CONDENSATION OF A LARGE-SCALE BUBBLE IN SUBCOOLED LIQUID: EXPERIMENTAL VERIFICATION OF THE SIMMER-III CODE

Transcription

1 THAS5: Fifth Korea-Japan Symposium on uclear Thermal Hydraulics and Safety Jeju, Korea, ovember 26-29, 26 THAS5-B7 CODESATIO OF A LARGE-SCALE BUBBLE I SUBCOOLED LIQUID: EXPERIMETAL VERIFICATIO OF THE CODE Koji Morita 1*, Tatsuya Matsumoto 1, Kenji Fukuda 1, Yoshiharu Tobita 2, Ikken Sato 2 and Hidemasa Yamano 2 1 Graduate School of Engineering, Kyushu University 744 Moto-oka, ishi-ku, Fukuoka , Japan Phone : , morita@nucl.kyushu-u.ac.jp 2 Advanced uclear System R&D Directorate, Japan Atomic Energy Agency 42 arita, O-arai, Ibaraki , Japan ABSTRACT Experimental verification of a reactor safety analysis code,, was performed for relatively shorttime-scale multi-phase flows with simultaneous phase transition. Two independent series of experiments were analyzed for transient behaviors of large-scale bubbles with condensation. One is a new series of experiments dedicated to fundamental phenomena of large-scale bubble condensation with noncondensable (C) gases. The other is a series of high-pressure experiments discharging nitrogen and low quality water, which was conducted by Purdue University in the late 197 s using the Omega facility with a 1/7-scale model of Clinch River Breeder Reactor. The characteristics of transient behavior of large-scale bubbles with condensation observed in the experiments were estimated through experimental analyses using, especially in terms of the effect of C gas on condensation. In addition, the code applicability was presented to the reactor accident conditions with multi-phase, multi-component hydraulic problems, among which vaporization and condensation, or simultaneous heat and mass transfer, play an important role. 1. ITRODUCTIO A concern with ability to reasonably predict consequences of even the most unlikely accident involving liquid-metal cooled fast reactors (LMFRs) has led to detailed studies of postulated core disruptive accidents (CDAs). This is because mechanistic simulation of an accident sequence during a postulated CDA is realized only by using a comprehensive computational tool that systematically models multi-phase thermal-hydraulic and neutronic phenomena occurring during what are known as the transition and expansion phases of CDA. In this area, the development of a reactor safety analysis code,, has been conducted at the Japan Atomic Energy Institute (JAEA) in collaboration with the Forschungszentrum Karlsruhe (FZK) in Germany, and the Commissariat à l'energie Atomique (CEA) which also includes the Institute de Radioprotection et de Sûreté ucléaire (IRS) in France (Tobita et al., 22). In the past several research groups performed experiments to study an expanding high-pressure gas or vapor bubble in a vessel filled with liquid. This is because the injection of a high-pressure gas into a stagnant liquid pool is a characteristic phenomenon during the expansion phase of CDA in LMFRs. Moszynski et al. (198) reviewed relevant experiments performed up to 198. In these experiments, a single- or twophase fluid (gas or superheated liquid) stored in a pressure vessel was discharged into a liquid pool by rupturing diaphragms, with which the liquid pool was separated from the pressure vessel. In this context, a new series of transient bubble behavior experiments was performed to verify the basic applicability of for relatively short-time-scale multi-phase flows with simultaneous phase transition. A system used in the present experiments is similar in form of a basic apparatus used in the past experiments mentioned above. In the present experiments, a large-scale single bubble of noncondensable (C) gas and steam mixture was discharged into a cylindrical pool filled with stagnant subcooled water. The characteristics of transient behaviors of large-scale bubbles with condensation observed in the experiments were also estimated through experimental analyses using. In the present study, one of the past experiments on the expansion of high-pressure bubble in a stagnant liquid pool was also analyzed. They are the experiments discharging nitrogen and low quality water in the Omega facility, which was conducted by Purdue University in the late 197 s using a 1/7-scale model of Clinch River Breeder Reactor (CRBR) (Theofanous et al., 1979; Saito et al., 1979; Simpson et al., 1981). Hereafter, we call them OMEGA tests. In addition, the code applicability to a reactor accident condition was presented for the expansion phase of CDA in a medium-sized sodium cooled fast reactor, where multi-component vaporization and condensation or simultaneous heat and mass transfer play an important role. 2. PHASE-TRASITIO MODELS OF is a two-dimensional, three-velocity-field, multiphase, multi-component, Eulerian, fluid-dynamics code coupled with a fuel-pin model and a space- and energydependent neutron kinetics model (Tobita et al., 22). The fundamental equations of the fluid dynamics are based on a so-called multi-fluid model. The code models the nonequilibrium phase-transition processes occurring at fluid interfaces. In particular, the multi-component vaporization/ condensation (V/C) model characterizes the V/C processes

2 associated with phase transition by employing heat-transfer and mass-diffusion limited models (Morita et al., 23). The physical model to represent the effect of C gas and multicomponent mixture on V/C processes is based on a study performed originally for SIMMER-II (Bohl et al., 199). The equations for this model were obtained by considering the quasi-steady, stagnant Couette-flow boundary layer, as shown Fig. 1, to relate the mass and energy fluxes to the overall forces driving heat and mass transfer. A multicomponent vapor mixture at bulk temperature T and mass fraction k, (k = 1 ) is under the phase transition on a liquid or solid phase, which is maintained at a constant temperature T o. Here, the conservation of each vapor species can be described based on multicomponent diffusion law given, for example, by Bird et al. (196). Assuming that mass diffusion due to thermal and pressure gradient is negligibly small, the mass-transfer rate k of vapor component k at the interface i, defined per unit volume and positive for condensation, is governed by k = a i k * k ( k,i k, ) + k,i j (1) * where a i is the interfacial area per unit volume, k k is the mass-transfer coefficient of vapor component k and k,i is the mass fraction of vapor component k at the interface. In the * presence of mass transfer, k k is modeled by (Bird et al., 196) k * k = j a i exp j a i k k 1 (2) It is noted that the right side of Eq. (1) includes both diffusive and convective contribution. The heat flow per unit volume at the interface should include contribution of heat conduction, bulk convection and diffusion. Thus, approximating the temperature gradients by overall heat-transfer coefficients, energy balance applied to the interface yields j i lg,j = a i h * g T ) + h o T o ) (3) * where i lg is the latent heat of vaporization, h g is the vaporside heat-transfer coefficient in the presence of mass transfer, h o is the liquid or solid-side heat-transfer coefficient and T i is the interface temperature. The effect of mass flow through the boundary layer on the vapor-side heat-transfer coefficient h g can be accounted for by Bird et al. (196): h * g = j c p,j a i exp j c p,j a i h g 1 (4) where c p is the specific heat capacity. Equation (3) represents that the heat flow at the interface equals the sum of the latent heat flow and the sensible heat flow through the interface. Equations (1) and (3) can be combined into the following single algebraic equation: k * j ( j,i j, )i lg,j + j,i i lg,j k * j ( j,i j, ) 1 j,i = h * g T ) + h o T o ) (5) A solution of Eq. (5) in terms of T i is evaluated by assuming a constant pressure through the boundary layer to the interface. The condensed phase at the interface is also assumed to be in saturated thermodynamic equilibrium with the vapor component, of which saturation pressure in the immiscible system is independent of its concentration in the condensed phase. T i T o ω k,i ω k, Liquid or solid phase p sat,k ) condensable-gas pressure at interface Fig. 1 Interface Γ k T Vapor phase p k, condensable-gas pressure p ng, C-gas pressure Basis of mass-diffusion process 3. BASIC EXPERIMETAL VERIFICATIO Basic applicability of to relatively short-timescale multi-phase flows with simultaneous phase transition was investigated by a new series of experiments, which is dedicated to fundamental phenomena of transient large-scale bubble condensation with C gas. 3.1 Experimental Procedure A schematic view of the experimental apparatus and overall arrangement including key dimensions are illustrated in Fig. 2. The basic apparatus consists of a pressure vessel and a cylindrical inner pool (inner diameter 31 mm and height 75 mm), which is surrounded by a larger quadrate outer pool. The confinement vessels of these pools are made of transparent acrylic resin. The top of the inner water pool was connected to a transparent pipe with an inner diameter of 57 mm. The inner pool was filled with stagnant water at room temperature up to the upper pipe with a height of around 2 ~ 3 cm from the top of the inner pool. The transient gas volume in the inner pool and the pressure vessel with the adiabatic region was determined from displacement of the water level in the upper pipe. The outer pool was also filled with water to improve visual observation inside the cylindrical inner pool by minimizing optical distortion, which is caused by the difference in refractive indexes between water and acrylic resin. The pressure vessel made of stainless steel was connected to the bottom of the inner pool through the adiabatic region. The exits of the adiabatic region and the pressure vessel were initially closed with upper and lower rupture disks, respectively. In the pressure vessel, whose effective capacity was about 6 cm 3, a mixture of steam and C gas was generated and pressurized by boiling water of 8. cm 3 using an electric heater. The adiabatic region giving an approximate volume of 15 cm 3 was initially evacuated down to several kpa by a vacuum pump. This can reduce the heat transfer from the pressure vessel to the inner pool during the pressurization of gas mixture. Two pressure sensors were installed to measure the pressure transients in the pressure vessel and at the top of the water pool. After the two rupture disks break, the gas mixture was

3 discharged from the pressure vessel into the water pool, expanding in the inner pool. Two high-speed cameras, both of which can record 4 frames per second, were used to record the movements of the gas phase in the inner pool and the water surface level changes in the upper pipe, respectively, as digital motion pictures. Two rupture disks used to seal the pressure vessel and the adiabatic region at the beginning of the experiment were diaphragms hand-made of polyester film. The lower disk was designed to rupture at a desired pressure, while the upper disk, which is thinner than the lower one, has strength only enough to stand water head. The break of the lower rupture disk was triggered when the pressure in the pressure vessel reaches the rupture limit of the diaphragm during the pressurization of the gas mixture by electric heating. After the lower rupture disk breaks, the initially pressurized gas mixture spurts into the adiabatic region. The upper rupture disk also breaks subsequently and then the gas mixture starts expanding inside the water pool. Main initial conditions of the experiments are shown in Table 1. Concentration of C gas in the mixture was taken as an experimental parameter as well as species of C gas. In the present experiments, nitrogen and xenon were used as a C gas mixed with steam because of difference in a mass-transfer resistance for the diffusion of steam. The experiments except for the case only using nitrogen gas were performed with similar rupture pressures of about.3 MPa. The water subcooling in the inner pool was in the range K. In the pressure vessel, the initial gas mixture with water boiling had a very high void fraction, which could suppress two-phase flashing during the gas expansion process. Table 1 Initial conditions of transient bubble behavior experiments C gas mixed with steam itrogen Xenon C gas concentration (mol%) Pressure in pressure vessel (MPa) Steam saturation temperature (K) Void fraction in pressure vessel (%) Pressure in adiabatic region (kpa) Water pool temperature (subcooling) (K) (116) (111) (19) (---) (117) (11) Water level in upper pipe (cm) Fig. 2 Schematic view of experimental apparatus for transient large-scale bubble condensation 3.2 Experimental Results Gas volume changes defined here correspond to the differences between the instantaneous gas volumes in the experimental processes and the initial gas volumes at the beginning of the experiments. Experimental data for the total gas volume changes in both the water pool and the pressure vessel with the adiabatic region were obtained from the water level changes in the upper pipe of the experimental setup. Figure 3 shows the experimental results of the gas volume changes of the four cases, in which different nitrogen-gas concentrations were used. Here, the time zero is defined by the triggering of the lower diaphragm rupture. The results of all cases show that once the lower and upper rupture disks break, the gas volume begins to increase with the expansion of the originally pressurized gas in the pressure vessel, while not more than a hundred millisecond the gas volume starts to decrease. This occurs because, as the gas expands, the pressure of the gas becomes lower than the atmospheric pressure due to the steam condensation and/or the inertia of the water slug moving upward. When the gas pressure decreases to a value that is smaller than the surrounding pressure, the gas will be compressed, and as a result its volume decreases. By comparing the experimental results between the four cases shown in Fig. 3 it can be seen that when the nitrogen-gas concentration is larger, the corresponding gas volume change, especially the maximum gas volume change, is much larger. The experimental results also show that the larger the nitrogen-gas concentration, the longer the period of the gas expansion. The effect of thermophysical properties of C gas on condensation was also investigated by comparison between steam-nitrogen and steam-xenon systems. Figure 4 indicates the experimental results of the gas volume changes in the cases using the C gas of around 1 and 2 mol%. As shown in Fig. 4 the expansion period of the gas in the steam-xenon

4 system becomes longer than that in the steam-nitrogen system under a similar concentration of C gas although the maximum gas volume change in the steam-xenon system is comparable to that in the steam-nitrogen. This might be because a mass-transfer resistance for the diffusion of steam toward the condensation interface in the steam-xenon system is larger than that in the steam-nitrogen one due to the larger molecular weight of xenon than that of nitrogen. Fig. 3 Fig. 4 Fig. 5 1, 5-5 C gas : itrogen.4 mol% 1.3 mol% 19.5 mol% 1 mol% -1, , 5 Experimental results of gas volume change: effect of nitrogen gas concentration itrogen 1.3 mol % 19.5 mol % Xenon 11.7 mol % 2.8 mol % Experimental results of gas volume change: effect of C gas properties Comparison of bubble shape change in water pool: effect of nitrogen gas concentration In Fig. 5, images of the gas phase in the water pool are shown for the cases using a different initial concentration of C gas at selected time intervals. In these figures, the time zero is defined by the instance when the gas mixture is released into the water pool. The experimental images obtained using a high-speed camera represent a 18 mm (width) 15 mm (height) region, of which base corresponds to the bottom of the water pool. As can be seen from Fig. 5, in the case using steam mixed with a very small amount of nitrogen gas the gas mixture discharged into the water pool disperses like bubbly flows due to rapid condensation, and forms no large-scale bubble as observed in the case using only nitrogen gas. In the case using steam mixed with nitrogen gas of 1.3 mol%, the images also indicate the similar behavior around the bubble surface, where dispersive mixing of the gas and liquid phases is observed. 3.3 Experimental Analyses In the calculations, five regions of the pressure vessel, the adiabatic region, the inner water pool, the upper pipe and the atmosphere were modeled to represent the corresponding experimental setup shown in Fig. 2 by a twodimensional cylindrical geometry. For the water pool region, 18 radial and 75 axial computational cells were used. Simulation results of the gas volume change and the pressure in the pressure vessel as the reference case using only nitrogen gas are indicated in Figs. 6 and 7, respectively, in comparison with the corresponding experimental results. well reproduces the experimental results of the transient behavior of nitrogen-gas bubble expansion. 1, 5 C gas : itrogen Exp: 1 mol% Fig. 6 Comparison of gas volume change between experimental result and the reference case using only nitrogen gas Pressure in pressure vessel (MPa) C gas : itrogen Exp: 1 mol% Fig. 7 Comparison of pressure in the pressure vessel between experimental result and the reference case using only nitrogen gas In the present system, during discharge of the gas mixture from the pressure vessel to the water pool, steam condensation can occur mainly on both the cold structure surface of the adiabatic region and the bubble surface in the subcooled water pool. On the other hand, simulation results using

5 suggested that much more condensation occurs at the gasliquid interface rather than at the gas-structure interface as the nitrogen-gas concentration increases, while the condensation at the gas-liquid interface is reduced due to a buildup of nitrogen gas at which condensation occurs. Therefore, here the inhibiting effect of C gas will be discussed for the condensation at the gas-liquid interface. Figure 8 shows the simulation result of the gas volume change in the case using nitrogen gas of.4 mol% in comparison with the corresponding experimental result. In this figures, two parametric results of the simulations are presented. One case considers the nitrogen-gas effect on condensation at both the gas-liquid and gas-structure interfaces. The other considers no nitrogen-gas effect at the gas-liquid interface. In the latter case, the condensation rate at the gas-liquid interface is evaluated by assuming that the interface temperature is equal to the saturation temperature of bulk vapor. As can be seen from this figure, regardless of fairly small amount of nitrogen gas, considering the nitrogen-gas effect underestimates the condensation rate. On the other hand, the simulation with no gas effect fairly well reproduces the experimental result represented by the gas volume change. The results of the gas volume change in the cases using nitrogen gas of 1.3 and 19.5 mol% are indicated in Fig. 9. Good agreements with the experimental results were also obtained in the simulations considering no nitrogen-gas effect at the gas-liquid interface. The present results suggest that under the present experimental conditions the nitrogen gas has less effect on steam condensation at the gas-liquid interface. 5-5 C gas : itrogen Exp.:.4 mol% itrogen-gas effect o nitrogen-gas effect -1, Fig. 8 Fig. 9 1, 5 Effect of nitrogen gas on condensation: nitrogen gas concentration of.4 mol% C gas : itrogen Exp.: 1.3 mol% o nitrogen-gas effect itrogen-gas effect Exp.: 19.5 mol% o nitrogen-gas effect itrogen-gas effect Effect of nitrogen gas on condensation: nitrogen gas concentration of 1.3 and 19.5 mol% Figure 1 shows the comparison of the simulation results in the cases using xenon gas of 11.7 and 2.8 mol%. The simulations considering no xenon-gas effect at the gas-liquid interface result in fairly better agreement with the experimental results. It seems that also represents the effect of thermophysical properties of C gas on steam condensation reasonably. 1, 5 Fig. 1 C gas : Xenon Exp.: 11.7 mol% o xenon-gas effect Xenon-gas effect Exp.: 2.8 mol% o xenon-gas effect Xenon-gas effect Effect of xenon gas on condensation: xenon gas concentration of 11.7 and 2.8 mol% The present numerical simulations suggest that there is less inhibiting effect of C gas on vapor condensation of largescale bubbles discharged into a stagnant subcooled liquid. This is a characteristic different from quasi-steady condensation of small-scale bubbles observed in our previous experiments (Suzuki et al., 23), where the condensation rate was reduced dominantly by C gas mixed with steam. In the present experiments with steam condensation, as can be seen from Fig. 5, dispersion of the gas mixture discharged into the liquid pool was accompanied by the steam condensation at the bubble surface. This might be relevant to one of the liquid entrainment mechanisms in the bubble expansion process, the condensation-induced instability or condensation shock (Moszynski et al., 198; Simpson et al., 1981). Such dynamic mixing of the gas and liquid phases induced by condensation would prevent a buildup of the C gas on the bubble surface. This results in the less inhibiting effect of the C gas on condensation. The comparison with our previous experiments using small-scale bubbles (Suzuki et al., 23) suggests that the characteristics of condensation observed in small-scale bubbles disappear in large-scale bubble, where local twophase mixing might not develop around its surface. 4. ITEGRAL EXPERIMETAL VERIFICATIO Integral code assessment is necessary for the code application to reactor safety analyses, where the pressure condition is significantly higher than that in the basic verification experiments described in the previous chapter. For this purpose, verified with the low-pressure experiments is extensively applied to the following reactor safety experiments. 4.1 Overview of the OMEGA Test The OMEGA test is a series of out-of-pile experiments conducted at Purdue University in the late 197 s to investigate mechanical energy mitigation mechanisms in upper sodium pool during a post-disassembly expansion phase (Theofanous et al., 1979; Saito et al., 1979; Simpson et al., 1981). In the experiments, high-pressure nitrogen or hightemperature mixture of two-phase water and steam was injected into a stagnant water pool from the bottom. The OMEGA facility is the 1/7-scale model of the upper sodium plenum of CRBR as shown in Fig. 11. The test section consists of the pressure vessel containing high-pressure nitrogen or water/steam mixture and the water pool with cover gas. The high-pressure fluid was injected through a blowdown

6 nozzle by breaking a rupture disk, which separates the pressure vessel and water pool regions. There exists another diaphragm over the rupture disk to hold the water and prevent the water from contacting the high-temperature pressure vessel. The mechanical strength of this diaphragm is negligibly small so that it ruptures instantaneously when the high temperature fluid reaches. The space in the blowdown nozzle between the rupture disk and the diaphragm can be evacuated to insulate the high-temperature pressure vessel and the water pool. The growth of bubble in the water pool has been recorded by a high-speed camera. The level of pool surface and the pressure in the cover gas have been also measured in the experiments. Fig. 11 representation of the OMEGA test Among the tests conducted in the OMEGA facility, run IV-12 is selected as the sample case for the present calculation. In this run, the high-temperature mixture of water and steam at 49 K was injected into the cylindrical pool vessel initially filled with stagnant water at 298 K. The initial pressure and void fraction of the mixture were MPa and 16.7%, respectively. The water pool was covered with atmospheric air (.1 MPa). The blowdown nozzle was also filled with atmospheric air. 4.2 Analyses As previously mentioned, can calculate the nonequilibrium phase-transition processes at interfaces where C gases are present. In this section, we will discuss the inhibiting effect of C gas on condensation in the highpressure vapor bubble as well as the code applicability Analytical Geometry and Conditions In the calculations, three regions of the pressure vessel, the blowdown nozzle and the water pool were modeled to represent the corresponding experimental setup shown in Fig. 11 by a two-dimensional cylindrical geometry. For the water pool region with the cover gas, 25 radial and 43 axial computational cells were used. It was assumed that the twophase fluid in the pressure vessel was mixed uniformly in thermal equilibrium. The diaphragm was assumed to rupture instantly with negligible flow resistance. The heat loss to the structure of the blowdown nozzle was also neglected Analytical Results In the experiment, the net gas-phase volume of expanding bubble equals to the decrease of the cover gas volume. Therefore, the change in the cover gas volume is compared with the experimental result as shown in Fig. 12. The comparison is made using cover gas volume ratio, which was defined as the decrease in void fraction of the cover gas region in the calculation. In the experiment, it was defined as the water slug interface increase fraction observed by the highspeed camera. It should be noted that the slug interface displacement measured in the experiment involves errors to some extent, judging from observed ripple interface and nonuniform bubble shape. The figure indicates that the transients of the cover gas volume calculated both with and without C gas effect agree with the experimental result well. The calculation with C gas effect increases slightly earlier than that without the effect. This is because the C gas suppressed the condensation in the vapor bubble, although the effect is rather small. Cover-gas volume ratio (-) Cover-gas pressure (MPa) Mechanical energy (MJ) Exp. without C-gas effect with C-gas effect Fig Comparison of cover gas volume ratio Exp. without C-gas effect with C-gas effect Fig. 13 Comparison of cover gas pressure Mechanical energy (total) Slug kinetic energy Cover-gas compression energy Mechanical energy (total) Slug kinetic energy Cover-gas compression energy Exp. Fig. 14 Comparison of mechanical energy between the experiment and the calculation with C gas effect The calculated cover gas pressures are compared with the experiment in Fig. 13. The pressure peaks of both the cases appear slightly later and become higher than the experiment, although the calculation with gas effect improves the pressure behavior slightly compared to the case without gas effect. The

7 major cause of this discrepancy between the calculation and experiment might be the entrainment (entrained droplet of water) in the bubble and the cover gas. It was noticed in the visual observation of the experiment, but is not modeled explicitly in the present calculation. In particular, taking into account the fact that the bubble growth behavior is well simulated by just before the pressure peak (about 2 ms), it is deduced that the entrainment in the cover gas region played an important role. Such an entrainment phenomenon and resultant heat transfer from gas to entrained droplets could take the heat from the cover gas, thereby reducing the cover gas pressure. This suggests the importance of mechanistic modeling of the entrainment phenomenon for better numerical simulations. In reactor safety assessment, mechanical energy, which is defined as the sum of coolant kinetic energy and compression energy of cover gas, is generally used as the measure of mechanical load on reactor vessel. Figure 14 compares the mechanical energy and its components, slug kinetic energy and cover-gas compression energy, between the experiment and the calculation considering C gas effect. It should be noted that errors in the measurement of slug displacement around and after 2 ms lead to some uncertainty in mechanical energy evaluated in the experiment. Considering this fact, we can say that can reasonably estimate the mechanical energy, although the peak of mechanical energy in the calculation is larger than that in the experiment. In addition, the peak of the cover gas pressure overestimated in the calculation seems to have a little effect on the mechanical energy. 5. APPLICATIO TO REACTOR-SCALE AALYSIS To investigate the code applicability to reactor-scale analysis, a core expansion phase analysis was performed following the initiating phase in an unprotected loss of flow event for a medium-sized sodium-cooled fast reactor. 5.1 Analytical Geometry and Conditions Figure 15 shows the analytical geometry, where an upper internal structure (control rod drive mechanism) and dip plates were neglected for simplification. In this figure, initial material distributions are distinguished by different colors. The radial shield and lower part of pin bundles were also not considered in the calculation. It was assumed that flow resistance in an upper core structure (upper pin bundle region) is small as done in the prototype reactor analysis on a conservative basis (Tobita et al., 1997). The initial core average fuel temperature and pressure are 34 K and 3.72 MPa (3.32 MPa for fission gas,.35 MPa for sodium vapor,.2 MPa for steel vapor and.2 MPa for fuel vapor pressure), respectively. The cover gas has the initial volume of 88.7 m 3 and pressure of.2 MPa. Fuel pins melt in the main part of the core center region initially, while they are still intact in the outer core region and the lower part of pin bundles. Fuels in the initial core consist of 61 % pellets, 35 % liquid and 4 % particles by volume. Stainless steel includes 36 % cladding, 42 % wrapper can, 1 % liquid and 21 % particles. 5.2 Analytical Results Transient material distributions in the region framed by broken lines in Fig. 15 are presented in Fig. 16 at the selected time intervals. The vapor bubble appears at around.6 s, and then expands in the upper sodium plenum due to successive supply of high-temperature vapor from the core. The heat exchange among core materials occurs in the upper plenum, resulting in phase transition such as sodium vaporization and fuel solidification, as can be seen in the figure at.2 s. The total V/C rates of sodium vapor occurring in the upper plenum are indicated in Fig. 17, where the positive and negative mass-transfer rates mean the condensation and vaporization rates, respectively. It can be seen that both vaporization and condensation processes work there. The net mass transfer rate indicates that vaporization is dominant at the early stage of the transient, and then condensation develops subsequently. A dominant mass-transfer path for vaporization is due to the interaction between liquid steel and liquid sodium. On the other hand, condensation is driven mainly by the vapor-sodium contact. Fig. 15 Fig. 16 Analytical geometry for reactor-scale analysis Vapor bubble expansion behaviors in reactorscale analysis The rates of sodium condensation by the vapor-sodium contact in the upper sodium plenum are compared between the cases with and without C gas effect in Fig. 18. The absolute value of the condensation rate in the case with C gas effect is smaller than that without gas effect. From this figure, the inhibiting effect of C gas can bee seen on sodium condensation to some extent.

8 The transients of mechanical energy are compared between the cases with and without C gas effect in Fig. 19. The difference in the mechanical energy between them is very small. This suggests that the inhibiting effect of C gas on condensation has a less impact on the mechanical energy release under the present rector conditions. Mass-transfer rate of sodium vapor (kg/s) 6, 4, 2, -2, -4, et mass transfer rate -6, Condensation by vapor/sodium contact Condensation by vapor/solid-particle contact -8, Vaporization by liquid-steel/sodium contact Vaporization by liquid-fuel/sodium contact -1, Fig. 17 Condensation rate of sodium vapor (kg/s) 6, 5, 4, 3, 2, 1, Time (s) Vaporization/condensation rates of sodium vapor in upper sodium plenum without C-gas effect with C-gas effect Time (s) Fig. 18 Condensation rate of sodium vapor in upper sodium plenum in reactor-scale analyses with and without C gas effect Mechanical energy (MJ) Fig without C-gas effect with C-gas effect COCLUSIOS Time (s) Mechanical energy in reactor-scale analyses with and without C gas effect In the present study, experimental verification of was performed for the transient behavior of large-scale bubble with condensation using the two independent series of experiments. In addition, the code applicability was investigated for the core expansion phase of CDA in an LMFR. The following results were obtained: The new series of experiments with a low-pressure largescale bubble and simulations indicated that the C gas has less inhibiting effect on condensation of large-scale bubbles, in which the gas and liquid phases are dispersively mixed without a buildup of the C gas at their surface. The experimental analysis of the OMEGA test, which was performed under high-pressure conditions, showed that the C gas has a little effect on the mechanical energy release. On the other hand, the present results suggest importance of mechanistic modeling of the droplet entrainment phenomenon for better numerical simulations. The reactor-scale analysis of the core expansion phase indicated that the C gas has the inhibiting effect on condensation to some extent in a multi-component system under a CDA condition. However, the simulation results also suggest that its effect has a less impact on the mechanical energy. Although cannot predict the buildup of the C gas at the interfaces, where condensation occurs, explicitly, the present study indicates that the code represents the condensation processes of large-scale bubbles in sufficient physical details. We expect that the present results will contribute to the improvement of reliability in accident analyses of LMFRs. REFERECES Bird, B. et al. (196). Transport Phenomena, John Wiley and Sons, ew York. Bohl, W.R. et al. (199). "SIMMER-II: A Computer Program for LMFBR Disrupted Core Analysis," LA MS, Los Alamos ational Laboratory. Morita, K. et al. (23). "Development of Multicomponent Vaporization/condensation Model for a Reactor Safety Analysis Code : theoretical modeling and basic verification," ucl. Eng. Des., 22, pp Moszynski, J.R. et al. (198). "Liquid Entrainment by an Expanding Gas or Vapor Bubble A Review of Experiments and Models," BL-UREG-2B343, Brookhaven ational Laboratory. Saito, M. et al. (1979). "The Termination Phase of Core Disruptive Accidents in LMFBRs," Int. Mtg. on Fast Reactor Safety Technology, Seattle, WA, USA, August Simpson, D.J. et al. (1981). "The Termination Phase of Core Disruptive Accidents in LMFBR. 198 Annual Report," PE , Purdue University. Suzuki, T. et al. (23). "Development of Multicomponent Vaporization/condensation Model for a Reactor Safety Analysis Code : extended verification using multi-bubble condensation experiment," ucl. Eng. Des., 22, pp Theofanous, T.G. et al. (1979). "The Termination Phase of Core Disruptive Accidents in LMFBRs. First Annual Progress Report," PE , Purdue University. Tobita, Y. et al. (1999). "Evaluation of CDA Energetics in the Prototype LMFBR with Latest Knowledge and Tools," 7th Int. Conf. on ucl. Eng. (ICOE-7), Tokyo, Japan, April Tobita, Y. et al. (22). "The Development of, an Advanced Computer Program for LMFR Safety Analysis," Joint IAEA/EA Technical Mtg. on the Use of Computational Fluid Dynamics Codes for Safety Analysis of Reactor Systems (including Containment), Pisa, Italy, ov