Nuclear Engineering and Design

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1 Nuclear Engineering and Design 258 (2013) Contents lists available at SciVerse ScienceDirect Nuclear Engineering and Design j ourna l ho me pag e: Numerical study on thermal stratification phenomena in upper plenum of LMFBR MONJU Makoto Shibahara, Takashi Takata, Akira Yamaguchi Graduate School of Engineering, Osaka University, 2-1 Yamada-Oka, Suita, Osaka, Japan h i g h l i g h t s Three-dimensional analysis of thermal stratification in the upper plenum of MONJU is conducted using a commercial CFD code, FLUENT. A comparison between the numerical simulation and the trip test of MONJU is made. This study shows that the interface of thermal stratification is influenced by the flow pattern in the upper plenum of MONJU. a r t i c l e i n f o Article history: Received 29 February 2012 Received in revised form 28 December 2012 Accepted 1 February 2013 a b s t r a c t The three-dimensional analysis of thermal stratification in the upper plenum of MONJU is conducted using the commercial CFD code, FLUENT ver Since the temperature gradient near the thermal stratification interface would cause thermal stress in the reactor components, it is important to understand the characteristics of thermal stratification for evaluating the structural integrity in the upper plenum. As the result of numerical analysis, it is understood that the interface of thermal stratification is influenced by the flow pattern in the upper plenum. After the jet from the core outlet impinges on the upper core structure, the hot fluid flows obliquely upward to the inner barrel under the 40% electric output and flow condition. On the other hand, the jet from the core outlet flows to the lower part of the upper plenum, and then cold fluid flows through the s after the turbine trip. Hence, the flow structure has changed from the initial condition as the flow rate and temperature of the core outlet decrease due to the turbine trip. It is considered that the flow path of the s has been maintained for a long duration since the thermal stratification changes the distribution of buoyancy forces Elsevier B.V. All rights reserved. 1. Introduction The knowledge of thermal stratification phenomena may be important for safety assessment in the upper plenum of some liquid metal fast breeder reactors (LMFBRs). The thermal stratification interface in the upper plenum of MONJU was observed due to the flow coasting down after the turbine trip (Doi et al., 1996). Since the temperature gradient near the thermal stratification interface could cause thermal stress in the reactor components, it is important to understand the characteristics of thermal stratification to evaluate the impact on the structural integrity in the upper plenum. In order to predict the temperature distribution after the turbine trip, many numerical analyses have been reported under various conditions. Ieda et al. (1990) conducted the analysis of the thermal stratification for water and sodium with 1/10 scaled models of the upper plenum using a multi-dimensional thermal hydraulic code, AQUA. And then, Doi and Muramatsu (1996) performed the Corresponding author. Tel.: address: takata t@see.eng.osaka-u.ac.jp (T. Takata). numerical analysis of the thermal stratification for the turbine trip test was conducted using AQUA. Since it was complicated to represent the reactor components such as the s of the inner barrel and the upper core structure (UCS) accurately, the empirical correlations for the pressure drop were given in the numerical analysis. Kimura et al. also carried out the numerical analysis using AQUA on the thermal stratification of plant dynamics test loop integrated with direct heat exchanger (PLANDTL-DHX) (Kamide et al., 1998; Nishimura et al., 2000) which has an upper plenum of 2 m diameter (Kimura et al., 1999). They evaluated the thermal stratification interface and clarified the effect of Richardson number and Reynolds number. Moreover, Ohno et al. (2007) conducted the numerical analysis for thermal stratification in the 1/10 scale model using FLUENT (ANSYS, 2012) and STAR-CD (STAR, 2012) to validate the commercial codes. They clarified the influence of mesh arrangement, discretized scheme, and various turbulence models on the thermal stratification. Although the axial temperature distribution and the thermal stratification interface were evaluated in the 1/10 scale model, the full scale model including the reactor components was not simulated. As a benchmark problem of the International Atomic Energy Agency (IAEA), Yoshikawa and /$ see front matter 2013 Elsevier B.V. All rights reserved.

2 M. Shibahara et al. / Nuclear Engineering and Design 258 (2013) Fig. 1. Schematic of the upper plenum. Minami (2008) summarized detailed engineering data of the turbine trip test under the 40% rated power operation of MONJU. Honda et al. (2010) performed the numerical analysis of the upper plenum of MONJU under the 40% rated power operational condition using the commercial code, FrontFlow/Red (Tubokura et al., 2009), which was developed at the University of Tokyo under the national project Revolutionary Simulation Software (RSS). As a parametric study on the thermal stratification in the upper plenum of MONJU, Sakamoto et al. (2010) carried out the numerical analysis using FLUENT code in order to investigate the influence of mesh arrangement on the thermal stratification. It was clarified that thermal stratification interface was affected by the mesh arrangement of the s. In this research, we focused on the flow dynamic structure in the upper plenum of MONJU after the turbine trip. The three-dimensional analysis of thermal stratification in the upper plenum of MONJU is conducted using the commercial CFD code, FLUENT ver (ANSYS, 2012). The purposes of this study are: (1) to clarify the flow pattern of cold sodium which flows to the upper plenum after the turbine trip and (2) to compare the axial temperature distribution and the temperature gradient with those of experimental data (Doi et al., 1996). Fig. 2. UCS overall view. flow guide tubes and honeycomb structure in Fig. 2 is assumed as the porous media body, which is applied with the porosity of 0.5. Fig. 3 shows the top view of the core subassemblies of MONJU (Yoshikawa and Minami, 2008). The entire core channels are divided into eighteen regions. Table 1 shows the flow channel ID at the core outlet. The core subassemblies of MONJU consist of inner diver core, outer diver core, neutron source, blanket, neutron 2. Outline of turbine trip test of MONJU In the turbine trip test of MONJU, the 40% rated power operational condition was set as an initial condition and the turbine trip was conducted manually. The steady-state condition means the 40% electric output and flow condition. When the turbine trip test started, the circulation pump stopped with a flow coasting down characteristic and a pony motor was activated. This operation after the turbine trip was defined as the transient condition. As the result of the turbine trip test, the flow rate of the primary coolant decreased with the flow coasting down characteristic and was kept to be approximately 10% of the rated operation under the transient condition. The temperature distribution in the axial direction of the upper plenum of MONJU was measured with thermocouple (T/C) plug installed in the upper plenum. The upper plenum consists of upper instrument structure (UIS), T/C plug, core barrel, three hot leg pipes (H/L), fuel handling machine (FHM) and inner barrel with the s as shown in Fig Numerical method 3.1. Analytical model Fig. 1 shows the analytical model of the upper plenum. In the numerical analysis, the upper core structure (UCS) such as fingers, Fig. 3. Top view of the core subassemblies.

3 228 M. Shibahara et al. / Nuclear Engineering and Design 258 (2013) Table 1 Flow channel ID at the core outlet. Core region Flow channel ID Inner core Outer core Blanket Neutron shielding FC Fig. 4. Cross-sectional view of the mesh arrangement. shielding, and control rods such as coarse control rods, fine control rods and backup control rods. Fig. 4 shows the cross-sectional view of the mesh arrangement with C-loop H/L for the analytical model. In the numerical analysis, the mesh number of approximately 5.11 million was determined based on the previous study (Sakamoto et al., 2010). To clarify the flow distribution in the s and overflow of the inner barrel, the fine mesh is arranged in the s and top of the inner barrel as shown in Fig. 5. Since it was complicated to model the flow holes on the inner barrel accurately, the pressure drop correlations were applied in the previous analysis (Doi and Muramatsu, 1996) as mentioned in Section 1. Since the thermal stratified layer would affect on the coefficient of pressure drop, the s have been represented by as built-geometry in this numerical analysis Boundary condition and numerical method The flow rate and the sodium temperature for the inlet boundary condition are consistent with the IAEA benchmark condition (Yoshikawa and Minami, 2008). Fig. 6 shows the typical flow rate of channels at the inlet boundary condition. As mentioned in Section 2, the flow rate from the core outlet decreases with time and kept approximately 10% of the rated operation. The turbulence intensity, I, is estimated to be 5% at the inlet boundary conditions, and Flow rate [kg/s] Nutron source N Coarse CR rods C Fine CR rods F Backup CR rods B Channel 1 (Inner core) Channel 6 (Outer core) Channel 9 (Blanket) Channel 12 (Neutron Shielding) Time [s] Fig. 6. Flow rate of channels after the turbine trip. the turbulent kinematic energy, k, is calculated as k = 3/2(U avg I) 2, where, U avg is the mean outlet velocity for each channel. And also, Fig. 7 shows the typical sodium outlet temperature of each channel (Yoshikawa and Minami, 2008). As shown in Fig. 7, the sodium Fig. 5. Mesh arrangement for (a) and (b) top of inner barrel.

4 M. Shibahara et al. / Nuclear Engineering and Design 258 (2013) Temperature [ o C] Channel 1 (Inner core) Channel 6 (Outer core) Channel 9 (Blanket) Channel 12 (Neutron Shielding) of gravity is considered in the numerical analysis. The basic conservation equations are discretized with finite volume method (FVM) in FLUENT code (ANSYS, 2012) The QUICK scheme is applied for convective terms so as to reduce numerical diffusion (Maekawa, 1990), and the time marching term is calculated by the second order Euler implicit method. The time step is 2 s based on a previous study (Sakamoto et al., 2010). The analysis time is ranged from 0 s to 600 s. Also, the SIMPLE method and standard k ε model are applied in the numerical analysis Result and discussion Time [s] Fig. 7. Sodium temperature of channels after the turbine trip. temperature decreases with time after the trip. The outlets of three H/L pipes are bounded by satisfying a constant pressure condition. The walls of reactor components are adiabatic in this numerical analysis. The density, isobaric specific heat, and kinematic viscosity of sodium are expressed by quadratic or cubic polynomial of sodium temperature (Sodium Technology Education Committee, 2005). The thermal conductivity of sodium is defined with a straight-line approximation polynomial, which depends on sodium temperature. Since the buoyancy force would affect stratified flow, the effect 4.1. Steady-state analysis Fig. 8 shows the velocity (a) and temperature (b) profiles under the steady-state condition. It can be seen from Fig. 8(a), the jet from the core outlet ascends through the porous media zone which is assumed as the UCS, and it impinges on UIS. Then, the hot fluid flows obliquely upward to the inner barrel. After impinging on the inner barrel, the flow is branched off into two directions; upward flow and downward flow. The upward flow along the inner barrel overflows the top of the inner barrel, and then it descends through the annular gap between the inner barrel and the reactor vessel wall. In the numerical analysis, the ratio of flow rate of s is evaluated to be 10% compared with the overflow rate of the inner barrel. Hence, the oblique flow comes to govern under the steady-state condition. On the other hand, as shown in Fig. 8(b), the temperature profile is not uniform at the lower side of the upper plenum. Fig. 8. Velocity and temperature profiles under the steady state condition. Fig. 9. Velocity profiles for the depth of (a) 5300 mm, (c) 4400 mm, and (d) 1000 mm from the liquid surface.

5 230 M. Shibahara et al. / Nuclear Engineering and Design 258 (2013) Fig. 10. Temperature profiles for the depth of (a) 5300 mm, (c) 4400 mm, and (d) 1000 mm from the liquid surface. This is because the coolant temperature of blanket fuel subassembly is lower than that of the inner core fuel subassembly, so that the lower temperature fluid descends along the core barrel and accumulates between the support plate and the lower s. Fig. 9 shows the velocity profiles at the depth of (a) 5300 mm, (b) 4400 mm, and (c) 1000 mm from the liquid surface. These correspond to the vertical cross-sectional view of figure in Fig. 8(a). The circumference direction of velocity profile is almost uniform at each depth from the liquid surface even though the reactor components such as the FHM and the T/C plug are installed, while there are circumferential variations in the annular flow regions due to the outlet pipes as shown in Fig. 7(b). Fig. 10 shows the circumference direction of the temperature profiles at the depth of (a) 5300 mm, (b) 4400 mm and (c) 1000 mm from the liquid surface. The circumference direction of temperature profiles is also uniform at each depth from the liquid surface. Hence, the asymmetric structure of the upper plenum has relatively little effect on the flow structure and temperature distribution under the steady-state condition Transient analysis Fig. 11 shows the flow pattern during transient operation ranged from 60 s to 600 s. The flow pattern has changed from the Fig. 11. Flow pattern during the transient operation ranged from 60 s to 600 s.

6 M. Shibahara et al. / Nuclear Engineering and Design 258 (2013) Fig. 12. Temperature profiles during the transient operation ranged from 60 s to 600 s. steady-state condition due to the flow coasting down. As shown in Fig. 11(a), the jet is spread to the upper plenum horizontally after it flows between UIS and the core outlet. Then, the jet flows to the support plate as shown in Fig. 11(b). It is considered that the inertial force of jet becomes smaller than the plugging effect, which means that the hot sodium acts as the plug due to the buoyancy force. After the coolant flows on the support plate, it flows to the lower s. For the transient operation ranged from 240 s to 600 s as shown in Fig. 11(c) (f), the flow pattern has not changed since the plugging effect is greater than the inertial force. Fig. 12 shows the vertical cross-sectional view of temperature profiles during the transient operation ranged from 60 s to 600 s. It can be seen from Fig. 12(a) that the cold sodium begins to penetrate the upper plenum since the core outlet temperature has reduced rapidly due to scram. The temperature variations result in density changes that affect the flow conditions. As shown in Fig. 12(b), the thermal stratification interface is formed near the upper flow holes as cold sodium fills with lower part of upper plenum. It seems that the thermal stratification is characterized by the flow distribution. Once the thermal stratification is established, it would last for a long time. Moreover, the axial temperature gradient near the interface of thermal stratification may be of significant concern as mentioned in Section 1. Fig. 12(c) (f) shows that the interface of thermal stratification ascends due to bottom-up flow from the core Fig. 13. Temperature profiles in horizontal cross section for the elapsed time of (a) 60 s, (b) 120 s, and (c)240 s at 5300 mm.

7 232 M. Shibahara et al. / Nuclear Engineering and Design 258 (2013) Annular gap Upper s Lower s Flow ratio [%] Time [s] Fig. 14. Temperature distribution under the transient condition. Fig. 15. Flow ratio of annular gap and s in the transient analysis. thermal stratification interface, the non-dimensional temperature T * is defined as following equation (Ohno et al., 2007): outlet since the flow rate through the s is limited. In other words, the position of thermal stratification interface moves slowly under the transient condition. As the result of numerical analysis of flow dynamic structure and thermal stratification in the upper plenum of MONJU, it is understood that the interface of thermal stratification is influenced by the flow pattern after the turbine trip. Fig. 13 shows the temperature profiles in horizontal crosssection at the depth of 5300 mm from the liquid surface, where the lower s exist, for the elapsed time of (a) 60 s, (b) 120 s, and (c) 240 s. These correspond to the vertical cross-sectional view of figures in Fig. 12(a) (c), respectively. As shown in Fig. 13(a), a uniform temperature distribution exists at 60 s. When the time is longer than 60 s, the temperature difference appears between the annular gap and the inside of the inner barrel since the coolant flows in the upper plenum as shown in Fig. 13(b). There are temperature variations inside the inner barrel at 120 s, however, since the T/C plug is far from the modalities of the temperature change, it seems that they have a little influence on the long-time temperature profiles at the T/C plug. As the thermal stratification interface rises with time, the coolant inside the inner barrel flows through the s as shown in Fig. 13(c). And then, after that the coolant in the inner barrel and coolant in the annular gap mix and flow toward H/L pipes. In short, the temperature distribution changes along with the radial direction, in contrast, it is uniform on the circumference direction under the transient condition. Fig. 14 shows the comparison of the axial temperature distributions between this analysis and the experimental data at the T/C plug (see Fig. 1). The symbols show the experimental data for the turbine trip under the 40% rated power operational condition while the solid lines show the numerical result. The numerical result shows good agreement with the experimental data except for the depth lower than 6000 mm under the initial condition of the turbine trip (0 s). After the turbine trip, the calculated temperature accords with the experimental data at the s, however, it is lower than that of the experiment above the upper s after 240 s. Therefore, we investigated the flow ratio of s to the annulus gap as shown in Fig. 15. As a result, it is found that the fraction of the flow rate through the s increases after the trip, while the flow rate of the annular gap decreases with time. However, since the flow ratio of annular gap remains about 18% in Fig. 15, the fluctuation of annular flow has a possibility to delay the raise of the thermal stratification layer. Moreover, it remains possible that the flow rate in s is lower in the numerical analysis than in the experiment. Fig. 16 shows the position of the thermal stratification interface after the turbine trip. In order to evaluate the position of the T = T T c (1) T h T c where T c is minimum temperature, and T h is maximum temperature of liquid surface at the trip for the T/C plug. When the value of T * become 0.5 both in the numerical analysis and the experiment, the position will be defined as the interface of thermal stratification. As shown in Fig. 16, it is understood that the interface of the thermal stratification raises as the coolant flows into the upper plenum. The trend of numerical values is consistent with the experimental data at the early time ranged from the turbine trip (0 s) to 240 s, while both the numerical result and the experiment curve after 240 s. The slope of numerical result is higher that of experiment, however, the numerical result shows similar curvature to the experiment. As shown in Fig. 11, there is no obvious change of the flow pattern after 240 s in the numerical result, while the interface of the thermal stratification would be not liner during s. Since the flow rate through the s increase with time as shown in Fig. 15, the thermal stratification layer would be affected by the bypass flow through the s. Fig. 17 shows the comparison of the axial temperature gradients between the numerical result and experimental data for the elapsed time of (a) 120 s, (b) 240 s, and (c) 600 s. Each temperature gradient (dt/dz) is determined from the T/C temperature distribution as shown in Fig. 14. As mentioned in Section 1, the evaluation of temperature gradient is important for determining the structural integrity of the reactor component such as the inner barrel. In the experiment, the temperature gradient is obtained from the [mm] Depth from liquid surface Numerical result Experimental data Time [s] Fig. 16. Thermal stratification interface after the turbine trip. Liquid surface Upper Lower

8 M. Shibahara et al. / Nuclear Engineering and Design 258 (2013) (a) 0.00 Time 120s Liquid surface 0.00 Time 240s Liquid surface Depth from liquid surface [m] Experimental data Numerical result Top of inner barrel Upper Lower Depth from liquid surface [m] Experimental data Numerical result Top of inner barrel Upper Lower Temperature gradient [ ºC/m] Temperature gradient [ºC/m] (c) 0.00 Time 600s Liquid surface Depth from liquid surface [m] Experimental data Numerical result Top of inner barrel Upper Lower Temperature Gradient [ºC/m] Fig. 17. Temperature gradients for the elapsed time of (a) 120 s, (b) 240 s, and (c) 600 s. readings from thermocouples and the arrangement. The symbols show the experimental data and numerical result, at each elevation defined by the arrangement of the thermocouple in the experiment. As shown in Fig. 17(a), the maximum temperature gradient in the numerical result is about 80 C lower than that of the experimental data. After 120 s, it appears that the maximum temperature gradient decreases due to the mixing effect on the thermal stratification. The numerical result is approximately consistent with the experimental data. Moreover, the maximum temperature gradient of the numerical simulation shows good agreement with that of experiment. In Fig. 17(c), the elevation of the maximum temperature gradient at 600 s is higher than that of experiment, although the maximum temperature gradient approximately accords with that of the experiment. As mentioned in Fig. 14, the thermal stratification interface is overestimated after 240 s, therefore, the elevation of the maximum temperature gradient becomes higher than that of experiment. And also, the spike of the numerical result was caused by a discrepancy in the temperature distribution at the top of the inner barrel as shown in Fig. 14. Moreover, since the interface of thermal stratification has risen at 600 s in the numerical result, the temperature change becomes lower than that of the experiment at the upper s. 5. Conclusions The three-dimensional analysis of thermal stratification in the upper plenum of MONJU is conducted using the commercial CFD code, FLUENT ver to analyze the flow dynamic structure on the thermal stratification phenomena. As a result, thermal stratification interface is influenced by the flow pattern in the upper plenum of MONJU. The flow structure of the upper plenum has changed from obliquely flow beyond the top of inner barrel to downward flow through the s after the turbine trip. The numerical result is consistent with that of the experimental data until 240 s, while the thermal stratification interface after 240 s ascends faster than that of experimental data. The rise of the thermal stratification layer would be affected by the flow ratio of the s to the annular gap. Therefore, it may be important to clarify a fluctuation of the flow ratio by the related boundary conditions of the s and annular flow. Acknowledgements Present study is the result of 2010 R&D core program for practical realization of fast breeder reactor by using MONJU entrusted to University of Fukui by the Ministry of Education, Culture, Sports, Science and Technology of Japan (MEXT). References ANSYS FLUENT, ANSYS, Inc., Technology/Fluid+Dynamics/. Fluid + Dynamics + Products/ANSYS + Fluent (accessed ). Doi, Y., Muramatsu, T., Investigation of analytical methods in thermal stratification analysis evaluation of flow rates through s for normal and scram conditions of 40% power operation with AQUA code. Power reactor and Nuclear Fuel Development Corporation, PNC TN (in Japanese). Doi, Y., Muramatsu, T., Kunoki, K., Thermal stratification tests in MONJU upper plenum (I) temperature distribution under normal and scram conditions with 40% power operation. Power Reactor and Nuclear Fuel Development Corporation, PNC TN (in Japanese). Honda, K., Ohira, H., Sotsu, M., Yhoshikawa, S., Thermal-hydraulic analysis of MONJU upper plenum under 40% rated power operational condition. In: Proc. Int. Topical Meeting on Nuclear Thermal-Hydraulics, Operation and Safety (NUTHOS-8). Ieda, Y., Maekawa, I., Muramatsu, T., Nakanishi, S., Experimental and analytical studies of the thermal stratification phenomenon in the outlet plenum of fast breeder reactor. Nucl. Eng. Des. 120,

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