety systems need to be considered in the computer model. Interactions between the systems may be complex and different to interpret. The thermal hydraulic input decks for the codes may be quite large because all primary circuit legs could have to be modelled if the effects of non-symmetric actions are to be assessed. The nodalisation may need to be fairly fine in certain components e.g. in the feed line, surge line and the cross-over legs. This will tend to imply shorter timesteps because many of the codes are semi-implicit and therefore timesteps are limited by the Courant limit For these reasons computational times may be expected to be longer in general than for unrecovered accident predictions. Inevitably transients where AM actions are considered involve phenomena and conditions on the fringe of current experience. Therefore empirical closure relations for the more mechanistic models may need to be regarded with caution and preferably validated against suitable data. For assessment of the consequences of possible accident management actions it may be necessary to perform a number of calculations and therefore input decks need to be as general and flexible as possible. Finally a reasonable degree of accuracy or confidence in the predictions is required. The feasibility of a procedure may depend on the time that it is initiated. It is also important to know the timescale of events in order to judge which accident management actions should be considered and carried out.

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As discussed much earlier in this book in Chapter 5, the leading thennal-hydraulic system codes use sophisticated five or six equation models with additional scalar fields for incondensables e.g. nitrogen, hydrogen and solutes e.g. boron. Systems can be readily built from building blocks or basic components e.g. pipes, volumes etc. Control and safety systems are modelled with logic switches which control e.g. water injection or release. Additionally many of the leading thennal-hydraulics codes are being coupled with melt progression and fission product release codes to provide a capability for beyond design basis and severe accident management simulation. Many of the front-line system codes include system-mimic sub-codes which provide immediate visual display of the system interactions. The more advanced of these can be run in the interactive mode. 24.4

Code Assessments

Many of the major thennal hydraulic experimental programmes have addressed issues and phenomena that are of relevance to AM procedures for managing transients. These thennal­ hydraulic programmes have been summarised earlier. Validation matrices of integral tests relevant to LOCA and transients have been compiled in a CSN I report [24.9] . The parameters in these matrices are the key phenomena and the available test facilities. Examples of significant phenomena include: natural circulation including reflux condenser mode asymmetric loop behaviour break flow phase separation, level fonnation and stratification emergency core cooling and condensation loop seal clearance pressuriser and surge line thennal-hydraulics non-condensable gas phenomena accumulator behaviour. A very detailed review of phenomena is given in a validation matrix for separate effects tests [24. 10] recently carried out by the CSNI. Natural circulation, horizontal flow stratification and loop seal clearance were all studied [24 . 1 1 ] in the LSTF facility, carried out as part of the ROSA-IV programme for PWR. Other tests for BWR were carried out in the ROSA-III programme [24. 12] .

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Accident Management

Many of the tests were concerned with depressurisation of both the primary and secondary side. Both secondary and primary feed and bleed AM measures were investigated. A small-break LOCA test was chosen for a CSNI-ISP (ISP-26) [24. 1 1] and changes in some codes were instigated as a result of various assessments. Loss of feed water tests with primary feed and bleed and also secondary bleed and feed were considered in the LOBI programme. The latter phenomenon was a feature of a test that was the subject of a 'blind' prediction exercise [24. 14] . The important phenomena exhibited included global mixture level formation, stratification, core dry out and also thermal dis­ equilibrium effects from accwnulator injection. These tests have been used for assessment of most of the major system codes including RELAP5 and CATHARE . PKL tests [24. 1 5] relevant to accident management have addressed loss of feed water, secondary feed and bleed, SG tube break with some coolant pumps running and other complex transients. The tests are particularly relevant to the transients and procedures in Gennan PWRs. Consequently they are also being used for validation of the ATHLET code. To validate the French AM procedures various tests are being commissioned in the BETHSY facility. AM tests [24. 16] so far include shut down with natural convection and a multiple failure transient involving a small cold leg break but assuming the HPIS system remains unavailable. The latter has been the subject of a CSNI International Standard Problem and it has been used for validation of many codes including RELAP5, ATHLET and CATHARE. Data from other facilities including LOFT and Semiscale have been used for assessment of the USNRC codes TRAC and RELAP for feed and bleed [24. 1 7] .

24.5

Summary - Preventative Accident Management

Preventati ve AM can have an important contribution towards reducing core melt frequency. Thermal-hydraulic calculations supported by appropriate experimental data are now in place. Best Estimate methods need to be used: further work is required in order to quantify uncertainties. The most effective AM measures are secondary feed and bleed procedures e.g. for transients such as loss of feed water.

24.6

Mitigative Accident Management

In preventative accident management the emphasis is in preventing potentially beyond design basis accidents reaching the stage of core degradation. However, in the unlikely event that core damage is reached, operators will require instructions as to what actions they should take. Ideally the operator would like to know the state of the reactor and in particular the state of the core. However much of the instrumentation may not survive if the conditions become too

323

Accident Management

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1 000 900 800

w �

700

w

600

2 I-

500 400

I I I

EXP E RI M E NT :

I l

--\

CALCU LATI O N

0

2000

4000

6000

8000

1 0000

TI M E (SECON DS)

M AXI M U M C LAD D I NG TE M P E RATU R E

FIGURE 24.2 BETHSY TEST 9.1b; SMALL BREAK TEST WITHOUT HPIS

severe [24.18] e.g. high temperatures are reached and there may be considerable uncertainty. There may be further uncertainty associated with the boundary conditions or lack of understanding of severe accident phenomenology [24 . 1 ] . There are little information on the progression to severe accidents in LWRs. T MI-2 provides the main source of data but it has taken many years to understand the complexities of this single accident. The remainder of this section considers uncertainties in core degradation behaviour, the interaction of core debris with coolant and structures and the overall behaviour of the primary circuit if the core is substantially degraded. The emphasis is on uncertainties that could impact on accident management decisions and the possible consequences of decisions.

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Accident Management

The main objective in any AM action is to develop AM strategies to prevent threats to the integrity of the containment building and to achieve cooling of the molten debris. In practice there are relatively few actions that can be taken when considering severe accident mitigation to prevent further core melting. Possible options include adding water to the degraded core or possibl y depressurising the system. However lack of understanding of the phenomena could mean that the consequences of such actions are detrimental rather than beneficial.

24.7

Water Addition to a Degraded Core

Even without water addition the uncertainties in phenomenology are in general greater, further on into the accident. The early phase stage up to the time of relocation of molten Zircaloy eutectic material is reasonably well understood and is characterised by a temperature excursion and hydrogen production associated with the strongly exothermic reaction of Zircaloy with steam. Accordingly code predictions are reasonable for the early phase but must be regarded as becoming increasingly more uncertain further on into the accident. The addition of water to an overheated core gives rise to even greater uncertainties. The understanding of phenomena is less and certainly code performance is worse. It is now established that the addition of water results in increased oxidation initially in the upper regions of the core (or bundle in the case of experiments) thus increasing temporarily the heat and hydrogen production. In some cases the increase may be due to previous steam starvation (i.e. the limiting of the oxidation kinetics by insufficient steam production). However, it may also be due to a spalling mechanism giving rise to a new metallic surface. The timing of substantial blockage formation is uncertain and it is notclear how coolable such blockages might be. The accident at TMI-2 provided evidence that prolonged heat up of the blockage continued even after the pumps had been turned on: the phases of materials found in the blockage were consistent with high temperatures having been attained. It is also known that the addition of water to a partially degraded core can cause fragmentation and destroy parts the core, particularly free standing highly oxidised fuel rods. It appears that this happened in TMI-2 [24.19] . Structural damage could also occur due to thermal shocking. The important issue from the point of view ofaccident management is not so much concerned with whether water should be added or not: without cooling core damage will inevitably continue. The question to answer is how it can be determined whether the core or blockage is coolable and how can the operator tell whether core damage has been arrested. 24.8

Primary Circuit Integrity

Natural circulation in the steam phase could provide a mechanism for failing the primary circuit in high pressure sequences such as Station Blackout sequences. The effects of natural

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circulation are difficult to predict for a number of reasons. As the steam is oxidised, it is converted to hydrogen and this may effect natural circulation because of stratification or reduction of heat transfer in the steam generation. The system code models tend to be based on one dimensional pipe models and the natural circulation flows observed in experiments are based on counter current flows within pipes. Finally fission products may plate out at vulnerable parts of the primary system, causing local heating which might increase the potential for failure but which also might reduce the potential for natural circulation. These two effects could be cancelling. A breach of the primary circuit would cause depressurisation and would eliminate concerns about a high pressure melt injection from the bottom of the vessel (following lower head failure) into the containment. This raises the question as to whether it might be advantageous to deliberately depressurise the primary circuit to avoid this melt ejection [24.20] . The disadvantages of depressurisation are that primary coolant inventory would be lost and that if the core is molten and eventually is released through the bottom of the vessel into water, then there might be an increased probability of a steam explosion.

24.9

Late Phase Melt Progression

The are considerable uncertainties associated with the coherency and paths of melt relocation into the lower head. The only data at realistic scale are from TMI-2 which showed that the major relocation route took place through the core bypass. The details on the melt relocation route which in tum affects the melt relocation rate are important because these mechanisms impact on the coherency of any melt water reaction in the lower head and therefore the potential for steam explosions. These parameters may be design specific since not all PWRs has the same internal structural geometry in the side and below core region. The uncertainties in these late phase melt relocation mechanisms impact on accident management decisions concerning depressurisation because of the implications regarding possible steam implosions.

24.10 Melt-Water Interactions in the Reactor Vessel

The extent of melt water reactions will depend on the rate of flow of melt, the overall quantities of melt involved and also the amount of water available in the lower head. For small intermittent flows the melt will be quenched as solid particles. For larger or greater quantity flows there may be a greater tendency to form melt pools. In TMI-2 the end state of the debris (which may not have been like the state of the debris immediately after relocation) consisted of particulate debris but also of larger more coherent debris regions.

326

Accident Management

24.11 Steam Explosions

The uncertainty over the likelihood of damaging steam explosions presents a major problem in taking decisions over accident management actions. This applies both to judgements on depressurisation and also adding water. Much of the research on steam explosions has concentrated on the energy releases of melt pouring into water and the potential for containment failure [24.2 1 ] . However from an accident management point of view the releases from pouring water over or on to melt poo ls are also of prime interest Concerning depressurisation there are opposing mechanisms which need to be considered. Depressurisation may increase the likelihood of a steam explosion: it seem s to be generally agreed that decreasing pressure increases the likelihood of triggering. Alternatively depressurisation would also decrease the contribution of pressurised material to the damage potential. On coolant addition it seems unlikely that if coolant were available that the choice would be made not to add it. However a rising water level could trigger a greater amount of melt to relocate to the lower head increasing the potential for a steam explosion. Similarly pouring water on a melt pool in the lower head could be a risk.

24.12 Failure of the Reactor Vessel

Possible failure mechanisms are that the vessel could undergo a major integral failure or an instrument penetration tube could be ejecte�. For reasons already discussed the main issue of interest in accident management is whether they could be an ejection of melt in a high pressure sequence following the failure of a tube weld. Whether this will occur will depend on the material properties of the materials at high temperature and the extent of vessel deformation. A key factor could be the coolability of debris beds sitting on the lower head. This is somewhat uncertain for beds with particles of irregular size and shape as were found in the TMI lower head [24.22] . The predictions of the coolability of the material following accident management action (e.g. coolant injection or depressurisation) are difficult 24.13 Threat to the Containment

The major concern is whether finally fragmented melt could be ejected into the containment, leading to rapid heat transfer and early failure. The main interest from the point of view of accident management is whether to avoid this phenomenon by intentional depressurisation of the primary circuit Uncertainties make this judgement difficult. High pressure melt ejection could be mitigative if direct containment heating did not occur. Alternatively if some action to depressurise

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327

involved an increased relocation of melt to the lower head and a penetration tube failed quickly then high pressure melt ejection could still occur.

24.14 Ex-Vessel MeltlWater Interactions

The interaction of melt with water in a reactor cavity is likely to be mitigative from the point of view of direct containment heating although a well tamped stream explosion could occur e.g. if the cavity was full of water. Whether this energy could result in a damaging missile is less clear and would depend on the strength of the surrounding concrete structure, the amount of water in the cavity and how well the vessel was secured to its pipework and surrounding structures. The accident management issue is whether the cavity should be flooded or not. 24.15 Molten Core/Concrete Interactions

The interaction of molten core with concrete is complex and is important because itcould lead to failure of the containment barrier and thus to the release of fission products. It also results in copious aerosol production which increases the source term to the containment. The main interest from the point of accident management concerns the effects ofadding water on the erosion rate of the concrete. The rate of erosion is not necessarily reduced by the presence of water [24.23] . 24.16 Summary - Mitigative Accident Management

There are clearly uncertainties in assessing the exact state of the core once severe accident conditions have occurred. The main consequences are that it is difficult to detennine when a particular action has been effective in meeting the objective. After water has been added how can it be detennined whether the material has been cooled? If there are uncertainties about whether a particular measure is likely to have beneficial or adverse consequences, it may be sensible to take action to avoid having to take a judgement on applying that particular measure or not An example here might be depressurisation of the circuit at an early stage in the accident progression when the core is still largely intact REFERENCES

24. 1

24.2

V Teschendorff, W Pointer, B Putter "Development and Assessment of Computer Codes for Accident Management, CSNI Specialist Meeting on Transient Two Phase Flow, Aix-en-Provence. April 6-8, 1992. E F Hicken, U Erven, E Kersting "Gennan Accident Management Program" Proc 1 5th Water Reactor Safety Infonnation Meeting, Gaithersburg, MD, Oct 26-29, 1987 Vol 6, pp 1 33- 148.

328 24.3

24.4

24.5

24.6

24.7 24.8 24.9 24. 1 0 24. 1 1

24. 12 24. 1 3

24.14

12. 1 5 24. 16 24. 1 7

24. 1 8 24. 19

Accident Management

OECDINENCSNI Specialist Meeting on Severe Accident Management Programme Development, Rome, Sept 23-25, 199 1 . Summary and Conclusions, Report NEN CSNI/R(92)6. W Korbach and E J Kersting "Prevention of Core Melting Under High Pressure with Bleed and Feed in PWRs", Proc. OECD Specialist Meeting on International Coolant System Depressurisation, Garching, F R Germany, June 12- 14 1989. CSNI Report 158 (1989) pp 225-243. L M Shotkin "Completion of Thermal-Hydraulic Code Development TRAC-PF1/ MOD2, RELAP5/MOD3, and TRAC-BF1/MOD 1", Proc. 1 8th WaterReactor Safety Information Meeting, Rockville, MD, Oct 22-24, 1990, Vol 4, pp 339-345. M-J Burwell, G Lerchl, J Miro, V Teschendorff, K Wolfert, ''The Thermal-hydraulic Code ATHLET for Analysis of PWR and BWR Systems" Proc. 4th Intern. Topical Meeting on Nuclear Reactor Thermal-Hydraulics (NURETH4), Karlsruhe, Oct 101 3 , 1989. Vol 2, pp 1234- 1239. F Barre and M Bernard "The CATHARE Code Strategy and Assessment" Nuclear Engineering and Design 124 (1990) 257-284. D Bestion "Capabilities and Limitations of Thermal-Hydraulic Codes" CSNI Specialist Meeting in Transient Two-Phase Flow, Aix-en-provence, April 6-8, 1 992. OECDINENCSNI, "CSNI Code Validation Matrix of Thermal-Hydraulic Codes for LWR LOCA and Transients", CSNI Report 1 32, March 1987. OECDINENCSNI, Separate Effects Test Matrix for Thermal-Hydraulic Code Validation NENCSNI!R(93) 14, December 1993. H Asaka and Y Kukita, "Intentional Depressurisation of Steam Generator Secondary Side during a PWR Small-Break Loss of Coolant Accident" Proc. 1st JS:ME-AS:ME Joint Inti. Conf. on Nucl. Engrg. (lCONE- 1), Tokyo, Nov. 4-7 1 99 1 , vol 2, pp 197202 (199 1 ) Y Kukita et a l "Intentional Coolant System Depressurisation: Experimental S tudies in the ROSA-III and ROSA-IV Programs", CSNI Report 1 58, 1989, pp 449466. Y Kukita et al "OECDINENCSNI International Standard Problem No 26 (lSP-26) ROSA-IV LSTF Cold-Leg Small-Break LOCA Experiment" Comparison Report, Japan Atomic Research Institute (JAERI) 199 1 . B Worth et al "Assessment of TRAC-PF 1 , RELAP5/MOD1 -EUR, CATHARE and DRUFAN Codes Against LOBI-MOD2 Test BL- 12", Proc. NURETH-4, Karlsruhe, 1989, Vol 1 , pp 190- 1 77. B Branc, R Mandl and H Watzinger "Feed and Bleed Experiments in the PKL Test Facility" CSNI Report 1 58 (1989) pp 383-402. P Clement, T Chataing, R Deruaz "PWR Accident Management Related Tests, NVRETH-6, Grenoble Oct 5-8, 1993. W E Driskwell, W E Hansen "Summary of ICAP Assessment Results for RELAP5/ MOD2 "Proc 16th Water Reactor Safety Information Melting, Gaithersburg, MD , Oct 24-27, 1988, Vol 4, pp 161- 167. W C Arcieri and D J Hanson "Instrumentation Availability for a Pressurised Water Reactor with Large Dry Containment, NUREG/CR-569 1 , March 199 1 . P Kuan, J L Anderson and E L Tolman ''Thermal Interactions During the Three Mile Island Unit 2, 2-B Coolant Pump Transient, Nucl. Tech. 87, 977, 1989.

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24.20 D J Hanson et al "Depressurisation as an Accident Management S trategy to Minimise the Consequences of Direct Containment Heating, NUREG/CR-5447, 1 990. 24.2 1 T G Theofanous et al "An Assessment of Steam Explosion-Induced Containment Failure, Nucl. Sci. Eng. 97, 259, 1987. 24.22 J M Broughton, P Kuan, D A Petti and E Tolman "A Scenario of the Three Mile Island Unit 2 Accident, Nucl. Sci. Eng. 87, 34� 1 989. 24.23 R E Blose, J E Gronager, A J Suo-Anttila and J E Brockmann "SWISS: Sustained Heated Metallic Melt/Concrete Interactions with Overlying Water Pools, NUREG/ CR-4727, June 1 989.

331

Chapter 25 ADVANCED REACTORS

25.1

Introduction

New design options are currently being considered to improve the safety of new generation nuclear plants. As a consequence various new reactor concepts with passive safety features are being developed. These are for a wide range of reactor types in LWR (Light Water Reactor), HTGR (High Temperature Gas-Cooled Reactor) and LMR (Liquid Metal Reactor). Particular attention has been paid towards developing advanced concepts for LWRs. Some of these developments are outlined in this chapter. Design requirements for various countries are given in [25 . 1 ] and [25.2]. Some of the proposed advanced reactors under consideration are similar to present operating reactors, retaining the same basic systems, but aiming to eliminate known shortcomings. Reactors in this class are often termed as 'Evolutionary' The intention in other designs is to remain close to existing designs but to take advantage of passive gravity driven systems for emergency core cooling and residual heat removal. Reactors in this class are tenned as 'Passive'. The final class ofreactors under consideration are based on significant departures from existing designs: these generally also exhibit passive and gravity driven features. These reactors are described as 'Revolutionary' The exposition here concentrates on the features of design that are specifically to improve safety. Other aspects which have considerable influence on design criteria, e.g. economy, operational features, maintenance ease, construction time etc are not considered.

25.2

Design Concepts

2S.2.1 Evolutionary Designs

The evolutionary designs are based on previous designs but include improvements that have been incorporated following accumulated experience with the previous design. Many of the improvements have originated from ideas emanating from a number of countries indicating the effect of international influences on design. Some of the current leading evolutionary designs are summarised below.

332

Advanced Reactors

PWRS

The APWR 1300 design by Westinghouse - Mitsubishi [25.3] is derived from the Westinghouse 1 300 MWe four loop PWR plant design. The power density is reduced by increasing the number of fuel pins for the some output and it has increased strength in the fuel assembly design. It has an increased pressuriser volume and increased containment volume. There are improved engineered safety features particularly in the high pressure injection residual heat removal and containment spray systems. Similar improvements have been put forward for the Westinghouse APWR 1000 design [24.5] . This concept have been derived from conventional 3-loop PWR design. The System 80+ [25.5 } design has been developed by Combustion Engineering and is an improved version of the System 80 PWR in operation in the USA. The core outlet temperature is somewhat lowered improving the safety margin. Here again the pressuriser volume has been increased in order to mitigate transients and there are improvements in the engineered safety features. There are also improvements in the containment design to mitigate severe accidents.

ACCUIII (I o f :!)

o PRI MARY CI RCU IT SAF ETY SYSTE M S

FIGURE 25.1 WESTINGHOUSE AP600

Advanced Reactors

333

The Konvoi 95+ [25.6] design is based on the Konvoi plants currently in operation in Gennany. The main safety improvements concern: improved preventative and mitigative accident management procedures and microprocessor control systems. A new French design N4+ is based on the current N4 PWR of Framatome. It aims to have an improved design against operator errors and an improved containment design to contain any loads arising from severe accidents. NPI, a subsidiary of Framatome and Siemens is designing a new common French-Gennan PWR. The NPI concept [25.2] aims to take advantage of the attributes of both reactor types, N4 and Konvoi. BWRs

This GE, Hitachi and Toshiba ABWR I 300 [25.2] plant is derived from the US BWR6 and the latest 1 100 MWe Japanese plant Some experience has been incorporated from Gennan and Swedish current BWR designs. The ABB design BWR 90 [25.7] is based on the latest ABB designed BWRs in Sweden. Major safety advancements include a pressure suppression containment with filtered venting and improved engineered safety systems. 25.2.2 Passive Designs AP600

The AP600 reactor [25.2] , [25.8] , [25.9] is based on proven PWR technology derived from standard 2 loop Westinghouse PWRs. Its main features include a lower power density core, passive safety injection, residual heat removal systems and containment cooling. The low power density core is larger with 145 ( 1 7 x 17) fuel assemblies than the standard design with 121 (16 x 16) fuel assemblies. It also has lower enrichment and improved DNB and LOCA margins. The primary circuit configuration consists of two hot legs, four cold legs and two steam generators. The pressuriser is larger than standard to provide improved transient control. The containment is constructed of steel and surrounded by a concrete shield. The AP600 passive safety systems, a reference design is shown in Figure 25. 1 , rely on gravity, natural circulation and compressed gas for operation. They do not depend on operators or active safety systems. There are three main components: Passive residual heat renewal (PRHR) Passive containment cooling system (pceS)

334

Advanced Reactors

Passive safety injection (PSI). Decay heat removal is supplied by a passive residual heat removal exchanger together with passive containment cooling. The heat exchanger resides in a large tank within the containment. There is sufficient water to absorb heat for about two hours before boiling occurs. Afterwards any steam produced would be condensed on the steel containment wall and drain back to the tank. The steel wall is cooled by water flowing by gravity from a tank contained in the shield building structure: the water fonns a water film over the structure. In addition air flows upward by natural circulation between the steel containment and the concrete shield. The containment is therefore cooled by heat conduction through the steel shell and the water fIlm by convective heat transfer and evaporation mass transfer from the film to the air and also by radiation from the film surface. The containment storage tank is sized to have sufficient water to last three days: after that time external sources of water could be provided from e.g. the rue protection system. The passive safety injection system is to provide inventory make-up in the event of LOCAs. There are three water sources that supply water injection in a passive manner. Two core make-up tanks (CMTs) are located inside the containment and supply water by gravity in the case of minor loss or leakage. Two accumulators, pressurised by compressed nitrogen supply additional water for larger LOCAs or drops in coolant system pressure. Finally the in-containment refuelling water storage tank supplies any make-up once the system is depressurised. SBWR

The Simplified Boiling Water Reactor (SGWR) [25.2] , [25. 10] is a 600 MWe design developed by General Electric and other foreign partners. The core has a lower power density compared with conventional BWRs. The reactor pressure vessel is high -24m in order to establish the natural convection core flow. There are no main coolant recirculation pumps. The quantity of water above the core is large which ensures the time to core uncovery is longer under accident conditions compared with conventional BWRs. There are three main engineered safety systems: Gravity Driven Core-cooling System (GDCS) Isolation Condenser System (ICS) Passive Containment Cooling System (PCCS). The GDCS and the water in the suppression pools provide water sufficient to flood the entire containment to one metre above the active core. The design objective is that the SBWR should not uncover in the event of a LOCA.

Advanced Reactors

335

The purpose of the ICS is to remove decay heat during reactor isolation or under transient conditions. The ICS is made up of three independent loops connected to the steam region and to the downcomer which transfer heat through heat exchangers immersed in water pools and which vent to the atmosphere. When the ICS is in operation, e.g. on a high pressure reactor trip, the condensate from the IC units drains into the downcomer. Any non-condensable gases that might build up in the condenser are routed to the suppression pool . The PeCS purpose is to remove heat under LOCA conditions. The design aims to provide heat removal for at least 72 hours without any operator intervention. In the PeCS there are two independent loops connecting the drywell to the passive containment cooling condenser which is located outside of the containment. The condenser receives a steam gas mixture. Condensate is drained back to the GDCS pool and non­ condensable gases are routed to the suppression pool in the wetwell. The PCCS is initiated by pressure increase in the drywell. Other Designs

Design research on advanced PWR containments is in progress in Germany [25. 1 1 ] , [25. 12] . Th e objective is to achieve significant reduction i n fission product release under severe accident conditions, including core melt, steam explosions and hydrogen deflagration. The containments under examination consist of a strong steel shell surrounded by a pre­ stressed concrete structure. They are designed to withstand high internal static and dynamic pressures and also impact loadings from vessel parts e.g. the vessel head. Beneath the core, it is proposed that a core catcher is installed to withstand high pressure and basement erosion. The core catcher is installed to withstand downward moving missiles e.g. the lower pressure vessel end cap. Cooling of the molten core is via water evaporation and by a water/vapour circulation system to transfer heat to the steel shell of the containment. This is dissipated finally by natural air convection through the passage between the steel shell and the outer concrete containment. New designs for both medium size PWRs and BWRs are under development by the Japanese e.g. the Mitsubishi MS-300 and MS-600 PWRs and the Hitachi HS BWR [25.2] . 25.2.3 Innovative Designs

The PIUS design [25.2], [25 . 13], [25. 14] concept has been developed over the past decade and verified by ABB Atom. The coolant pumps circulate hot low-boron concentration water but the core is also openly connected in a natural-circulation circuit to a large pool of highly borated water. Under normal operation there is a hydraulic balance between the primary circulation loop and the pool , boron water ingress will occur under transient conditions should this balance be disturbed and automatic shut down occurs.

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336

Another concept has been developed by a consortium involving the partners ABB, Combustion Engineering, Rolls Royce, AEA Technology and Stone and Webster. The Safe Integral Reactor (SIR) [25. 1 5] is a small integral reactor with passive safety systems. The design is integral in the sense that all the major components, core, steam generators, reactor coolant pumps and pressuriser are all contained in a single pressure vessel. The SIR reactor includes various engineered safety systems: Emergency Coolant Injection System (ECIS) Secondary Condensing System (SCS) Safety Depressurisation System (SDS) Passive Containment Cooling System (PCCS). In summary safety related design features may be categorised under the following main headings: Enhancement in system reliability Improved opportunity for Accident Management both for preventative and mitigative measures Inherent or passive features in place of powered systems. 25.3

New Phenomena

Thennal-hydraulic phenomena relevant to current LWRs are discussed extensively in [25. 1 6], [25 . 1 7] and [25 . 1 8]. The evolutionary designs are not expected to exhibit new thennal-hydraulic phenomena. However new kinds of phenomena and also new accident scenarios have been postulated for the passive and innovative designs. Many of these designs rely on depressurisation of the primary loop in order to allow gravity driven injection systems to discharge water into the core and avoid the risks of high pressure core melting. Low pressure phenomena tend therefore to be characteristic of advanced plants. In many designs the primary circuit system is closely coupled with the containment. Computational tools need therefore to be developed to predict the integral response. There may be other phenomena associated with new systems. There may be different and unexpected interaction between systems particularly if these are now passive compared with active systems in current generation plant New components may require special models. These may be new risk dominant accident sequences that need to be considered.

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337

Most LWRs utilising passive decay heat removal systems employ water boiling and steam condensation for transport of heat from the core to the ultimate heat sink. In these designs. boiling is the common heat transfer mode in the core region to remove the decay heat. However the routes to the ultimate heat sink differ between designs. There are generally two main stages: the transfer of heat from the core by natural circulation to a condenser system. the transfer of heat from the condenser system to the ultimate heat sink. The main objective in safety studies is to assess the effectiveness of the different systems in respect to certain key licensing questions? Will a particular system initiative correctly and reliably from a range of different accident conditions? Is there sufficient decay heat removal capacity? Will a system operate as intended? Condenser Pools

Most condenser pools are sized to provide a large heat sink and contain large volumes of water. Heat may be supplied to a relatively small volume of water and the heat transfer may be degraded if there is poor thermal mixing or stratification. Both condensation within heat exchanger tubes or direct contact condensation may occur depending on the design. Both these phenomena could be adversely affected by the presence of non-condensables. Non-Condensables

Non-condensable gases are important in ALWR safety assessments for various reasons. The primary circuit is closely coupled with the containment which will be full of air or possibly inert gas. Condensing processes in the containment will depend on the effects of non­ condensables in wall heat transfer and with possible thermal stratification. Non-condensables will also affect the efficiency of direct contact condensation processes e.g. as in an isolation condenser. The accumulators are pressurised with nitrogen gas. The effects on the primary system response and the impact of gravity driven ECCS are not get properly understood. Gravity Driven ECCS Phenomena

There are a number of questions and uncertainties associated with the depressurisation of the primary circuit to allow gravity driven injection systems to operate. Repressurisation following the operation of ECCS could reduce further coolant injection. The interaction of the various passi ve safety systems amongst themselves and also with non-safety systems may be difficult to predict.

Advanced Reactors

338 Natural Circulation

Natural circulation, particularly in coupled systems, is a complex phenomenon and often exhibits unexpected behaviour. It is a design objective in many advanced plants that systems should be simplified but in many designs the circuits are still relatively complicated. The presence of non-condensables will also affect natural circulation. Natural circulation flows may not initiate or continue as expected if flow regimes change e.g. due to becoming two­ Qhase. Small Driving Forces

Circulation between various parts of the primary circuit and containment system depends in many cases on small driving forces, once the pumps are switched off. Valves may not operate as intended. Failure of valves to open might prevent coolant injection, failure to close could result in loss of inventory.

25.4

Experiments

The AP600 design development programme [25. 19] , [25.20] , [25.21 ] aims to provide specific data for safety analyses, to confirm the construction and systems operation and to provide final design detailed information for all unique system features and components. The tests fall into three general categories: 1.

Passive Core Cooling System tests

2.

Passive Containment Cooling System tests

3.

Component Design Verification tests.

The Passive Core Cooling System tests are for verification of the AP600 passive safety systems. The series of Passive Containment Cooling System tests are concerned with heat passage through the containment to the ultimate heat sink. The Component Design Verification tests are feasibility tests concerning the use of canned motor reactor coolant pumps in the primary circuit, the operability of a fixed in-core detector and the flow mixing characteristics in the reactor vessel. Programmes ofPassive Core Cooling System tests include: long-term cooling tests at Oregon State University and high pressure integral systems tests, large scale, full height, full pressure integral system tests (SPES-2) conducted the ENEL/ENEA's SIET facilities in Piacenza, Italy. Other experiments are taking place in the Japan Atomic Energy Research Institute (JAERI) owned Rig of Safety Assessment (ROSA) Large-Scale Test Facility (LSTF) [25.22]. Since natural circulation is a generic phenomenon, many of the data available on a range of reactor designs are relevant to both PWR and BWR This is particularly true for data relevant .

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339

to accident analysis, where two phase flow regimes may exist, and the differences between PWRs and BWRs (from the coolant phase change point of view) become less important. The Paul Scherrer Institute has recently initiated the major new experimental and analytical programme ALPHA, which is aimed at understanding the long-tenn decay heat removal and aerosol questions for the next generation of Passive Light Water Reactors [25.23] . The ALPHA project currently includes four major items: the large-scale, integral system behaviour test facility PANDA, which will be used to examine multidimensional effects of the SBWR decay heat removal system; an investigation of the thennal hydraulics of natural convection and mixing in pools and large volumes (LINX); a separate effects study of aerosol transport and deposition in plena and tubes (AIDA); while fmally, data from the PANDA facility and supporting separate effects tests will be used to develop and qualify models and provide validation of relevant system codes. The LINX programme is concerned with an investigation of natural circulation and mixing phenomena in single- and multi-phase/multicomponent systems in large pool s. The areas of interest and investigation include the mixing of hot and cold liquids in open pools, the mixing and energy distribution within liquid pools resulting from the submerged injection (venting) of steam and gas mixtures, and the mixing of steam, nitrogen and possibly other gases in large, interconnected volumes. The AIDA programme is mainly relevant to severe accident conditions where fission products in the fonn of aerosols may escape from the RPV into the various compartments of the reactor containment The condenser units may be subjected to aerosols. The possible fonnation of an aerosol layer on the inside tube surface may affect the heat removal characteristics of the system, and plugging of some tubes may substantially degrade the condenser efficiency and impact on long-tenn pressurisation of the containment. The GIST (GDCS Integrated Systems Test) is an 1/508 scale test conducted by General Electric to study the Gravity Driven Cooling System for the SBWR. This test established the feasibility of the G DCS as an Emergency Core Cooling System for the SB WR during low power and low flow conditions. These tests were completed in 1989 and were also used to qualify the TRACG code [25.25] . GIRAFFE is a full height 1/400 scale integral systems test facility at the Toshiba Kawassaki Laboratories in Japan. Full height volume scaled separate effects tests on condenser perfonnance [25.26] , [25.27] and nitrogen venting and, [25.28], integral system perfonnance tests were perfonned to demonstrate the feasibility of IC/PCCS. The PANTHERS tests to be conducted at S IET, Italy are prototype heat exchanger tests for both the IC, a half unit (single module) is to be tested, and for the PCCS, a duel module (full unit) test is planned. These tests are steady state component tests for mechanical qualification of the condenser designs, and will also provide local heat flux data for multi-tube heat exchangers.

Advanced Reactors

340

The UNIVERSITY tests conducted at University of California, Berkeley (VCB) and Massachusetts Institute of Technology (MIT) are full length single tube condensation tests designed to develop heat transfer correlations in the presence of significant amounts of non­ condensables.

25.5

Accident Assessments

LOCAs

The LOCA perfonnance of the AP600 Passive Safety Systems have been assessed by Kemper (Westinghouse) et al [25. 19]. Calculations were perfonned for a range ofbreak sizes and locations to fmd the worst set of conditions for depressurisation of the reactor coolant system. The main objective was to establish that the core remained covered. The resulting break spectrum study indicated that only the double-ended guillotine shear of the direct vessel injection line (a line that feeds the emergency core coolant flow into the vessel) resulted in a momentary core uncover. For all other small-break cases, the core remained covered as the reactor coolant system depressurised. The passive safety systems provided sufficient mass flow to the reactor vessel such that even under conservative assumptions, the core remained covered and in a coolable state. LOCA studies have also been carried out by Fisher [25.30] and Bianconi [25.3 1 ] . Intact Circuit Faults

Figure 25.2 shows some RELAP5 calculations to investigate the effectiveness of the passive cooling systems following an intact circuit fault in a representative advanced PWR with similarities to AP600. A principal feature is the loss of cooling capacity of the steam generators. The analysis is relevant to initiating faults such as station blackout, loss of feedwater or a steam generator tube rupture. Passive cooling circuits inevitably would be isolated by valves during nonnal operation to ensure that the pond cooling capability remains at its maximum at any time. At the detection of a fault the valves would be opened to pennit natural circulation to take place. There is no 'artificial' means to ensure the circulation starts. Circuits and initial states are designed to encourage circulation to start. In the design considered here this involves preventing reverse flow which would occur until pump run down is complete (effectively 10 see s after the start of the transient). ....

Once the valves in the circuit are opened no further engineered action takes place. In the long tenn decay heat is ultimately transferred by natural circulation processes to the atmosphere. The calculations are for a simple representation of the primary circuit consistent with including all the main components likely to play a significant role in an intact circuit fault transient The main components of the circuit with the Passive Heat Removal System (pHRS) are the cold legs, vessel, hot legs, the pressuriser surge line, the cooling pond and heat exchanger.

Advanced Reactors

2.0 1 .8 1 .6 ro

1 .4

6

1 .2

a..

w

� =>

"" "" w

� a..

1 .0 0. 8 0.6 0.4 0.2 0.0

0

5000 1 0000 1 5000 20000 2 5000 3 0000 TI M E (SECON DS)

(a) PRESS U R E 450 400 3 50 300 VI C'I �

2 50 200 1 50 1 00 50 00

5000 1 0000 1 5000 20000 2 5000 3 0000 TI M E (SECO N D S)

(b) CI RCU IT F LOW FIGURE 25.2 PASSIVE RESIDUAL HEAT REMOVAL

34 1

342

Advanced Reactors

The calculation is shown continued to 3()()()()s. Apart from some oscillation in the first 500s of the transient, a relatively smooth pool level decrease is predicted. Falling pressure, even in the absence of a break in the circuit is observed. There is a rapid fall in the core temperature to about 425K in 5000s . The temperature fall is then much less rapid and is controlled by the saturation temperature in the passive removal heat exchanger (PRHRX) circuit. Thus even with a relatively simple model it is possible to establish the main features of the PRHRX circuit behaviour.

25.6

Computer Code Requirements

Passive decay heat removal prediction in advanced plants places new challenges on computer codes [25.32] , [25.33] . Systems codes have been exercised largely and validated for forced flow high pressure accident conditions where the flows are predominately one-dimensional. In advanced plants engineered systems (e.g. pumps) are being replaced by passive gravity driven systems. The system codes require validation under these conditions and also for a range of pressures. The effectiveness of heat sinks depends on three dimensional flow phenomena often in the presence of incondensables e.g. within a containment or in a condenser pool . These flows require modelling capabilities beyond those of the current generation system codes. The need for further code developments for modelling of advanced plants is being recognised by the code development sponsors. The latest version of system codes include improved condensation modelling in the presence of incondensables. The need for improved numerical techniques particularly for low pressure buoyancy driven flow has been recognised and new methods are under development Later versions of the codes are being developed to also include certain specific engineered safety features. In order to determine containment pressures a coupled primary circuit/containment code capability is required. Integrated primary circuit/containment codes are under development. A key area of weakness in current computational capabilities concerns modelling the condenser pools. The effects of three dimensional mixing cannot be modelled adequately and problems have been encountered with representation of poo ls with system codes. Computational Fluid Dynamics Codes are needed to investigate pool mixing phenomena. These codes may need to be adapted for multiphase flow conditions but they include an appropriate three-dimensional modelling capability. These codes can then in tum be used to validate the system codes. An area of concern in some designs is the system response under small break LOCA conditions. Adverse interaction between the gravity driven systems could result in failure to depressurise and therefore failure of the gravity driven systems to inject coolant. Codes require validation to ensure system responses are predicted correctly.

Advanced Reactors

343

Scoping calculations carried out to investigate plant response under intact fault conditions show that for both representative primary and secondary condensing systems, natural circulation does initiate and continue to be effective even under two phase flow conditions. The timescale for boil off of all available water is long, typically -10 hours or more for most current designs.

REFERENCES

25. 1

25.2 25.3 25.4

25.5

25.6

25.7 25.8 25.9 25. 10 25. 1 1

25. 1 2

25. 1 3 25. 14 25. 1 5 25. 16

Proceedings of an International Workshop on the Safety of Nuclear Installations of the Next Generation and Beyond. Sponsored by IAEA and ANL . Chicago, III, USA, August 28-3 1 , 1989. IAEA Technical Committee Meeting on "Progress in Development and Design Aspects of Advanced Water-Cooled Reactors", 9-12 September 199 1 , Rome, Italy. Status of Advanced Technology and Design for Water Cooled Reactors: LWR IAEA­ lECI:><X 479, page 79-86. APWR 1000 Westinghouse Summary Document, available through: Mr H J Bruschi, Director, Nuclear Plant Programs, Westinghouse Electric Corporation, PO Box 355, Pittisburgh, Pennsylvania 1 5230 USA. System 80+TM Design Certification by G A Davis. Combustion Engineering Company Document, available through: Combustion Engineering, Inc. 1 ()()() Prospect Hill Road, Windsor, Conneticut, 06095 USA. German Design Features and Safety Goals for the Next Generation of PWRs. P. Meyer. IAEA Technical Committee Meeting to Review Safety Features of New Reactor Design, Vienna, Nov 1 1 - 1 5, 199 1 . A Rastas, C S undqvist"AdvancedLight Water Reactors: A Finnish-S wed ish Proposal" European Nuclear Conference 1990, Lyon, France. R Vijuk, H Bruschi "AP600 Offers a Simpler Way to Greater Safety, Operability and Maintainability" Nuclear Engineering International, Nov 1 988. H Bruschi, R Vijuk "AP600 Plant Design Meeting the Industry Needs" 1990 International Joint Power Generation Conference Oct 2 1 -25, Boston MA, USA. R J McCandless, J R Redding "SB WR: The Key to Improved Safety Performance and Economics" Nuclear Engineering International, November 1 989. H Hennies, G Kessler, J Eibl "Improved Containment Concept for Future PWRs" 5th International Conference on Emerging Nuclear Energy Systems, Karlsruhe, July 36, 1989. H Hennies, G KeBler "Improved Containment and Core Catcher for Future LWRs and FBRs" 6th International Conference on Emerging Nuclear Energy Systems, Monterey, USA, June 1 6-2 1 , 1 99 1 . K Hammerz, L Nilsson, T Pedersen, Ch Pind ''The PIUS Pressurised Water Reactor: Aspects of Plant Operation and Availability" Nuclear Technology, Vo1 9 1 , J ul Y 1990. "PIUS - A New Generation of Nuclear Power Plants" Company Documentation, available through: ABB Atom, S - 72163 Vasteras, Sweden. Nuclear Energy, 199 1/30. No 2, April 199 1 . CSNI Group of Experts "CSNI Code Validation Matrix of Thermalhydraulic Codes for LWR LOCA and Transients" CSNI Report 1 32, Paris (F), March 1 987.

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25. 17 CSNI Group of Experts: "Thennohydraulics of Emergency Core Cooling in Light Water Reactors" CSNI Report 1 6 1 , Paris (F), Oct 1 989. 25. 1 8 US NRC: "Compendium ofECCS Research for Realistic LOCA Analysis" USNRC Report NUREG 1 230, Washington (USA), Dec 1 987 25. 1 9 C L Caso "Putting AP600 to the Test" January/February 1 993 ATOM 426. 25.20 D Squarer et al ''Test Program in Support of the Westinghouse Advanced Light Water Reactor, Westinghouse Electric Corporation, Research and Development Centre, 1 3 10 Beulah Road, Pittsburgh, PA 1 5235, USA, 1 988. 25.21 M M Corletti et aI, "AP600 Passive Residual Heat Removal Heat Exchanger Tests" ANS Transactions 1 988. 25.22 M G Ortiz et al "Investigation of the Applicability and Limitations of the ROSA Large-Scale Test Facility for AP600 Safety Assessment" NUREG/CR-5853, EGG2670, December 1992. 25.23 P Coddington et al, ALPHA. ''The Long Term Passive Decay Heat Removal and Aerosol Retention Programme", Technical Committee Meeting on Thermohydraulics of Cooling Systems in Advanced Water-Cooled Reactors Villigen, Switzerland, 2527 May 1 993. 25.24 P F Billig, A J James , E Lumini, "SBWR Passive Core and Containment Cooling Test Program", ANP 1992, Tokyo, Japan, November 1992. 25.25 M D Alagmir, J G M Anderson, A I Yang and B S Shiralker, "TRACG Prediction of Gravity-Driven Cooling System Response in the SBWR/GIST Facility LOCA Tests", ANS Transactions 62, pp 665-668, November 1 990. 25.26 H Nagasaka, K Yamada, M Katoh and S Yokobori, "Heat Removal Tests of Isolation Condenser Applied as a Passive Containment Cooling System," 1 st JSME/ASME Joint International Conference on Nuclear Engineering, Tokyo, Japan, November 47, 199 1 . 25.27 S Yokobori, H Hagasaka, T Tobimatsu, System Response Test of Isolation Condenser Applied as a Passive Containment Cooling System", 1 st JSME/ASME Joint International Conference on Nuclear Engineering, Tokyo, Japan, November 47, 199 1 . 25.28 K M Vierow et ai, "BWR Passive Containment Cooling System b y Condensation Driven Natural Circulation", 1 st JSME/ASME Joint International Conference on Nuclear Engineering, November 1 99 1 . 25.29 R M Kemper et aI, The LOCA Perfonnance o f the AP600 Passive Safety Systems, pp 601 -603. Transactions of the American Nuclear Society (United States). International Meeting on Fifty Years of Controlled Nuclear Chain Reaction: Past, Present, and Future. Chicago, IL (United States). 1 5-20 November 1992. 25.30 J E Fisher, Large Break LOCA Calculations for the AP600 Design, NURETH 5, Salt Lake City, USA, September 1 992. 25.3 1 G Bianconi et al, Small Break Calculation for the AP600 Plant and the SPES-2 Experimental Test Facility, NURETH 5, Salt Lake City, USA, September 1992. 25.32 S M Sloan, Code Assessment Studies of RELAP5 in Support of AP600 Thennal­ Hydraulic Analysis, NURETH 5, Salt Lake City, USA, September. 25.33 R Riemke, RELAP5/MOD3 AP600 Problems, July 6-9, 1 993 RELAP5 International Users Seminar, Boston Massachusetts.

345 INDEX

Absorber rods I I , 1 6, 88 -materials interactions 88 Accident assessments 21, 297, 340 -advanced reactors 340 -LWRs 2 1 , 297 Accident classification 21 Accident management 3 17 -assessments 321 -code requirements 320 -containment 326 -degraded core 324 -ex-vessel melt/water interactions 327 -failure of the reactor vessel 326 -late phase melt progression 325 -melt water interactions 325, 327 -mitigative 323, 327 -preventative 3 1 8, 322 -primary circuit integrity 324 -steam explosions 325 -threat to the containment 326 -total loss of feed water 3 19 -water addition 324 Accident scenarios 2 1 , 284 -BWR 23 -classification 21 -event categories 22 -containment bypass 30 -intact circuit 2 1 , 26, 3 1 -large break LOCA 24, 25 -LOCA 24, 28, 3 1 -loss of cooling 21 -PWR 21 -severe accidents 26 -small break LOCA 24, 26 -Three Mile Island 284 Accident sequence symbols 29 Advanced reactors 33 1 -accident assessments 340 -code requirements 342 -design 33 1 -evolutionary 33 1 -experiments 338 -innovative 335 -passive 333

-phenomena 336 Aerosols 163, 1 93, 2 1 5 -agglomeration 1 64 -bubble collapse 195 -bubble rise 197 -condensation 195, 2 1 8 -containment 2 1 5 -deposition 163, 2 1 8 -disequilibrium 200 -entrainment 1 94 -evaporation 2 1 8 -gas composition 196 -mass transport 1 98 -material entrainment 194 -mechanical production 200 -mechanisms 193 -particle size 196 -production 193, 1 94, 1 95 -scrubbing 219 -size distribution 216 -sources 2 1 9 -vaporisation 1 9 5 Agglomeration 216 ALARP 4 ATWS 301 Blockage 101 -formation 101 -melt progression paths 104 -melt release models 1 03 -melt relocation 104 Boiling water reactors 14, 23 -containment 16, 207 -cooling systems 1 7 -core 16 -design data table 9 -emergency core cooling 17 -large break LOCAs 25 -safety systems 17 -severe accidents 31 -small break LOCAs 26 -steam cycle 16 -vessel 16 Bottom reflood models 1 3 1 -fluidisation 1 32

346 -lumped parameter models 1 3 1 -quenching models 1 3 1 -sophisticated models 132 -status of modelling 1 36 Boundary conditions 1 1 1 -steam explosions 1 1 1 Bubble behaviour 190 -aerosol production 195 -bubble collapse 195 -rise phenomena 1 97 Calculational models 47, 63, 73, 95, 1 07, 127, 14 1 , 1 53, 169, 1 8 1 , 193, 203 , 2 1 5 , 227 -aerosol 193 -cavity 169 -containment aerosol 2 1 5 -containment thennal-hydraulics 203 -core/concrete interaction 1 8 1 -debris coolability 127 -debris interactions 1 4 1 -fission products 1 53 -heat transfer 63 -mechanical 73 -melt progression 95 -steam explosions 107 -thennal-hydraulics 47 -thennophysical 227 Cavity 1 69 -debris coolability 1 74 -debris fonnation 1 69 -debris mixing 169 -debris transport 173 -hydrogen production 170 -phenomena 169 -steam explosions 1 7 1 -uncertainties 176 -uncoolable debris 175 Chemistry 161 , 1 88 -core/concrete 188 -fission products 1 6 1 Choked flow 60 Cladding 67 -ballooning 74 -energy transfer 67 -materials interactions 86 -oxidation heating 70

Cladding behaviour 35, 252 -codes 252 Cladding oxidation heating 70 Component heat transfer 63 -cladding energy transfer 67 -conduction in structures 63 -fuel energy transfer 67 -radiation 70 -under reflood conditions 65 Computer code requirements 342 -advanced reactors 342 Computers 3 Computer codes 20, 249 -cladding 249 -containment 252 -fuel 249 -hydraulics 249 -integrated 253 -mechanistic codes 255 -separate effects 258 -separate phase 250 -severe accident 253 -system 249 Code assessments 265, 32 1 -accident management 321 -code validation 265 Code categories 47 Code validation 265 -core degradation 273 -fission product 277 -intact circuit 266 -LOCA 268 -natural circulation 276 -severe accident 273 -thennal-hydraulics 265 Codes 249 Computer code requirements 249, 320 -accident management 320 Concrete 1 8 1 -ablation 1 87 -chemical interactions 1 88 -core debris interactions 1 8 1 Condensation 195 Conduction 63, 65, 2 1 1 -reflood 65 -structures 63, 2 1 1

347 Constitutive relations 54 -choked flow 60 -flow regime maps 54 -interphase drag 54 -interphase mass transfer 57 -reflood heat transfer 57 -turbulence modelling 58 -wall friction 56 -wall heat transfer 56 Containment 1 3, 16, 30, 43, 203 -aerosol size 216 -aerosol sources 219 -aerosols 2 1 5, 2 1 8 -agglomeration o f particles 2 1 6 -burning models 208 -BWR processes 207 -bypass sequences 30 -codes 252 -compartments 204 -condensation 2 1 8 -conduction 2 1 1 -decay heat 221 -deposition 2 1 8 -energy transfer 209 -engineered safety features 2 1 1 , 2 1 6, 222 -evaporation 2 1 8 -fission products 2 1 5 , 216, 2 1 8, 219 -gas burning 208 -heat conduction 21 1 -integral experiments 43 -mass transfer 209 -material properties 208 -phenomena 43, 204 -radionuclide behaviour 220 -scrubbing 219 -sprays 224 -thennal-hydraulics 204, 205 -transfer rates 222 Containment behaviour 3 12 Containment bypass 309 Control rod 1 1 , 16, 88, 234 -materials interactions 88 -thennophysical properties 234 Control rod cladding 1 1 , 16, 235 -thennophysical properties 235 Control rod eutectic 88, 238

-materials interactions 88 -thennophysical properties 238 Convective transport 1 59 Core concrete interactions 4 1 , 1 8 1 , 240, 327 -accident management 327 -aerosol production 194 -ablation 1 87 -bubble behaviour 1 90 -chemistry 188 -crust 1 86 -energy conservation 1 90 -energy transfer 1 89 -heating 183 -integral experiments 4 1 -mass transfer 189 -melt/concrete interface 1 85 -phenomenology 1 8 1 -pool 183 -surface 1 85 -thennophysical properties 240 Core degradation 273 , 303 -code validation 273 -containment bypass 309 -interfacing LOCA 309 -plant studies 303 -shutdown accidents 309 -small break LOCA 307 -station blackout 303 Core fragmentation 101 Core melt programmes 36, 227 -integral experiments 36 -separate effects data 227 Core meltdown 95 -mechanistic models 97 -parametric models 95 Critical heat flux 1 34 -steam explosions 134 Crust behaviour 1 86 Damage potential 1 1 9 Debris bed 40, 127 -behaviour at vessel failure 149 -behaviour of penetrations 147 -bottom reflood 1 3 1 , 136 -coolability models 127 -critical heat flux 1 34 -experiments 128

348 -integral experiments 40 -interactions 14 1 -interactions with vessel 142 -interactions with water 144 -material release from core 142 -models 13 1 -phenomena 128 -quench 1 34 -top reflood 1 34, 1 37 -vessel failure 149 -vessel response 146 Debris in cavity 169 -coolability 174, 1 75 -formation 169 -mixing 169 -transport 1 73 -uncertainties 1 76 Debris in lower head 14 1 , 29 1 -Three Mile Island 29 1 Decay chains 221 Decay heating 162, 22 1 -fission products 162 Design data 1 5 , 1 8 Design concepts 33 1 -advanced reactors 331 -evolutionary 33 1 -innovative 335 -passive 333 Direct containment heating 42, 176 -integral experiments 42 Disequilibrium two-phase models 49 Drift flux model 49 Energy transfer 1 89, 209 -containment 209 -core/concrete 1 89, 1 90 Engineered safety features 2 1 1 , 222 Entrainment 194 Equilibrium conditions 1 96, 200 Event categories 22 -examples 23 Experiments 33, 120, 338 -advanced reactors 338 -steam explosions 120 Failure of the reactor vessel 326 Field equations 5 1 , 63 -heat transfer 63

-thennal-hydraulics 5 1 Final state-Three Mile Island 288 Fission product transport 3 10 Fission products 39, 1 53, 215, 277 -aerosols 163 -agglomeration 164 -behaviour in fuel 155 -chemistry 16 1 , 162 -code validation 277 -containment 2 1 5 -convective transport 1 59 -debris 165 -decay heating 162 -deposition 163 -equilibrium chemistry 162 -extra-granular processes 1 58 -integral experiments 39 -intra-granular processes 1 57 -non-equilibrium chemistry 1 62 -processes 1 57, 1 58 -release 39, 294 -severe accident 1 54 -transport 39, 1 59 Flow regime maps 54, 55 Fluidisation models 132 Friction 56 Fuel 67 -energy transfer 67 -FP behaviour in fuel 1 55 -heat generation 67 -materials interactions 86 -oxidation 9 1 -thennophysical properties 227 Fuel behaviour 35, 252 -codes 252 -integral experiments 35 Fuel cladding eutectic 86, 237 -materials interactions 86 -thennophysical properties 237 Fuel pin cladding 1 1 , 16, 86, 23 1 -materials interactions 86 -thennophysical properties 23 1 Fuel rod boundary conditions 70 Gap conductance 68 Gas burning models 208 Gas composition 1 96

349 Heat conduction 63 -reflood 65, 66 -structures 63, 2 1 1 Heat generation 67 -core/concrete interaction 1 83 -fuel 67 Heat transfer 63 -boundary conditions 70 -cladding 67 -conduction 63 -fuel 67 -gap 68 -oxidation 70 -melt/concrete 1 85 -pool 1 83 -reflood 65 -surface 185 High pressure melt ejection 42, 176 -integral experiments 42 Homogeneous flow 49 Homogeneous model with slip 49 Hydraulics 249 Hydrogen 83, 170 -Zircaloy oxidation 83 -cavity 1 70 Intact circuit faults 2 1 , 26, 3 1 , 266, 298 -boiling water reactors 23, 3 1 -code validation 266 -plant studies 298 -pressurised water reactors 2 1 , 26 -severe accidents 26 Integral experiments 33 -cladding behaviour 35 -containment 43 -core melt 36 -core/concrete 4 1 -debris bed 40 -direct containment heating 42 -fission product 39 -fuel behaviour 35 -materials 36 -melt/water interaction 4 1 -melt ejection 42 -natural circulation 39 -structural 36 -thennal-hydraulics 33

Integrated codes 253 Interfacing LOCA 309 Interphase drag 54 Interphase mass transfer 57 Large break LOCA 303 Late phase melt progression 325 Light water reactors 2, 9, 14 -boiling water reactors 14 -pressurised water reactors 9 Liquefaction 99 -control rods 99 -fuel rods 100 LOCAs 24, 268, 301 -BWRs 25, 3 1 -large break 25 -small break 26 -severe accidents 3 1 -code validation 268 -PWRs 24, 28 -large break 24, 303 -small break 24 -severe accidents 28 -SGTR 301 Loss of feed water 298 Lower head 29 1 , 292 -debris in TMI -2 29 1 -failure analysis in TMI-2 293 -state of TMI-2 292 Lumped parameter models 1 3 1 , 203 -containment 203 -debris coolability 1 3 1 Major technical areas 3 Mass transfer 57, 189 -core/concrete 189 -interphase 57 Mass transport 198 -containment 209 Material entrainment 194 Material properties 208 Materials behaviour 36, 142 -integral experiments 36 -release from core 142 Materials interactions 83 -absorber rod 88 -burnable poison 90 -control rod 88

350 -fuel and cladding 86 -fuel oxidation 9 1 -Inconel grid 90 -kinetics 85 -oxidation 83, 90, 9 1 -steel oxidation 90 -structures oxidation 90 -temperature regimes 89 -Zircaloy interactions 90 -Zircaloy oxidation 83 -Zircaloy/uranium dioxide 86 -Zircaloy/zirconium dioxide 87 Mechanical aerosol production 200 Mechanical phenomena 73, 1 16 -modelling 73 -steam explosions 1 16 Mechanical models 73, 75, 76 -clad ballooning 74 -criteria 78 -design characteristics 77 -structures 76 Mechanistic codes 255 Melt progression 95, 290 -phenomena 95 -Three Mile Island 290 Melt progression models 95 -blockage 101 -blockage melt release 103 -blockage melt relocation 1 04 -control rods 99 -core fragmentation 101 -fuel rods 100 -liquefaction 99 -mechanistic 97 -melt progression paths 104 -molten pools 103 -parametric 95 Melt/water interactions 4 1 , 325 -ex-vessel 327 -integral experiments 4 1 -reactor vessel 325 Mitigative accident management 323 Mixing 58, 1 12 -molten debris 1 12 -turbulent 58 Modelling in TMI-2

Modelling requirements 5, 21 -capabilities 5 -scenarios 2 1 Molten pools 103 Natural circulation 39, 276 -code validation 276 -integral experiments 39 Non-condensable gases, 83, 170, 243 Nuclear power 1 Nuclear safety 4, 5 -modelling requirements 5 -research objectives 4 Phases 49, 109 -steam explosion 109 -two-phase flows 49 Physical phenomena 47, 63, 73, 95, 107, 127, 141, 1 53, 169, 1 8 1 , 193, 203, 215, 227, 336 -advanced plants 336 -aerosol 193 -cavity 169 -containment aerosol 2 1 5 -containment thennal-hydraulics 203 -core/concrete interaction 1 8 1 -debris coolability 127 -debris interactions 14 1 -fission products 1 53 -heat transfer 63 -mechanical 73 -melt progression 95 -steam explosions 107 -thennal-hydraulics 47 -thennophysical 227 Plant studies 297 -ATWS 301 -containment behaviour 3 1 2 -containment bypass 309 -core degradation 303 -fission product 3 10 -intact circuit 298 -interfacing LOCA 309 -large break LOCA 303 -LOCA 301 , 303, 307 -loss of feed water 298 -reactivity transients 300 -severe accidents 303, 3 1 0, 3 12

351 -SGTR 30 1 -shutdown accidents 309 -small break LOCA 307 -station blackout 303, 3 1 0, 3 1 2 -steam line break 299 -Three Mile Island 287 Pressurised water reactors 9, 2 1 , 24 -containment 1 3 -coolant system 9 -core 1 1 -design data 1 5 -large break: LOCAs 24 -pressure vessel 1 1 -safety systems 1 3 -severe accidents 26 -small break LOCAs 24, 25 -steam generators 9 -vessel 1 1 Preventative accident management 3 1 8, 322 -total loss of feed water 3 19 Primary circuit integrity 324 Prominent safety studies 2 Quenching models 1 3 1 , 1 3 5 Radiation models 70, 7 1 Radionuclide behaviour 220 Reactivity Transients 300 Reactor 9 -containment 13, 16 -BWR 16 -PWR 13 -coolant system (PWR) 9 -core 1 1 , 16 -BWR 16 -PWR 1 1 -pressure vessel 1 1 , 1 6 -BWR 16 -PWR 1 1 -safety systems 13, 17 -BWR 1 7 -PWR 1 3 -steam cycle (BWR) 1 6 Reflood heat transfer 57 Safety studies 2 Safety systems 1 3, 17 Separate effects codes 258 Severe accident codes 253

-integrated 253 -mechanistic 255 -separate effects 258 -system 255 Severe accidents 26, 1 55, 273 -boiling water reactors 3 1 -intact circuit 3 1 -LOCAs 3 1 -code validation 273 -core degradation 273 -fission product release 1 55 -fission product transport 277 -natural circulation 276 -pressurised water reactors 26 -containment bypass 30 -intact circuit 26 -LOCAs 28 -PWR accident symbols 29 SGTR 301 Shutdown accidents 309 Small break LOCA 307 Sprays 224 Station blackout 303, 3 10, 3 1 2 S team cycle 1 6 Steam Explosions 1 07, 120, 325 -boundary conditions 1 1 1 -calculations 1 2 1 -cavity 1 7 1 -damage potential 1 19 -expansion 1 10 -experiments 120 -initial conditions 109 -mechanical energy 1 16 -mixing 109, 1 1 2 -phases 1 09 -propagation 1 1 0 -triggering 109, 1 1 5 Steam line break 299 Steel oxidation 90 Structural behaviour 36, 73 -integral experiments 36 -modelling 73 Structural temperatures 63, 287 -component heat transfer 63 -Three Mile Island 287 Structures 76, 239

352 -mechanical modelling 76 -thennophysical properties 239 System codes 249, 255 System components 1 8 1 Thennal-hydraulics 33, 47, 203, 252, 307 -cladding codes 252 -code categories 47 -code validation 265 -constitutive relations 54 -containment 203 -containment codes 252 -containment bypass 309 -equations 50 -fuel codes 252 -intact circuit 266 -integral experiments 33 -interfacing LOCA 309 -interphase drag 54 -interphase mass transfer 57 -LOCA 268 -maps 54 -model classification 49 -physical phenomena 48, 204, 205 -reflood heat transfer 57 -shutdown accidents 309 -small break LOCA 307 -station blackout 303 -system codes 249 -Three Mile Island 285 -turbulence 58 -two-phase flow 49 -equations 50 -wall heat transfer 56 Thennophysical properties 70, 227 -cladding 23 1 , 235 -control rod 234 -control rod eutectic 238 -control rod cladding 235 -core concrete materials 240 -fuel 227 -fuel cladding eutectic 237 -fuel pin cladding 23 1 -gases 243 -structural materials 239 -water/steam 24 1 Threat to the containment 326

Three Mile Island 283 -accident scenario 284 -conclusions 294 -debris 29 1 -failure analysis 293 -final state 288 -fission product release294 -fission product transport 294 -lower head 29 1 , 292, 293 -lower head failure 293 -lower head debris 29 1 , 292 -melt progression 290 -modelling 294 -structural temperatures 287 -system analysis 285 -thennal-hydraulics 285 -upper vessel temperatures 287 Top reflood models 1 34 -critical heat flux 1 34 -quench front limitation 1 35 -status of modelling 1 37 -upper bed quenching 1 34 Transfer rates 222 Transport of FPs in TMI-2 294 Triggering 1 1 5 -steam explosions 1 1 5 Turbulence modelling 58, 59 Two-phase flow models 49, 50 -disequilibrium 49 -drift flux 49 -equations 50, 5 1 , 53 -field equations 5 1 -homogeneous 49 -homogeneous with slip 49 Uncertainties 176 -debris behaviour 1 76 Upper bed quenching 1 34 Upper vessel temperatures 287 Vaporisation 1 95 Vessel 146 -behaviour of penetrations 147 -debris interactions 145 -failure potential 146 -penetrations 147 -response to temperature 146 -thennal-hydraulic regions 12

353 Vessel failure 149, 169 -cavity phenomena 169 -debris behaviour 149 Vessel internals 1 1 , 16, 77, 142 -BWR 16 -debris interactions 142 -PWR 1 1 , 77 Wall friction 56 Wall heat transfer 56 Water 144, 324 -addition to a degraded core 324 -debris interactions 144 Water/steam properties 24 1 Zircaloy reactions 83 -absorber rods 88 -burnable poison 90 -Inconel 90 -kinetics 85 -oxidation 83 -kinetics 85 -phenomena 83 -temperature regimes 89 -uranium dioxide 86 -zirconium dioxide 87