An Experimental Study on Melt Fragmentation, Oxidation and Steam Explosion During Fuel Coolant Interactions

An Experimental Study on Melt Fragmentation, Oxidation and Steam Explosion during Fuel Coolant Interactions

LOUIS MANICKAM

Doctoral thesis KTH Royal Institute of Technology Engineering Sciences Department of Physics Division of Nuclear Power Safety Stockholm, Sweden, 2018

AlbaNova University Center Roslagstullsbacken 21 TRITA-SCI-FOU 2018:54 10691 Stockholm ISBN 978-91-7873-057-5 Sweden

Akademisk avhandling som med tillstånd av Kungliga Tekniska Högskolan framlägges till offentlig granskning för avläggande av teknisk doktorexamen i fysik den December 20, 2018 i Sal FA31, AlbaNova universitetscentrum, Stockholm.

Tryck: Universitetsservice US AB I

Abstract

Nordic type boiling water reactors (BWRs) adopt reactor cavity flooding as a severe accident mitigation strategy (SAMS) to achieve core melt fragmentation and long-term cooling of decay heat generating core debris. The qualification of this SAMS needs to address two main severe accident issues: debris bed coolability and steam explosion.

Since the coolability of a debris bed is determined by the bed’s properties including debris particle’s size distribution and morphology as well as the bed’s configuration and inhomogeneity, it is important to investigate the mechanisms of melt jet breakup and resulting fragmentation in water which affect debris bed’s properties. Hence, the first part of this thesis is concerned with characterization of melt jet breakup and resulting debris particles. A series of jet breakup experiments have been conducted in small scale with simulant binary oxide melt mixtures of WO3-Bi2O3, WO3-ZrO2 and Wood's metal. The experiments reveal significant influence of melt superheat, water subcooling, melt jet diameter and material properties on debris size and morphology. Specifically, transition in debris size and morphology is found to occur at a specific water subcooling range. The difference in debris properties at varied melt release conditions is attributed to the competition between liquid melt hydrodynamic fragmentation and thermomechanical fracture of quenched particles.

The second part of this thesis work is dedicated to provide a new understanding of steam explosion (SE) with the support of small-scale experiments at the level of droplets. Self- and externally-triggered SE experiments are conducted with simulant binary oxide melt mixtures in the temperature range of 1100 to 1500°C. The dynamics of steam explosion process is recorded using a sophisticated simultaneous visualization system of videography and X-ray radiography. Further, the influence of melt composition on steam explosion is summoned. The results reveal that a droplet of eutectic composition is more explosive than a droplet of non-eutectic composition since latter may form a mushy zone which thereby limits the amount of melt actively participating in a steam explosion. To reduce the temperature difference between simulant melt and corium, investigation was extended to perform high temperature

(˃2000°C) melt experiments. For this purpose, steam explosion of a molten Al2O3 droplet was investigated, and the experimental results confirmed that Al2O3 melt can undergo spontaneously triggered steam explosion at a high melt superheat and high subcooling. Within the context the effects of melt superheat and water subcooling were obtained.

The third part of this thesis is concerned with the oxidation of metallic melt representing unmixable metallic liquid of molten corium, which interactions with water can be spatially and chronologically separated from the oxidic corium FCI. The objective of the study is to provide new insights into the characteristics of oxidation of Zr droplet falling in a water pool through a series of small-scale experiments. The dynamics of droplet and bubbles were recorded by high-speed cameras, and the spatial distributions of the elements in the quenched droplet (debris) were acquired by Energy- Dispersive X-Ray Spectroscopy (EDS). The results have shown noticeable influence of generated hydrogen and oxidation heat on droplet behavior and cooling rate. Water subcooling had significant influence on oxidation kinetics, and the oxygen content of the solidified particle increased with decreasing subcooling. Incomplete oxidation of Zr happened before melt crystallization and cooling down in all experiments.

Keywords: Severe accident, fuel-coolant interactions, melt jet breakup, debris properties, steam explosion, oxidation, high-speed visualization.

III

Sammanfattning

Nordiska typer av kokvattenreaktorer (BWR) antar översvämning av den primära reaktorkaviteten som en begränsningsstrategi för hantering av svåra haverier (SAMS) med syfte att fragmentera härdsmältan och uppnå en långsiktig resteffektnedkylning. Då kylbarheten av en partikelbädd beror på avfallspartiklarnas storlek och morfologi likväl dess geometri och inhomogenitet är det viktigt att undersöka smältastrålbrytningens mekanismer samt dess efterföljande spridning av droppar då detta påverkar bäddens egenskaper. Således behandlar den första delen av denna avhandling karaktären på smältastrålbrytningen och härssmältpartiklarna som resulteras av detta. En experimentserie med simulant binär oxid smältblandning bestående av WO3-Bi2O3, WO3-ZrO2 och Wood's metal har visat att faktorer som överhettad smälta, underkylt vatten, smältastrålens diameter och materialegenskaper har signifikant påverkan på partikelstorlek och morfologi. Mer specifikt har det fastställts att en förändring av partikelegenskaper och morfologi sker med specifika variationer av underkylningstemperatur. Variationen i partikelegenskaper vid varierande tillstånd av smältutsläpp från reaktortanken kan tillskrivas mängden hydrodynamisk fragmentering av smältan kontra utbrytning från stelnade droppar på grund av temperatur.

Den andra delen av denna avhandling ägnas åt att, genom småskaliga experiment, skapa ny kunskap kring ångexplosioner samt att undersöka smältasammansättningens påverkan på ångexplosioner. Experimenten är utförda med simulant binäroxidsmältblandning på mellan 1100 - 1500 °C. Dynamiken i ångexplosionen är registrerad genom ett sofistikerat system som samtidigt filmar och röntgar processen. Resultaten visar att en droppe med eutektisk sammansättning är mer explosiv än en droppe med en icke eutektisk sammansättning då den senare kan skapa en grötig zon som begränsar mängden smälta som är en aktiv del av ångexplosionen. För att reducera temperaturglappet mellan simulantsmälta och corium utökades undersökningen till att innefatta smältaexperiment på hög temperatur (>2000°C). För att göra detta undersöktes en droppe av smält Al2O3. Resultatet av detta experiment visade att smälta av Al2O3 kan genomgå en spontant utlöst ångexplosion vid en hög överhettad smälta och hög underkylning. Effekten av smältöverhettning och vattenunderkylning har därmed mätts inom givna ramar.

Den tredje delen av denna avhandling behandlar oxidering av metallkomponenten (Zr) i corium, då oxideringen kan inverka på ångexplosioner och partikelegenskaper. Målet med denna studie är att, genom en serie av småskaliga experiment, finna ny kunskap om egenskaperna hos oxiderade Zr droppar som fallit ner i en vattenbassäng. Dropparnas dynamik och bubblorna spelades in med en höghastighetskamera, och

IV information om den spatiala fördelningen av beståndsdelarna i den nerkylda droppen(partiklarna) erhölls genom Energy- Dispersive X-Ray Spectroscopy (EDS). Resultaten har visat att hela droppens yta blev suboxiderad med ökande grad av oxidering från mitten av droppen till det yttre lagret. Underkylt vatten hade en signifikant påverkan på oxideringen och partikelns syremängd ökade med minskad underkylning.

Sökord: svåra haverier, smälta stråle fragmentering, ångexplosionen, slaggformationens kylbarhet, oxidering, hög hastighet visualisering. V

List of Publications

Journal papers included in this thesis: I. L. Manickam, P. Kudinov, W. Ma, S. Bechta, D. Gishchenko, On the influence of water subcooling and melt jet parameters on debris formation, Nuclear Engineering and Design, 309: 265-276, 2016. II. L. Manickam, S. Bechta, W. Ma, On the fragmentation characteristics of melt jets quenched in water, International Journal of Multiphase Flow, 91: 262-275, 2017. III. L. Manickam, G. Qiang, W. Ma, S. Bechta, An experimental study on the intense heat transfer and phase change during melt and water interactions, Experimental heat transfer, DOI: 10.1080/08916152.2018.1505786, 2018. IV. J.A. Zambaux, L. Manickam, R. Meignen, W. Ma, S. Bechta, S. Picchi, Study on thermal fragmentation characteristics of a superheated alumina droplet, Annals of Nuclear Energy, 119:352-361, 2018.

Respondent’s contribution to the papers included in the thesis: The respondent (Louis Manickam) was involved in elaboration of ideas, approaches and methodologies in all the above-mentioned papers. The respondent as the first author’s contribution includes conceptual design, performing experiments, post-test analysis and writing of papers (Paper-I, II and III). As a co-author (Paper-IV), the respondent contribution includes conceptual design, performing experiments, post- test analysis in collaboration with the first author, and helped in writing the paper.

Other papers produced during PhD studies: I. S. Thakre, L. Manickam, W. Ma, A numerical simulation of jet breakup in melt coolant interactions, Annals of Nuclear Energy, 80: 467-475, 2015. II. R.C. Hansson, T.N. Dinh, L. Manickam, A study of the effect of binary oxide materials in a single droplet vapor explosion, Nuclear Engineering and Design, 264: 168-175, 2013. III. G. Qiang, L. Manickam, W. Ma, S. Bechta, Investigation on quenching of a high-temperature spherical particle in sea water. Proc. of NUTHOS-12, Qingdao, China, October 14-18, 2018. IV. J.A. Zambaux, L. Manickam, R. Meignen, W. Ma, S. Bechta, S. Picchi, Study on thermal fragmentation characteristics of a superheated alumina droplet, Proc. of ERMSAR-17, Warsaw, Poland, May 16-18, 2017. V. L. Manickam, S. Thakre, W. Ma, An experimental study on void generation around a hot metal particle quenched into water pool, Proc. of NURETH-16, Chicago, U.S.A, Aug 30 –Sep4, 2015. VI. L. Manickam, S. Thakre, W. Ma, S. Bechta, Simultaneous visual acquisition of melt jet breakup in water by high speed videography and radiography, Proc. of NUTHOS-10, Okinawa, Japan, December 14-18, 2014.

VII. S. Thakre, L. Manickam, W. Ma, “A numerical simulation of jet breakup in melt-coolant interactions”, Proc. of NUTHOS-10, Okinawa, Japan, December 14-18, 2014. VIII. L. Manickam, P. Kudinov, S. Bechta, The influence melt parameters and water subcooling on debris formation, Proc. of NURETH-15, Pisa, Italy, May 12 -17, 2013.

VII

Acknowledgements

First and foremost I would like to express my deep gratitude to my supervisors, Prof. Weimin Ma, thank you for the constant motivation, support, teaching, advice, patience and guiding me professionally throughout this entire period at nuclear power safety department. Without you, this would not have been possible.

Prof. Sevostian Bechta thank you for giving this opportunity. Thank you for all the various ideas, insights, suggestions and teaching only with which such an experimental work was possible.

I am sincerely thankful to late Prof. Sehgal. You have motivated, supported every time and always wanted me to do experiments with prototypic material. I hope someday, I will be able to do it. I have learnt a lot from you. Thank you.

Thank you Dr. Hansson for all the initial training with MISTEE single droplets during the early stages. I really had very tough times with these small droplets especially at high temperatures. Thank you Dr. Komlev for the support with SEM, discussions and insights into material science. Thank you Dr. Alexander for all the support, help, discussions and advice with the experiments. Thank you. Dr. Aram for the support with the experiments. Thank you Dr. Pavel for the advice and technical suggestions on small-scale fragmentation experiments. Thank you Per Skold, for all the technical help and discussions about the experiments. Thank you Dr. Dmitry for all the suggestions.

Thank you my colleagues/friends Dr. Sachin, Huimin, Dr.Youjia, Qiang, Per, Dr. Li, Dr.Ravi for being supportive as well as being very friendly otherwise. Thank you Walter for all the valuable comments on the thesis. Thank you Sean for the help and support with administrative work and also for being very friendly. I would also like to thank the immense support of all my colleagues at NPS.

Thank you Tucklex, every single thing means a lot! Thank you Mike for showing the way to Sweden. This time in Sweden has taught me several things in life. Thank you Jonitta for being kind during very hard times with my health. Thank you Arul, Lal, Amal, Murray for all the splendid time and happy moments. Thank you Amma, Appa, Arun and all cousins for the unconditional love.

I would like to thank the research grants and support from Swedish Nuclear Radiation Protection Authority (SSM), Swedish Power Companies, Nordic Nuclear Safety Program (NKS), and Swiss Federal Nuclear Safety Inspectorate (ENSI) under the Accident Phenomena of Risk Importance (APRI) project and Severe Accident Facilities for European Safety Targets (SAFEST) project.

Table of Contents

Abstract ...... I Sammanfattning...... III List of Publications ...... V Acknowledgements ...... VII 1. Background ...... 9 1.1. Fuel-coolant interactions ...... 11 1.2. Objectives and tasks ...... 13 2. Debris formation ...... 15 2.1. Introduction ...... 15 2.2. Experimental methods and conditions ...... 16 2.2.1. MISTEE-JET facility ...... 16 2.2.2. JEBRA facility ...... 17 2.2.3. Melt simulants and test conditions ...... 18 2.3. Effect of experimental conditions on jet breakup ...... 20 2.4. Effect of experimental conditions on debris properties ...... 23 3. Steam explosion ...... 34 3.1. Introduction ...... 34 3.2. MISTEE-LT study of steam explosion ...... 40 3.2.1. Corium simulant materials ...... 42 3.2.2. Observation of steam explosion ...... 43 3.2.3. Effect of melt composition on steam explosion ...... 48 3.3. MISTEE-HT study of steam explosion ...... 52 3.3.1. MISTEE-HT facility ...... 52 3.3.2. Steam explosion experiments with alumina melt ...... 56 4. Oxidation of metallic melt ...... 64 4.1. Introduction ...... 64 4.2. MISTEE-OX droplet oxidation experiment ...... 66 4.3. Experimental results ...... 68 4.3.1. Bubble analysis ...... 68 4.3.2. SEM/EDS analysis ...... 72 4.4. Discussions...... 79

5. Summary ...... 87 6. References ...... 91

Appendix A (Paper I)

Appendix B (Paper II)

Appendix C (Paper III)

Appendix D (Paper IV)

List of Figures

Figure 1.1: Defence in depth approach: Barriers and levels of protection…………… 10 Figure 1.2: Illustrative sketch depicting a severe accident scenario in Nordic type BWRs…………………………………………………………………………………………………………. 12 Figure 2.1 Schematic of the MISTEE-Jet facility…...... 17 Figure 2.2: Schematic of JEBRA facility………………………………………………………….. 18 Figure 2.3: Snapshots depicting fragmentation dynamics of molten Wood's metal melt in water pool under non-boiling conditions……………………………………………… 21 Figure 2.4: Jet breakup length as a function of nozzle diameter...... 22 Figure 2.5: Jet breakup length as a function of jet penetration velocity (a) and comparison with correlations for jet breakup length (b)……………………………………. 23 Figure 2.6: Fragmentation of molten eutectic WO3-Bi2O3 (∆Tsuperheat ≈ 130K) under varied water subcooling...... 25 Figure 2.7: Fragmentation of molten eutectic WO3-ZrO2 (∆Tsuperheat ≈ 160K) under varied water subcooling…...... 26 Figure 2.8: Debris size distribution for WO3-Bi2O3 (a) and WO3-ZrO2 (b) as a function of water subcooling…………………………………………………………………………. 27 Figure 2.9: Mass averaged median particle diameter as a function of water subcooling…...... 27 Figure 2.10: Debris particle morphology as a function of water subcooling …...... 28 Figure 2.11: Mass fraction of fractured particles as a function of water subcooling...... 28 Figure 2.12: Comparison of mass averaged median particle diameter and predictions by fragmentation theories...... 30 Figure 2.13: Mass averaged median particle diameter as a function of nozzle diameter for simulant and prototypic binary oxide materials...... 30 Figure 2.14: Mass averaged median particle diameter as a function of initial jet diameter (a) and dimensionless mass averaged median particale diameter as a function of jet penetration velocity (b)...... 31 Figure 3.1: Illustration of the distinct phases of a steam explosion sequence during a severe accident. Melt material: WO3-Bi2O3, Tmelt: 135K, Tcoolant: 82K...... 35 Figure 3.2: Thermal fragmentation models proposed in (Kim & Corradini, 1988)(above) and (Ciccarelli & Frost, 1994)(below)...... 37 Figure 3.3: MISTEE phenomenology of melt fine fragmentation. Schematic interpretation of fine fragmentation (above) and Matched images acquired by videography and X-ray radiography...... 38 Figure 3.4: Schematic diagram of the MISTEE-LT test facility (above) and a snapshot of the facility (below)...... 41 Figure 3.5: Molten WO3-CaO droplet: a) displacement below water pool surface and velocity; and b) aspect ratio during quench in water pool. (Mmelt ≈1.3g ∆Tsup ≈ 115K and ∆Tsub ≈ 79K)...... 43

Figure 3.6: Simultaneous visuals and X-ray radiography images recorded for a single droplet of WO3-CaO undergoing explosion in water (Mmelt ≈1.2g ∆Tsup ≈ 116K and ∆Tsub ≈ 77K)...... 44 Figure 3.7: Image processing procedure for quantification of droplet dynamics... 46 Figure 3.8: Melt and vapour interrelated equivalent diameter history during steam explosion (a) and work done by the expanding vapour bubble (b)...... 46 Figure 3.9: Effect of melt deformation ratio (MD*) on the conversion ratio for experiments with binary oxide mixture of WO3-CaO melt at high water subcooling conditions (Mmelt ≈1 to 1.5g ∆Tsup ≈ 100 to 200K and ∆Tsub ≈ 75 to 85K)...... 47 Figure 3.10: Cumulative size distribution of solidified fragments acquired from thermal fragmentation experiments with WO3-CaO melt compared to fragment size shown in (Nelson L. , et al., 2001) and (Abe, et al., , 2002)...... 48 Figure 3.11: Time sequence photography (above) and X-ray radiography (below) of a single droplet binary oxide mixture of WO3-CaO melt at eutectic composition (75:25 mol%, Mmelt ≈1g ∆Tsup ≈ 100 K and ∆Tsub ≈ 80K)...... 50 Figure 3.12: Time sequence photography (above) and X-ray radiography (below) of a single droplet binary oxide mixture of WO3-CaO melt at non-eutectic composition (67:33 mol%, Mmelt ≈1g ∆Tsup ≈ 100 K and ∆Tsub ≈ 80K)...... 50 Figure 3.13: Difference in conversion ratio as a function of melt deformation ratio with binary oxide mixture of WO3-CaO melt at Eutectic and non-Eutectic composition at varied melt superheat (∆Tsup ≈ 100 and 200K)...... 51 Figure 3.14: Binary oxide mixture of WO3-CaO melt at eutectic and non-eutectic composition undergoing non-equilibrium solidification...... 52 Figure 3.15: Schematic of the MISTEE-HT test facility (above) and a snapshot of the facility (below)...... 55 Figure 3.16: Top-view of melt droplet aerodynamically plugged inside crucible by constant purge of inert gas in MISTEE-HT furnace...... 56 Figure 3.17: Energy conversion ratio as a function of water subcooling for corium and alumina melts from KROTOS experiments (Ursic, e al., 2012)...... 57 Figure 3.18: Time sequence video images recorded for a single droplet of Al2O3 undergoing explosion in water (Mmelt ≈0.8g ∆Tsup ≈ 163K and ∆Tsub ≈ 82K).. 59 Figure 3.19: Effect of melt superheat on the conversion ratio for experiments with Al2O3 melt at high and moderate water subcooling conditions (Mmelt ≈1g and ∆Tsub ≈ 55 to 85K)...... 59 Figure 3.20: Spontaneous steam explosion observed for the self-triggering case (Mmelt ≈1g, ∆Tsuperheat ≈ 215K and ∆Tsub ≈ 81K)...... 60 Figure 3.21: Time sequence snapshots depicting vapour film evolution under low subcooling conditions (Mmelt ≈0.9g ∆Tsup ≈ 116K and ∆Tsub ≈ 14K)...... 61 Figure 3.22: Time resolved equivalent vapour evolution...... 62 Figure 4.1: Schematic of the MISTEE-HT test facility with modifications for single droplet oxidation studies...... 67 5

Figure 4.2: Time sequence raw snapshots depicting interaction between single droplet of zirconium melt and water at ∆Tsub ≈85K...... 69 Figure 4.3: Time sequence raw snapshots depicting interaction between single droplet of zirconium melt and water at ∆Tsub ≈45K...... 69 Figure 4.4: Temporal velocity profile of the melt droplet in water at varied degree of water subcooling...... 70 Figure 4.5: Image processing methodology for the estimation of the temporal distribution of hydrogen (left). Labels 1 and 2 depict typical individual hydrogen bubbles (right)...... 71 Figure4.6: Hydrogen generation history during Zr melt and water interactions... 71 Figure 4.7: Evolution of oxygen content in melt drop estimated from hydrogen generation data...... 72 Figure 4.8: SEM images showing microstructure of solidified Zr-droplet cross section for all cases studied under high subcooling conditions…...... 75 Figure 4.9: Oxygen concentration profile from the surface to the bulk of the quenched drops after the tests at high subcooling...... 76 Figure 4.10: SEM images showing microstructure of solidified Zr-droplet cross section for all cases studied under medium subcooling conditions...... 78 Figure 4.11: Zirconium-Oxygen phase diagram [Massalski, 1990]...... 80 Figure 4.12: Phase composition analysis of HSUB-1 and MSUB-1 samples...... 81 Figure 4.13: O/Zr atomic ratio as a function of particle diameter...... 83 Figure 4.14: Time sequence raw snapshots depicting interaction between alumina melt and water (∆Tsup ≈1ooK, ∆Tsub ≈15K)…...... 84

List of Tables

Table 2.1: Thermophysical properties of simulant melt materials ...... 19 Table 2.2: Test conditions for study of debris formation behaviour ...... 19 Table 3.1: Thermal fragmentation experimental studies ...... 39 Table 3.2: Thermophysical properties of simulant melt materials selected for SE studies ...... 42 Table 3.3: Test conditions for study of steam explosion with binary oxide simulant material ...... 42 Table 3.4: Test conditions for study of binary oxide composition effects ...... 49 Table 3.5: Methods used for high temperature melt droplet generation and examples of their application...... 53 Table 3.6: Parametric effects observed in KROTOS tests ...... 57 Table 3.7: Test conditions for MISTEE-HT Al2O3 tests...... 58 Table 4.1: Experimental works on the effect of chemical augmentation on FCI ..... 65 Table 4.2: Conditions for MISTEE-OX droplet tests...... 68

Nomenclature

Cp specific heat, J/(kg·K) D diameter, m Lbrk breakup length, m Eo entrainment coefficient fps frames per second g gravitational acceleration, m2/s J fluid superficial velocity, m/s k thermal conductivity, W/(m·K) t time, s T temperature, °C u velocity, m/s p pressure, MPa M mass, kg 2 Wea Weber number (푊푒푎 = 휌푐푉 퐷⁄휎)

Greek letters λ wavelength, m ρ density, kg/m3 σ surface tension, N/m µ dynamic viscosity, Pa·s

Subscripts p particle c coolant d droplet j jet v vapor liq liquidus sol solidus e effective i internal

List of Acronyms

BWR Boiling water reactor

CRGT Control rod guide tube

DAS Data acquisition system

EDS Energy dispersive X-ray spectroscopy

FCI Fuel-coolant interactions

IAEA Internation atomic energy agency

IGT Instrumentation guide tube

JEBRA Jet breakup

KTH Kungliga Tekniska Högskolan

LWR Light water reactors

MCCI Molten corium-concrete interactions

MISTEE Micro interactions in steam explosion energetics

NCDC Natural convection driven coolability

PWR Pressurized water reactor

RPV Reactor pressure vessel

SAM Severe accident management

SEM Scanning electron microscopy

WDS Wavelength dispersive X-ray spectroscopy

1. Background

Nuclear power is a dominant source of energy, which contributes to approximately 14% of the world’s electricity production. Concerns over greenhouse gas emissions from using fossil fuels as well as diminishing resources of fossil fuels have made nuclear energy the most attractive option to meet a significant portion of the worldwide energy demands. Approximately, 448 nuclear reactors are currently deployed for electricity generation and approximately 57 units are under construction worldwide [IAEA PRIS, 2018]. The vast majority of the units currently deployed are light water reactor designs among which are Pressurized Water Reactors (PWRs) and Boiling Water Reactors (BWRs) that contribute to about 87% of the total worldwide electricity production by nuclear power plants. PWRs are the most common type of reactors that accounts to approximately 70% followed by the BWRs that account to approximately 17% used worldwide. In Sweden, eight nuclear power plants are currently under operation out of which half of the reactors are BWRs and another half of the reactors are PWRs. The eight nuclear power plants account to approximately 40% of the total electricity production in Sweden [World nuclear association, 2018].

In the past decade, many countries have shown specific interest to exploit further the use of nuclear power by power upgrades, life-extension and new-builds. The environmentally friendly resource of nuclear energy has some dissenting aspects including disposal of radioactive wastes and risk of accidents with radiological consequences. More specifically, in the aftermath of the 2011 Fukushima-Daichi accidents in Japan, public skepticism has certainly reduced the anticipated extent of nuclear power expansion to a slower pace. However, safe and secure operation of the existing power plants is currently the top priority. The field of nuclear safety aims to mitigate consequences related to undue radiation hazards by achieving proper operating conditions; establishing, maintaining and assessing safety measures; and assessment of risks associated with an unintended accident.

From the early stages of nuclear power-plant development, safety goals are established. Nuclear power plants are designed in a way that i) nuclear reaction is controlled; ii) cooling systems are controlled to protect nuclear fuel melting; and iii) storage of spent nuclear fuel is controlled. Further, steel-reinforced concrete containment is erected to contain radioactive material release during an accident. Development of safety systems and accident management strategies have since matured with redundant safety features compared to the initial stages. However, assessment of nuclear power plant vulnerability to severe nuclear accidents as the one in Fukushima-Daichi and the effectiveness of safety systems adopted to mitigate nuclear accidents is an essential part of nuclear safety.

The concept of defense in depth approach is used in nuclear safety. The concept enables a sequential approach at five levels to prevent and mitigate the consequences of an accident at various stages (Figure 1.1). In general, all nuclear power plants in operations today are equipped with redundant safety features to control design- based failures that are less probable but can occur. As a very unlikely event, several equipment may fail altogether with design-based accident sequence implying the fourth level of defense in depth approach. Such a severe accident scenario is a highly improbable event; however, the outcomes can be catastrophic. Typically, a severe accident may involve severe core degradation, fission product release to the reactor, containment or environment. The containment is the final barrier of defense. To mitigate containment failure, several severe accident management strategies (SAMs) are adopted in nuclear power plants [Sehgal, 2006]. SAM is an essential component of the defense in depth approach. For example, according to the SAMs adopted for Nordic type BWR installations, during the fourth level of defense in depth approach, the ex-vessel cavity in the Nordic type BWRs are flooded with subcooled water to protect the cables penetrating the containment base-mat from molten corium interactions and early containment over-pressurization by generation of non-condensable gases by molten corium-concrete interactions (MCCI). However, in some LWR designs, the ex-vessel cavity is left dry on purpose. Saying so, the ex-vessel cavity can get flooded by activation of containment sprays and leakage of the primary cooling system [Sehgal, 2006].

Figure 1.1: Defence-in-depth approach: Barriers and levels of protection.

The flooded cavity in a Nordic type BWR (a deep pool of subcooled water pool typically 7 to 12m deep) is expected to achieve fragmentation and quenching of the pouring melt into particulate debris that can pile-up in the bottom of the cavity and cooled by natural water circulation through open porosity (Figure 1.2). A severe accident can be considered as mitigated only when a coolable state of corium is achieved in a long term without significant release of radioactive fission products into the environment. Generation of non-condensable gases and a long-term natural 11 cooling cycle is established when temperature of the melt drops below 1000°C. However, long-term coolability of the core melt cannot be assured as a dry out may occur and the debris can re-melt leading to MCCI [Sehgal, 2006]. Hence, a qualitative and quantitative assessment of the adopted severe accident management strategy is crucial for nuclear safety. Further, the interaction between core melt and water may also involve a strong steam explosion that can result in a early containment failure. It is thus also inevitable to address the concern of steam explosion loads in nuclear safety analysis.

1.1. Fuel-coolant interactions

Interactions between reactor core melt and a liquid representing reactor coolant is generally termed as a “Fuel-Coolant Interaction (FCI)”. FCI is an issue which is not only of interest to the nuclear industry but also spans to various other industries such as liquefied natural gas industries, foundries, chemical industries, metallurgical industries, paper production industries etc. (See, for example [Epstein, 2000] [Bertram, et al., 2007] [Ekenes, 2007] [Epstein, 2009]). The issue of FCI is of paramount importance considering the catastrophic consequences of an explosive interaction (steam explosion). However, not all the interactions will result in a steam explosion. Steam explosion is expected to occur only under some specific conditions of melt and water that is still not well understood.

The FCI phenomena hold specific importance to nuclear safety analysis where the stakes are high. Typically, during the late stages of a severe nuclear accident, when the core has severely degraded, a large mass of molten corium (which in general case represents a mixture of molten fuel ((U, Zr) O2-x) and structure materials (steel and Zr)) may relocate into the lower plenum of the RPV and interact with the residual water there. The relocated material can form particulate debris in the lower plenum and even involve a steam explosion (In-vessel FCI). However, the failure of RPV (alpha-mode) due to in-vessel steam explosion is concluded to be less probable with little significance to the overall risk [NUREG, 1996]. On insufficient cooling of the particulate debris, the debris can re-melt and fail the RPV. Upon failure of the RPV, the molten materials flow into the reactor cavity. The initial conditions of melt ex- vessel relocation depend upon the location & size of the breach, pressure in the RPV, etc.

In reactor cavity filled with water, the falling melt jets break up into discrete droplets due to hydrodynamic and thermal effects, undergo a steam explosion or solidify without a steam explosion, sediment and form a debris bed at the reactor cavity bottom (Ex-vessel FCI). A sketch illustrating an ex-vessel FCI scenario is shown in Figure 1.2. Formation of a debris bed in the basemat is paramount to terminate the accident since the porous media has more chances to be cooled down to ambient

12 conditions. However, long-term coolability of the debris has to be achieved for successful termination of the severe accident. The steam explosion energetics and debris bed coolability depend upon several inter-related mechanisms: including melt jet breakup, melt droplet behaviour (e.g., deformation/fragmentation in water), solidification and settlement (sediment/packing) of debris particles. The jet breakup process forms the initial conditions for melt-droplet evolution in the pre- mixture region (indicated in Figure 1.2.). The sizes of the melt droplets depict the rate of effective heat transfer to water, and consequently the quench rate and void build-up in the pre-mixture. Quench rate and void build-up are important parameters that govern steam explosion intensity [Meignen, et al., 2014]. The jet breakup and droplets quench process directly affect the debris bed characteristics (size distribution and morphology of debris particles as well as bed configuration), which are crucial to debris bed coolability since they determine the effective diameter of debris particles [Li & Ma, 2011] and the effective porosity of the debris bed [Ma & Dinh, 2010] employed in debris bed coolability analysis. Hence, it is important to investigate the individual mechanisms to qualify the effectiveness of the severe accident management strategy and successful termination of the severe accident.

Figure 1.2: Illustrative sketch depicting a severe accident scenario in Nordic type BWRs [Basso, 2017]. 13

1.2. Objectives and tasks

This thesis is aimed at complementary addition to the overwhelmingly progressive experimental research database in the field of nuclear safety with a specific focus on FCI. In this work, the broader as well as detailed aspects of FCI phenomenology is experimentally studied with an interest in i) improving understanding, ii) clarifying existing hypothesis and providing data useful for development or validation of models implemented in state of art FCI codes. The general objective of the thesis is to focus on the various aspects of FCI including debris formation, steam explosion and oxidation. These general set of objectives is accomplished by dealing with a set of tasks and subtasks listed as follows,

Experimental investigation of debris formation phenomena (Addressed in Section 2 (Paper I-II)).  To upgrade and use the MISTEE facility platform at KTH for study of jet breakup and debris formation phenomena at small scale.  To develop a facility for a relatively large-scale study of FCI mechanisms with a low temperature simulant metal melt that does not involve steam explosions.  To provide qualitative and quantitative data for the development of models used in analysis of debris formation and coolability by performing systematic separate effect tests.

Experimental investigation of steam explosion phenomena. (Addressed in Section 3 (Paper III-IV))  To perform separate effect experimental studies for study of influential parameters and clarification of relevant hypothesis formulated based on large-scale experimental findings.  To upgrade and develop the MISTEE-HT facility at KTH to perform experiments with high temperature melts and thereby reduce the gap between simulant and prototypic conditions.

Experimental investigation of oxidation behavior of metal melts during FCI. Addressed in Section 4  To perform separate effect study on oxidation behavior of metal drops in water.

The forthcoming sections in this thesis provides details of the specific tasks and a summary of respective research results.

2. DEBRIS FORMATION

2.1. Introduction

With a general focus on developing a physical understanding of the thermal- hydraulic processes that govern debris formation mechanism, several dedicated FCI studies have been carried out [Cronoberg, et al., 1974], [Johnson, et al., 1974], [Ginsberg, 1985], [Gabor, et al., 1988], [Schneider, et al., 1992], [Chu, et al., 1995], [Kondo, et al., 1995], [Dinh, et al., 1999], [Bang, et al., 2003], [Piluso, et al., 2005], [Magallon, 2006], [Abe, et al., 2006], [NEA., 2007] [Mathai, et al., 2015], [NEA., 2014] [Lu, et al., 2016] [NEA., 2018]. These research works helped characterize the FCI process by dissecting the mechanisms of jet breakup and debris bed formation. Generally speaking, capillary instabilities in the case of small jets, and Rayleigh- Taylor & Kelvin Helmholtz instabilities in the case of large jets are considered as the dominant mechanisms of jet breakup [Wang, et al., 1989], [Berg, et al., 1993], [Burger, et al., 1995]. To investigate the effect of melt-coolant physical properties on jet breakup, Dinh, et al. [1999] performed several experiments with varied melt simulants (e.g., Cerrobend, water) and coolant simulants (paraffin, water). Their study deemed prominent effect of melt-coolant density ratio and no effect of melt- coolant viscosity ratio on jet breakup behavior. To clarify effects of jet parameters on breakup under boiling conditions, Abe, et al. [2004], Abe, et al. [2006], Matsuo, et al. [2006] and Matsuo, et al. [2008], conducted experiments by quenching 400 g of molten U-alloy78 to simulate molten corium behavior in water. The results suggested the diameter of the jet affected the breakup length while jet velocity was found to be insignificant. In some FARO experiments, Magallon, [2006], studied debris bed formation behavior by pouring large masses (≈180 kg) of corium prototypic mixtures (UO2-ZrO2) into water. The results show agglomeration of melt (cake formation) can occur if the melt reaches the bottom of the water pool in liquid state without complete fragmentation and sufficient coolaing. The masses of agglomerates were shown to decrease as a function of increased water pool depth and the mean particle diameter increased at saturation conditions. Similarly, to develop a comprehensive database on debris agglomeration characteristics, Kudinov, et al. [2010] and Kudinov, et al. [2013] performed experiments with corium simulant materials and water system. The results have shown that larger fraction of agglomerated debris are formed if the water pool depth is reduced [Kudinov, et al., 2013].

While substantial progress in understanding of the basic factors phenomena has been achieved, thanks to the combined experimental and analytical efforts, there is a lack of separate effect experimental studies that can be used for model development and validation. For a complex phenomenon such as FCI, attaining a comprehensive database is a challenging task. Typical for large scale experiments

(starting from few kilograms of molten materials), quantification of individual parameters is difficult especially when one considers the wide range of parameters involved and the number of experiments that are necessary. The aim of this work (described in details in Paper I-II) is to provide qualitative and quantitative data for the development of models used in analysis of debris formation and coolability by performing systematic separate effect tests.

2.2. Experimental methods and conditions

2.2.1. MISTEE-JET facility

The MISTEE (Micro interactions in steam explosion experiments) set-up is originally constructed and intensively used for steam explosion study using droplets of different molten materials [Hansson, 2010]. In the present study, the facility is modified to enable experiments involving small jets in the diameter range of 5 to 10 mm. Figure 2.1, shows schematic of the MISTEE-Jet facility. The melt is prepared in a ceramic crucible heated by an induction furnace (6 kW, 80-180 kHz) with a maximum achievable temperature of approximately 1500°C. Melt temperature is monitored by a C-type thermocouple (measurement uncertainty of ±4.5°C) mounted from the inside of the crucible. The crucibles accommodate a hole in the bottom which is covered by a plug during melt preparation. The size of hole can be varied to attain the initial diameter of melt jet. Upon attaining the desired temperatures, the plug is raised by pneumatic operation so that the bottom hole of the crucible is uncovered for melt ejection. The melt is ejected through the bottom hole into a rectangular plexi-glass tank (500mm×75mm×75mm) filled with water. The depth of the water pool is 465 mm. A high-speed cinematography system comprising of a CMOS digital camera (Redlake HG50LE) and X-ray radiography is used to record the dynamic interactions. The X-ray radiography system comprises of a continuous X-ray source tube (Philips MCN 321 – max. 320 keV), an X-ray converter and an image intensifier coupled to a high speed camera (Lightning RDTPlus camera) with capability of acquiring images at 5000 fps.

Pneumatic plug set-up

Crucible Induction heating – Melt generation chamber

Figure 2.1: Schematic of the MISTEE-Jet facility.

2.2.2. JEBRA facility

The JEBRA (Jet Breakup) facility is developed at KTH with an intention to study breakup and debris formation characteristics of metallic melt at a lager scale than in MISTEE-Jet experiments. Wood's metal is used in the present study. Figure 2.2, presents the schematic of the JEBRA facility. The experimental arrangement mainly consists of a melting chamber, a discharge system with a valve and nozzle, a confined water pool chamber and a fast visual acquisition system. The melt generation system is a stainless steel container heated by two band heaters (240 V, 14 A) with a capability to provide up to 90 kg of molten metal. The temperature of the melt is measured at multiple levels inside the melt generation container by K-type thermocouples (measurement uncertainty of ±2.2°C). The nozzle underneath the container is changeable to provide the desired initial diameter of melt jet. The test section is a rectangular plexi-glass tank which is 0.5m × 0.5m × 2.5m in dimensions. The depth of the water pool is 2.4m. A pneumatically controlled ball valve arrests the melt inside the container. The valve and nozzle structures are heated by independent rope heaters for temperature control. A debris catcher is assembled at the bottom of the test section. The dynamic interactions are visualized through a

18 high-speed camera (Lightning RDT Plus camera) with capability of acquiring images at 5000 fps.

Figure 2.2: Schematic of JEBRA facility.

2.2.3. Melt simulants and test conditions

The purpose of using different simulant materials is to attain possible effects of thermo-physical properties of different materials on melt jet breakup and debris formation. Since the melting temperature of Wood's metal (≈70°C) is well below the boiling point of water, intensive (boiling) heat transfer between melt and coolant is negligible; the corresponding study can therefore be conceived to investigate the effective contribution of hydrodynamics in jet breakup characteristics. Eutectic mixtures of WO3-Bi2O3 and WO3-ZrO2 are also used in this study (see, Table 2.1. for properties) as corium simulant materials. Melting temperatures (Tliquidus = 870 °C 19 and 1231°C respectively) of the simulant binary oxide materials are accessible and provide higher flexibility and better control of the tests than that of corium melt (Tliquidus ~ 2800°C to 3200°C). It is also shown that the morphology of the WO3- Bi2O3 and WO3-ZrO2 debris show resemblance to those obtained with prototypic compositions [Kudinov, et al., 2010], [Kudinov, et al., 2013].

Table 2.1: Thermophysical properties of simulant melt materials

Parameter Wood's metal WO3-Bi2O3 WO3-ZrO2 UO2-ZrO2 (27:73 mol%) (74:26 mol%) (70:30 mol%)

Tliq (°C) 170 4870 41231 22811 ρ (kg/m3) 19730 36724 35580 27121.6 σ (N/m) 11 30.18 30.12 20.8 k (W/mK) 118.8 - - 23.601

Cp (J/KgK) 1168 - - 2680.7

Properties referenced from 1 [Bang, et al., 2003], 2[Song, et al., 2006], 3[Journeau, 2010], 4[Chang, et al., 1966]. Component mixture ratios are eutectic. Properties correspond to the melting point (Tliq) of each material.

To simplify experiments and measurements/visualization, most of the experiments are performed at the small-scale MISTEE-Jet facility. However, a few tests are carried out with Wood's metal at the large-scale JEBRA facility, to incorporate the effects of spatial scales and also to establish a rather fair comparison with other relatively large-scale experiments. Table 2.2 presents the experimental conditions.

Table 2.2: Test conditions for study of debris formation behaviour Parameter Experiment series 1 Experiment series 2 Experiment series 3 Component Bi O :WO (27:73 mol%) WO :ZrO (74:26 mol%) Wood's metal mixture 2 3 3 3 2 Composition (-) Eutectic Melt mass (kg) ≈0.2 to 0.5 ≈0.5 to 70 T liq (°C) ≈870±5 ≈1231±5 73±2 T superheat (K) ≈100 to 200±5 ≈20 to 40±3

∆T subcool (K) ≈10 to 90±3 ≈70 to 90±3

Dj (mm) ≈5 to 10 ≈5 to 30

2.3. Effect of experimental conditions on jet breakup

The sequential snapshots of the melt-coolant interaction under non-boiling conditions are presented in Figure 2.3. For convenience, we define time t=0ms when the melt jet leading edge first contacts the water pool surface, and assume jet diameter Dj equal to the diameter of the crucible’s bottom hole in MISTEE-Jet or the nozzle in JEBRA. For the cases of low jet penetration velocity (Uj ≈ 1 .5 m/s, Wea ≈ 10 to 40; Figure 2.3a-b), momentum exchange between the jet and water causes significant deceleration of the jet, thereby resulting in an increased surface area at the leading edge (for example, see Figure 2.3b, t=19, 30ms). The jet leading edge resembles an inverted mushroom-like shape, which in general arises due to strong enough deceleration forces that overcome the surface tension forces, which keeps the jet surface intact. The snapshots (Figure 2.3b, t=19, 30ms) pictorially depict the concept of Rayleigh-Taylor type instabilities that can occur at the plane interface between two fluids of different densities accelerated towards each other. Followed by this initial deformation, the jet lateral surface undergoes wavy undulations (Figure 2.3b, t=63, 95ms) while fragments are gradually stripped from the jet leading edge thereby undergoing a coarse breakup. A similar behavior for the jet leading edge deformation was also observed in the series of experiments conducted by Dinh, et al. [1999] with paraffin oil and water system.

For the other cases with relatively higher penetration velocity (Uj ≈ 3m/s, Wea ≈ 180 to 320; Figure 2.3c-d), the leading edge decelerates followed by severe stripping of fragments from the jet lateral surfaces. Stripping of fragments from jet side causes erosion leading to thinning of melt jet and eventually breakup. The observed stripping mechanisms resembles the concept of Kelvin-Helmholtz type instabilities, which may occur as a result of velocity difference across the interface of the fluids. The stripped fragments existing in the flow can be expected to undergo further hydrodynamic decomposition into fragments of even smaller size. General observations indicate that wave stripping in the form of Kelvin-Helmholtz type of instabilities that prevail on the jet lateral surface is the dominant mechanism for jet breakup of Wood's metal melt under higher Uj conditions.

Figure 2.3: Snapshots depicting fragmentation dynamics of molten Wood's metal melt jet in a water pool under non-boiling conditions.

The jet breakup length is the distance travelled by the melt jet in water pool until its leading edge fragments into a cloud of droplets. In other words, it is the length of the coherent jet penetrating into the water pool. The influence of jet diameter (denoted by Dj) on jet breakup length is shown in Figure 2.4. One can see that the jet breakup length increases with increasing jet diameter. Under varied experimental conditions, the jet breakup lengths seem to follow a linear trend and show a good agreement with those of Matsuo, et al. [2008] with Ualloy-78, KTH experiments with Cerrobend and FARO experiments with UO2-ZrO2 [Pohlner, et al., 2013]. The jet breakup length is a function of density ratio of melt to coolant as shown by the correlations of Taylor, [1940] (Equation 2.1), Epstein & Fauske, [2001] (Equation 2.2) and Matsuo, et al. [2008] (Equation 2.3). Figure 2.5a depicts the effect of jet penetration velocity (denoted by Uj) on jet breakup length. It can be observed that the jet breakup length seems to be rather constant. However, if experimental uncertainty is taken into account in the observed deviations, one can say that the jet penetration velocity has no significant influence on the jet breakup length, as reflected by the correlations of Taylor, [1940], Epstein & Fauske, [2001] and Matsuo, et al. [2008], where the ratio of jet breakup length to jet diameter is only proportional to the square root of 휌푗⁄휌푐 (Figure 2.5b). 퐿 𝜌 0.5 푏푟푘 = 5.3 ( 푗) (2.1) 퐷푗 𝜌푐 퐿 1 𝜌 0.5 푏푟푘 = ( 푗) (2.2) 퐷푗 2퐸표 𝜌푐 퐿 𝜌 0.5 푏푟푘 = 7.52 ( 푗) (2.3) 퐷푗 𝜌푐

1.2

0.8

(m) 0.6

brk L Woods metal (Uj:1-3.5m/s) 0.4 WO3-Bi2O3 (Uj:1-2m/s) WO3-ZrO2 (Uj:1-2m/s) FARO (UO2-ZrO2) (Uj:4.5-7m/s) 0.2 KTH (Cerrobend) (Uj:3.6m/s) Matsuo (U-alloy78) (Uj:1-5m/s) Linear (-) 0 0 0.02 0.04 0.06 0.08 0.1 Dn (m) Figure 2.4: Jet breakup length as a function of nozzle diameter

50 Woods Metal (Dj:5-30mm) WO3-Bi2O3 (Dj:5-10mm) 40 WO3-Bi2O3 (Dj:7mm) KTH (Cerrobend) (Dj:25mm)

j 30 FARO (UO2-ZrO2) (Dj:40-75mm)

/ D / brk

L 20

10 a 0 0 2 4 6 Uj (m)

Woods metal (Uj:1-3.5m/s)

1 WO3-Bi2O3 (Uj:1-2m/s)

WO3-ZrO2 (Uj:1-2m/s)

KTH (Cerrobend) (Uj:3.6m/s) (m) FARO (UO2-ZrO2) (Uj:4.5-7m/s)

brk 0.5 L Matsuo corr…

Taylors corr… b 0 Epsteins corr… 0 0.05 0.1 0.15 0.2 0.5 Dj (휌j /휌c ) Figure 2.5: Jet breakup length as a function of jet penetration velocity (a) and comparison with correlations for jet breakup length (b).

2.4. Effect of experimental conditions on debris properties

Snapshots of a molten binary oxide mixture of WO3-Bi2O3 and WO3-ZrO2 quenched into a water pool at varied degrees of water subcooling are shown in Figure 2.6 and Figure 2.7 respectively. As the coherent jet penetrates the water column, it immediately forms a vapor film that restricts direct contact between the melt and subcooled water. The snapshots reveal significant undulations on the jet surface. The molten jet decelerates and spreads sideways. The breakup is noticed at the melt jet leading edge. Snapshots of a molten binary oxide mixture of WO3-Bi2O3 shows trails of very fine dust released from the melt during quenching. The dust cloud significantly obstructs the view of jet breakup pattern. However, the snapshots of molten binary oxide WO3-ZrO2 does not show traces of smoky dust except under high subcooling conditions (Figure 2.7). The snapshots depict significant influence of water subcooling on the cooling of the molten droplets. At low water subcooling, the fragments seemed to reach the bottom of the water pool at high temperature (Figure 2.7) while at high water subcooling the fragments started to cool and solidify soon after detachment from the jet (Figure 2.7). It can also be observed that at high

24 subcooling the solidifying fragments are broken down further into smaller fragments.

The effect of water subcooling on the particle size distribution of WO3-Bi2O3 and WO3-ZrO2 is shown in Figure 2.8. The effect is quite apparent. The fraction of small size particles increases at high subcooling conditions. The plot in Figure 2.8a shows that the transition in particle size occurs in rather narrow range of water subcooling. “Fine” fragmentation regime is observed at 80 K to 60K subcooling with no significant variation of the particle size. A transition of the fragmentation regime is observed from 60 K to 40K. Large particle sizes are observed at subcooling below 40 K. No significant changes in the particle size are observed up to 10K of water subcooling. It is remarkable that at ~10K subcooling the size of the particles noticeably increases. Similar to WO3-Bi2O3, the plot for WO3-ZrO2 shows a transition in particle size, however, at a higher subcooling ~70 K (Figure 2.8b).

The mass median particle diameter (Dm) (i.e. the size that corresponds to 50% on the CDF of particle diameter) is plotted as a function of water subcooling in Figure 2.9. A comparison with the DEFOR and FARO data shows a good agreement between different tests (if other conditions such as melt superheat are similar) with corium simulant WO3-Bi2O3. The size of the debris decreases as subcooling increases. There are also apparent changes in debris morphology (Figure 2.10). For water subcooling between 10 K and 40K the mass fraction of fractured particles is ~ 85 to 95%. The fraction of fractured particles increases in the transition regime up to 97% and even further (Figure 2.11).

Figure 2.6: Breakup of molten eutectic WO3-Bi2O3 jet (∆Tsup ≈ 130K) in a water pool at varied subcooling.

Figure 2.7: Breakup of molten eutectic WO3-ZrO2 jet (∆Tsuperheat ≈ 160K) in a water pool at varied subcooling. 27

100 a ∆Tsubcool≈80K (average)

80 ∆Tsubcool≈70K (average)

∆Tsubcool≈60K (average) 60 ∆Tsubcool≈50K (average)

40 ∆Tsubcool≈40K (average)

∆Tsubcool≈30K (average) 20 ∆Tsubcool≈20K (average)

Cumulative mass fraction (%) fraction mass Cumulative D =5mm 0 n ∆Tsubcool≈10K (average) 0 1 2 3 Debris size (mm) 100 b ∆Tsubcool≈ 80K (average) 80 ∆Tsubcool≈ 70K (average)

60 ∆Tsubcool≈ 60K (average)

40 ∆Tsubcool≈ 50K (average)

20 ∆Tsubcool≈ 40K (average)

Dn=7mm 0 ∆Tsubcool≈ 30K (average) Cumulative mass fraction (%) fraction mass Cumulative 0 1 2 3 4 5 6 Debris size (mm) Figure 2.8: Debris size distribution for WO3-Bi2O3 (a) and WO3-ZrO2 (b) as a function of water subcooling.

8 7 MISTEE Jet (WO3-Bi2O3, Dn=5mm, ∆Tsuperheat≈ 130K) 6 MISTEE Jet (WO3-Bi2O3, Dn=10mm, 5 ∆Tsuperheat≈ 130K) 4 MISTEE Jet (WO3-ZrO2, Dn=7mm, ∆Tsuperheat≈ 160K)

(mm) 3 m m DEFOR (WO3-Bi2O3, Dn=10mm, D 2 ∆Tsuperheat≈ 100-130K ) 1 FARO (UO2-ZrO2, Dn=100mm, 0 ∆Tsuperheat≈ 90-200K) -10 10 30 50 70 90

∆Tsubcool (K) Figure 2.9: Mass averaged median particle diameter as a function of water subcooling.

Figure 2.10: Debris particle morphology at varied water subcooling.

100

80 MISTEE Jet (WO3-Bi2O3, Dn=5mm, ∆Tsuperheat≈ 60 130K)

40 MISTEE Jet (WO3-ZrO2, Dn=7mm, ∆Tsuperheat≈ 20 160K)

Mass fractionof fractured particles(%) 0 0 20 40 60 80 100 ∆T (K) subcool Figure 2.11: Mass fraction of fractured particles as a function of water subcooling. 29

The experimental results clearly depict significant differences in the level and modes of fragmentation at high and low subcooling conditions. The indicative snapshots of the jet breakup mechanism have shown no significant differences at varied water subcooling. Hence, a mechanism other than primary fragmentation (jet breakup) can be considered responsible for the transition in debris morphology. More specifically, the observations of fractured particles reveal significance of cracking induced by thermal stresses. The molten droplets, detached from the jet are surrounded by a vapor film that drastically reduces heat transfer from the molten droplet to water. As the droplet cools and solidifies, the surface temperature drops and stable film boiling can no longer be maintained. Transition from film to nucleate boiling, involving very high heat transfer rates, reduces particle temperature within a very short span of time. The rapid quench process produces significant temperature gradients between the center of the solidifying droplet and the surface resulting in a thermal shock with high tensile stresses. As a result, the solidifying particles undergo fracture. At a higher degree of water subcooling the transition from film to nucleate boiling occurs at higher particle surface temperature and leads to higher heat fluxes and respective stresses. This also suggests that the size of solidified debris may not be representative for the sizes of liquid droplets formed after detachment from the jet. It can also be clearly noted that the debris sizes of WO3-ZrO2 are considerably larger than WO3-Bi2O3 debris (Figure 2.10 and 2.11) indicating the effect of melt composition. The fraction of the fractured particles is also lesser compared to what is observed with WO3-Bi2O3. These observations suggest that WO3-ZrO2 material has higher resistance to thermal shock compared to WO3-Bi2O3. This can be explained if WO3-Bi2O3 has a lower thermal conductivity and fracture stress. The observation on debris morphology shows similarities to the observations made by Tyrpekl [2012] for UO2-ZrO2 mixture. The debris particles acquired after KROTOS experiments show formation of broken particles with sharp edges and full rounded debris. The analysis showed that approximately 90% of the particles blow 1mm size had an angular morphology (typically broken particles with sharp edges) that can be attributed to solid fracture of melt during quenching. However, it is also instructive to note that the KROTOS experiment with UO2-ZrO2 underwent a mild steam explosion whether the MISTEE-Jet experiments presented in this section did not involve steam explosion.

From the visual evidences, the Rayleigh-Taylor (R-T) type instabilities and Kelvin- Helmholtz (K-H) type instabilities can be considered as the mechanisms responsible for jet fragmentation. Further fragmentation of the droplet formed during the jet breakup process can be depicted by the critical Weber number theory. A comparison of median debris size acquired from experiments with the fragmentation theories is shown in Figure 2.12. The median debris size for WO3-Bi2O3 material at a superheat range of ≈ 130K is seemingly in agreement with the theories. However, at a higher superheat (∆Tsup≈ 330K) the debris size seems to be larger compared to the

30 theoretical predictions. A large fraction of the debris actually originates not from fragmentation of liquid (described by the theories) but from fracture of solid debris. Comparison of the experimental data for WO3-ZrO2 melt material (that mostly undergoes liquid fragmentation) demonstrates that the predictions underestimate the size of debris.

100 MISTEE Jet (WO3-Bi2O3, Dn=10mm, ∆Tsuperheat≈ 130K) MISTEE Jet (WO3-Bi2O3, Dn=5mm, ∆Tsuperheat≈ 130K) MISTEE Jet (WO3-Bi2O3, Dn=5mm, ∆Tsuperheat≈ 330K) MISTEE Jet (WO3-ZrO2, Dn=7mm, ∆Tsuperheat≈ 160K) Critical Weber number (WO3-Bi2O3) Critical Weber number (WO3-ZrO2) K-H (WO3-Bi2O3) K-H (WO3-ZrO2) 10 R-T (WO3-ZrO2)

R-T (WO3-Bi2O3)

(mm) m m 2 D 휌 푢 퐷 푊푒 = 푐 푑 1 푐 σ

휌푗 + 휌푐 2π σ푗 휆퐾−퐻 = 2 휌푗휌푐푢

3 σ푗 휆 = 2π 푅−푇 ൫휌 −휌 )푔 0.1 푗 푐 0.1 1.1 2.1 Relative velocity (m/s) Figure 2.12: Comparison of mass averaged median particle diameter and predictions by fragmentation theories.

6 MISTEE Jet (WO3-Bi2O3, 5 ∆Tsuperheat≈ 130K, 4 ∆Tsubcool≈ 10-30K)

(mm) 3

m m DEFOR (WO3-Bi2O3,

D ∆Tsuperheat≈ 100-130K, , 2 ∆Tsubcool≈ 7-30K ) 1 FARO (UO2-ZrO2) 0 1 10 100 D (mm) n Figure 2.13: Mass averaged particle median diameter as a function of nozzle diameter for simulant and prototypic binary oxide materials.

Figure 2.13 shows the effect of jet diameter on debris size of binary oxide simulant and prototypic materials. Slight discrepancies seem to be well within the experimental uncertainties. The mass median particle diameter (Dm) from the current experiments with low water subcooling (0 to 30K) is compared with the data for loose debris from DEFOR and FARO experiments. It is remarkable that ~10 31 times different jet diameter and the use of simulant or prototypic melt materials have no significant effect on the minimum values of the average particle size (~2- 3 mm) obtained in different experiments. The maximum value of the mass median particle size increases from ~3 to ~4.5mm as the nozzle diameter increases. However, increase of the maximum mass median particle after nozzle diameter exceeds 20 mm is only marginal. It is instructive to note that the average size of the debris has a profound effect on debris coolability (e.g. [Yakush, et al. 2013a], [Yakush, et al. 2013b].

10 2.5 MISTEE-Jet (Woods metal, Tmelt: 90-110°C, a ∆Tsubcool: 75-90K, Uj:1-2m/s) 2 8 JEBRA (Woods metal, Tmelt: 90-110°C, ∆Tsubcool:

85-90K, Uj:2.5-3.5m/s) )

j Matsuo et al (2008) JEBRA (Ualloy-78, Tmelt: 270-

mm 6 1.5 (

/ D / 300°C, ∆Tsubcool: 30-40K, Uj:1.5-6m/s)

m m

D D 4 1 MISTEE-Jet (Woods metal, Tmelt: 90-110°C, ∆Tsubcool: 75-90K, Uj:1-2m/s) JEBRA (Woods metal, Tmelt: 90-110°C, ∆Tsubcool: 0.5 2 85-90K, Uj:2.5-3.5m/s) b Matsuo et al (2008) JEBRA (Ualloy-78, Tmelt: 270- 300°C, ∆Tsubcool: 30-40K, Uj:1.5-2m/s) 0 0 0 10 20 30 40 0 2 4 6 Dj (mm) Uj (m/s) Figure 2.14: Mass averaged particle median diameter as a function of initial jet diameter (a) and dimensionless mass averaged median particle diameter as a function of jet penetration velocity (b).

In Figure 2.14, the median diameters of quenched Wood's metal debris particles are compared with the results obtained from the experiments performed by Matsuo, et al. [2008]. which involves Ualloy-78 (ρ≈8200kg/m3, Tliq ≈78.8°C) fragmentation in water under boiling conditions. It can be observed that the median debris size increases as a function of jet thickness. The size distribution of the metallic debris particles showed that the average size of the particles increases as a function of jet diameter and decreases with an increasing velocity. However, the average size of oxidic fragments did not seem to be significantly affected by jet diameter and velocity. The distinct results indicated that the hydrodynamic mechanism was overridden by the prominence of thermal mechanism that instigates fracture of oxidic droplets when subject to rapid quench rate. The rate of fracture or in other words the prominence of the thermal mechanism was governed by water subcooling as well as the material properties that determine the magnitude to which a material can resist fracture. A clear distinction in debris morphology between metallic and oxidic melts was observed. Metal debris was mostly rough and elongated with no internal porosities. With more pronounced effect of thermal fracture, the binary oxide material WO3-Bi2O3 mostly comprised of rough edged broken particles (The

32 mass fraction of fractured particles is ˃80 % even under low subcooling conditions), with a very small amount of smooth particles having internal cavities.

3. Steam explosion

3.1. Introduction

An extraordinary phenomenon as steam explosion or vapour explosion is very important in the context of reactor safety analysis since a high energy explosion can result in a catastrophic event of an early containment failure. It is expected that the FCI occurs quite quiescently, involving melt fragmentation at the scale of some millimetres as shown in the previous chapter. However, under some conditions, the FCI can escalate into an event called steam explosion, in which, due to a very fine fragmentation of the melt premixed with water, the melt enthalpy is rapidly and massively transformed into mechanical energy of expanding vapour. This type of potentially very energetic phenomenon leads to the formation of shock waves and can endanger the structures and the containment, which is the last physical barrier isolating radioactive materials from the environment during ex-vessel phase of SA.

Classically, the sequence of events leading to a steam explosion during FCI is conceptualized to occur in four distinct phases: pre-mixing, triggering, propagation and expansion. Figure 3.1 depicts a case where WO3-Bi2O3 melt undergoes steam explosion because of spontaneous triggering in MISTEE-Jet experiments. Initially the falling melt jet breaks up into molten droplets typically in the range of several millimetres forming a so-called pre-mixture with water. The melt droplets undergo film boiling which restricts direct contact between melt and water. The vapour film acts as a thermal resistance reducing heat transfer from melt to water (Figure 3.1a- c: pre-mixing phase). Following an triggering event e.g. pressure wave or spontaneous self-triggering e.g. due to transition in boiling regimes, the vapour film surrounding the melt droplet is destabilized and melt-water direct contact is established at localized regions (Figure 3.1d-e: triggering phase) leading to the fine fragmentation of melt droplet. Fine fragmentation and subsequent rapid heat transfer between the finely fragmented melt and water produces a high pressure shock wave that destabilizes the vapour films surrounding the adjacent melt droplets in the pre-mixture thereby propagating the process (Figure 3.1f-h: propagation phase) and expanding against the inertial constraints of the surrounding (expansion phase).

Figure 3.1: Illustration of the distinct phases of a steam explosion sequence during a FCI event observed in MITE-Jet experiments. Melt material: WO3-Bi2O3, ∆Tsup ≈ 135K, ∆Tsub ≈ 82K.

The most common approach to understand the thermal hydraulic process of steam explosion research is to simulate the hypothetical scenario of FCI in nuclear reactors by performing out-of-pile experiments at a relatively large-scale or small-scale by pouring hot melt into water. The experiments that involve pouring a relatively large quantity (˃1kg) of melt mixture into a water pool are typically considered “larger- scale” experiments. Such relatively large-scale experiments have been performed with prototypic melt composition as well as high temperature simulant melts such as Alumina. Some of such notable experimental programs include KROTOS (CEA) [Huhtniemi, et al., 1997], FARO (JRC, Ispra) [Magallon & Huhtniemi., 2001], PREMIX (KIT) [Albrecht, et al., 2001], ZREX/ZRSS (ANL) [Cho, et al., 1988], TROI (KAERI) [Song, et al., 2005], FITS (SNL [Buxton & Benedick., 1980]), ALPHA (JAERI) [Yamano, et al., 1995], ECO (FZK) [Cherdron, et al., 2005] etc. It is instructive to note that there is a significant difference in melt mass between the various experimental programs that involve melting of protypic and simulant materials. For example, the KROTOS experiments involve upto 5 kg of prototypic melt composition and simulant melt interacting with water whereas the FARO experiments involve up to 200 kg of prototypic melt composition interacting with water. The phenomena of pre-mixing and propagation/expansion can be studied in these integral experiments. Several interesting insights have been made for example it was found in KROTOS experiments that molten alumina produced strong steam explosions compared to that of molten corium mixture which produced weaker explosions [Huhtniemi, et al., 1999] giving rise to the so-called “material effect” on steam explosion. Such experimental programs are used widely to complement the development and validation of state-of-art FCI computer codes such as TEXAS [Chu & Corradini., 1989], MC3D [Berthoud & Valette., 1994], JASMINE, IDEMO [Carachalios, et al., 1983], CHYMES [Fletcher & Thyagaraja., 1991], ESPROSE [Medhekar, et al., 1991], CULDESAC [Fletcher, 1993] etc. that analyse the steam explosion phenomena. State-of-the-art reviews on experimental and analytical studies of steam explosion can be found in Corradini, et al. [1999], [Fletcher, 1995] and Meignen, et al. [2014] [Shen, et al., 2018] and status of steam explosion research can be found in [NEA., 2007], [NEA., 2014], [NEA., 2018].

Fine fragmentation behaviour of melt droplets is the key mechanism that forms the fundamental of a steam explosion phenomenon. For this purpose, study of fragmentation mechanism at the level of droplets (“small-scale” experiments usually involving melt up to 20 grams) has also been a popular approach used by several researchers in the past decades. Single droplet experiments may not simulate the realistic conditions related to propagating and expanding events however can produce insights into the mechanisms involved in melt fragmentation and also provide useful data on the limiting as well as promoting mechanisms of steam explosion.

Melt fine fragmentation is initiated locally by some triggering events that are still not very well known and understood. Several mechanisms have been identified and studied, among which, either a hydrodynamic or a thermal mechanism. The hydrodynamic mechanism is jointly driven by surface tension and shear force due to the relative flow of the droplet with respect to surrounding water, typically limited by a critical Weber number: 𝜌 푢2퐷 푊푒 = 푙 푑 (3.1) 푐 𝜎 Where 휌푙 is the density of water, 푢 is the melt-water relative velocity, 퐷푑 is the droplet diameter and σ is the surface tension of the droplet. The Weber number criteria determines the conditions for different regimes of a hydrodynamic fragmentation that include (i) vibrational breakup (We < 12); (ii) bag breakup (12 < We < 50); (iii) bag-and-stamen breakup (50 < We < 100); (iv) sheet stripping (100 < We < 350); (v) wave crest stripping (We > 350); and (vi) catastrophic breakup [Pilch & Erdman., 1987], [Gelfand, 1996]. A recent analysis of droplet breakup regime by direct numerical simulation is provided in [Castrillon, et al., 2015].

Figure 3.2: Thermal fragmentation models proposed in Kim & Corradini. [1988], (above) and Ciccarelli & Frost. [1994] (below).

The thermal mechanism however is largely related to the destabilization of vapour film that surrounds the melt droplet so that direct contact between water and melt is established at some localized region, driving the fine fragmentation of the melt. The mechanism of thermal fine fragmentation has been extensively studied in the past decades ([Kim & Corradini., 1988], [Ciccarelli & Frost., 1994]). Considering the very specific physical conditions involved with such fast transient multiphase phenomena, visual information of the process has been the predominant requirement to understand the mechanisms involved. A classic fragmentation theory, proposed in Kim & Corradini [1988], suggests that the vapour film surrounding the melt droplet collapses on exposure to a pressure wave and as a result of Rayleigh-Taylor type instabilities that arise on the interface, micro jets of water penetrate into the melt droplet and vaporize causing disintegration of the melt droplet similar to what can be observed in cavitation phenomena (Figure 3.2, above). Later on, based on the visual evidences, obtained by flash X-ray and regular photography, it is proposed in Ciccarelli & Frost. [1994], that, on destabilization of

38 the vapour film, fine filaments of melt protrude and the drop surface becomes highly convoluted establishing quasi contact between water and melt at localized region driving the fine fragmentation process (Figure 3.2, below).

The phenomenological description of fragmentation mechanism observed in early MISTEE experiments with tin melt Hansson, [2010] is consistent with some aspects of the classic theories aforementioned. It is observed using a simultaneous visualization by normal videography and X-ray radiography that the melt deforms or in other words pre-fragments on destabilization of the vapour film due to the passage of a pressure wave (Figure 3.3b: consistent with proposal in Ciccarelli & Frost. [1994]. The coolant then seems to entrain the already pre-fragmented melt, where the entrapped coolant rapidly evaporates leading to fine fragmentation of the melt droplet (Figure 3.3c: consistent with proposal in Kim & Corradini. [1988] and Buchnan, [1974]. Tyrpekl [2012] performed a detailed analysis of MISTEE tin debris resulting after steam explosion by SEM/EDS and estimated that the final porosity is in the range of approximately 52%. Further it is shown that the void inside the debris particles are formed mainly by open porosities that can be attributed to micro water jets penetrating the droplet surface consistent with what is proposed in Kim & Corradini. [1988].

Figure 3.3: MISTEE phenomenology of melt fine fragmentation. Schematic interpretation of fine fragmentation (above) and Matched images acquired by videography and X-ray radiography.

Several experimental works have been conducted at the level of droplets to study the fragmentation behaviour using different materials (Table 3.1 presents an example of the conventional thermal fragmentation experimental studies performed in the past decades and the respective focus parameters). Predominantly tin, aluminium and cerrolow have been used as the melt materials for study of steam explosion at small scale considering control over experimental conditions at lower temperatures. Some experiments however involved relatively high temperature materials including Iron oxide and silicon rich alloys [Nelson, et al., 2001]. The experiments using low and medium-temperature melts suffer from the combined effect of material and 39 temperature scaling. Study of droplet-scale steam explosion behaviour of high temperature melts (˃2000°C) that can closely simulate the prototypic conditions is however not accomplished. Despite substantial improvements in understanding the fine fragmentation mechanisms and heat transfer process over the years it is still not clear why certain conditions for melt and water are prone to strong steam explosion while other conditions result in no or weak steam explosions. Nevertheless, improvements in physical understanding and thereby improvement of mechanistic modelling is necessary.

Table 3.1: Thermal fragmentation experimental studies Melt Program/Author/ Melt # temperatures Focus of the study Reference components (°C) Effect of melt temp., water 1 [Dullforce, et al., 1976] Tin 232 temp., droplet falling depth. Silicon, [Nelson & Duda 1981] Ferrosilicon, Effect of melt temp., water [Nelson & Duda 1982], Aluminium, 1414, 1200, 2 temperature, trigger [Nelson, et al., 1995], Alloyed 660, 1565 intensity, system pressure [Nelson, et al., 2001] Aluminium, Iron oxide Hydrodynamic Vs thermal [Ciccarelli & Frost., 3 Cerrolow, Tin 177, 232 fragmentation, 1994] fragmentation mechanism Effect of non-condensable 4 [Akiyoshi, et al., 1990] Tin 232 gases [Matsumura & Nairai., Effect of melt temp., water 5 Tin 232 1996] temp., droplet falling depth. Tin, Lead, Zinc, 232, 328, Effects of melt temp., water 6 [Abe, et al., 2002] Aluminium 420, 660 temp.,trigger magnitude MISTEE [Park, et al., Tin, Mechanism of 7 2005], [Hansson, et al., WO3-Bi2O3 232, 870, 1135 fragmentation, effect of 2010], [Hansson, WO3-CaO. non-condensable gases 2010] Mechanism of SIGMA-2000 10 Steel 1500 fragmentation, Imposition [Chen, et al., 1999] of high pressure fields Mechanism of 11 [Board, et al., 1974] Tin 232 fragmentation [Takashima & Lida., Spontaneous steam 12 1995] [Takashima, Tin 232 explosion, Mechanism of 2008] fragmentation [Kouraytem, et al., Fields metal, 13 60, 232 Explosion behaviour 2016] Tin` Effect of material MISTEE WO -CaO, WO - 1135, 1250, composition, water 14 3 3 Current work ZrO2, Al2O3, Zr, 2054, 1855 temperature, melt superheat, metal oxidation.

This part of the thesis work (described in details in Paper III) falls into the category of small-scale experiments where approximately 1 to 3 grams of hot melt is dropped into a subcooled water pool followed by steam explosion triggering or quenching without steam explosion triggering. The idea of this experimental work is not to focus on the classic study of fine fragmentation mechanism rather to focus on the limiting mechanisms of steam explosion and clarification of the hypothetical formulations put forth in steam explosion research. The latter part (described in details in Paper IV) of this work shows an attempt to bring the parameters for single droplet experiments closer to that of prototypic conditions to a practical extent achievable.

3.2. MISTEE-LT study of steam explosion

Figure 3.4 shows the schematic of the test facility named as Micro Interactions in Steam Explosion Experiments using Lower Temperature melt (MISTEE-LT). The melt generation system is an induction furnace (260 V, 40 A, 80-180kHz). Around ≈ 1 g of melt material is loaded in a tungsten crucible which consists of a hole at the centre of the bottom of the crucible. The crucible is housed inside a chamber where argon gas is purged to maintain inert conditions. A well-controlled pneumatic plug system is utilized to cover the bottom hole of the crucible during melt preparation. A C-type thermocouple (measurement uncertainty of ±4.5°C) is mounted from the inside of the crucible to measure the melt-temperature. On achieving desired temperature of melt in the crucible, the pneumatic plug is released to uncover the bottom hole in the crucible so that the droplet falls by gravity into a rectangular plexi-glass tank (180mm×130mm×150mm) containing water. The water temperature is monitored by K-type thermocouples (measurement uncertainty of ±2.2°C). A weak pressure pulse generated by an external triggering system aligned underneath the water tank disturbs the discharged melt droplet underwater. The pressure wave generating system is a piston set-up. A trigger hammer driven by rapid discharge of a capacitor bank (3 capacitors of 400Vdc and 4700mF each) impacts on the piston to generate a weak pressure pulse of up to 0.15 MPa. The interactions between the melt and water is recorded by a fast synchronous visualization system that is aligned with a capability to record simultaneously X-ray and photographic image. It consists of a CMOS digital camera (Redlake HG50LE), with an acquisition rate up to 20,000 fps for photography; and a continuous X-ray source tube (Philips MCN 321 – max. 320 keV), an X-ray converter and an image intensifier coupled to a high speed camera (Lightning RDTPlus camera), for radiographic imaging with an acquisition rate up to 8,000 fps. The camera system was calibrated with a phantom in order to obtain the spatial resolution. Typical spatial resolution for the tests range between 0.1 and 0.2 mm/pixel at an uncertainty range of approximately up to ±4 pixels.

The MISTEE tests have been designed to investigate a possible triggering mechanism of steam explosion, which is the so-called thermal fragmentation. This specific phenomenon leads to the fast fragmentation (some milliseconds) of a hot liquid drop to fine fragments in stable hydrodynamic conditions (i.e. Weber number << 10). Nevertheless, the phenomenon often needs a small perturbing event, e.g. a small pressure perturbation. It is noticed that, for a typical molten metal (or corium) drop of 5 mm, a pressure wave perturbation of 1 bar leads to a Weber number of about 0.05, whereas hydrodynamic instability (We ~10) is reached with perturbations of the order 15 bars (assuming Δ푃 = 휌푉푐).

Figure 3.4: Schematics of the MISTEE-LT test facility (top) and snapshot of the facility view (below).

3.2.1. Corium simulant materials

Several simulant materials are found to be suitable candidates for MISTEE study of steam explosion. To simulate the composition of oxidic corium melt, binary oxide materials including WO3-CaO, WO3-ZrO2 and Al2O3 are considered in MISTEE experiments. Binary oxide simulant materials with a temperature range between 800 and 1300°C have been widely used in DEFOR experiments and MISTEE-Jet experiments considering the similarities observed in debris morphology with UO2- ZrO2. Such moderate temperature melt simulants are advantageous for systematic studies with well defined parameters and control. However, it is not possible to simulate all aspects of corium phenomenology especially at high temperatures. Properties of select simulant materials are shown in Table 3.2. The first series of experiments have been conducted with molten binary oxide mixtures of WO3-CaO and WO3-ZrO2 in eutectic composition at the MISTEE-LT facility. Table 3.3 shows the test conditions. The use of simulant material Bi2O3-WO3 is neglected considering the peculiar characteristic of smoky trace in droplet trajectory that hinders the visualization of fine fragmentation process.

Table 3.2: Thermophysical properties of simulant melt materials selected for SE studies 2 Parameter Sn WO3-Bi2O3 WO3-ZrO2 WO3-CaO Al2O3 UO2-ZrO2

Tliq (°C) 232 9870 81231 81135 72054 2811 ρ (kg/m3) 7290 16724 15580 75490 3,43054 7121.6 σ (N/m) 0.574 10.18 10.12 70.17 3,40.65 0.8 k (W/mK) 67 - - - 56 3.601

Cp (J/KgK) 210 - - 6533 1400 680.7 Properties referenced from 1[Journeau, 2010], 2[Song, et al., 2006], 3[Rasmussen & Nelson 1972], 4[Elyutin, et al., 1973], 5[Frano, et al., 2015], 6[Haraldsson, 2000], 7[Schneider, 1979], 8[Chang, et al., 1966], 9[Speranskaya, 1970]. Component mixture ratios are eutectic. Properties correspond to the melting point (Tliq) of each material.

Table 3.3: Test conditions for study of steam explosion with binary oxide simulant material Parameter Experiment series 1 Experiment series 2

Component mixture WO3:CaO (75:25 mol%) WO3:ZrO2 (74:26 mol%) Composition (-) Eutectic Melt mass (g) ≈0.8 to 1.5

Melting temperature (T liq) (°C) ≈1135±5 ≈1231±5 Melt superheat (K) ≈100 to 350±5 ≈200±5 Water subcooling (K) ≈75 to 85±3 Nozzle diameter (mm) ≈5

3.2.2. Observation of steam explosion

Driven by gravity the molten droplet prepared in the furnace falls into the transparent tank filled with water. Once the melt droplet penetrates the water pool surface it emerges with a long trail of non-condensable gas which detaches from the droplet rear at ≈ 1.8ms (Figure 3.5a). After an initial acceleration into the water pool, the droplet starts to decelerate and attains terminal velocity at ≈100ms. As the droplet descends further into the water pool, the vapour bubble enshrouding the droplet remains wobbly, but no detachment of bubbles from the droplet rear is observed. However, the vapour film thickness seems to undergo periodic oscillations, which can be inferred from the aspect ratio (the ratio of the distance between the vertical peripheries and the horizontal peripheries) of the bubble over the quench period (Figure 3.5b). Hydrodynamic fragmentation is expected to occur for a Weber number above the value of 18 Matsuo, et al. [2008]. With a Weber number, ≈ 50 during the initial stages of quenching (for current experiments) it can be assumed that development of oscillations and hydrodynamic decomposition of the melt droplet into a few larger fragments is probable during this initial stage. However, within 45ms into the water pool, the droplet attains a Weber number range ≈7 and seems to be stable without any decomposition. A droplet elliptical in shape with a wobbling interface is observed to evolve in the water pool (Figure 3.6a depicts the simultaneous visual of the droplet dynamic by regular photography and X-ray radiography). To investigate the thermal explosion phenomena, the droplet is exposed to the pressure wave only after ≈300ms at a Weber number range of ≤ 10.

200 100 4 droplet displacement a Dragging b 150 We 80 3 of gases 100 60

a 2 We

50 40 AR Distance (mm) Distance 0 20 1

-50 0 0 0 100 200 300 0 100 200 300 Time (ms) Time (ms) Figure 3.5: Molten WO3-CaO droplet: a) displacement below water pool surface and velocity; and b) aspect ratio during quench in water pool. (Mmelt ≈1.3g, ∆Tsup ≈ 115K and ∆Tsub ≈ 79K).

At ≈ 200mm into the water pool the vapour film (black coloured region in Figure 3.6a surrounding the melt droplet (Mmelt ≈1.2 g) (red coloured region in Figure 3.6a) is destabilized by an external pressure wave (Figure 3.6b). Upon destabilization of the vapour film a partial melt-water contact is initiated at the surface of the melt droplet resulting in melt swelling and subsequent vapour film growth during a period of up to 3.2ms (Figure 3.6b-c). The over grown vapour film then collapses

44 within 0.8ms due to condensation leading to complete melt-water contact (Figure 3.6d) followed by fine fragmentation of the melt droplet and subsequent explosive evaporation of water (Figure 3.6e-f).

Figure 3.6: Simultaneous visual and X-ray radiography images recorded for a single droplet of WO3-CaO undergoing explosion in water (Mmelt ≈1.2g ∆Tsup ≈ 116K and ∆Tsub ≈ 77K). Here time, t=o corresponds to the instance of direct melt- water contact. 45

To obtain the dynamics of the convoluting melt droplet, an image processing procedure is adopted to quantify relevant information from the images acquired by the high speed visualization system (Figure 3.7). The images are acquired at a frame rate of up to 20,000 fps using the regular videography and at a frame rate of up to 8000 fps using the X-ray radiography system. The acquired sequence of raw images is treated to reduce the noise and segmented to obtain the projected area of the droplet, droplet velocity and shape. Melt droplet displacement as well as the aspect ratio of a melt droplet when exposed to pressure wave is depicted in Figure 3.7. The average velocity of the droplet is 0.56 m/s and the droplet appears to be nearly spherical prior to the disturbance by pressure wave.

A process of bubble growth and collapse can be observed in Figure 3.8a. The equivalent diameter (D) of the vapour bubble projected area shows an initial growth rate of up to 2 times the initial diameter (D0) of the droplet on submission to the pressure pulse (-2.5ms0ms) where the melt droplet is disintegrated into several micrometer sized fragments and dispersed in the water pool. An estimation of the pressure inside the vapor bubble can be made using the classic Rayleigh-Plesset equation (Equation 3.2) for bubble dynamics [Hansson, et al., 2009].

3 2 1 2𝜎 4휇푅̇ 푅푅̈ + 푅 = [(푝(푅) − 푝∞) − − ] (3.2) 2 𝜌푙 푅 푅

The work done by the expanding bubble (W) can be estimated by calculating the internal pressure using the Rayleigh-Plesset equation (Equation 3.3). The accumulative work is thus estimated as approximately 1.4J (Figure 3.7b).

푅 3 3 2 2 2𝜎 4휇 푊(푡) = 4휋휌푙 ∫ [푅 푅̈ + 푅 푅̇ + 푅 + 푅푅̇ ] 푑푅 (3.3) 푅0 2 𝜌푙 𝜌푙

Where, 푅 is the bubble radius, 푅̇ the bubble radial velocity, 푅̈ the bubble radial acceleration, 휌 the density, 휎 the surface tension and 휇 the dynamic viscosity. By considering that all the energy transferred from the drop goes into PdV type work done by the expanding bubble. The conversion ratio ( 휂(푡) ) is estimated at approximately 0.18 by Equation 3.4.

푊(푡) 휂(푡) = 0 (3.4) 퐸푚

0 Where the total available thermal energy (퐸푚) within the melt drop is estimated by:

0 퐸푚 = 푚푑(퐶푝(푇푚 − 푇푙) + 퐻푓푢푠푖표푛) (3.5)

Figure 3.7: Image processing procedure for quantification of droplet dynamics.

6 3 a b 5 2

4 (J)

cumulative D/D0

2 W 1

0 0 -4 -3 -2 -1 0 1 2 3 4 -4 -3 -2 -1 0 1 Time (ms) Time (ms) Figure 3.8: Vapour equivalent diameter history during steam explosion (a) and work done by the expanding vapour bubble (b).

An experimental quantification of the mechanisms that initiate the fine fragmentation of melt during the thermal process is far from realization considering the short time spans. However, by utilizing the unique features of simultaneous visualization capabilities at MISTEE facility it is substantiated in Hansson, et al. [2009], that the explosion energy conversion efficiency depends upon the rate at which the melt droplet can deform during the initial phase of the interaction or in other words the ability of the melt droplet to deform on exposure to a pressure wave. The insight was based on the data acquired by performing several experiments with single droplets of molten tin that underwent an explosive interaction. A rationale to the observation of melt deformation is the growth rate of instabilities on the melt 47 surface due to the dynamic pressure differences. To explore further the melt deformation mechanism during the initial phase of the interaction we compare results obtained from the experiments with a relatively high temperature binary oxide mixture of WO3-CaO at eutectic composition (75:25 mol %, Tliquidus ≈ 1135±5 °C). Here we define the melt deformation rate (MD*) as the ratio of melt projected area prior exposure to pressure wave (for example, see Figure 3.8a, t= -4.65ms) to the melt projected area during the initial phase of melt-water direct contact (for example, see Figure 3.8a, t=0ms). The specific observation is made possible by acquisition of simultaneous X-ray visuals that show the evolution of melt during the intense interactions. The effect of melt deformation rate on the estimated thermal to mechanical conversion ratio (휂(푡)) for experiments with WO3-CaO is presented in Figure 3.9. The results indicate an increased conversion ratio for conditions related to an increased melt deformation rate thereby showing a good agreement with the previous results acquired with tin.

1.2 Linear ( )

0.8 y = 0.468x - 0.6544

cycle cycle CRc (%) 0.4

nd 2

0 0 1 2 3 MD* Figure 3.9: Effect of melt deformation ratio (MD*) on the conversion ratio for experiments with binary oxide mixture of WO3-CaO melt at high water subcooling conditions (Mmelt ≈1 to 1.5g ∆Tsup ≈ 100 to 200K and ∆Tsub ≈ 75 to 85K).

The particles collected from the experiments with WO3-CaO melt experiments were sieved for size distribution by subjecting to a set of sieve screens in the range between 38µm and 1700 µm. Considering the very micron size range of the debris it was not possible to collect the debris appropriately in most of the cases. However, it was obtained for some experiments. Typical, mass averaged size distribution of fragmented debris is shown in Figure 3.10.

100 Silicon (Nelson et al., 2005) 80

60 WO3-CaO (ΔTsubcool≈ 80K, 40 ΔTsuperheat≈ 100K)

20 Tin (ΔTsubcool≈ 70K, ΔTsuperheat≈ 318K)

Cummulative mass Cummulative mass (%) fraction 0 (Abe et al., 2002) 0 400 800 1200 1600 2000 Fragment size (μm) Figure 3.10: Cumulative size distribution of solidified fragments acquired from thermal fragmentation experiments with WO3-CaO melt compared to fragment size shown in Nelson, et al. [2001] and Abe, et al. [2002].

3.2.3. Effect of melt composition on steam explosion

To verify the hypothesis put forth in Dinh, [2007] and Dinh, et al. [2008], that material composition can exhibit important differences in steam explosion intensity as a result of underlying solidification mechanisms, an experimental study is undertaken in MISTEE facility. The hypothesis delineates the physical mechanisms that govern differences in solidification mechanism for melts of eutectic vs non- eutectic composition. Given the presence of mushy phase (formation of microstructures rich in one component) for a non-eutectic composition in the cooling curve of a binary system: the notion is that formation of mushy phase can directly influence the melt deformation during the initial stages of steam explosion micro interactions leading to resilience in melt fine fragmentation. However, the eutectic composition of binary melt possesses coexistence of the solid and liquid phases of the binary components intimately mixed thereby excluding the presence of a mushy phase.

Motivated to understand the aforementioned effect of binary melt composition on the fine fragmentation behaviour, MISTEE experiments were performed with a binary oxide mixture of WO3-CaO melt at a eutectic composition (75:25 mol %, Tliquidus = 1135±5°C) and at a non-eutectic composition (67:33 mol%, Tliquidus = 1344±5°C, Tsolidus = 1135±5°C) (See, Table 3.4 for test conditions).

Snapshots acquired from experiments with WO3-CaO at a eutectic and non-eutectic composition with an initial melt superheat of 100K is shown in Figure 3.11 and 3.12. Melt-freeze-crust during the fine fragmentation of a non-eutectic composition can be clearly seen from the images (see, indication of crust in Figure 3.12) thereby reducing the melt mass that actively participated in the fine fragmentation process. This condition is observed to be active only within a melt superheat range of 100K. 49

The experiments with a high melt superheat (∆Tsup ≈ 200 K) however showed no apparent differences between the eutectic and non-eutectic tests where both the composition underwent a complete fine fragmentation. The results clearly indicate the effect of melt superheat on steam explosion of a binary oxide composition. It can also be seen in Figure 3.13 that the experiments involving a low melt superheat (∆Tsup ≈ 100 K) shows a clear divergence between the energetics and melt deformation of the eutectic and non-eutectic tests, i.e. non-eutectic tests led to a milder interaction whereas the eutectic tests regularly led to a complete fine fragmentation.

Table 3.4: Test conditions for study of binary oxide composition effects Parameter Experiment series 1 Experiment series 2

Component mixture WO3:CaO WO3:CaO Component molar fraction 75:25 mol% 67:33 mol% Melt mass (g) ≈0.8 to 1.5 (T liquidus) (°C) ≈1135±5 ≈1344±5

Melt superheat (Tmelt-T liquidus) (K) ≈100 to 315±5 Water subcooling (K) ≈75 to 85±3 Nozzle diameter (mm) ≈5

A plausible explanation for the observed differences can be related to the phase change of a non-eutectic binary oxidic melt droplet (i.e. mushy phase) that occurs at a specific temperature range determined by the solidus and liquidus line, Given a relatively high cooling rate, both due to radiative heat transfer and film boiling over a fairly prolonged period after release, as well as, the melt-liquid contact, the droplet's initial cooling during the first vapour film cycle forms a mushy layer on the non-eutectic melt droplet, whereas the eutectic melt are kept mostly in its liquid state. As one considers high cooling rates, non-equilibrium phase change might lead to differing solidification paths which will depend on the material ‘impurities’ that serve as solidification nucleation sites in an undercooled liquid, similar to what is suggested in Dinh, et al. [2008]. Investigation of non-equilibrium solidification as a mechanism that governs the eutectic/non-eutectic composition effect is outlined in Figure 3.14. While it is known that solidification is initiated as the melt undercools below the solidus temperature, this meta-stable behaviour is expected to be more pronounced in eutectic materials [Dinh, 2007]. In contrast, non-eutectic melts form a mushy phase as soon as the melt gets cooled below the liquidus temperature, since microstructures rich of one component, work as nucleation sites that initiate the solidification process.

Figure 3.11: Time sequence photography (above) and X-ray radiography (below) of a single droplet binary oxide mixture of WO3-CaO melt at eutectic composition (75:25 mol%, Mmelt

≈1g ∆Tsup ≈ 100 K and ∆Tsub ≈ 80K)

Figure 3.12: Time sequence photography (above) and X-ray radiography (below) of a single

droplet binary oxide mixture of WO3-CaO melt at non-eutectic composition (67:33

mol%, Mmelt ≈1g ∆Tsup ≈ 100 K and ∆Tsub ≈ 80K)

To understand the impact on debris microstructures, debris samples collected from MISTEE tests were subject to physical and chemical analysis by means of scanning electron microscopy (SEM). While the SEM images show certain differences, e.g., presence/absence of gas and shrinkage porosity, no significant findings about eutectic and non-eutectic fragment microstructures could be inferred from the SEM images. However, it is instructive to note that the non-conclusive debris analysis might not apply to the corium system where cooling is far more intense (see [Min, et al., 2006] [Tyrpekl, et al., 2015] for chemical analysis of corium particles). The MISTEE WO3–CaO tests are about two times lower in temperature level, and hence over an order of magnitude lower in heat flux than in reactor prototypic corium; thus with respect to corium melt undercooling, MISTEE experiments using medium- 51 temperature binary-oxidic melt suffer the combined effect of material and temperature (cooling rate) scaling. Although tentative, the experimental results are found to align with the proposition that non-eutectic melt (that forms high-viscosity mushy phase during cooling) offers a less effective interaction. The effect is not active in high-superheat tests, in which the interaction is initiated before the non- eutectic melt can be brought into a mushy state. The effect then can be observed only at a very specific narrow temperature range when the mushy phase is active in non- eutectic melt while the eutectic melt is still in liquid state.

1.2 1.2 Eutectic Eutectic 1 1 Non-Eutectic Non-Eutectic 0.8 0.8

0.6 0.6

0.4 0.4

2nd cycle CRc 2nd cycle (%) CRc 2ndcycle CRc (%) 0.2 0.2 ∆Tsup≈200 to 315K ∆Tsup ≈ 100 to 120K ∆Tsub ≈ 80K ∆Tsub ≈ 80K 0 0 1.5 2 2.5 3 3.5 0.9 1.4 1.9 2.4 2.9 MD* MD* Figure 3.13: Difference in conversion ratio as a function of melt deformation ratio with binary oxide mixture of WO3-CaO melt at Eutectic and non-Eutectic composition at varied melt superheat (∆Tsup ≈ 100 and 200K).

Figure 3.14: Binary oxide mixture of WO3-CaO melt at eutectic and non-eutectic composition undergoing non-equilibrium solidification.

3.3. MISTEE-HT study of steam explosion

3.3.1. MISTEE-HT facility

To perform experiments with high temperature melts, a series of modifications were done in the MISTEE facility. For the purpose of operation at high temperature, various designs of the facility (called MISTEE-HT) and feasibility studies of prototypes have been conducted for high-temperature melt preparation and melt droplet delivery. After a series of testing, qualification/calibration of the designs and prototypes, which were necessary to develop the infrastructure, the furnace developed and utilized in the previous MISTEE experiments is replaced with a completely new design. A new methodology is developed to generate single droplets of molten materials at high temperature (˃1300°C).

Table 3.5: Methods used for high temperature melt droplet generation and examples of their application.

Melting Melt arrest Temp: Major Examples of Drawbacks Reference method method range advantage application

Ultra-high Aerodynamic temperatures levitation Material [Winborne, Laser beam No discharge (2000°C or using property et al., 1976] heating capability above) can be diverging measurements easily conical nozzle realized Fragmentation of dense liquid Material Laser beam Aero-Acoustic [Weber, et drops, No property heating levitation 2000°C al., 1994] discharge measurements or above capability Ability to levitate large Material [Garnier & Induction Gas-film No discharge mass property Portard, heating levitation capability samples measurements 1987] (upto 200g) [Germany Induction Electromagne Restricted to Patent No: heating tic levitation metals 42204, 1923] Corrosion, uncertain welding of [Haraldsson, Induction Pneumatic Upto FCI droplet plug and 2000], heating plug 1500°C experiments delivery at crucible) high temp.

While droplet generation techniques have advanced significantly for cold fluids (for examples see, [Fang, et al., 2008] and [Luo, et al., 2012]), generation of high temperature melt droplets, pose several challenges. However, several improvements have been achieved in generating melt droplets within a temperature range of 1000°C (for example see [Rumschoettel, et al., 2017]). Choice of heating techniques, crucible materials, aggressive chemical reactions, wettability, orifice design, ejection of the droplet in the intended trajectory etc. are some of the major complexities associated with the generation of ultra-high temperature melt droplets. Some methodologies are available to prepare ultra-high temperature melt droplets while being chemically inert. The methodology is also successfully implemented in various research applications. For example, aerodynamic levitation of melt droplet by inert gas purge and gas film levitation methods are utilized for thermophysical property measurements (see for example [Grischenko & Piluso, 2011]). However, the available design methodologies make ejection of droplets to an intended space practically impossible. The most conventional technique used for melting the sample is laser heating since it provides control over the energy input spatially and temporally [Bottomley, et al., 2015]. A summary of the available methods used for high temperature melt droplet generation and examples of their application is summarized in Table 3.5. One of the widely used methodology for ejection of single melt droplets is the pneumatically operated plug ejection methodology. The

54 methodology has been successfully adapted for a temperature range of 1400°C in several research facilities including MISTEE-LT. However, the disadvantage of the technique is hot welding of the plug and crucible at high melt temperatures. For controlled melt generation and discharge to intended space a methodology of aerodynamically plugging the droplet by constant purge of inert gas is developed while retaining induction heating of crucible. After a series of proof of concept tests with metal melts and oxidic melts, the sophisticated design is developed to generate high temperature melts. Figure 3.15 presents the scheme of aerodynamic plugging methodology adapted in MISTEE-HT facility. The melting chamber is a high frequency induction furnace (20kW, 50 to 250 kHz). Concentric tubes of advanced ceramics (porous zirconia, alumina and magnesia) surround the tungsten crucible to minimize heat losses. The crucible temperature is monitored by the use of a C- type tungsten-rhenium thermocouple mounted from the side of the crucible. The melt prepared in the tungsten crucible is aerodynamically plugged by constant purge of inert gas through the crucible nozzle. An example of aerodynamically plugged melt droplet is shown in Figure 3.16. A fast acting three-way valve is used to control open/close operation for gas flow through the nozzle and also to isolate the furnace from the water pool chamber. On achieving the desired temperatures in the crucible, the fast acting three-way valve is released thereby discharging the melt droplet into the water pool chamber. The droplet temperature before entering the water pool is monitored by a fast two-color pyrometer with a spectral range between 1.4 and 1.8 µm and uncertainties on the temperature measurements around ±12°C. The water pool chamber is a rectangular plexi-glass tank (180 mm*130 mm* 250 mm). K-type thermocouples are used to monitor water temperature at two levels. The triggering system is retained from the MISTEE-LT facility.

Figure 3.15: Schematic of the MISTEE-HT test facility (above) and a snapshot of the facility (below)

Figure 3.16: Top-view of melt droplet aerodynamically plugged inside crucible by constant purge of inert gas in MISTEE-HT furnace.

3.3.2. Steam explosion experiments with alumina melt

A significant number of out-of-pile steam explosion experiments at a relatively large- scale have been conducted with various corium simulant and prototypic materials with an aim to discern the thermal-hydraulic process and assess steam explosion as a safety issue in nuclear power plants. Al2O3 [Huhtniemi, et al., 1999], Fe3O4 [Buxton & Benedick., 1978], a mixture of Al2O3 and Fe components [Cherdron, et al., 2005], ZrO2 [Song, et al., 2002], a mixture of UO2, ZrO2, Zr, SS components [Kim, et al., 2008], Zr- ZrO2, Zr-SS components [Cho, et al., 1988], UO2, ZrO2, Zr, U, Fe2O3 [Hong, et al., 2013] etc. have been used in experimental programs. The experimental works have shown significant scatter in data on the triggerability and explosion yield for various kinds of simulant and prototypic materials used in the experiment. In general, it seems that single component oxides or metal melts are prone to energetic explosions compared to that of a binary or multi-component system. Pondering over the experimental data leads us to assume that material specific properties are responsible for the observed differences in triggerability and explosivity. Such an assumption stems for example from the differences observed between Al2O3 and UO2-ZrO2 in KROTOS experimental program, leading to the largely discussed “material effect” research on steam explosion.

In KROTOS tests [Huhtniemi, et al., 1999], Al2O3 showed characteristics of undergoing spontaneous steam explosion at subcooled conditions independent of melt superheat. The tests performed at saturated conditions had a suppression effect on the spontaneous trigger ability, however an external initiating event (triggering) was sufficient to result in an energetic interaction. Of all the tests with prototypic corium melt compositions; there was no indication for an occurrence of spontaneous steam explosion. An external trigger was required to initiate the event. The triggered interactions however were mild compared to that of the tests with Al2O3. Some major observations from KROTOS experimental program is summarized in Table 3.6. In general, the explosion conversion ratio was an order of magnitude higher for tests with aluminum oxide melts compared to that of a corium composition (Figure 3.17). 57

Table 3.6: Parametric effects observed in KROTOS tests Material Experimental Explosion characteristic CR (%) conditions Al2O3 High water subcooling Spontaneous energetic interactions Up to 2.5

Low water subcooling Explosion trigger was required. Resulted in energetic interactions. Melt super heat (150K to Independent of melt superheat the 750K) interactions were energetic. Elevated pressure (0.1 to Independent of initial pressure the 0.375 MPa) explosions were energetic. No suppression. UO2-ZrO2 Water subcooling Explosion trigger was required. The Up to 0.15 interactions were not as energetic as for Al2O3 melt. Elevated initial pressure Comparatively mild explosions only (0.2 to 0.4 MPa) after triggering externally.

Figure 3.17: Energy conversion ratio as a function water subcooling for corium and alumina melts from KROTOS experiments [Ursic, et al., 2012].

It is now largely admitted that the difference in steam explosion intensity and characteristic is a result of difference in premixing characteristics between the materials. Visual observations from KROTOS tests have shown that the highly dense corium melts generally produced smaller fragments than the less dense alumina melts. The subsequent variations in solidification trends between molten fragments of semi-transparent alumina and opaque corium [Dombrovsky & Dinh., 2008] and level of integral void fraction in the pre-mixture region are described as major influential parameters [Ursic, et al., 2012], [Meignen, et al., 2014]. It is also considered that, aluminum oxide melts exhibit an ability to dissolve steam and thereby establish direct melt-coolant contact resulting in spontaneous triggering of steam explosion [Piluso, et al., 2005]. In general, the experimental results have shown peculiarity of aluminum oxide melt both physically and chemically, which

58 directly influences the steam explosion characteristic. A number of long-standing hypothesis thus supports a complex issue as material effects on steam explosion.

To complement the investigation of material effects on steam explosion behaviour, droplet scale experiments with Al2O3 melt is considered (see, Table 3.7. for test conditions). The ability of Al2O3 melt to undergo thermal fragmentation and trigger mechanism involved in fine fragmentation is thus investigated in MISTEE-HT facility. Within the scope, the role of melt superheat and water subcooling on the triggerability and fine fragmentation is pursued.

Table 3.7: Test conditions for MISTEE-HT Al2O3 tests.

Component Al2O3 Melt mass (g) ≈0.8 to 1.5 Melting temperature (Tmelt) (°C) ≈2054±6 1 Melt superheat (K) ≈60 to 250±12 Water subcooling (K) ≈10 to 85±3 Nozzle diameter (mm) ≈5 1Referenced from (Schneider S. , 1979)

An example of the droplet and surrounding vapour evolution during the experiment is shown in Figure 3.18. A first visual analysis of the fine fragmentation process at high melt temperatures (˃2000°C) show remarkable similarities in behaviour to what was observed for other materials previously studied in MISTEE-LT and also consistent with the observations of Nelson & Duda, [1981] and Nelson & Duda, [1982]. After meeting the pressure wave, the droplet and the surrounding vapour exhibits two or three successive cycles of isotropic expansion and collapse. In what follows after time t=0 ms corresponds to the beginning of the second cycle which is the most important and expected to be the beginning of the fine fragmentation process.

Within the scope, the effect of melt superheat is considered. The results do not seem to show any specific trend (Figure 3.19). The superheat seemed to be always sufficient to initiate fine fragmentation of alumina melt. For all those cases, the conversion ratio evaluated from Equation 3.4, by considering the maximum equivalent diameter during the second expansion/collapsing cycle ranges between 0.1% and 0.7% for a water subcooling range of 90 to 45K. 59

Figure 3.18: Time sequence video images recorded for a single droplet of Al2O3 undergoing explosion in water (Mmelt ≈0.8g ∆Tsup ≈ 163K and ∆Tsub ≈ 82K).

cycle cycle (%) CRc

nd 2

0 20 70 120 170 220

∆Tsuperheat (K) Figure 3.19: Effect of melt superheat on the conversion ratio for experiments with Al2O3 melt at high and moderate water subcooling conditions (Mmelt ≈1g and ∆Tsub ≈ 55 to 85K).

However, spontaneous steam explosion was observed at a higher melt superheat of 215K and a water subcooling of 81K. Figure 3.20, shows the time sequence snapshots of the explosion given by the camera. The steam explosion seemed energetic compared to what was observed for the triggered explosion cases. Since the fragmentation process is not entirely captured by the camera, it is not possible to evaluate the equivalent bubble diameter after a given time and the conversion ratio. The equivalent diameter can be assessed approximately up to the middle of the second cycle. Compared with other triggered cases with similar subcooling and melt charge conditions, the equivalent bubble diameter is increased by nearly 42% when the explosion occurs. This spontaneously exploding case shows that alumina can exhibit spontaneous steam explosion under certain conditions, and thereby strengthens the hypothesis of thermal explosion as an initiating event.

Figure 3.20: Spontaneous steam explosion observed for the self-triggering case (Mmelt ≈1g, ∆Tsup ≈ 215K and ∆Tsub ≈ 81K).

The mass median fragment size acquired from the alumina tests that exploded and at high subcooling conditions is in the range between 100 and 200µm. It is also worth noting that, compared with the other materials previously studied in MISTEE, alumina gives the smallest fragment size. For instance, for eutectic WO3-CaO, D50 is approximately 250 µm. This smaller size observed for the fragments with alumina is the result of a more energetic explosion. It is likely to be due to the higher temperature, but the different material properties also play a role.

Water subcooling strongly influences the vapour production in the droplet vicinity. It is clear from Figure 3.21 that low subcooling leads to a higher aspect ratio as a vapour wake appears at the rear of the droplet. The vapour film and wake can be identified around the droplet. Due to hydrodynamic forces, the wake breaks up and a detached bubble can be observed (see the second image from the figure). The size of this detached bubble decreases as it is condensed (from the second image to the fourth). As the detached bubble nearly disappears before leaving the frame of the camera, it seems that it is only made of vapour. Thus, at the observation position, there seems to be no more of the air entrapped by the alumina droplet. The larger aspect ratio under low subcooling conditions also signifies that a larger volume of vapour enshrouds the melt droplet. Therefore, the vapour film is more stable. It explains the suppression of steam explosion observed for those conditions. Therefore strengthening the notion that low water subcooling has a suppression effect on steam explosion observed for example in Corradini [1981] and [Hohmann, et al., 1995].

Bubble Condensing detachment bubble Figure 3.21: Time sequence snapshots depicting vapour film evolution under low subcooling conditions (Mmelt ≈0.9g ∆Tsup ≈ 116K and ∆Tsub ≈ 14K).

The volume of void produced for low subcooling cases can be assessed from the void area projected on the camera images. The evaluation is made by assuming that the droplet and attached vapour wake form an ellipsoid. The global volume is thus calculated using the maximum vertical and maximum horizontal dimensions of the droplet and surrounding vapour shape. The volume of the alumina droplet alone is similarly evaluated in order to get the actual volume of vapour. With this method, the void volume can be expected to be overestimated as the droplet and wake shape is not really elliptic (Figure 3.21). Nevertheless, it

62 is a simple way to give a rough estimation of the volume of void produced by film boiling for low water subcooling.

The ratio between the volume of void and the volume of the alumina droplet is presented over time in Figure 3.22, for two low subcooling cases. The volume of void produced by the film boiling around the hot alumina droplet is quite substantial. It varies between twice and more than eight times the volume of the droplet. Over time, the void volume periodically increases as the wake is formed and suddenly decreases as bubbles are detached. The detachment frequency is nearly constant and around 35 ms. For the case with water subcooling 26K (blue curve on Figure 3.22), the droplet was subjected to the pressure wave. While no steam explosion was observed under those low subcooling conditions, variations of the calculated volume can be observed. These oscillations come from the method used to calculate the void volume since the vapour film and wake are perturbed and subsequently deformed by the pressure wave.

16 ∆Tsubcool: 14 K ∆Tsubcool: 26 K 12

Vvoid/Vdrop 4

0 346 376 406 436 466 496 526 Time (ms) Figure 3.22: Time resolved equivalent vapour evolution.

Experiments with alumina melt in MISTEE-HT facility is the first step in achieving high temperature single droplet steam explosion experiments thereby reducing the gap between simulant and prototypic materials. The steam explosion suppression effect of low water subcooling is thereby also substantiated. MISTEE-HT experiments with alumina melt, confirm the mechanism of spontaneous steam explosion observed in KROTOS tests. MISTEE experimental data are useful for validation of numerical models on thermal fragmentation phenomena (An example of which is shown in Paper IV). Progressing towards understanding the so-called “material effect” phenomena in steam explosion, experimental study in MISTEE-HT facility with prototypic corium compositions are necessary to understand the differences in steam explosion behaviour between simulant and prototypic materials.

4. Oxidation of metallic melt

4.1. Introduction

Zircaloy oxidation, which produces hydrogen and chemical heat, is encountered either during core degradation when steam flows through the hot cladding surfaces of fuel rods or in a molten fuel coolant interaction (FCI) when molten corium (mixture of oxides and metals) falls into a water pool (e.g., in the lower head). While extensive investigations have been carried out for the former, little work has been done for the latter. This motivates the present study, i.e., to understand and characterize zirconium oxidation phenomena that occur during FCIs and can affect steam explosion risk and energy.

In general, the metal can react chemically with water vapor exothermically producing Zirconium oxide and hydrogen (Equation 4.1) Zr+ 2H2O(g)→ ZrO2+ 2H2(g) + 6.47 MJ/kg Zr at 1900 °C) (4.1) Melt oxidation during FCI can be considered as a major source for the void produced in the pre-mixture. It can also be expected that the change of metallic melt composition due to increase in oxygen content can affect the physical-chemical behavior of melt during crystallization process. Further, generation of non- condensable gas (hydrogen) that envelops the melt droplet provides greater stability against direct melt-water contact. It is known that the non-condensable gases in the vapor film shall suppress the condensation of vapor, which is an essential parameter for the occurrence of a steam explosion.

While the above mentioned can be considered as steam explosion limiting effects of melt oxidation, some dissenting aspects of melt oxidation can also be mentioned which includes build-up of an explosive atmosphere through excess hydrogen production, increased melt thermal energy content and delayed solidification through heat release. These are important mechanisms that contribute in enhancing the steam explosion strength [Meignen, et al., 2014], [Meignen, et al., 2014b], [Leskovar, et al., 2016]. Apart from such logical propositions, a mechanistic understanding of melt oxidation behavior and influence during the course of a FCI is rather abstruse.

The most comprehensive and systematic experimental work to date on zirconium oxidation during FCI was performed at Argonne National Laboratory (ANL) [Cho, et al., 1998]. Approximately 1 kg of Zr-ZrO2 or Zr-SS melt mixtures were poured into subcooled water to assess the extent of chemical augmentation during a FCI. The major experimental parameters were water subcooling and the zirconium content in the melt. Results show that the rate of oxidation is independent of the zirconium content added to the melt while the rate of oxidation increased at a lower water 65 subcooling. For explosive cases, the rate of oxidation was estimated up to 99% and the energetic of the explosion was observed to increase with added Zirconium content indicating the vital role of chemical augmentation on steam explosion energetics (Table 4.1).

Interpretation of results on chemical augmentation from TROI and KROTOS tests performed with added zirconium metal to prototypic materials are not straight forward considering the stochastic integral effects of a steam explosion. Generally speaking, contradicting results were obtained where the experiments with added zirconium content (UO2-ZrO2-Zr) results in the least energetics recorded in the whole series of experiments conducted in the frame of OECD SERENA project [Leskovar, et al., 2016]. Understandably, one may not assume the low energetics can be related to chemical augmentation considering several integral effects associated with a phenomenon as FCI. Further, FARO experiment (L11) with added Zr melt showed that agglomeration of melt is not observed whereas a repetitive test without Zr under otherwise almost similar conditions (L14) has shown agglomeration of 20% of initial melt mass indicating the effect on melt fragmentation and debris formation.

Table 4.1: Experimental works on the effect of chemical augmentation on FCI

Experimental Results Facility Material conditions Rate of oxidation was independent of Zr content Zr content (0 to 100 (≈1 to 2 moles of hydrogen production for all ZREX, ZRSS wt.%) cases) premixing Rate of oxidation decreased with increase in SS/Zr and [Cho, et al., subcooling ZrO2/Zr 1998] a. ≈1 to 2 moles of hydrogen at 80K and 4 to 5 moles of hydrogen at 30 to 20K b. Up to 15% oxidation at 80K and up to Subcooling (≈80K) 25% oxidation at 10K water subcooling ZREX , ZRSS Conversion ratio increased with increase in Zr explosion [Cho, SS/Zr and content (≈0.5% at 0 wt.% Zr content and ≈3.2% et al., 1998] ZrO2/Zr Zr content (0 to 100 at 100 wt.% Zr content ) wt.%) No agglomeration of melt with added FARO Saturated water, 76.2 zirconium. UO2- [Annunziato, et wt.% UO2, 19.2 wt. ≈3.2 g of hydrogen produced per kg of ZrO2-Zr al., 2000] %ZrO2 and 4.6 wt. % fragmented melt, which is the highest of all Zr FARO tests. 62.8 wt.% UO2, 13.5 wt. %ZrO2, 12.6 wt. % Explosivity of partially oxidized corium seemed UO2- TROI [Kim, et Zr & 11.1% SS and 61 higher than oxidized corium. However, one of ZrO2-Zr- al., 2008] wt.% UO2, 16 wt. the two tests did not undergo steam explosion SS %ZrO2, 12.2 wt. % Zr even when triggered. & 10.8% SS

Zr was completely oxidized (no traces of KROTOS UO2- 80.1 wt.% UO2, 11.4 wt.% ZrO2, 8.5 wt.% metallic Zr in debris). ≈4.7 g of hydrogen [NEA., 2014] ZrO2-Zr Zr, produced per kg of fragmented melt.

Apart from the integral experiments, small-scale separate effect tests with zirconium in molten state at the scale of droplets is not available elsewhere. However, a series of experiments were performed by Baker.L., & Just.L., [1962], by pulse electrical heating with the condenser-discharge method where the energy stored in the condensers were delivered to a circuit consisting of specimen zirconium wires (25mm long and a diameter of 0.76 and 1.52 mm respectively) immersed under water at room temperature. The melt temperature is estimated based on electrical measurements that determined the energy input per gram of the specimen assuming that uniform adiabatic heating had occurred. It is shown that the initial melt temperature had a significant effect on oxidation rate. The runs involving temperatures above the melting point resulted in breakup of the wire into discrete spherical particles of varied size. Hydrogen generated from the reaction was collected and the reaction rates were estimated. Further at a temperature range above 2600°C, ignition with explosive pressure rise was observed. Oxidation of up to 18% is estimated for non-ignition cases for room temperature water. The cases that underwent ignition however had a reaction rate (mass% of Zr that was oxidized to form ZrO2) up to 70%. Considering the scarcity of experimental data on Zr melt oxidation, the experiments reported in Baker.L., & Just.L., [1962], will be used as reference for our current study.

There is scarcity of experimental data on the oxidation of molten zirconium while quenching under water. New data at the scale of droplets is needed on to improve our understanding of oxidation phenomena and improvement of oxidation models implemented in FCI codes. A recent review on the status of steam explosion understanding and modeling [Meignen, et al., 2014] has also indicated the vitality of performing small scale experiments with liquid metal melts at very high temperatures to investigate oxidation phenomena.

4.2. MISTEE-OX droplet oxidation experiment

MISTEE-OX study is the first step in advancing our current state of knowledge on corium melt oxidation behavior during FCI. MISTEE-HT facility with modifications is thus utilized to generate single droplets of pure zirconium melt followed by rapid quenching in subcooled water (see, Figure 4.1). The crucible is made of graphite with a hole in the center of the bottom. Zirconium melt is suspended inside the crucible by constant purge of argon gas through the crucible nozzle. The temperature of the crucible is measured by a C-type thermocouple (measurement uncertainty of ±1%) 67 mounted from a hole on the crucible wall. A three-way valve isolates the furnace from the water pool. A longer rectangular plexi-glass tank (500mm×75mm×75mm) is arranged below the furnace. On attaining the desired melt temperature, the melt droplet is allowed to free-fall by opening the passage to the water pool while simultaneously arresting gas purge into the crucible. The entire mass loaded in the crucible is not delivered depending if some material is lost in the furnace walls. A high speed camera (Red lake HG50LE) with acquisition rate up to 5000fps is used to record the interaction between melt and water. The melt droplet is captured in a debris catcher suspended inside the water pool. Two sets of experiments were conducted to study the influence of water subcooling on the oxidation behavior of zirconium melt during quenching in water (see, Table 4.2).

Figure 4.1: Schematic of the MISTEE-HT test facility with modifications for single droplet oxidation studies.

Table 4.2: Conditions for MISTEE-OX droplet tests. Temperature of Melt Water Initial mass Water pool ID # melt in the superheat subcooling, of Zr melt, g temperature, °C crucible, °C (∆Tsup), K (∆Tsub), K HSUB run 1 HSUB run 2 15 85

HSUB run 3 2 2005 150 MSUB run 1 55 45 MSUB run 2

4.3. Experimental results

4.3.1. Bubble analysis

Single droplets of molten zirconium metal quenched in water depict a pattern of repetitive bubble growth and detachment along the zirconium melt droplet rear periphery (Figure 4.2). The trail of bubbles left behind the droplet trajectory do not seem to condense even under the high subcooling conditions indicating that the bubbles are mostly filled with non-condensable gas (hydrogen). A first-hand look at the recorded images suggests a number of "cyclic oxidation processes" of heat/hydrogen generation and subsequent bubble detachments from the drop during its free-fall in water. The visuals clearly indicate a very specific pattern related to zirconium oxidation. Initially, the flow of water vapor around the melt, oxidizes the melt instantly releasing heat and hydrogen (inferred by the melt glow at t=90ms,235ms, 365ms in Figure 4.2). Growth of the bubble filled with non- condensable gas reduces steam supply to the drop surface and melt oxidation/hydrogen generation (inferred by the decrease in melt glow at t=159ms,329ms in Figure 4.2). On reaching the threshold gas bubble size for the current drop velocity in water, size, the gas bubble detaches from the droplet rear, allowing fresh water vapor access to the drop surface. Based on the discussion above, it would be logical to mention here that the melt shall dynamically increase in temperature as a result of heat released due to oxidation while also being simultaneously quenched. The specific observations provide valuable insight that the oxidation process does not stop as soon as the bubble is filled with hydrogen restricting access to water vapor but the hydrogen bubble detaches allowing continuous but periodic flow of fresh water vapor.

Snapshots depicting the progress in interaction between water and zirconium melt under medium water subcooling conditions (MSUB) is shown in Figure 4.3. A long trail of non-condensable gas bubbles in the trajectory of the droplet can be noted similar to that of the high subcooling conditions. Typical velocity profile of the melt 69 droplets during quenching under high and medium water subcooling is shown in Figure 4.4. A significant difference in the velocity of the droplets can be observed at the varied water subcooling conditions. Under high subcooling conditions the melt droplet reaches the catcher arrangement much more rapidly compared to that of the medium subcooling conditions.

Figure 4.2: Time sequence raw snapshots depicting interaction between single droplet of zirconium melt and water at ∆Tsub ≈85K (HSUB run 3).

Figure 4.3: Time sequence raw snapshots depicting interaction between single droplet of zirconium melt and water at ∆Tsub ≈45K (MSUB run 2).

0.60

0.55

0.50

0.45

0.40

0.35 ∆Tsub:45K, ∆Tsup:150K Melt droplet velocity (m/s) (m/s) velocity droplet Melt ∆Tsub:85K, ∆Tsup:150K 0.30 0.100 0.200 0.300 0.400 0.500 Time (s) Figure 4.4: Temporal velocity profile of the melt droplet in water at varied degree of water subcooling.

To estimate the hydrogen mass released from the exothermic chemical reaction, we adapt a simple methodology where we study the gaseous wake in the trail of the melt droplet. By making a very simplified assumption that the water vapor in the bubbles detached from the melt droplet condense completely when reaching the surface layer of the water pool and the bubble contains purely hydrogen at STP, we estimate the volume of hydrogen generated during the interaction at approximately 50mm below the water level. However, it is instructive to note that the resolution of the high speed visualization system which is the function of the number of pixels in the sensor is inversely proportional to the recording speed and hence the recorded video is limited to approximately 5s. Further high speed imaging usually suffers from low light collectiveness. Combined with the necessity to neutralize light emitted from the high temperature melt, the images become dark when the light intensity reduces from the melt droplet as a result of heat loss. Digital image analysis is employed to extract the bubble parameters from the acquired sequence of raw images. As a general scheme, a series of procedures including background subtraction, noise reduction, edge detection, morphological opening is done to obtain the projected area of the bubbles (Figure 4.5). Based on the assumptions, the kinetics of hydrogen released during the interaction is estimated for the cases suitable for analysis (Figure 4.6). The time scale in Figure 4.6 corresponds to the time of bubble detachment from the droplet. Based on the bubble analysis, the O/Zr atomic ratio is estimated to be approximately up to 0.3 under medium subcooling conditions and approximately up to 0.07 under high subcooling conditions (O/Zr atomic rati0 equal to 2 corresponds to fully oxidized zirconium – ZrO2) (Figure 4.7). A significant increase in oxidation under medium subcooling conditions can be clearly observed.

Figure 4.5: Image processing methodology for the estimation of the temporal distribution of hydrogen (left). Labels 1 and 2 depict typical individual hydrogen bubbles (right)

90 80

(mL) 70 Zr HSUB Run 1 2 (∆Tsub:85K, ∆Tsup:150K) 60 50 Zr HSUB Run 3 40 (∆Tsub:85K, ∆Tsup:150K) 30 20 Zr MSUB Run 2 (∆Tsub:45K, ∆Tsup:150K ) Cumulative volume of H of volume Cumulative 10 0 0 1 2 3 4 Time (s) Figure 4.6: Hydrogen generation history during Zr melt and water interactions.

0.35

0.3 Zr HSUB Run 1 (∆Tsub:85K, ∆Tsup:150K) 0.25

0.2 Zr HSUB Run 3 0.15 (∆Tsub:85K, ∆Tsup:150K)

0.1

O/Zr atomic ratio ratio atomic O/Zr Zr MSUB Run 2 0.05 (∆Tsub:45K, ∆Tsup:150K )

0 0 1 2 3 4 Time (s) Figure 4.7: Evolution of oxygen content in melt drop estimated from hydrogen generation data.

4.3.2. SEM/EDS analysis

SEM images depicting the microstructure of solidified zirconium melt droplet under high water subcooling (HSUB run 1-3) conditions are presented in Figure 4.8. The scanning electron microscope (Hitachi S-3700N) combined with Energy-dispersive X-ray spectroscopy (Bruker`s XFlash Detector 4010) technique is used to perform electron-excited X-ray microanalysis of the solidified droplet. The analysis reveals stratified layers of varying oxygen content. Visual analysis of presented microstructures allows us to see oxygen-enriched (as lighter) and metal-enriched (as darker) areas. Outer layer (k-area: crust region/melt-vapour interface) can be clearly observed as the most oxygen rich region (see, for example, label a2 in Figure 4.8). Further, in the bulk towards the centre of the droplet, the concentration of oxygen can be observed to decrease. Figure 4.9 depicts the radial stratification in O/Zr atomic ratio through the depth of the droplet for all the cases studied under high subcooling conditions. The entire region of the droplet is neither fully oxidized nor un-oxidized. Three or four distinct concentric layers with varied concentration of oxygen can be identified in the droplet. The O/Zr atomic ratio values were obtained by the scan of a relatively larger area. For example, for layers k, l and m (in case of HSUB run 1), scanning areas correspond to thickness of the respective layers. Length of scanning area was in the range 0.5-1.5 mm for different layers. Layers boundaries were determined visually, based on the change of microstructural patterns exhibited within each layer of the droplet (see areas k, l, m and n under label a1, b1 and c1 in Figure 4.8). The EDS technique employed cannot detect the difference in O/Zr atomic ratio over the respective layer and hence can be assumed that O/Zr atomic ratio is evenly distributed in the specific layer.

O/Zr atomic ratio at the outer surface estimated by EDS is in the range of approximately 0.8 and 0.9 (Figure 4.9). The outer surface with the highest concentration of oxygen is found to have a constant thickness in the range of approximately 200 to 300µm constantly distributed through the entire surface area of the droplet. The innermost layer of the droplet is almost evenly oxidized where the O/Zr atomic ratio is approximately 0.6. However it can be observed for case HSUB#1, that the innermost layer shows a significantly reduced oxygen concentration where the O/Zr atomic ratio is approximately 0.2. The SEM image for case HSUB#1, shows a well formed crack at the outer surface (see label a1 in Figure 4.8) which probably could have acted as a mechanical resistance for oxygen migration towards the inner surfaces.

SEM images depicting the microstructure of solidified zirconium melt droplet under medium water subcooling (MSUB run 1-2) conditions are presented in Figure 4.10. On the first look, it is quite obvious that the shape of the droplets show irregular morphology compared to the cases of high subcooling conditions where the droplets resemble a morphology close to that of a spheroid. The observations indicate that the droplets could have still been liquid under medium subcooling conditions that allow spreading of the melt in the catcher compared to that of the droplet under high subcooling condition where the droplet could have already formed a crust limiting the shape to that of a spheroid. The images depict a similar pattern of melt oxidation where the outermost layer shows the highest degree of oxygen concentration in the droplet. However, the thickness of this outer layer is not constant as the case was for high subcooling conditions. The thickness of the layer ranges between 180 and 320µm. Similar to high subcooling cases, the oxygen content is stratified from the outer surface to towards the center of the droplet (see label a, b, c for cases MSUB Run 1 and MSUB Run 2 in Figure 4.10). However, formation of concentric evenly distributed layers through the entire surface of the droplet is not observed as was the case for high subcooling conditions. O/Zr atomic ratio at the outermost surface is approximately in the range between 1.1 and 1.5 which is almost up to 75% oxidation compared to high subcooling conditions where it is up to 45% oxidation. It can also be noted that the O/Zr atomic ratio on the inner most surface is in the range between 0.85 and 1 which is almost up to 50% oxidation for medium subcooling conditions compared to high subcooling conditions where the O/Zr atomic ratio is up to 0.7 which is almost up to 35% oxidation.

It can be clearly observed that there is a significant difference between bubbles hydrogen analysis and SEM/EDS analysis. SEM/EDS analysis shows increased oxidation compared to that of hydrogen analysis. It is instructive to note that the results obtained from SEM/EDS should be treated with caution considering the uncertainties involved in estimating a light element as oxygen. Use of other techniques/measurements is required for an accurate estimation. It should also be

74 noted that the use of SEM/EDS in corium analysis it is shown that the uncertainty in quantifying oxygen is not random rather systematic. Although SEM/EDS can have significant uncertainties, the method can provide reliable information on relative oxygen concentration.

Figure 4.8: SEM images showing microstructure of solidified Zr-droplet cross section for all cases studied under high subcooling conditions.

2 O/Zr atomic ratio 1.8 HSUB Run 1 1.6 1.4 1.2 k 1 l 0.8 m 0.6 0.4 0.2 n 0 0 200 400 600 800 1000 1200 1400 Distance from surface (µm)

2 O/Zr atomic ratio 1.8 HSUB Run 3 1.6 1.4 1.2 1 k l 0.8 m 0.6 0.4 0.2 0 0 200 400 600 800 1000 1200 1400 Distance from surface (µm) Figure 4.9: Oxygen concentration profile from the surface to the bulk of the quenched drops after the tests at high subcooling.

Figure 4.10: SEM images showing microstructure of solidified Zr-droplet cross section for all cases studied under medium subcooling conditions. 79

4.4. Discussions

Underwater molten Zr droplets oxidation is a highly non-equilibrium process, therefore phase transitions taking place during oxidation process can be characterized only partially based on the Zr-O phase diagram (Figure 4.11, phase diagram adapted from [Massalski, 1990]). The typical observed zones across droplets cross-section are shown in Figure 4.12. There are two characteristic zones, which can be observed in the crust region and bulk region as well: i) oxygen-enriched areas (H1, H2, M1, M2) and ii) zirconium-enriched areas (H3-H5, M3, M4). Compositions of areas enriched by oxygen for all studied samples (see for example areas H1 and H2 for HSUB-1 or M1 and M2 for MSUB-1 - Figure 4.12) belong to two- phase region of the Zr-O phase diagram, which includes α-Zr(O) (with maximal content of dissolved oxygen) and ZrO2-x as equilibrium phases. But observed areas look like one-phase region, have uniform microstructure and any evidences of phase difference were not detected. This is typical for all studied samples, which have oxygen-enriched areas covering a range of 46-60 atomic percent of oxygen. Two different phases were detected for Zr-enriched areas (see for example areas H3 and H4 for HSUB-1 or M3 and M4 for MSUB-1 - Figure 4.13), what can be explained by formation of α-Zr(O) and β-Zr(O) solid solutions. Maximal registered content of oxygen dissolved in α-Zr(O) phase was approximately 15 at. %.

As mentioned earlier, oxidation (O/Zr atomic ratio) of Zr-droplet is quite uniform for the whole volume (Figure 4.10). It means that oxygen mass transfer from the surface to the bulk volume is very fast, which is typical for mass transfer in liquid state. In the case of liquid droplet, oxygen transfer is not limited by diffusion through the oxidized-Zr layer like in case of Zr solid material oxidation where formation of ZrO2 surface layer can be clearly observed (see, for example [Yang, et al., 2013]). Based on the EDS analysis, the total oxygen concentration of the obtained droplet is about 40-60 at. %. According to phase diagram (Figure 4.11), this composition region can be reached by two ways: 1) Below the solidus line by increase of oxygen content and undergoing following phase transitions in solid state: β-Zr(O) → α-Zr(O) → (α-Zr(O) + ZrO2) 2) Above the liquidus line by increase in oxygen content and quenching of liquid phase with corresponding composition: Liquid → (α-Zr(O) + ZrO2) Due to very short time (few seconds) when Zr droplet, in principle, is available for oxidation it is extremely unlikely that oxidation could have occurred in solid phase, according to reasons were mentioned above. Therefore, it is reasonable to assume, that oxidation occurred only in liquid state. It means, that as a result of oxidation reaction the temperature of melt could have increased simultaneously with the increase of oxygen concentration in the droplet before the actual rapid solidification. Observed oxygen concentration shows that the melt superheat could have potentially increased by approximately 200 K under high subcooling conditions. The

80 oxygen concentration on the inner surface of the droplet near the center (core) is approximately in the range up to 40 at.% O. indicating the core of the droplet could have remained at a relatively lower temperature compared to that of the crust region. Similarly, under medium subcooling conditions, maximum oxygen content observed in the crust region is in the range up to approximately 60 at.% O. indicating that the melt superheat could have potentially increased by approximately 400 K. The oxygen concentration on the inner surface of the droplet near the center (core) is approximately in the range up to 50 at.% O. indicating a similar pattern to high subcooling conditions. The aforementioned hypothesis is a rather simplified thought; however the actual mechanisms can be far more complex and requires further research.

It is important to note, that for HSUB samples, probably, whole oxidation process (oxygen transfer for almost 3mm droplet radius) occurred during the droplet fall in the water pool, i.e. ~0.5 s. This process needs to consider kinetic of droplet oxidation, but it will add additional complicity for experiment procedure – relatively precise temperature measurement, what is not trivial part for high temperature and high speed process.

Figure 4.11: Zirconium-Oxygen phase diagram [Massalski, 1990].

crust

100 100

00 00

bulk

50 50

50 50 μ

MSUB run 1 MSUB run sample m

HSUB run1 sample run1 HSUB

H2 M

H1,H2

M2 M1

Figure 4.12: Phase composition analysis of HSUB-1 and MSUB-1 samples.

A significant increase in oxidation under medium subcooling conditions can be clearly observed based on both bubble analysis as well as SEM/EDS analysis. The observations indicate that only a limited extent of oxidation is possible under high subcooling conditions while the oxidation capacity increases at decreased degree of water subcooling. A simplified rational to the observation is the limitation in vapor flow to the melt considering the combined effects of rapid solidification of the melt and rapid condensation of water vapor under high subcooling conditions. The vapor flowing to the melt is rapidly broken down, where oxygen diffuses inside the melt droplet and the bubble surrounding the melt droplet is filled with hydrogen hindering water vapor transport. Under high subcooling conditions, the rapid formation of solid crust prevents further the transport of water vapor to the fresh metal. For a lower degree of water subcooling, water vapor condenses at a comparatively slower rate and since the energy required to vaporize water will be comparatively less, the melt shall tend stay in liquid state for a prolonged period increasing the capacity for oxidation.

O/Zr atomic ratio in the quenched drops estimated based on the hydrogen generation data obtained from bubble analysis of MISTEE-OX experiments is compared with the experimental data available in Baker.L., & Just.L., [1962] (Figure 4.13). Final Zr oxidation characterized by the average O/Zr atomic ratio in the quenched drops as a function of particle diameter is plotted. The particle diameter corresponds to mean particle diameter estimated from the discrete spherical fragments collected after experiments in Baker.L., & Just.L., [1962]. O/Zr atomic ratio can be observed to decrease with an increased particle diameter. Although not a fair comparison, MISTEE-OX experimental data seems to show a good agreement. 83

1.6 Experimental data involving liquid Zr (Baker.L., & Just.L., 1962), ∆Tsub:80K 1.4 Theoretical curve proposed in Baker.L., & Just.L., 1962 (based on experiments involving solid and liquid Zr oxidation) MISTEE-OX, ∆Tsub:85K 1.2 MISTEE-OX, ∆Tsub:45K 1 cases involving an explosive pressure rise (Tmelt˃2600°C) 0.8

0.6 O/Zr atomic ratio atomic O/Zr

0.4 Extrapolated

0.2

0 0 1000 2000 3000 4000 5000 6000 7000 Mean particle diameter (µm) Figure 4.13: O/Zr atomic ratio as a function of particle diameter.

An observation that needs to be stressed here is that the zirconium melt droplets do not seem to undergo spontaneous steam explosion even under high subcooling conditions indicating the stability of the vapor and gas film surrounding the melt droplet. The integral experiments, performed at ANL did not also involve spontaneously triggered steam explosions [Cho, et al., 1988]. In the current experiments with Zr melt, a cyclic process of hydrogen generation and release is observed, leaving behind a trail of bubbles in the droplet trajectory even under high subcooling conditions (Figure 4.2). For a very general comparative purpose, see Figure 4.14, where a droplet of molten alumina is discharged into water at almost similar test conditions. The melting point of alumina (2050°C) are comparable to that of a zirconium. Alumina droplets falling into highly subcooled water (∆Tsub: 80K) do not show formation of large trail of bubbles indicating rapid condensation of the generated water vapor except during the initial phase of entering into the water pool where the entrapped non-condensable gases detach from the melt droplet periphery (at t=79ms in Figure 4.14). Non condensable gases have substantial effects on steam explosion triggering as well as the overall explosion yield since the pressure of the non-condensable gases within the vapor film shall reduce condensation of vapor which is an essential process that dictates vapor explosion occurrence [Corradini, et al., 1988], [Akiyoshi, et al., 1990]. Non condensable gas release from Zr droplet oxidation reaction under water shows that steam explosion triggering can be suppressed even under high subcooling conditions whether alumina droplets can be relatively easily triggered under similar conditions. The cyclic process of bubble

84 release also indicates that the Zr droplet is left with a thinner film during bubble detachment promoting the effective participation Zr melt droplet in steam explosion. For a pre-mixture with many droplets of discrete sizes undergoing varied cycles of bubble growth and departure it can be rational to consider that the active mass that can effectively participate in steam explosion is considerably reduced. However it will be very useful to clarify the threshold bubble volume required to suppress steam explosion triggering in further work for a given trigger strength to quantify the time scales under which the bubble cannot be compressed or in other words triggered.

Figure 4.14: Time sequence raw snapshots depicting interaction between alumina melt and water (∆Tsup ≈1ooK, ∆Tsub ≈15K).

Oxidation influence on FCI is considered the most uncertain part in severe accident analysis. The investigation has produced several interesting insights into Zr melt oxidation behavior during FCI. The interactions between Zr melt and water is recorded by high-speed camera. The recorded images have shown periodic release of hydrogen bubbles in the water pool from melt droplet rear periphery. For conditions of melt and water temperature studied in the current experiments have shown the interactions to be benign without a spontaneously triggered steam explosions. However, specific tests with varied melt superheat is required to identify the thermal interaction zone if any.

The analysis has shown a significant influence of water subcooling. O/Zr atomic ratio increases with decrease in subcooling. Further, it is found from SEM/EDS analysis that the solidified Zr droplet is neither fully oxidized nor un-oxidized. The maximum observed O/Zr atomic ratio in the outer surface is approximately up to 0.9 for high subcooling conditions and approximately up to 1.5 for medium subcooling 85 conditions. SEM/EDS analysis of a droplet with an internal pore has shown formation of a crust layer on the pore surface with comparable O/Zr atomic ratio to the outer surface.

Based on the EDS analysis, the total oxygen concentration of the obtained droplet is about 40-60 at. %. This region on phase diagram can be reached only through the quenching of liquid droplet with corresponding composition. Given the initial melt superheat is 150K, the corresponding composition to 40-60 at. % of the phase diagram could have only been attained by increase in melt temperature during the interaction. It is thus hypothesized that the melt temperature could have potentially increased by approximately up to 200K under high subcooling conditions and up to 400K under medium subcooling conditions as a result of the chemical reaction. However, the hypothesis requires further clarification typically through relatively precise temperature measurement during the quench process in water pool.

Considering the complexity of the phenomena, scarcity of experimental data and importance of such experimental data for model development, it is vital to follow-up this experimental work (typically perform experiments involving a wide range of conditions including melt temperature, droplet diameter etc.) with improved measurements (i.e. temperature measurement during melt droplet quenching in water and use of techniques other than EDS for oxygen measurement for example Wavelength dispersive x-ray analysis (WDS) or carbo-thermal reduction) and also implement the use of metal mixtures (for example Fe-Zr).

5. Summary

This thesis work was motivated to reduce the uncertainties in understanding of the fuel-coolant-interactions (FCI) which may occur during a severe accident of light water reactors. The main outcomes were important insights into some specific aspects of the FCI phenomena, including melt jet breakup, debris characterization, steam explosion and oxidation of metallic melt.

The first part of the thesis was focused on melt jet/droplet fragmentation and the resulting debris properties during FCI. To delineate influential parameters on debris properties, an extensive set of experiments were performed at the MISTEE-Jet facility and JEBRA facility. In the MISTEE-Jet experiment, small jets of molten binary oxide materials (WO3-Bi2O3 or WO3-ZrO2) in the range of jet diameter up to 10mm were quenched in a subcooled water pool. It was found that a transition in debris size and morphology (mostly smooth to mostly fractured) occurred at a narrow range of moderate water subcooling. For the WO3-Bi2O3 melt, the transition occurred between 40K and 60K water subcooling at a melt superheat of 130K. For the WO3-ZrO2 melt, the transition occurred at 70K to 80K water subcooling at a melt superheat of 130K. The mass fraction of fractured particles increased with increasing water subcooling, indicating the difference in levels and mode of fragmentation. The average debris size increased with increasing melt superheat. The experimental results also indicated that the mass fraction of fractured particles was considerably less for WO3-ZrO2 melt than for WO3-Bi2O3 melt. The variation of melt jet diameter between 5 and 10mm was also investigated, and the results showed that the jet breakup length increased linearly with its diameter, but no apparent difference was observed in debris size distributions. Interestingly, a comparison with DEFOR and FARO experiments with approximately 10 times variation in jet diameter have shown no substantial differences on the minimum values of average particle size (in the range of 2 to 3 mm). Further experimental study was performed on the JEBRA facility which had larger diameter (10 to 30 mm) of molten Wood's metal jet to avoid boiling effect on fragmentation since Wood's metal has a melting point below the boiling point of water. The experimental results using Wood's metal showed that both jet breakup length and debris size increased with increasing jet diameter. The distinct trend in particle size of metallic melt from oxidic melt indicated that the hydrodynamic mechanisms are overridden by thermal mechanisms in the case of binary oxide materials.

The second part of this thesis was focussed on steam explosion. Experimental studies of single droplet steam explosion were carried out on the MISTEE test facilities. The experiment did not only produce insights into the mechanisms involved in steam explosion, but also provided useful data on the limiting/promoting mechanisms of steam explosion. The steam explosion

experiment was first conducted with lower melting-point binary oxides WO3-CaO on the MISSTEE-LT facility. High-speed simultaneous visuals of normal videography and X-ray radiography depicting melt droplet fragmentation into fine fragments and bubble dynamics were recorded. A process of bubble growth and collapse in two to three cycles with low melting-temperature melt was observed. The X-ray images showed a process of initial melt deformation or melt swelling on submission to pressure pulse was followed by an intense increase in the melt surface area and disintegrating of the melt droplet into several micro meter sized fragments dispersed in the water pool. The rate of melt deformation was directly proportional to the energetics of the explosion, and very fine fragments of melt were produced. A mass averaged median diameter of the fragments was obtained as ≈ 200 µm by conventional sieving of the fine fragments acquired after the experiment. Spontaneously triggered steam explosions have not been observed with experimental conditions at melt superheat up to 350K and subcooling up to 85K. The effect of binary oxide melt composition on steam explosion was then investigated by using binary oxides WO3-CaO at eutectic composition (75:25 mol %, Tliquidus = 1135 °C) and non-eutectic composition 67:33 mol%, Tliquidus = 1344 °C, Tsolidus = 1135 °C), respectively. The high-resolution images obtained during the experiment clearly depicted the difference in steam explosion between the varied compositions at a specific superheat of 100K. It was found that the non-eutectic composition formed a solid crust layer thereby restricting the amount of melt that can actively participate in steam explosion. The difference however was diminishing at a higher degree of melt superheat.

To reduce the temperature gap between the simulant melt and corium. MISTEE-HT was developed to be able to melt high melting-point materials. The new development included a technique of aerodynamic levitation of a droplet by purging inert gas through a hole underneath the crucible during induction heating. The molten alumina with a melting point ˃2000°C was chosen for the steam explosion experiment on the MISTEE-HT facility. The high-speed visuals showed droplet and bubble dynamics in two or three cycles similar to what was observed with lower melting-point melt materials. Though fine fragmentation of melt was observed for a water subcooling range of 90 to 45K, it was supressed at low subcooling conditions 10 to 30K. Under low subcooling conditions, the void surrounding the melt droplet increased, and thereby causing the suppression effect. Spontaneous explosion was observed at a high melt superheat (215K) and a high water subcooling (81K). The specific case was an indication that alumina melt can exhibit spontaneous steam explosion under certain conditions.

The third part of this thesis was focussed on oxidation of metallic melt falling into a water pool. The quenching process of Zr droplets in a subcooled water pool was recorded by a high-speed video camera. The obtained images showed a cyclic growth and departure of non-condensable bubbles due to oxidation of the Zr droplet with 89 surrounding steam. Upon the oxidation, the generated hydrogen was added to the steam film that formed due to film boiling and encapsulated the melt droplet. With continued oxidation, the film was expanding but the oxidation rate was demising due to deleting of steam in the film, until detachment of bubble from the droplet’s rear; afterwards the next cycle of oxidation began. Based on the SEM/EDS analysis, it was found that the entire surface area of the debris (quenched droplet) was sub- oxidized, i.e.,. neither fully oxidized nor pure metal was observed. It was also found that the oxygen content was radially stratified in three to four concentric circles. The outermost circular layer of the debris that spans only about 250µm contained the highest concentration of oxygen where O/Zr atomic ratio was approximately up to 0.9. The oxygen content reduced toward the central layer of the debris. The other inner concentric layers were however comparatively thicker than the outermost layer. Water subcooling had a prominent effect, and the case with a moderate subcooling (∆Tsub ≈45K) had a significant increase in the extent of oxidation, compared to that of high subcooling conditions (∆Tsub ≈85K). Considering the importance of such experimental data for model development, it is important to extend this experimental work.

6. References

Abe, Y., Nariai, H., & Hamada, Y. (2002). The trigger mechanism of vapor explosion. Journal of Nuclear Science and Technology, 39, 845–853. Abe, Y., Kizu, T., Arai, T., Nariai, H., Chitose, K., & Koyama, K. (2004). Study on Thermal-hydraulic behavior during molten material and coolant interaction. Nuclear Engineering and Design, 277–291. Abe, Y., Matsuo, E., Arai, T., Nariai, H., Chitose, K., Koyama, K., & Itoh, K. (2006). Fragmentation behavior during molten material and coolant interactions. Nuclear Engineering and Design. Akiyoshi, R., Nishio, S., & Tanasawa, I. (1990). A Study on the Effect of NonCondensable. International journal of heat and mass transfer, 33. Albrecht, G., Cherdron, W., Imke, U., Jacobs, H., Kaiser, A., Schütz, W., & Will, H. (2001). Experiments on Melt-Coolant Interaction Processes, including Steam Explosions. Structural Mechanics in Reactor Technology (SMIRT-16). Washington DC: International & American Associations for Structural Mechanics in Reactor Technology. Annuziato, A. (2014). FARO test L33, Quick look report, Technical note no: I.00.111, JRC Ispra, Italy. Baker.L., Just.L. (1962). Studies of metal-water interactions at high temperature – III. ANL report No:6548. Argonne national laboratory, Illinois. Bang, K., Kim, J., & Kim, D. (2003). Experimental study of melt jet breakup in water. Journal of Nuclear Science and Technology, 807–813. Basso, S. (2017). Particulate debris spreading and coolability. The Royal Institute of Technology, Physics. Stockholm: Doctoral Thesis. Berg, E., Burger, M., Cho, S., & Schatz, A. (1993). Modelling of the breakup of melt jets in liquids for LWR safety analysis. Nuclear Engineering and Design, 419– 429. Berthoud, G., & Valette, M. (1994). Development of a multidimensional model for the premixing phase of a fuel–coolant interaction. Nuclear Engineering and Design, 149, 409-418. Bertram, M.D., Hubbard F.R. & Pierce, D.C. (2007). Scrap inspection requires ingenuity and management. Light Metals. Board, S., Farmer, C., & Poole, D. (1974). Fragmentation in thermal explosions. International Journal of Heat and Mass Transfer, 17, 331-339. Bottomley, P.D.W., Farthing, M., Wiss, T., Cremer, B., Boshoven C., Lajarge, P., Rondinelaa, V. (2015). Investigations of the melting behaviour of the U–Zr– Fe–O system. Journal of Nuclear Science and Technology, 52, 1217-1225.

Buchanan, D. (1974). A model for Fuel-Coolant Interactions. Journal of Physics D: Applied Physics, 7. Burger, M., Cho, S., Berg, E., & Schatz, A. (1995). Breakup of melts as pre-condition for premixing: modeling and experimental verification. Nuclear Engineering and Design, 215-251. Buxton, L., & Benedick, W. (1978). Steam explosion efficiency studies. SAND-79- 1399. NUREG/CR-0947. Buxton, L., & Benedick, W. (1980). Steam explosion efficiency studies: Part II corium melts. NUREG/CR-1746. Castrillon, S., Meignen, R., Picchi S. (2015). A two-scale approach for modeling the corium melt fragmentation during fuel-coolant interaction. NURETH-16, Chicago, IL, August 30-September 4. Carachalios, C., Bürger, M., & Unger, H. (1983). A transient two-phase model to describe thermal detonations based on hydrodynamics fragmentation. Proceedings International meeting on LWR Severe Accident Evaluation, Cambridge, (pp. 1-8). Cambridge. Chang, L., Scroger, G., & Phillips, B. (1966). A1 kaline-Earth Tungstates: Equilibrium and stability in the M-W-0 Systems. Journal of the American ceramic society, 49. Chen, X., Luo, R., Yuen, W., & Theofanous, T. (1999). Experimental simulation of microinteractions in large. Nuclear Engineering and Design, 163–178. Cherdron, W., Huber, F., Kaiser, A., & Schutz, W. (2005). ECO Steam explosion experiments documentation and evaluation of experimental data. Karlsruhe: Wissenschaftliche Berichte. Cho, D., Armstrong, D., & Gunther, W. (1988). Experiments on interaction between zirconium containing melt and water. NUREG/CR-5372. Cho, D., Armstrong, D., & Gunther, W. (1998). Experiments on Interactions Between Zirconium-Containing Melt and Water. Argonne, IL: U.S. Nuclear Regulatory Commission. Chu, C., & Corradini, M. (1989). One-dimensional transient fluid model for fuel/coolant interaction analysis. Nuclear Science and Engineering, 101, 48- 72. Chu, C., Sienick, J., Spencer, B., Frid, W., & Wenhielm, G. (1995). Ex-vessel melt- coolant interactions in deep water: studies and accident management for Swedish BWRs. Nuclear Engineering and Design, 159–213. Ciccarelli, G., & Frost, D. (1994). Fragmentation mechanisms based on single drop steam explosion. Nuclear Engineering and Design, 109-132. Corradini, M., Kim, B., & Oh, M. (1988). Vapor explosions in light water reactors: A 93

review of theory and modeling. Progress in Nuclear Energy, 1-117. Cronoberg, W., Chawla, T., & Fauske, H. (1974). A thermal stress mechanism for the fragmentation of molten UO2 upon contact with sodium coolant. Nuclear Engineering and Design, 30, 434-443. Dinh, T., Bui, V., Nourgaliev, R., Green, J., & Sehgal, B. (1999). Experimental and Analytical Studies of Melt Jet-Coolant Interactions: A synthesis. Nuclear Engineering and Design, 299–327. Dinh, T. (2007). Material property effect in steam explosion energetics: Revisited. The 12thInternational Topical Meeting on Nuclear Reactor Thermal Hydraulics (NURETH-12). Pittsburgh, Pennsylvania. Dinh, T., Hannson, R., & Kudinov, P. (2008). On Solidification Mechanism that Govern the Effect of Binary Melt Composition on Steam Explosion Energetics. Transaction of American Nuclear Society. Dombrovsky, A., & Dinh, T. (2008). The effect of thermal radiation on the solidification dynamics of metal oxide melt droplets. Nuclear Engineering and Design, 238, 1421–1429. Dullforce, T., Buchnan, D., & Peckover, R. (1976). Self-triggering of small scale fuel- coolant interactions. J. Phys. D. Appl. Phys. Elyutin, V., Mitin, B., & Nagibin, Y. (1973). Properties of liquid aluminium oxide. Fizika Aerodispersnykh Sistem, 104-109. Ekenes, J.M. (2007). Training for preventing molten metal explosion in aluminium cast houses. Light Metals. Epstein, M., & Fauske, H. (2001). Applications of the turbulent entrainment assumption to immiscible gas-liquid and liquid-liquid systems. IChemE, 453– 462. Epstein, S.G. (2000). Recycling aluminium safely. Scrap, 57. Epstein, S.G. (2009). Molten metal explosions are still occuring. Light metals. Fang, M., Chandra, S., & Park, B. (2008). Building three-dimensional objects by deposition of molten metal droplets. Rapid Prototyping Journal, 14, 44-52. Fletcher, D., & Thyagaraja, A. (1991). The CHYMES coarse mixing. Progress in nuclear energy, 31-61. Fletcher, D. (1993). Propagation investigations using the CULDESAC. Proc. OECD/NEA/CSNI Specialists Meet. on Fuel Coolant Interaction. Santa Barbara. Fletcher, D. (1995). Steam explosions triggering: A review of theoretical and experimental investigation. Nuclear Engineering and Design, 155, 27-36. Frano, R., Aquaro, D., Scaletti, L., & Olivi, N. (2015). Characterization of the thermal

conductivity for ceramic pebble beds. Journal of Physics: Conference Series. Gabor, J., Purviance, R., Aeschlimann, R., & Spencer, B. (1988). Breakup and quenching of molten metal fuel in sodium. American nuclear society, safety of next generation power reactors meeting. Seattle, Washington. Gelfand, B. (1996). Droplet breakup phenomena in flows with velocity lag. Progress in Energy and Combustion Science, 22, 201-265. Ginsberg, T. (1985). Liquid jet breakup characterization with application to melt- water mixing. National Heat Transfer Conference, Denver, CO. ANS Proceedings. Granier, J., & Potard, C. (1987). In Proc. 6th Eur. Symp. Micrograv (pp. 421–425). Bordeaux, France: ESA SP-256. Grischenko. D, & Piluso, P. (2011). Recent progress in the gas-film levitation as a method for thermophysical properties measurements: application to ZrO2- Al2O3 system. High Temperature-High Pressures, 40, 127-149. Hannson, R., Park, H., & Dinh, T. (2009). Simultaneous high-speed cinematographic and X-ray radiographic imaging of an intense multi-fluid interaction with rapid phase changes. Experimental Thermal and Fluid Science, 33, 754-763. Hansson, R. (2009). Triggering and Energetics of a Single Drop Vapor Explosion: The role of Entrapped Non-Condensable Gases. Nuclear Engineering and Technology, 41. Hansson, R. (2010). An Experimental Study on the Dynamics of a Single Droplet Vapor Explosion. Stockholm: PhD thesis, KTH-School of Engineering Sciences. Haraldsson, H. (2000). Breakup of jet and drops during premixing phase of fuel coolant interaction, PhD thesis. The Royal Institute of Technology (KTH). Hong, S.W., Piluso, P., Leskovar, M. (2013). Status of OECD-SERENA project for the resolution of ex-vessel steam explosion risks. Journal of Nuclear Power Engineering, 7, 423. Huhtiniemi, I., Hohmann, H., & Magallon, D. (1997). FCI experiments in the corium/water system. Nuclear Engineering and Design, 177, 339-349. Huhtiniemi, I., Magallon, D., & Hohmann, H. (1999). Results of recent KROTOS FCI tests: alumina versus corium melts. Nuclear Engineering and Design, 379- 389. IAEA PRIS database. (2018, Februvary). Retrieved from IAEA: www.iaea.org Jin, X. (2017). Smelt droplet-water interaction in the kraft recovery boiler dissolving tank. PhD thesis, University of Toronto, Department of Chemical Engineering and Applied Chemistry, Torono. 95

Johnson, T., Baker, L. J., Pavlik, J., & Spencer, B. (1974). Large scale molten fuel sodium interaction experiments. Proceedings of fast reactor safety meeting. Journeau, C. (2010). PLINIUS FP6 Transnational Access to the Prototypic Corium Platform PLINIUS. Cadarache, France: Research Report of EURATOM 6th Framework Programme. Kim, J.H., Zhao, & X., Xiao, P. (2014). In situ measurement of stresses and phase compositions of the zirconia scale during oxidation of zirconium by Raman spectroscopy. Oxidation of Metals, 81, 331-343. Kim, B., & Corradini, M. (1988). Modeling of Small-Scale Single Droplet Fuel/Coolant Interactions. Nuclear Science and Engineering, 98, 16-28. Kim, J., Min, B., Park, I., Kim, H., & Hong, S. (2008). Steam explosion experiments using partially oxidized corium. Journal of mechanical science and technology, 22. Kitamura, N., Makihara, M., Hamai, M., Sato, T., Mogi, I., Awaji, S., . . . Motokawa, M. (2000). Containerless Melting of Glass by Magnetic Levitation Method. Japanese Journal of Applied Physics, 39. Kondo, S., Konishi, K., Isozaki, M., Imahori, S., Furutani, A., & Brear, D. (1995). Experimental study on simulated molten jet-coolant interactions. Nuclear Engineering and Design, 73-84. Kouraytem, N., Li, E., & Thoroddsen, S. (2016). Formation of microbeads during vapor explosions of Field’s metal in water. PHYSICAL REVIEW E, 93. Kudinov, P., Karbojan, A., Vilanueva, W., & Tran, C. (2013). Agglomeration and size distribution of debris in DEFOR-A experiments with Bi2O3–WO3 corium simulant melt. Nuclear Engineering and Design, 284-295. Kudinov, P., Karbojian, A., Ma, W., & Dinh, T. (2010). The DEFOR-S experimental study of debris formation with corium simulant materials. Nuclear Technology, 219-230. Leskovar, M., Centrih, V., & Uršič, M. (2016). Analysis of oxidation influence on steam explosion energetics. Annals of Nuclear Energy, 90, 441-449. Li, L., & Ma, W. (2011). Experimental characterization of effective particle diameter of a packed bed with multi-diameter spheres. Nuclear Engineering and Design, 241, 1736-1745. Li, L., & Ma, W. (2011). Experimental characterization of the effective particle diameter of a particulate bed packed with multi-diameter spheres. Nuclear Engineering and Design, 1736-1745. Lu, Q., Chen, D., & Li, C. (2016). Visual investigation on the breakup of high superheated molten metal during FCI process. Applied Thermal Engineering, 962-975.

Luo, J., Qi, L., Zhong, S., & Li, H. (2012). Printing solder droplets for micro devices packages using pneumatic drop-on-demand (DOD) technique. Journal of Materials Processing Technology, 2066-2073. Ma, W., & Dinh, T. (2010). The effects of debris bed’s prototypical characteristics on corium coolability in a LWR severe accident. Nuclear Engineering and Design, 240, 598-608. Ma, W., & Dinh, T.-N. (2010). The effects of debris bed’s prototypical characteristics on corium coolability in a LWR severe accident. Nuclear Engineering and Design, 598–608. Magallon, D. (2006). Characteristics of corium debris bed generated in large-scale fuel–coolant interaction experiments. Nuclear Engineering and Design. Magallon, D., & Huhtiniemi, I. (2001). Corium melt quenching tests at low pressure and subcooled water in FARO. Nuclear Engineering and Design, 204, 369- 376. Massalski, T.B, (1990). Binary alloy phase diagram. ASM international, Materials park, ISBN 0-87170-403-X. Mathai, A., Sharma, A., Anandan, J., Malarvizhi, B., Das, S., Nashine, B., & Chellapandi, P. (2015). Investigation of fragmentation phenomena and debris bed formation during core meltdown accident in SFR using simulated experiments. Nuclear Engineering and Design. Matsumura, K., & Nariai, H. (1996). Mechanisms of vapor explosion for a tin and water system. Journal of Nuclear Science and Technology, 33. Matsuo, E., Abe, Y., Chitose, K., Koyama, K., & Itoh, K. (2008). Study of jet breakup behavior at core disruptive accident for fast breeder reactor. Nuclear Engineering and design, 238, 1996-2004. Medhekar,, S., Abolfadl, M., & Theofanous, T. (1991). Triggering and propagation of steam explosions. Nuclear engineering and design, 126, 41-49. Meignen, R., Picchi, S., Lamome, J., Raverdy, B., Castrillon , S., & Nicaise, G. (2014). The challenge of modeling fuel–coolant interaction: Part I – Premixing. Nuclear Engineering and Design, 280, 511-527. Meignen, R., Raverdy, B., Buck, M., Pohlner, G., Kudinov, P., Ma, W., . . . Ivanov, I. (2014). Status of steam explosion understanding and modelling. Annals of Nuclear Energy, 125–133. Meignen, R., Raverdy, B., Picchi, S., & Lamome, J. (2014). The challenge of modeling fuel–coolant interaction: Part II – Steam explosion. Nuclear Engineering and Design, 528-541. Min, B., Song, J., Park, Y., & Kim, J. (2006). A physical and chemical analysis of fast quenched particles of UO2 and ZrO2 mixture. Journal of Nuclear Materials, 358, 243–254. 97

Muck, O. (1923). Germany Patent No. 42204. NEA (2007). OECD research program on fuel-coolant interactions-Steam explosion resolution for nuclear applications SERENA final report, NEA/CNSI/R. NEA (2014). OECD research program on fuel-coolant interactions-Steam explosion resolution for nuclear applications SERENA –Phase 2 final report, NEA/CNSI/R. NEA (2017). Status report on Ex-vessel steam explosion, NEA/CNSI/R. Nelson, L. (1995). Steam explosions of single drops of pure and alloyed. Nuclear Engineering and Design, 155, 413-425. Nelson, L., & Duda, P. (1981). Steam explosion experiments with single drops of iron oxide. NUREG CR-2285. Nelson, L., & Duda, P. (1982). Steam explosions of molten iron oxide drops: easier initiation at small pressurizations. Nature, 296, 844–846. Nelson, L., Brooks, P., Bonazza, R., Corradini, M., Hildal, K., & Bergstrom, T. (2001). Steam explosions of single drops of molten silicon rich alloys. Proceedings of the Ninth International Ferro alloys Congress (INFACON 9). Quebec. Nelson, L., Brooks, P., Bonazza, R., Corradini, M., Hildal, K., & Bergstrom, T. (2001). Steam explosions of single drops of molten silicon-rich alloys. Proceedings of the Ninth International Ferroalloys Congress (INFACON 9) (pp. 338-351). Quebec City: Ferroalloys Association, Washington, DC. Park, H., Hansson, R., & Sehgal, B. (2005). Fine fragmentation of molten droplet in highly subcooled water due to vapor explosion observed by X-ray radiography. Experimental Thermal and Fluid Science, 29, 351-361. NUREG (1996). A reassessment of the potential for an alpha mode containment failure and a review of the current understanding of broader fuel-coolant interaction issues. Pilch, M., & Erdman, C. (1987). Use of breakup time data and velocity history data to predict the maximum size of stable fragments for acceleration-induced breakup of a liquid drop. International journal of multiphase flow, 13, 741-757. Piluso, P., Trillon, G., Oh, H., & Magallon, D. (2005). Material effect and steam explosion at high temperature (T>2300 K). Internation Journal of Thermophysics, 1095-1113. Pohlner, G., Buck, M., & Rahman, S. (2013). Modeling and validation basis of the JEMI code. Rasmussen, J., & Nelson, R. (1972). Surface Tension and Density of Molten Al2O3. Journal of the American Ceramic Society, 54, 398-401. Rumschoettel, D., Griebel, B., Irlinger, F., & Lueth, C. (2017). A fast pneumatic droplet generator for the ejection of molten alunminium. Pan Pacific

Microelectronics Symposium. Pan pacific. Schneider, J., Marciniak, M., & Barclay, G. (1992). Breakup of metal jets penetrating a volatile liquid. 5th International Topical Meeting on Reactor Thermal Hydraulics (NURETH-5). Salt Lake City, LIT, : ANS Proceedings. Schneider, S. (1979). Cooperative Determination of the Melting Point of Alumina. Pure and Applied Chemistry, 117-122. Sehgal, B. (2006). Stabilization and termination of severe accidents in LWRs. Nuclear Engineering and Design, 1941-1952. Shen, P., Zhou, W., Louis, N., Journeau, C., Piluso, P., Liao, Y. (2018). Corium behaviour and steam explosion risks: A review of experiments. Annals of Nuclear Energy, 121, 162-176. Song, J., Park, I., Chang, Y., Shin, Y., Kim, J., Min, B., Kim, H. (2002). Experiments on the interaction of molten ZrO2 with water using TROI facility. Nuclear engineering and design, 213. Song, J., Kim, S., & Kim, H. (2005). On some salient unresolved issues in severe accidents for advanced light water reactors. Nuclear Engineering and Design, 235, 2055-2069. Song, J., Park, I., Chang, Y., Shin, Y., Kim, J., Min, B., Kim, H. (2006). The effect of corium composition and interaction vessel geometry on the prototypic steam explosion. Annals of Nuclear Energy, 1437-1451.

Speranskaya, E. (1970). The Bi2O3‐WO3 System. Neorg. Mater, 149-151.

Takashima, T. (2008). Observations of Initiation Stage of Spontaneous Vapor Explosions for Droplet Scale. Heat Transfer—Asian Research, 37. Takashima, T., & Lida, Y. (1995). A Study on the Mechanism of Spontaneous Vapor Explosions with Single Molten Tin Drops and Water. JSME International Journal Series B Fluids and Thermal Engineering, 38, 114-120. Taylor, G. (1940). G I Taylor Scientiﬁc Papers. Cambridge University Press. Tyrpekl, V. (2012). Effet matériaux lors de l'interaction corium-eau: analyse structural des débris d'une explosion vapeur et mécanismes de solidification. France: PhD thesis. CEA, Cadarache. Tyrpekl, V., Piluso, P., Bakardjieva, S., Niznansky, D., Rehspringer, J., Bezdicka, P., Dunge, O., (2015). Prototypic corium oxidation and hydrogen release during the fuel-coolant interaction. Annals of Nuclear Energy, 75, 210-218. Ursic, M., Leskovar, M., & Mavko, B. (2012). Simulations of KROTOS alumina and corium steam explosion experiments: Applicability of the improved solidification influence modelling. Nuclear Engineering and Design, 246, 163– 174. 99

Wang, S., Blomquist, C., & Spencer, B. (1989). Modelling of thermal and hydrodynamic aspects of molten jet/water interaction. 26th National Heat Transfer Conference (pp. 225-232). Philadelphia: ANS. Weber, J., Hampton, D., & Merkley, D. (1994). Aero-acoustic levitation: a method for containerless liquid-phase processing at high temperatures. Review of Scientific Instruments, 65, 456–465. Winborne, D., Nordine, P., Rosner, D., & Marley, N. (1976). Aerodynamic levitation technique for containerless high temperature studies on liquid and solid samples. Metallurgical Transactions B, 7, 711–713. World nuclear association. (2018). Retrieved from World nuclear association: http://www.world-nuclear.org/ Yakush, S., Kudinov, P., & Lubchenko, N. (2013a). Coolability of heat-releasing debris bed. Part 1: Sensitivity analysis and model calibration. Annals of Nuclear Energy, 52, 59-71. Yakush, S., Kudinov, P., & Lubchenko, N. (2013b). Coolability of heat-releasing debris bed. Part 2: Uncertainty of dryout heat flux. Annals of Nuclear Energy, 52, 72-79. Yamano, N., Maruyama, Y., Kudo, T., Hidaka, A., & Sugimoto, J. (1995). Phenomenological studies on melt-coolant interactions in the ALPHA program. Nuclear Engineering and Design, 155, 369-389. Yang, F., Zhao, & X., Xiao, P. (2014). In situ measurement of stresses and phase compositions of the zirconia scale during oxidation of zirconium by Raman spectroscopy. Oxidation of Metals, 81, 331-343.