Review of thermal-hydraulic researches in severe accidents in light water reactors

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1 Journal of Nuclear Science and Technology ISSN: (Print) (Online) Journal homepage: Review of thermal-hydraulic researches in severe accidents in light water reactors Isao Kataoka To cite this article: Isao Kataoka (2013) Review of thermal-hydraulic researches in severe accidents in light water reactors, Journal of Nuclear Science and Technology, 50:1, 1-14, DOI: / To link to this article: Published online: 21 Dec Submit your article to this journal Article views: 1560 Citing articles: 6 View citing articles Full Terms & Conditions of access and use can be found at

2 Journal of Nuclear Science and Technology, 2013 Volume 50, No. 1, 1 14, 50TH ANNIVERSARY INVITED REVIEW Review of thermal-hydraulic researches in severe accidents in light water reactors Isao Kataoka* Department of Mechanical Engineering, Osaka University, 2-1 Yamadaoka, Suita, Osaka , Japan (Received 27 September 2012; accepted final version for publication 15 November 2012) The Tohoku Region Pacific Coast Earthquake and subsequent severe accident (SA) in Fukushima Daiichi Nuclear Power Station caused unprecedented disaster in Japan. Before this accident, considerable researches on SAs had been carried out in Japan. However, unfortunately, such researches could not prevent the accident due to the unexpected huge Tsunami. However, the researches on SAs become more and more important in order to make clear the causes of the accident in Fukushima and improve the safety of nuclear power plants in Japan. In view of this, review on researches on thermal hydraulics in SAs in light water reactors was carried out. Important thermal-hydraulic phenomena in SAs were identified. Research activities on each phenomenon were surveyed mainly based on the articles published in Journal of Nuclear Science and Technology of Atomic Energy Society of Japan. Keywords: severe accident; thermal hydraulics; LWR 1. Introduction A huge earthquake attacked northeast Japan on 11 March Subsequently, large Tsunami attacked Fukushima Daiichi Nuclear Power Plant and caused severe accident (SA). As a result, a large amount of radioactive material was released and unprecedented nuclear disaster occurred in Fukushima prefecture. Accident managements could not prevent the core damage of nuclear power plants in Fukushima nuclear power plants, although considerable researches had been carried out in Japan for several decades before this accident. Many research activities have been carried out in analytical and experimental researches in universities, research institutes, and industries. Surveillance on SA researches has been carried out in many research committees such as Special Research Committee on Thermal Hydraulic of Severe Accidents (Atomic Energy Society of Japan, AESJ) [1], Research Committee on Severe Accident (AESJ), IVR-AM (In Vessel Retention-Accident Management) Research Committee (Nuclear Power Engineering Corporation, NUPEC), Severe Accident Research Committee, Reactor Safety Research Committee (Japan Atomic Energy Research Institute, JAERI), and Advisory Committee on Severe Accident (Japan Nuclear Energy Safety Organization, JNES). International conferences and meetings on SA were held in Japan by AESJ, NUPEC, JNES, and other institutes (e.g., Technical Meeting on Severe Accident and Accident Management for Nuclear Power Plants, March 2006, Tokyo, Japan, and SARJ (Workshop on Severe Accident Research, Japan), ). Research efforts are still continued in order to make clear the causes and events of Fukushima accident. Moreover, SA researches become more and more important in order to improve the safety of nuclear power plants in Japan. In view of this, in this article, a review of research activities on thermal hydraulics in SA in light water reactors (LWRs) was carried out. Survey of researches was mainly focused on Japanese works reported in Journal of Nuclear Science and Technology published by AESJ. The SAs consist of sequence of thermal-hydraulic events. As a result, reactor core is damaged and melted. If the pressure vessel failure occurs, a part of molten core is transferred into containment vessel. Furthermore, if the containment vessel is damaged, a large amount of radioactive materials is released into environment as Fukushima accident. Conceptual representation of SA is shown in Figure 1 [1]. The SAs are initialed by internal events such as loss of coolant accident (LOCA) or external events such as station blackout due to huge earthquake and/or tsunami. If the cooling of reactor core is not successful, reactor core is exposed to vapor and temperature of fuel rod rapidly increases due to decay heat. When the temperature of Zircalloy clad exceeds 9008C, * kataoka@mech.eng.osaka-u.ac.jp Ó 2013 Atomic Energy Society of Japan. All rights reserved.

3 2 I. Kataoka zirconium water reaction occurs and fuel rods start to collapse due to high inner pressure. Further increase of fuel temperature causes melting of fuel. Then, reactor core is damaged. Some of solid fragments of clad and fuel fall onto the lower head of pressure vessel. Melted fuel and clad also fall onto the lower head. This phenomenon is called relocation of molten core. If there is water in the lower head, molten fuel, and clad might violently interact with water and explode. It is called vapor explosion or fuel coolant interaction (FCI). Finally, molten and solid debris of reactor core are located in the lower head of pressure vessel. If there is water above the molten core, it is cooled by natural convection and/or boiling heat transfer. If the cooling of molten core is not adequate, the bottom of pressure vessel melts at the guide tube of control rod (boiling water reactor, BWR) and/or instrumentation tube (BWR, pressurized water reactor, PWR) and molten core falls into the containment vessel. In order to prevent the failure of the bottom of pressure vessel, there is a SA management called in-vessel retention (IVR). In this management, pressure vessel is cooled externally by water spray or flooding of bottom of containment vessel by water as shown in Figure 2 [2]. The researches of basic thermal-hydraulic mechanisms of boiling heat transfer at the outer surface of the bottom of pressure vessel have been carried out. If the failure of the bottom of pressure vessel (melt through) occurs, the molten core is ejected into the bottom of containment vessel in two ways. If the pressure of pressure vessel is very high, molten core is ejected very rapidly into containment vessel as a jet (high pressure melt ejection, HPME) and followed by vapor flow. Molten core is dispersed and transported into whole area of containment vessel as shown in Figure 3. Due to the high temperature of molten core, temperature, and pressure of containment vessel atmosphere rapidly increase and exceed the design pressure of containment vessel. This event is called as direct containment heating (DCH) which threatens the integrity of containment vessel. If the pressure of pressure vessel is not high, molten core falls onto the floor of bottom of containment vessel through the break in the pressure vessel. When the floor of containment vessel is dry, a pool of molten core is formed in the bottom of containment vessel. However, if the water exists in the bottom of containment vessel, there is a possibility of vapor explosion. The bottom of containment vessel consists of thick concrete wall. A pool of molten core reacts with concrete and erosion of concrete wall occurs. This Figure 1. Conceptual representation of SA [1]. Figure 2. Conceptual representation of IVR [2]. Figure 3. Conceptual representation of DCH.

4 Journal of Nuclear Science and Technology, Volume 50, No. 1, January event is called as molten core concrete interaction (MCCI). If a pool of molten core is adequately cooled by the water above it, the erosion of concrete wall stopped at certain depth in concrete wall. However, if the molten core is not sufficiently cooled, concrete wall of containment vessel is finally penetrated by molten core. The temperature of molten core is very high. Therefore, it reacts with steam or water and generates hydrogen. The hydrogen is accumulated in the containment vessel. If containment atmosphere is not inert and hydrogen concentration exceeds certain value, there is a possibility of hydrogen explosion in the containment vessel. If the hydrogen leaks from containment vessel and is accumulated in the reactor building, hydrogen explosion also occurs in reactor building as seen in Fukushima accident. In order to analyze and simulate above mentioned events, computer codes were developed. The Modular Accident Analysis Program (MAAP) and MELCOR codes were developed in United States and widely used in SA analyses in Japan. The ASTEC code is developed in Europe. The SA codes were also developed in Japan. The SAMPSON code was developed by NUPEC (now The Institute of Applied Energy, IAE) and THALES code was developed by JAERI (now Japan Atomic Energy Agency, JAEA). These codes are also used in the analyses of SA including Fukushima accident. In the following sections, review was made on the Japanese researches on above mentioned events and simulation codes mainly reported in Journal of Nuclear Science and Technology. experiments were adopted for verification of the multivelocity field model and vaporization/condensation model. Analysis by this module agreed well with the experimental data. The integrated function of this MCRA module simulating melting/freezing of a fuel rod and molten fuel is demonstrated in Figure 4. Relocation was also studied in relation to liquidmetal-cooled reactor [4]. Basic mechanisms are same as those in LWR. To simulate the behavior of molten metal mixed with solid particles, a computational framework was developed using the finite volume particle (FVP) method and the distinct element method (DEM). For mixed-flow calculations, FVP was coupled with DEM. A three-dimensional computer code developed for solid liquid mixture flows was validated by a series of pure and mixed-melt freezing experiments using a low-melting-point alloy. 2. Relocation of molten core If the reactor core is not sufficiently cooled, temperature of fuel increases and finally breaks. Some of fuel and clad disintegrated in a form of solid fragment and fall into the bottom of pressure vessel and others are melted and flow down as a form of liquid droplet or film. Since there are complicated structures in reactor core, the behavior of these solid fragments and liquid of fuel and clad is quite complicated. However, the knowledge of these phenomena (relocation of molten core) is quite important in SA analyses. Therefore, various models of relocation were developed and separate effect experiments were carried out to validate these models. The molten core relocation analysis (MCRA) module is developed by NUPEC as one of the modules of SA analysis code SAMPSON [3]. This module simulates the relocation of molten core in SA mechanistically. Flow regimes, interfacial area, momentum exchange, and phase change models are considered in this module. These models are verified by separate effect test experiments. The fluid dynamic model was verified by nitrogen gas bubbling through water in Leung s experiment. The JRC-ISPRA KROTOS-37 Figure 4. Melting/freezing model of MCRA [3]. (a) Melt contacts its particles and then each changes into the other. (b) Melt contacts structure or crust and then structure or crust melts or melt freezes into crust. (c) Melt contacts particles and then each changes its phase. (d) Melt contacts crust and then crust melts or melt freezes into crust. (e) Melt contacts crust on a rod and then crust melts or melt freezes into crust.

5 4 I. Kataoka In the analyses of relocation of molten core, thermal properties of molten and mixed core materials are required to be known and research on the thermal properties of molten core was carried out [5]. The specific heat capacity, thermal expansion coefficient, thermal diffusivity, and melting temperature were measured or estimated on the core debris samples of the Three Mile Island Unit 2 (TMI-2) reactor and simulated debris (SIMDEBRIS), which had chemical composition and porosity similar to the TMI-2 debris. The thermal diffusivity of the TMI-2 debris is as low as 10 25% of UO 2 at room temperature but is comparable above 1500 K. The melting temperature of SIMDEBRIS is about 2840 K, which is equivalent to the liquidus temperature of (U, Zr)O 2. Other core materials less than 10% in weight have no influence on the melting temperature. The debris coolability analysis model and simulation module were developed in order to predict more mechanistically the safety margin of reactor pressure vessels (RPVs) in a SA [6,7]. Debris three-dimensional natural convection is simulated with simultaneous spreading, melting, and solidification using the debris spreading cooling model and the pressure vessel wall failure is evaluated. Debris spreading is solved by the free surface calculation method using the height function. The results are compared with a water spreading experiment with reasonable agreements. The capability with energy transportation and solidification analysis is also verified by comparing with a thermal-hydraulic experiment for spreading of a stainless steel (SS) melt. The comparisons show good agreement. The model of the three-dimensional natural convection with simultaneous spreading, melting, and solidification is also tested. Fission products release during core damage and relocation of molten core is also an important subject in SA. In relation to this phenomenon, the Phebus FPT1 test at Institute for Radiological Protection and Nuclear Safety, IRSN experiments were carried out as International Standard Problem (ISP)-46. This experiment has been analyzed by the IMPACT/SAMPSON code [8]. The results for an examination of bundle degradation and fission products release were reported. Temperature changes of fuel, cladding, and control rod were well predicted. Accumulated hydrogen generation due to Zr/steam reaction differed only about 3% from the test result. Overall, good agreement was obtained for the fuel relocation and an accumulation of debris in the analysis. Analysis of enhanced diffusion due to degraded fuel well simulated of release behaviors of Xe, Cs, I, and Te. Diffusion analysis through singlecrystal grain well predicted the release behavior of Mo, Sb, Tc, Ru, and Ba observed in the Phebus FPT1 test. Behaviors of core melt and relocation in Fukushima accident were analyzed [9,10]. Analyses are performed of the first core melt behavior of the Unit 1, Unit 2, and Unit 3 reactors as well as the re-melt (melt again) behavior. The analysis is based on the total energy balance in the core. In the analysis, the total energy vs. temperature curve is developed for each reactor, which is based on the estimated core material inventory and material property data. The heat source is the decay heat of fission products and actinides together with reaction heat from the zirconium steam reaction. Core damage starting at 1200 K is shown, by fuel cladding burst due to gap pressure. With the cladding burst, Kr, Xe, and I are released into the RPV. Core material melting started at 1500 K, due to the control rod blade melting and relocation downward due to B4C/SS eutectic liquefaction in the actual accident situation. Analysis results on the major events in each reactor are summarized in Table Vapor explosion Vapor explosion is composed of several physical processes as schematically shown in Figure 5 [11]. At first, the high temperature molten material is coarsely mixed in low temperature liquid such as water as droplets. Film boiling occurs at the surface of droplets and droplets are covered with vapor film. This process is called as coarse mixing. In the next step, the vapor film around the droplet is collapsed by internal or external disturbances. Then, high temperature droplet surfaces directly contact with water and violent nucleate boiling occurs at the water molten material interface without solidification. This causes atomization of the high temperature molten material and violent nucleate boiling occurs at surface of fine droplets resulting in generation of pressure pulse. This process is called as triggering. This pressure pulse is propagated in water and the rest of droplets are also finely atomized and violent nucleate boiling occurs at all droplet surfaces. This process id called as propagation. Ultra-fast heat transfer and vapor generation cause rapid expansion of vapor volume, and very large pressure pulse with high propagation speed is generated. This process is called as expansion. If vapor explosion occurs, pressure vessel, containment vessel, and other structures in nuclear power plant may suffer from serious damages. There are many experimental and analytical investigations about the each process of vapor explosion. Experimental and analytical researches are carried out using small- and large-scale experiments [11 23]. The self-triggering mechanism of vapor explosions was investigated analytically and experimentally using molten tin and water [12]. A simple droplet system consisting of a hot-liquid droplet in a pool of cold liquid is considered. In order to model the selftriggering mechanism, it is assumed that instability in the vapor/cold-liquid interface produces a collapse of the vapor film. To investigate the stability of instability in a vapor film, a linear stability analysis was carried out. There was a region of film stability in the cold-liquid temperature where spontaneous vapor

6 Journal of Nuclear Science and Technology, Volume 50, No. 1, January Table 1. Analysis results on major events in Fukushima Daiichi nuclear reactors. Event Unit-1 Unit-2 Unit-3 Top of core uncovering 11 March 16:50 a 14 March 16:20 b 13 March 03:29 Bottom of core uncovering 11 March 19:31 14 March 18:22 13 March 07:46 Core damage starts (1200 K): 11 March 17:42 14 March 18:48 13 March 05:18 fuel cladding burst Core material melting starts (1500 K): 11 March 18:03 14 March 19:18 13 March 05:56 B4C/SS eutectic liquefaction; control rod meltdown Runaway Zr-steam reaction starts 11 March 18:03 No reaction 13 March 05:56 U02 melting starts (3113 K) 11 March 19:05 14 March 22:34 13 March 07:17 UO2 melting terminates and core collapses 11 March 19:31 14 March 23:27 13 March 07:46 Core material relocates to lower plenum 11 March 19:31 14 March 21:10 b 13 March 10:02 b 14 March 23:15 13 March 11:55 Hydrogen generated before 453 kg (31% oxidized) kg (61% oxidized) lower plenum relocates RPV bottom breaks 11 March 21:00 c 14 March 21:19 b 13 March 14:10 b Re-melting of core materials March March 21 March 01:25, March Places where rapid increase in dose rate was observed, which is thought to be a consequence of the re-melting IF, 2F, Yamagata, Mito IF, Niigata IF, 2F, Takahagi, Mito, Tokyo Fuel material distribution Most in D/W, significant part in RPV Note: a Assumed. b Analysis of measured data. c Inferred based on the analysis. Most in D/W, significant part in RPV Most in D/W, significant part in RPV Figure 5. Processes of vapor explosion [11]. (1) Coarse mixing, (2) triggering, (3) propagation, and (4) expansion. explosions did not occur. This model is validated by comparison of those with large-scale experiments for water and molten core materials in LWRs [13]. The mechanism of explosions triggered at the bottom of a pool of liquid (i.e., base-triggered explosion) was experimentally investigated by dropping tin into water [14]. The effect of the tin temperature, water temperature, water depth, and distance between the walls at the bottom of the pool was studied. The experiments showed that basetriggered explosions occurred at the bottom surface of the tank when water temperature was near saturation temperature. The occurrence of base-triggered explosion was also affected by the water depth. The interfacial behavior between high temperature molten liquid and low temperature water is experimentally investigated by using a molten material droplet and external pressure pulse [15]. It is indicated

7 6 I. Kataoka that spontaneous vapor explosion hardly occurs in high temperature water near saturation temperature since vapor film is stable. The vapor explosion can occur even in high temperature water near saturation temperature when external pressure pulse is applied to high temperature molten material. Vapor explosion cannot occur when the interfacial temperature between the molten material and water is lower than the material melting temperature. The possibility of the vapor explosion can be judged by the interfacial temperature and the molten material temperature. The results are applied to the large-scale experiments using uranium dioxide. The possibility of the vapor explosion of the uranium dioxide and water under the present LWR operational condition is extremely unlikely [16,22]. The criteria are also applicable to the case of containing metal component. A detailed analytical model to explain the vapor film collapse was developed to evaluate the conditions of self-triggering vapor explosions [11]. A simple correlation for the stability boundary is proposed by simplified model. The difference in cold-liquid temperature at the stability boundary is less than 1 K when the condensation heat transfer coefficient is over 10 4 W/m 2 K and the hot-liquid temperature is lower than 20008C. The unit sphere concept was adopted to predict the triggering stage of vapor explosions for coarse mixtures composed of hot-liquid droplets, cold liquid, and its vapor [17]. The triggering occurred at smaller water subcooling for alumina droplets and water than corium droplets. Vapor explosions were suppressed when the ambient pressure was elevated up to approximately 0.5 MPa. A new fragmentation model is proposed with three kinds of time scale for modeling instant fragmentation, spontaneous nucleation fragmentation, and normal boiling fragmentation [18]. The energetics of ex-vessel vapor explosion is investigated based on different fragmentation models. A higher pressure peak and a larger mechanical energy conversion ratio are obtained by spontaneous nucleation fragmentation. A smaller energy conversion ratio results from normal boiling fragmentation. When the delay time in thermal fragmentation model is near 0.0 ms, the pressure propagation behavior tends to be analogous with that in hydrodynamic fragmentation. If the delay time is longer, pressure attenuation occurs at the shock front. The high energy conversion ratio (44%) is obtained in a small vapor volume fraction. The results are consistent with FCI experiments with alumina melt. The propagation and expansion stages of vapor explosion are numerically simulated based on both hydrodynamic and thermal fragmentation mechanisms [19]. The thermal fragmentation model gives much higher pressure peak and higher energy conversion ratio than the hydrodynamic fragmentation. A low energy conversion ratio is expected in a large-scale vapor explosion because the hydrodynamic fragmentation is dominant when the pressure wave becomes strong. Fundamental experiment was performed on the vapor explosion with a mass of grains of certain particle sizes which simulate the molten fuel fragments [20]. In case of using water as cold liquid, boiling pressure showed the oscillations of higher frequency than 100 Hz with particle sizes ranging mm. The initial temperatures of grains and water showed little effect on generating oscillations in this test. The containment failure probability due to exvessel steam explosions was evaluated for Japanese BWR and PWR model plants [21]. A stratified Monte Carlo technique was applied for the evaluation of steam explosion loads using steam explosion simulation code JASMINE. Steam explosion is assumed in the pedestal area or in the suppression pool of a BWR model plant with a Mark-II containment, and in the reactor cavity of a PWR model plant. The generation of steam explosion loads and the containment failure were assumed at the penetration in the containment. The conditional containment failure probability (CCFP) was based on the failure of molten core retention within the reactor vessel, relocation of the core melt into the water pool without significant interference, and a strong triggering at the time of maximum premixed mass. The obtained mean and median values of the CCFP were (mean) and (median) for the BWR suppression pool case, (mean) and (median) for the BWR pedestal case, and (mean) and (median) for the PWR cavity case. A simplified method to evaluate the premixing molten core mass is proposed, based on the molten jet breakup length and settling/cooling dynamics [22]. The method was applied to KROTOS experiments at JRC Ispra with alumina and corium melts, as well as to a prototypic geometry of PWR. The calculation on the KROTOS showed that the corium melt makes a premixture containing much less fraction of molten droplets than alumina, explaining that corium is hard to make a steam explosion. The calculation with PWR ex-vessel conditions showed that the particle mass increased much slower than the total molten core ejected to the water pool. Sensitivities on the assumed molten core particle size, the void fraction, and the jet breakup length were examined in the PWR case. With vapor explosions, a large degree of numerical diffusion may occur, and local information may be lost. Pseudo-diffusions which act as high temperature heat sources affects component distributions. Numerical divergence is also often caused by fluctuation of these high-temperature components. Two treatment methods are proposed to avoid these problems [23]. One method is a dispersed component method, which was developed for dispersed components. The other is a multi-region scheme, where a calculation domain is

8 Journal of Nuclear Science and Technology, Volume 50, No. 1, January divided into a multiphase mixture region and a singlephase continuum region. The effectiveness of these treatments is demonstrated by numerical simulations. The interaction of molten metal drop and coolant is numerically analyzed to investigate the mechanism of fragmentation in vapor explosion [24]. The numerical study is carried out by using a droplet simulation code based on multiphase thermal-hydraulic model. Several computational techniques are also implemented to improve the efficiency and stability of the numerical scheme. The results show that the growth of spike on the molten metal drop surface is similar to that observed in experiment. The result suggests that quick growth of spikes is the essential mechanism of fragmentation, which is caused by Taylor instability. 4. In-vessel retention As a result of core melt and relocation, finally the molten core is accumulated at the bottom of pressure vessel forming a pool of molten core. If this molten core pool and lower head of pressure vessel are not adequately cooled, lower head of pressure vessel breaks at the guide tube of control rod and/or instrumentation tube. Then, molten core is ejected or fallen into the containment vessel and possibility of release of radioactive material to environment increases. Therefore, it is quite important to prevent the break of lower head of pressure vessel and to retain the molten core in bottom of pressure vessel. In view of this, as a SA management, it was proposed to cool the outer surface of lower head pressure vessel by water spray or flooding by water. If the cooling is sufficient, the integrity of pressure vessel is kept and molten core is retained in the bottom of pressure vessel. This accident management is called as IVR. Conceptual representation of IVR is shown in Figure 2. Usually, molten core is considered to be stratified in two through four layers. In Figure 2, the case of two layers is shown. Low density liquid metal (structural material) is located in upper layer and high density metal oxide (UO 2 fuel) is located in lower layer. The configuration of this stratification determines the heat flux distribution in lower head of pressure vessel and success of failure of IVR. When the lower head of pressure vessel is immersed in water in containment vessel, boiling occurs at the outer surface of lower head. Boiling heat transfer and limit of heat removal (critical heat flux, CHF) from the downward facing hemisphere are quite important in IVR. Also, in molten core pool, natural convection occurs in the layer of molten oxide due to the decay heat of fuel. Natural convection heat transfer and hydrodynamic behavior of molten core are the other important subjects in the analyses of IVR. Furthermore, if there is water above molten core pool, molten core debris is cooled by various ways. Water penetrated into narrow gaps in molten core debris or cracks in the crust of molten core. Heat transfer to water in such configurations is also quite important in coolability of molten core in IVR. Researches on these subjects have been carried out. The heat transfer of the inversely stratified molten corium in the lower vessel, which was experimentally demonstrated in the MASCA project (Organization for Economic Co-operation and Development, OECD/ The Nuclear Energy Agency, NEA joint project), was analyzed [2]. For the oxide layer, turbulent models of the k e and the large eddy simulation were examined through the French Alternative Energies and Atomic Energy Commission, CEA BALI test. A melt-solidification model was incorporated for the metal layer analysis. An analysis under an inversely stratified configuration shows that the peak heat flux from the corium does not exceed the critical heat flux of the flooded vessel. However, the heat flux focuses at the top of the lower metal layer because it is under the thermally stable condition, and this focusing would be a new challenge for the IVR. A thermodynamic corium database was developed and stratification of molten corium was analyzed [25]. The database consists of U Zr Fe O C B (fission product, FP oxides) system. Fundamentally, data were obtained from existing databases, such as Scientific Group Thermodata Europe, SGTEs. The liquid phase data were reconstructed based on the ionic model, and lacking data including excess energies were assessed to be consistent with existing phase diagrams. Liquids temperatures measured under OECD RASPLAV project were analyzed with the database. An analysis of corium under a SA condition was carried out and demonstrates that the database gives an improved method based on thermodynamics to analyze the corium stratification. The molten corium stratification tested in the OECD MASCA project was analyzed with thermodynamic database and the database was effective for the stratification analysis [26]. The MASCA test shows that the molten corium can be stratified with the metal layer under the oxide in the lower head of the reactor vessel. This stratification is caused by the increased density of the metal layer attributed to a transfer of uranium metal. Thermodynamic equilibrium calculations with the database for the corium U Zr Fe O B C FPs system agree with the MASCA test. Boron carbide influences on thermodynamic properties and phase separation of molten corium were estimated with U Zr Fe O B C FPs thermodynamic database [27]. The liquid temperature of the oxide for the typical corium was estimated to increase by 1008C with B 4 C addition when the corium included up to 10 wt% Fe. On the other hand, the liquid temperature was hardly changed when the corium included 50 wt% Fe. The interaction temperature between the steel and the corium with B 4 C was estimated at 1130 K. The estimated temperature is over 200 K higher than the criterion temperature where the vessel loses its structural strength. Other thermodynamic influences of B 4 C

9 8 I. Kataoka were also estimated as not having a negative impact on the IVR. Research on IVR for fast breeder reactor is also carried out [28] of which knowledge is useful in analyses of IVR in LWR. In the Japan Sodium Cooled Fast Reactor design, elimination of severe power burst events in the core disruptive accident is intended as an effective measure to ensure retention of the core materials within the reactor vessel. The design strategy is to control the potential of excessive void reactivity insertion and to exclude core-wide molten-fuel-pool formation by introducing an inner duct. The effectiveness of these measures is evaluated based on existing experimental data and computer simulation. Phenomenological consideration of general characteristics and preliminary evaluations for the long-term material relocation and cooling phases gave the perspective that IVR would be attained with appropriate design measures. A numerical method for the thermal-hydraulic phenomena in a narrow flow passage is developed to evaluate the gap cooling capability [29]. Based on a drift flux model, the two-dimensional gas liquid two-phase flow in the annular and hemispherical heated narrow flow passages is modeled. Experiment on thermalhydraulic phenomena in the heated narrow flow passage is performed. The critical heat flux data is obtained from measurement of the heating surface temperature. Counter-current two-phase flow is reproduced by the numerical analysis appropriately. The critical heat flux is well predicted based on flooding model. Validity of the newly developed numerical method is demonstrated through comparison of the experimental critical heat flux with the existing data in the gap width ranging from 0.5 to 5 mm and the pressure ranging from 1 to 50 bar. Experimental studies supposing in relation to IVR are made on the pool boiling heat transfer from hemispherical downward facing heating surface [30,31]. This experiment is the down-sized simulating experiment of external cooling by cooling the lower head of pressure vessel, which is the hemispherical downward heating surface with penetrations rods. The characteristics of pool boiling heat transfer from downward heating surface and obstacle effects of penetrations were investigated. Water under the atmospheric pressure is used as a coolant, and molten lead is used as a substitute of molten core. Direct observation by using digital video camera makes clear the behavior of the vapor film or bubbles at the various boiling region such as a film boiling, transient boiling, and nucleate boiling regions. Time series of temperature and heat flux during the quenching using thermocouples were also measured to get the boiling curve, and effects of obstacles and inclination angle of heating surface on characteristics of boiling heat transfer were evaluated. Experiments were carried out by changing the subcooling of the flooding water and the size of the vessel. The direct observations by using the digital video camera were performed and made clear the special characteristics taking place near the heating surface during the quenching process. The measurements for the temperature distribution, the wall superheat, and surface heat flux by using the 15 thermocouples placed inside of the vessel wall were also carried out to make clear the boiling heat transfer during the quenching process. Boiling transitions around the downward hemispherical heating surface with and without obstacles (simulating penetration) are shown in Figures 6 and Direct containment heating If the failure of the bottom of pressure vessel (melt through) occurs, the molten core is ejected into the bottom of containment vessel. In some scenario of SA initiated by total station blackout (denoted as TMLB ), pressure of primary loop and pressure vessel remains very high due to failure of depressurization. In such circumstance, molten core is ejected very rapidly into containment vessel as a form of jet (HPME). After the ejection of molten core, a large amount of high pressure vapor flows into the containment vessel. Molten core is dispersed into fine droplets due to the jet impingement onto the floor of containment vessel and entrainment for the liquid film due to the high speed vapor flow. These fine droplets of molten core are transported from the cavity to whole area of containment vessel. Since the temperature of molten core droplets is very high, containment vessel atmosphere is rapidly heated resulting in rapid pressure increase of containment vessel. Furthermore, high pressure and high temperature vapor from pressure vessel also increase the temperature and pressure of containment vessel. Therefore, temperature and pressure of containment vessel soon exceed the design limit which causes the destruction of containment vessel. This event is called as DCH. In NUREG-1150, the probabilistic safety assessment (PSA) report, this event is first considered as one of the important events of SA. After that, a lot of experimental and analytical researches have been carried out in US, Europe, and Japan. From the results of these researches, DCH is unlikely to occur if the pressure of pressure vessel is successfully depressurized under 2 MPa. This value was adopted as a safety guideline of DCH in Japan. A multi-component thermal-hydraulic model of the DCH was developed to predict the pressure loads to the containment during HPME from the RPV [32]. Using this model, parametric analyses of DCH in BWR Mark-II containment were carried out to identify the parameters and mechanisms controlling the pressure load history to the containment. The analysis showed that the peak pressure is strongly influenced by the heat source by chemical reaction between metals in debris and steam in the containment, which is dependent on the initial mass of steam, size of

10 Journal of Nuclear Science and Technology, Volume 50, No. 1, January Figure 6. Boiling phenomena on downward facing surface [30]. (a) Film boiling; (b) transition boiling, early; (c) transition boiling, later; (d) nucleate boiling, high heat flux; (e) nucleate boiling, medium heat flux; and (f) nucleate boiling, low heat flux. Figure 7. Boiling phenomena on downward facing surface with obstacles [30]. (a) Film boiling, early; (b) film boiling, later; (c) transition boiling; (d) nucleate boiling, high heat flux; (e) nucleate boiling, medium heat flux; and (f) nucleate boiling, low heat flux. debris particles, and fraction of suspended debris particles. The analytical model of DCH process was described and discussions were made on the mechanisms that control the pressure loads to the containment on the bases of sensitivity calculations on various parameters. During a station blackout of PWR, the pump seal will fail due to loss of the seal cooling. This particular transient LOCA sequence analyzed for Surry plant showed that the depressurization due to the pump seal LOCA would result in early accumulator injection and subsequent core cooling which lead to the delay of RPV meltthrough [33]. The analysis was performed with SCDAP/RELAP5 to evaluate this scenario shown in the analyses. The calculated results were compared with the similar experimental studies of JAERI s

11 10 I. Kataoka ROSA-IV program. The loop seal clearing would occur and cause a slight delay of accident progression. It is unlikely that the accumulator injection, which leads to the delay of RPV meltthrough by approximately 60 min, is initiated automatically. An intentional depressurization using power operated relief valve, PORVs is recommended for the mitigation of the accident consequences. The present SCDAP/ RELAP5 analyses did not show significant delay of accident progression. In DCH, mechanisms of dispersal of molten core and droplet size are the most important. Then, separate test of DCH basic phenomena was carried out [34]. To study the mechanisms of the corium dispersion phenomenon, a test facility of 1:10 linear scale for Zion PWR geometry is constructed. Experiments are carried out with air water and air-woods metal simulating steam and molten core materials. The physical process of corium dispersion is studied in detail through various instruments, as well as with flow visualization at several locations. Comprehensive measurements are obtained including the liquid jet velocity, liquid film thickness, and velocity transients in the test cavity, gas velocity, and velocity profile in the cavity, droplet size distribution and entrainment rate, and the fraction of dispersed liquid in the containment building. These data are of great importance for better understanding of the corium dispersion mechanisms. In relation to such separate effect tests of molten core dispersal, mechanistic model of droplet diameter and its distribution were developed and correlations were proposed and applied to the analyses of DCH [35]. 6. Molten core concrete interaction If the failure of the bottom of pressure vessel (melt through) occurs and the pressure of pressure vessel is not so high or the break is not so large, DCH does not occur and the molten core falls onto the floor of the bottom of containment vessel and forms a pool. This pool can be stratified in two layers of molten metal and molten oxide fuel as shown in Figure 8. The molten core reacts with concrete wall of the bottom of Figure 8. Conceptual representation of MCCI [36]. containment vessel and erodes the concrete wall. This event is called as MCCI and quite important in maintaining the integrity of the containment vessel. Due to the reaction between molten core and concrete, decomposed gas is generated. Therefore, complicated two-phase natural circulation occurs in the pool of molten core. The two-phase flow behavior and mass and heat transfer characteristics between molten core and concrete are quite important in analyzing MCCI. Researches on these phenomena were carried out. The analysis for the WITCH/LINER experiments was performed to investigate the heat transfer characteristics between the gas-agitated steel melt and the vertical surface [36]. The applicability of heat transfer correlations for a gas agitated fluid system was examined through the numerical analysis of the onedimensional heat conduction taking into account the crust formation due to the solidification of steel melt. The heat transfer correlation developed by Konsetov was modified for the application to fluids with low Prandtl number. The constant in the modified correlation was empirically determined by the experiments under churn-turbulent two-phase flow regime. The modified Konsetov correlation predicted the heat transfer characteristics observed through the experiments in an acceptable level. The FCI and MCCI have been studied experimentally in COTELS project as a joint study between NUPEC (Japan) and National Nuclear Center, NNC (Republic of Kazakhstan) using one of the testing complexes at NNC [37 41]. The testing complex includes three experimental facilities SLAVA, LAVA, and LAVA-M for debris coolability tests. Three types of experiments were carried out. To get the molten corium, the electric induction melting furnace (EMF) was used. The EMF produced 60 kg of corium containing UO 2, SS, Zr, and ZrO 2. The temperature of the produced melt was about 3200 K. The melt was discharged into the water pool in Test A or onto the concrete trap in Test B/ C. The corium in the concrete trap was heated in Test B/ C by another induction melt heater. Prior to main Test A and Test B/C, several supporting experiments were conducted [37]. Integrity of graphite crucible with TaC sheet during producing UO 2 corium was confirmed experimentally. The induction melt heater was calibrated and the efficiency for the induction heater of LAVA-M facility was determined as 47%. The thermal conductivity and thermal diffusivity of concrete up to about 1073 K, and melting and solidification points of corium components were determined experimentally. Discharge corium behavior, using UO 2 corium, was also observed by high speed cameras in Test 01. Test B/C in COTELS project was performed to investigate FCI and MCCI under water injection onto molten corium [38]. A 60 kg UO 2 corium mixture was fallen into concrete trap. The molten corium in the trap was reheated to simulate decay heat. When subcooled

12 Journal of Nuclear Science and Technology, Volume 50, No. 1, January water was injected onto the molten debris, steam explosion was not observed. Debris was cooled about 20 min after water injection and MCCI was suppressed. This favorable cooling was attributed to the existence of remaining porosity inside corium. Cross section of concrete trap along with solidified debris tested in COTELS Test B/C was structurally investigated [39]. In 6 tests out of 10 tests, particulate debris bed was formed above continuous ingot debris. The size distribution of the particulate debris was well correlated by Rosin Rammler equation. Large amount of smallest diameter particles was obtained. The upper region of the solidified debris included more concrete compositions. The concrete erosion depth, concrete degradation condition, and the structure of solidified debris were evaluated to clarify the basic difference between COTELS and former tests results. Concrete erosion depth was less than that observed in MACE, WETCOR, and SWISS tests. The major differences of COTELS results compared with the former test results were absence of strong adhesion of crust to melt trap side wall, water penetration into debris through both eroded side wall and channels inside ingot debris, absence of large void inside ingot debris, and formation of pebble bed below ingot debris. All of these promoted the suppression of MCCI. A series of tests were performed to examine influences of coarse aggregates in concrete, side wall ablative nature, and debris properties on the MCCI progression [40,41]. A UO 2 mixture or SS was gravitationally slumped into a concrete trap. The test results indicated that coarse aggregates mixed in concrete act as a resistance to the MCCI progression. No crust formation was confirmed by that unmolten coarse aggregates relocated by buoyancy force to the upper layer above the SS layer. An analytical code, COCO, is being developed for the evaluation of the MCCI progression and the ex-vessel debris coolability. Analysis of one of the COTELS MCCI tests implied that COCO code has a good capability to reproduce transient thermal response of the concrete trap. 7. Hydrogen explosion As a result of hydrogen generation due to metal steam interaction in relocation of molten core and MCCI, hydrogen explosion may occur in containment vessel and reactor building if the atmosphere is not inert. If the hydrogen explosion occurs, containment vessel and rector building will suffer from serious damages resulting in release of considerable amount of radioactive material to environment. Hydrogen explosion in SA is usually analyzed and predicted by codes. These codes are required to predict the distribution of hydrogen concentration and propagation of pressure wave in containment vessel and other structure of nuclear power plant. Researches on development of such codes and analyses of hydrogen explosion were carried out. Thermal and mechanical loads on the containment vessel due to vapor explosion were evaluated [42]. In order to predict hydrogen explosion, a two-dimensional numerical analysis code COMA has been developed using the explicit finite volume method and non-staggered mesh scheme. A new numerical method has been developed to treat a large pressure discontinuity at the wave front with the minimum numerical diffusion. The method was validated through a comparison with the hydrogen detonation tests in a cylindrical tube performed by NUPEC and Brookhaven National Laboratory, BNL. The calculated results agreed with observed overall effects of hydrogen concentration and initial temperature on detonation wave propagation velocities. Analyses were made on the hydrogen explosion in the Hamaoka Nuclear Power Station Unit-1 [43,44]. A pipe rupture occurred in the steam condensing line of the residual heat removal system. The explosion of the hydrogen accumulated in the pipe was considered to be the cause of the rupture. Hydrogen and oxygen, generated by radiolysis of reactor water, accumulated at the downstream end were distributed along the accumulated region. At the boundary between the accumulated noncondensable gas and the steam regions, temperature and concentration fluctuated due to operation of the high pressure core injection, HPCI valve. The three-dimensional hydrogen combustion behavior was first solved with the one-step irreversible overall reaction model. The ignition point was given at the upstream boundary surface of the non-condensable gas region. The temperature and concentration distributions in the pipe, which were obtained from the analysis, were given as the initial conditions. The analysis result showed that the detonation pressures in the straight pipes were about 120 MPa and the peak pressures at the elbows were times higher than those in the straight pipes, due to reflection and overlapping of the pressure waves. Then, a threedimensional dynamic response of pipe deformation was analyzed, with the time transients of the pressure distribution as boundary conditions inside the pipes. The result showed that the strain at the elbow exceeded the critical strain and the pipe ruptured there. This was generally consistent with the result of actual pipe deformation observed after the accident. Hydrogen explosion occurred in Fukushima accident in reactor building of Unit 1, Unit 3, and Unit4 [9,45]. Review was made on these events but the cause and reason of these events have not been completely clarified yet. 8. SA analysis code Since SA is a sequence of many complicated events, the safety evaluation and prediction of SA are usually carried out by computer codes. Such SA analysis codes