VALIDATION OF CODE ASTEC WITH HERMES-HALF EXPERIMENTAL RESULTS

Transcription

1 X VALIDATION OF CODE ASTEC WITH HERMES-HALF EXPERIMENTAL RESULTS Andrea Bachratá ČESKÉ VYSOKÉ UČENÍ TECHNICKÉ V PRAZE Fakulta Jaderná a Fyzikálně Inženýrská Katedra jaderných reaktorů 1. Introduction The TMI- accident and later Chernobyl accident underlined the need for common safety principles for all countries and all types of nuclear power plants. Although severe accidents are of very low likelihood and might be caused only by multiple failures, accident management is implemented for managing their course and mitigating their consequences. In case of such an accident, the core melts and relocates to the lower head. After melt relocates into the residual water in the lower head its fragmentation and quenching will eventually lead to partially fragmented and cooled debris in water called debris bed. The water will evaporate and the fragmented debris will heat-up by residual heat and re-melt to form a volumetrically heated molten pool. As a consequence, high temperatures of molten pool might endanger the integrity of the reactor pressure vessel wall (see Figure 1). Figure 1: TMI accident 1979

2 In severe accidents with core melting, one of the main goals of accident management is the corium localization and stabilization. To fulfil this requirement, two approaches are possible: In-vessel melt retention; Ex-vessel retention. The in-vessel retention is considered to be the most effective severe accident management strategy for corium localisation and stabilisation. The higher advantage of this strategy is that the radioactive corium stays beyond the second barrier. On the other hand, there is a danger of vessel failure. Consequently, thermal explosion could be caused if the high-temperature corium contacts with water in reactor cavity. The main idea of IVR consists in flooding the reactor cavity to submerge the reactor vessel up to its support structures. The PWR lower head containing the melt pool is cooled from outside, which keeps the outer surface of the vessel wall cool enough to prevent vessel failure. The decay heat is transferred through the RPV wall to the water and then to the atmosphere of the containment of NPP. Necessary condition for successful application of invessel retention concept is that: The primary system is depressurised; The thickness of the reactor wall, that remains cold, is sufficient to withstand the weight of molten pool in lower head. The nd necessary condition could be fulfilled only for heat flux values lower than critical heat flux value [1]. This could limit the applicability of the in-vessel retention concept only for low and medium power reactors. The in-vessel retention as a severe accident management strategy is considered at a design stage of low and medium power reactors (AP600, AP1000). This concept is being implemented also at the operated NPP (Loviisa in 000 and 00 in Finland, VVER440/13 in Czech Republic, Slovak Republic and Hungary: in process). Moreover, there is an effort to extend the in-vessel retention concept also for high-power reactors such as Korean Advanced Power Reactor APR1400 (in combination with in-vessel core catcher). In order to prove the applicability of in-vessel retention concept for APR1400, a T- HERMES program has been performed at Korea Atomic Energy Research Institute (KAERI). The HERMES-HALF experiment has been performed to observe and evaluate the two-phase natural circulation phenomena and finally to propose enhanced designs for the coolant inlets and the upper steam venting slots in the reactor vessel insulation of APR1400 []. The technical proposal of in-vessel retention strategy should be based on experimental as well as analytical support. A large number of computer codes are used in safety analysis. Nowadays, a European computer code ASTEC is being developed jointly by French IRSN and German GRS with a strong international support. The validation and application of this code is an important part of the SARNET project that began in 00, and its successor SARNET that began in 009. The aim of SARNET (Network of Excellence) is to resolve the

3 most important remaining uncertainties and safety issues for enhancing the safety of existing and future NPPs. The work presented in this thesis concerns the validation of stand-alone CESAR module of the integral ASTEC code. The CESAR module is devoted to analysis of front-end thermalhydraulic phenomena of severe accidents. The CESAR module can be applied either in standalone mode or in integral application coupling with other ASTEC modules. The aim of the presented work is to prove the capability of stand-alone CESAR module in modeling of twophase flow natural circulation in a specific geometry (in a curved channel and a large annular space around the vessel). Firstly, the hydraulic part of CESAR module was validated with HERMES-HALF experimental results. In the HERMES-HALF experiment, a two-phase flow in external reactor vessel cooling (ERVC) loop was not generated by a direct heating method but by a nonheating method. For this non-heating experiment, an equivalent air flow rate was injected by the air supply system instead of producing steam on the hot RPV surface. Therefore, the HERMES-HALF is meaningful as a hydraulic test for CESAR module and not directly for IVR []. Secondly, the calculations were performed using thermal-hydraulic part of CESAR module. A heat transfer through the wall was prescribed as a boundary condition. Consequently, the two-phase flow in half-scaled geometry of APR1400 was induced due to the heat transfer and vapour formation as expected in real reactor application. These calculation results with heat cannot be compared with experimental results. On the other hand, it could demonstrate the representativeness of HERMES-HALF experimental results that were obtained using air injection. The first limits to critical heat flux will be presented using two different correlations.. Severe accident mitigation at APR1400 The first advanced power reactor APR1400 is going to be finished in Korea in 013 [3]. As for a design of the APR1400 insulation system, there are 61-ICI penetrations and 4-shear keys that are installed in the reactor vessel. Because of a conical shape of the insulation, a minimum gap region (8.9 cm) exists between the reactor vessel and the insulation near the shear key as is shown on Figure. Figure : APR1400 reactor vessel and insulation

4 As for the severe accident mitigation, the in-vessel retention concept is implemented into a design of APR1400. A United States-Korean International Nuclear Energy Research Initiative (INERI) project has been performed to determine the feasibility of IVR concept at APR 1400MWe [4]. First approach for successful IVR was investigated by applying the in-vessel core catcher to provide an engineered gap between the relocated core materials and the waterfilled reactor vessel (see Figure 3). The design of this in-vessel core catcher relies on several mechanisms to enhance IVR, such as retention and dilution of decay heat in the relocated core materials, and heat transfer through the lower surface of the core catcher via narrow gap cooling. Figure 3: In-vessel core catcher for APR1400 Second approach for successful IVR was investigated by applying the external cooling of the reactor vessel (see Figure 4). For a natural circulation flow path, it is suggested that water inlets and venting slots will be installed on the insulation wall. A detailed insulation design is ongoing. The aim is to increase heat removal rate by two-phase natural circulation in a case of IVR concept via streamlining the coolant flow around the reactor vessel. To suggest optimal insulation design (inlet and outlet area) for effective external reactor vessel cooling of APR1400, HERMES-HALF experiment was performed by KAERI and will be shortly presented in a Section below. Figure 4: Externally flooded reactor vessel cavity

5 3. HERMES-HALF experiment The T-HERMES program has been performed in KAERI (Korea Atomic Energy Research Institute). The objectives of the T-HERMES program were to observe and evaluate the two-phase natural circulation phenomena under an external reactor cooling of the APR1400. The T-HERMES program consists of five parts: T-HERMES-SCALE, T- HERMES-SMALL, T-HERMES-1D, HERMES-HALF and T-HERMES-CFD []. In this Section, only the HERMES-HALF experiment will be shortly presented. The HERMES (Hydraulic Evaluation of Reactor cooling Mechanism by External Selfinduced flow) -HALF was a large-scale air injection experiment. The experiment has been performed to measure the two-phase natural circulation mass flow rate through the annular gap between the reactor vessel and the insulation material. In the HERMES-HALF experiment, a two-phase flow was generated by a non-heating method. For this non-heating experiment, an equivalent air flow rate was injected by the air supply system, while under the heating experimental condition, the vapour bubbles are naturally generated in the heated channel. The HERMES-HALF hydraulic test is an interesting experiment for CESAR validation because of large-scale and curved channel geometry with annular space around the vessel (see Figure 5). Figure 5: The HERMES-HALF test facility, visualization of void fraction

6 3.1 Description of HERMES-HALF facility The facility is a half scaled-down reactor vessel and insulation part which is prepared by utilizing the results of a scaling analysis proposed by Cheung to simulate APR1400 reactor and its insulation system [5]. The main parameters are summarized in Table 1. Table 1: Geometry scaling by F. B. Cheung Names APR1400 HERMES-HALF Scale RPV radius of the hemispherical section RPV radius of the annular section Radius of the bottom insulation plate.538 m ( ).574 m ( ).193 m (86.34 ) Location or minimum gap Minimum gap thickness Annular gap thickness Water level from the bottom of RPV ICI nozzle diameter m (3.5 ) m (8.5 ) m (81.17 ) m (3.01 ) 1.69 m 1/ 1.87 m 1/ m 1/ m 1 / m 1 / m 1/ m 1/ In the paragraphs below, the basic parameters of the test facility and of the air supply system are summarized. These data are then used to model the input data for ASTEC simulations summarized in Section Parameters of the test facility A schematic diagram of the HERMES-HALF experimental facility is shown in Figure 6. For maximizing the natural circulation flow, water inlets and outlet ports exist in the insulation. These openings represents enter and exit position of water. To simulate the water inlets, there are 3 holes in the central part, and 35 holes in the circumferential part of the water inlet plate (Figure 6). Each hole of 75 mm in diameter could be plugged. The investigated range for the inlet area is: m (see Table ). The shape of each outlet is rectangular, the horizontal size is 0. m and the vertical is m. The different outlet areas are tested: and 0.15 m (see Table ).

7 Water inlet area Figure 6: HERMES-HALF experimental facility Andrea Bachratá Page 7 Master thesis

8 3.1. Parameters of the air supply system In the HERMES-HALF experiment, a two-phase flow was generated by a non-heating method. The air bubbles with the same volumetric generation rates as the vapours were injected directly. The air was injected to the gap region through 141 air injectors installed on the lower head reactor vessel wall. The air injection diameter from each air injector was 70 mm. The injected air flow-rate could be calculated using the following equation: where. Va is a volume flow-rate of air, Aw is the outer vessel area, dissipated to the vessel wall, V. a " w q.aw = (1) h. ρ fg g " q w. Aw is the residual power h fg water enthalpy change of liquid to vapour and ρ g is density of vapour. The profile of air flow injection corresponds with the profile of heat flux distribution calculated by MAAP4 (Figure 7) [7]. The volume flow-rate of air towards each air distributor was controlled and measured by an air control valve and an air flow meter. Due to the blower capacity limitation [6] the air flow rate was limited to 4%. The air flow is defined as:. V. Va.100 = flow rate[%] The 100% of heat flux through the vessel wall is equal to 8379 m 3 /h of volume flowrate of air []. That represents a residual power dissipated to the vessel wall about 6.15 MW (half scale experiment). The 10% of heat flux through the vessel wall is simulated by 10% of air flow rate which is equal to m 3 /h. () Figure 7: Heat flux distribution along the lower head vessel wall by using the MAAP4 Andrea Bachratá Page 8 Master thesis

9 3. Experimental results In the HERMES-HALF experiment, the circulation mass flow rate, void fraction and local pressure were measured. Two-phase natural circulation flow was visualized []. The aim of the experiment was to evaluate the cooling capability in a case of APR1400 and optimisation of its insulation design. The effect of the air injection mass flow rate, water inlet area and water outlet area on water circulation mass flow rate was studied (see Table ). The experimental results are available in []. These results will be compared with the CESAR calculations (see Section 5.). From the HERMES-HALF experimental results, the natural circulation flow rate can be generated up to 00kg/s by adjusting the inlet and outlet areas of the insulation. Table : Objectives and results of HERMES-HALF tests TEST OBJECTIVE RESULT INLET OUTLET from m to 0.15 m m and 0.15 m To suggest optimal insulation design for effective external reactor cooling in APR1400 The maximum mass flow rate was reached for inlet area 0.15 m and outlet area 0.15 m (~00 kg/s) [3] Air flow from 10% to 4% of total To evaluate the cooling capability by the natural circulation flow under ERVC At higher air flow rate, there is a stagnation in circulation mass flow rate The critical heat flux (CHF) at the upper position of the RPV wall was assumed to reach about MW/m using the KAIST and SULTAN experimental correlations of circulation mass flow rates with the wall CHF value (see Figure 8) []. Generally, KAERI stats that if the maximal heat flux through the RPV submerged in water is below 1.3MW/m (value estimated by MAAP4 code: Figure 7 [7]), there will be a certain margin to the critical heat flux MW/m providing that the mass flow rate is about 00kg/s. In this case, the external reactor cooling as a severe accident management strategy seems to be applicable for the APR1400 NPP. Andrea Bachratá Page 9 Master thesis

10 HERMES-HALF Range Figure 8: CHF variations according to the mass flux and water subcooling Andrea Bachratá Page 10 Master thesis

11 4. ASTEC V1.3 Code ASTEC (Accident Source Term Evaluation Code) is an integral source term code for the simulation of severe accidents in Light Water Reactors. The ASTEC code is playing a central role in SARNET (6 th EU Framework Programme) in order to progressively become the reference European integral code for analysis of severe accident. The version V1.3 was released in December 007. The version V1.33 is yet available e.g. in CEA Grenoble. The code is jointly developed by IRSN and GRS since more than 10 years. The aim of the code is to simulate an entire severe accident sequence from the initiating event through core melting and relocation to fission product release out of the containment. The ASTEC code consists of several computational modules (see Figure 9), devoted to analysis of specific problems (e.g. thermal-hydraulics, core degradation, fission product release and transport, etc.). ASTEC V1 CONSTITUTIVE MODULES CPA Thermalhydraulic & Aerosol behavior in containment MEDICIS/WEX Corium/Concrete Interaction IODE Iodine in containment SOPHAEROS Aerosol & FP vapor behavior in RCS SYSINT Safety system management ISODOP Isotope treatment CESAR Thermal. In RCS DIVA Core degradation ELSA FP release CORIUM Corium behavior in containment RUPUICUV Corium ejection & Entrainment in cont. DYNAMIC MEMORY CESAR DIVA ELSA SOPHAEROS RUPUICUV MEDICIS/WEX CPA IODE ISODOP SYSINT Figure 9: Calculation modules of ASTEC code Andrea Bachratá Page 11 Master thesis

12 5. Simulations of HERMES-HALF experiment The stand-alone CESAR module of the ASTEC code can be used to simulate thermalhydraulic behaviour in a flooded cavity in case of severe accident with core melting. The objective of this work is to demonstrate the capability of CESAR module of the ASTEC code to calculate two-phase natural circulation flow in large-scaled and curved channel geometry. The validation of the code with the experimental results presented in this Section represents an important step in application of ASTEC code to ERVC phenomena. 5.1 Modeling of HERMES-HALF experiment in CESAR To simulate HERMES-HALF experiment, the stand-alone CESAR module of the ASTEC code V1.33 was used. Because of a non-heating experiment, the stand-alone CESAR module of the ASTEC code with an injection of incondensable gas was used. In sections below, the main parameters of the CESAR input-deck will be summarized. The calculated results will be later compared with the experimental results in Section Geometry of a cooling channel In the CESAR input deck, the hydraulic channel is modeled by the objects called VOLUMES [8]. The volumes are connected by JUNCTIONS into flow path (e.g. circulation system). The real dimensions of HERMES-HALF experimental facility were respected. The main diameters were taken from the Figure 6 and Table 1. Other parameters were calculated. The geometry used in CESAR input deck is shown in Figure Volumes in the input deck In the CESAR input deck, the cooling channel was divided into different VOLUMES [8]. The cooling channel is represented by volumes INTERM to TOP. Volumes INTERM to POOLTOP represent the lateral channel. The nodalization is shown in Figure 1. Andrea Bachratá Page 1 Master thesis

13 0.780 m m 1.87 m 4.58 m Outlet,065 m 3.61 m,5 m m [0,0] R1.69 m 0.18 m m m m m 0,05 m m Inlet m m Figure 10: Geometry of HERMES-HALF for the input deck of CESAR Andrea Bachratá Page 13 Master thesis

14 Modeling of cooling channel along the vessel wall ( riser ) For each VOLUME, the elevation, hydraulic diameter and the volume were calculated. The values Z and R1 (R) (see Figure 1) were found using stand-alone DIVA module meshing of vessel wall, to obtain the geometry almost hemispherical geometry [13]. The elevation of the centre of the volume was calculated as follows: 1 Elev center = Z h (3) The hydraulic diameter of the volume was calculated as follows: D _ hdown + D _ hup D _ h avg = D _ h = * gap thickness down / up (4) The volume was calculated as follows (see Figure 11): V1 V V = 1 V1 = πh 3 1 V = πh The results of the calculations are summarized in Table 3. ( A1 + A1A + A ) ( R + R R + R ) (5) R A RPV wall Cooling channel R 1 A 1 Figure 11: Calculation of volumes in cooling channel along the vessel wall Modeling of lateral channel ( downcommer ) The elevation of each volume in the lateral channel is the same as the elevation of volumes in the cooling channel. The hydraulic diameter is equal to *r where r = 0.78 m. The volume is calculated as follows: 1 V = π r h (6) Andrea Bachratá Page 14 Master thesis

15 The results of the calculations are summarized in Table Junctions in the input deck The connections between VOLUMES in CESAR input deck are called JUNCTIONS. The junctions were modeled between two neighbouring volumes. The elevation and the cross section area of the junction were specified. The elevation of the junction is taken from Table 3 (parameter Z). The cross section area of the junction was calculated for the INTERM to TOP volumes as follows (see Figure 11): and for INTERM to POOLTOP volumes: The results are summarized in Table 4. S S = S S 1 1 = πa = πr S (7) V S = (8) h In the HERMES-HALF experiment the inlet and outlet areas for water path were varied. In CESAR input deck, this fact was modeled changing the cross section area of the junction (enter: INLETCHA, exit: OUTCH, see Table 4). Andrea Bachratá Page 15 Master thesis

16 TOP POOLTOP h LOWA LATA LOW9 LAT9 [0,0] A LOW8 LAT8 A1 LOW7 LAT7 R Z LOW6 LOW5 LOW4 R1 LOW3 LOW LOW1 LOW0 LAT6 LAT5 LAT4 LAT3 LAT LAT1 LAT0 INTERM INTERM Figure 1: Nodalization used in CESAR input deck Andrea Bachratá Page 16 Master thesis

17 Table 3: Parameters of the VOLUMES for the CESAR input deck Name Z [m] h [m] R1 [m] R [m] A1 [m] A [m] VOLUME1 [m3] VOLUME [m3] VOLUME [m3] Elev center [m] D_h avg [m] INTERM -1,3190 1,5500 0,0000 0,0000 1,0970 1,0970 5,8560 0,0000,980 -,094,194 LOW0-1,690 0,0500 0,0000 0,0000 1,0970 1,0970 0,1889 0,0000 0,0945-1,94,194 LOW1-1,450 0,040 0,0000 0,476 1,0970 1,0970 0,0907 0,0015 0,0446-1,57 1,9464 LOW -1,170 0,0730 0,476 0,4856 1,0970 1,0970 0,758 0,0319 0,10-1,085 1,4609 LOW3-1,0550 0,1170 0,4856 0,7050 1,0970 1,0970 0,441 0,1317 0,155-1,1135 1,0035 LOW4-0,8973 0,1577 0,7050 0,8973 1,0970 1,0970 0,5959 0,3194 0,1383-0, ,5917 LOW5-0,7550 0,143 0,8973 1,000 1,0970 1,0970 0,5377 0,411 0,0633-0,8615 0,765 LOW6-0,4856 0,694 1,000 1,170 1,0970 1,60 1,1788 1,0178 0,0805-0,603 0,167 LOW7-0,1800 0,3056 1,170 1,560 1,60 1,4400 1,7540 1,4148 0,1696-0,338 0,74 LOW8 0,0000 0,1800 1,560 1,870 1,4400 1,4400 1,170 0,9138 0,191-0,09 0,337 LOW9 0,9635 0,9635 1,870 1,870 1,4400 1,4400 6,734 5,011 0,6311 0,4818 0,306 LOWA 1,970 0,9635 1,870 1,870 1,4400 1,4400 6,734 5,011 0,6311 1,4453 0,306 TOP 3,61 1,3340 1,870 1,870 1,4400 1,4400 8,6858 6,9381 0,8740,5940 0,306 Name h [m] VOLUME [m3] Elev center [m] D_h [m] INTERM 1,55 1,4800 -,094 1,56 LAT0 0,05 0,0478-1,94 1,56 LAT1 0,04 0,09-1,57 1,56 LAT 0,073 0,0697-1,085 1,56 LAT3 0,117 0,1118-1,1135 1,56 LAT4 0,1577 0,1506-0,976 1,56 LAT5 0,143 0,1359-0,86 1,56 LAT6 0,694 0,573-0,603 1,56 LAT7 0,3056 0,919-0,338 1,56 LAT8 0,18 0,1719-0,09 1,56 LAT9 0,9635 0,903 0,4818 1,56 LATA 0,9635 0,903 1,4453 1,56 POOLTOP 1,334 1,740,594 1,56 Andrea Bachratá Page 17 Master thesis

18 Table 4: Parameters of the JUNCTIONS for CESAR input deck Name Downstream volume Upstream volume S [m] D_h [m] Elev [m] INLETCHA LOW0 INTERM 0,0044-0,15 0,075-1,319 OW0 LOW1 LOW0 1,889,194-1,69 OW1 LOW LOW1 1,84 1,699-1,45 OW LOW3 LOW 1,655 1,3-1,17 OW3 LOW4 LOW3 1,315 0,784-1,055 OW4 LOW5 LOW4 0,868 0,399-0,897 OW5 LOW6 LOW5 0,443 0,154-0,755 OW6 LOW7 LOW6 0,30 0,180-0,486 OW7 LOW8 LOW7 0,559 0,368-0,180 OW8 LOW9 LOW8 0,718 0,306 0,000 OW9 LOWA LOW9 0,655 0,306 0,964 OUTLETCH TOP LOWA 0,655 0,306 1,97 OUTCH POOLTOP TOP 0,075-0,15 0,60,115 POOLJ LATA POOLTOP 0,955 1,560 1,97 POLAT9 LAT9 LATA 0,955 1,560 0,964 POLAT8 LAT8 LAT9 0,955 1,560 0,000 POLAT7 LAT7 LAT8 0,955 1,560-0,180 POLAT6 LAT6 LAT7 0,955 1,560-0,486 POLAT5 LAT5 LAT6 0,955 1,560-0,755 POLAT4 LAT4 LAT5 0,955 1,560-0,897 POLAT3 LAT3 LAT4 0,955 1,560-1,055 POLAT LAT LAT3 0,955 1,560-1,17 POLAT1 LAT1 LAT 0,955 1,560-1,45 POLAT0 LAT0 LAT1 0,955 1,560-1,69 INLETCH3 INTERM LAT0 0,955 1,560-1,319 INLETCH INTERM INTERM,900 1,690 -, Coefficients for pressure losses For each junction, the coefficient for pressure losses should be prescribed. Generally, in continuous and smooth tubes (e.g. downcomer in CESAR input deck of HERMES) the proportionality coefficient should be zero. In the experiments like SULTAN and ULPU these coefficients were virtually zero [9]. In the CESAR input deck, the effect of grid model in the JUNCTION between INTERM and LOW0 was respected (see Figure 13). The explanation of using coefficient different from zero between INTERM and LOW0 can be found in literature [10]. There can be found a way how to calculate the coefficient for the flow through the grid. In HERMES-HALF experiment, the thickness of the plate with the inlet holes was 10 mm. The hydraulic diameter of one hole was 75 mm. The coefficient of free section F 0 /F 1

19 should be taken into account. The F 0 represents the surface of free section in the grid. The F 1 represents the cross section area of canal before the grid. In our case, this coefficient reaches the values from 0.00 (=0.0044/1.889) to (=0.15/1.889). Consequently, the coefficient for pressure losses is from to (for l/d h = 10/75 = 0.13). These values are corresponding to the flow velocity before the grid. These values should be recalculated for the flow velocity in the inlet using the equation: So ASTEC = ζliterature S 1 ζ (9) Where S 0 is the cross section area of the junction declared in CESAR input deck, and S 1 is the cross section area before the grid (1.889 m ). 10 mm LOW0 INTERM Figure 13: Grid model Table 5: Grid model in the junction between INTERM and LOW0 Cross section area of junction [m ] F 0 /F 1 z literature (for l/d H = 0.13) z ASTEC The grid model has the strongest effect on the CESAR calculations of circulation mass flow rate. The coefficients of pressure losses calculated using literature [10] seem to be overestimated. There is an assumption that the coefficients in grid model should be somewhere between 1 and.7. The coefficients in Table 5 correspond to uniform distribution of opening holes. In the HERMES-HALF experiment, the inlets were distributed at the circumstance. As for the calculation results presented in this work, the proportionality coefficient for the grid model was set to 1.

20 5.1. The source of incondensable gas The parameters of the air supply system of the HERMES-HALF facility were summarized in Section The air injection into the gap between the RPV and insulation was simulated in CESAR by hydrogen injection into the volumes. At the present time only hydrogen as an incondensable gas is available in CESAR. The gas injections are modeled with Source boundary conditions and by keeping the experimental volume flow rate. The volume flow rate of air during the HERMES-HALF experiment could be calculated using equation (1). It can be found in the literature [], that the volume air flow rate 8379 m 3 /h should correspond to 100% of heat transfer through the RPV of APR1400 halfscaled geometry in a case of severe accident. Using the equation (1) and Figure 7, the distribution of the air flow rate into volumes LOW1-LOW8 could be calculated as is summarized in Table 6. That means, that the air flow rate 837.9m 3 /h (simulation of 10% of total heat flux) will be injected into the different volumes in CESAR using the per-cent partition that is summarized in a table below. Table 6: Redistribution of the air flow rate into different volumes Volume LOW1 LOW LOW3 LOW4 LOW5 LOW6 LOW7 LOW8 % Redistribution In the CESAR input deck, hydrogen was injected into different volumes instead of air. If the volume mass flow rate of air and hydrogen is the same, the effect of air injection and hydrogen injection is similar. The natural circulation is caused due to different hydrostatic pressures ~ gh ρ ρ ) in the volumes INTERM-TOP (with bubbles) and ( water air / hydrogen INTERM-POOLTOP (without bubbles). Even if the density of air and hydrogen is not same, it is negligible compared to the density of water. 5. Calculation results of HERMES-HALF in CESAR The objective of CESAR calculations was to evaluate its capability of two-phase flow modeling in a specific geometry of an external reactor vessel cooling circuit. In this section, the results of CESAR calculations of HERMES-HALF experiment will be presented. The calculations of water circulation mass flow rate will be compared with the experimental results. The effect of water inlet/outlet area and of hydrogen injection mass flow rate will be discussed.

21 5..1 Water circulation mass flow rate for outlet area 0.15 m In this section, the calculation of water circulation flow rate with a change of hydrogen injection mass flow rate from 5% to 100% of total will be presented. The CESAR calculations were performed for different inlet areas ( m ) and the outlet area 0.15 m. The results are presented in figures below. In the Figure 14, Figure 15 and Figure 16, we can see the effect of the hydrogen injection mass flow rate for the highest values of the inlet area in the investigable range. For low values of the flow rates, the circulation mass flow rate increases with the hydrogen injection mass flow rate. The increase in the circulation mass flow rate is up to ~5% hydrogen injection mass flow rate. From the HERMES-HALF experiment it follows that the circulation mass flow rate depends roughly linearly on the hydrogen injection. But the experiments were performed only for relatively narrow range of air injection (<5%). When the air injection became higher, the pressure losses in riser and in its outlet part should increase in a quadratic way with coolant velocity. Consequently, the circulation mass flow rate should not increase linearly with increasing of air (hydrogen in ASTEC) injection to 100% but should became saturated or even decreased. In the Figure 17 to Figure 0, we can see the effect of the hydrogen injection on circulation mass flow rate for the smaller values of the inlet area in the investigable range. For a small inlet area, the hydrogen injection mass flow rate has almost no effect on water circulation mass flow rate when it is greater than 40-50%. The aim of this work was to demonstrate the capability of stand-alone CESAR module of the ASTEC code to calculate two phase natural circulation flow. The results of CESAR are close to the experimental results on the experimental range. The difference between the CESAR results and the experimental results is about 10%.

22 HERMES-HALF ASTEC Ver.133 Circulation Mass Flow Rate[kg/s] Experiment Air (I.A:0.15m) ASTEC H (I.A:0.15m) Flow Rate[%] Figure 14: Circulation mass flow rate for outlet area 0.15 m and inlet area 0.15m 160. HERMES-HALF ASTEC Ver Circulation Mass Flow Rate[kg/s] Experiment Air (I.A:0.0795m) ASTEC H (I.A:0.0795m) Flow Rate[%] Figure 15: Circulation mass flow rate for outlet area 0.15 m and inlet area m

23 160. HERMES-HALF ASTEC Ver Circulation Mass Flow Rate[kg/s] Experiment Air (I.A:0.044m) ASTEC H (I.A:0.044m) Flow Rate[%] Figure 16: Circulation mass flow rate for outlet area 0.15 m and inlet area 0.044m 160. HERMES-HALF ASTEC Ver Circulation Mass Flow Rate[kg/s] Experiment Air (I.A:0.065m) ASTEC H (I.A:0.065m) Flow Rate[%] Figure 17: Circulation mass flow rate for outlet area 0.15 m and inlet area 0.065m

24 80. HERMES-HALF ASTEC Ver Circulation Mass Flow Rate[kg/s] Experiment Air (I.A:0.013m) ASTEC H (I.A:0.013m) Flow Rate[%] Figure 18: Circulation mass flow rate for outlet area 0.15 m and inlet area 0.013m 55. HERMES-HALF ASTEC Ver Circulation Mass Flow Rate[kg/s] Experiment Air (I.A:0.0088m) ASTEC H (I.A:0.0088m) Flow Rate[%] Figure 19: Circulation mass flow rate for outlet area 0.15 m and inlet area m

25 30. HERMES-HALF ASTEC Ver Circulation Mass Flow Rate[kg/s] Experiment Air (I.A:0.0044m) ASTEC H (I.A:0.0044m) Flow Rate[%] Figure 0: Circulation mass flow rate for outlet area 0.15 m and inlet area m 5.. Water circulation mass flow rate for outlet area m In this section, the calculation of water circulation flow rate for hydrogen injection mass flow rates from 5% to 100% of total value will be presented. The CESAR calculations were performed for different inlet areas ( m ) and the outlet area m. The results are presented in figures below. In Figure 1 and Figure, we can see the effect of hydrogen injection mass flow rate on the circulation mass flow rate for large inlet areas. We can see that the hydraulic behaviour in the cooling channel is similar as it was in a case of 0.15 m. However, the maximum circulation mass flow rate reaches lower values (~ 135 kg/s compared to ~155 kg/s for 0.15 m outlet area). For low values of the gas flow rates, the circulation mass flow rate increases almost linearly with the hydrogen injection mass flow rate. After the injection about 5% of total value, the increased pressure losses take effect (they depend quadratic on flow velocity), and the circulation mass flow rate became saturated. Moreover, there can be even seen a small decrease in a circulation mass flow rate. In Figure 3 and Figure 4, we can see the circulation mass flow rate in a case of small inlet area and outlet area m. In this case, the circulation mass flow rate stops to increase from about 40-50% of hydrogen injection mass flow rate. For the investigated experimental range (10-0% flow rate) the ASTEC results are in agreement with the experimental results.

26 HERMES-HALF ASTEC Ver.133 Circulation Mass Flow Rate[kg/s] Experiment Air (I.A:0.0795m) ASTEC H (I.A:0.0795m) Flow Rate[%] Figure 1: Circulation mass flow rate for outlet area m and inlet area m 140. HERMES-HALF ASTEC Ver Circulation Mass Flow Rate[kg/s] Experiment Air (I.A:0.044m) ASTEC H (I.A:0.044m) Flow Rate[%] Figure : Circulation mass flow rate for outlet area m and inlet area 0.044m

27 60. HERMES-HALF ASTEC Ver Circulation Mass Flow Rate[kg/s] Experiment Air (I.A:0.0088m) ASTEC H (I.A:0.0088m) Flow Rate[%] Figure 3: Circulation mass flow rate for outlet area m and inlet area m 30. HERMES-HALF ASTEC Ver Circulation Mass Flow Rate[kg/s] Experiment Air (I.A:0.0044m) ASTEC H (I.A:0.0044m) Flow Rate[%] Figure 4: Circulation mass flow rate for outlet area m and inlet area m

28 6. Simulations of HERMES-HALF with heat The capability of CESAR module to calculate the natural circulation flow in a typical ERVC geometry with a hemispherical vessel and a large annular space was demonstrated in previous Section. Because of a half-scaled geometry of HERMES-HALF experiment, the results of this experiment are not directly applicable for full-scale APR1400 NPP real situation. Furthermore, the bubbles formation is due to air injection instead of heat transfer from RPV wall. Therefore the bubbles do not appear in the same place, as it would occur in a case of heat transfer. In HERMES-HALF, the bubbles of air are injected into the hemispheric part. In a case of experiment with heat, the first bubbles would be created in the upper part of the channel (because of decrease of hydrostatic pressure and thus, decrease of saturation temperature). Since the heating method is very difficult and expensive for a large-scale and threedimensional spherical test section, the non-heating method of an air injection was decided for HERMES-HALF experiment []. In this section, the representativeness of this non-heating experiment will be demonstrated. 6.1 Modeling of HERMES-HALF with heat in CESAR The thermo-hydraulic behaviour in HERMES-HALF geometry can be predicted using CESAR module of the ASTEC code V1.33. The geometry in CESAR input deck was presented in Section 5.1. In the calculations presented in Sections below, the natural circulation is induced due to heat transfer and vapour formation Heat transfer modeling To simulate a heat transfer to the CESAR volumes, the walls that belong to these volumes should be modeled. In the CESAR input deck, the walls that belong to volumes: LOW1-LOW8 (see Figure 1) were modeled. The heat transfers is modeled with Heat boundary conditions by keeping the heat transfer distribution predicted by MAAP code calculations (see Figure 7 [7]). Using this boundary condition in CESAR, it is not possible to prescribe directly the heat flux through the wall to the fluid in unit W or W/m. The heat transfer for the CESAR input deck should be prescribed using heat transfer coefficient h. S and the temperature T in (see Figure 5). The user should prescribe these two parameters to simulate the heat flux through the wall that he knows (e.g. from DIVA module of ASTEC code or from another code, from experimental measurements ).

29 Q bc Q Q wf T in h in T w Tw Tp e Figure 5: The heat transfer through the CESAR wall As for the HERMES-HALF calculations with heat, the heat flux was prescribed using the heat flux distribution obtained by MAAP calculation (see Figure 7) [7]. The heat transfer coefficient for CESAR input deck was calculated using equation (10). The heat flux in equation (10) was brought into a harmony with temperature T in and order to obtain the desired heat flux values. ( hs ) wall element in in wall in T (see Figure 5) in w Q = (10) T T There has been an effort to model a slowly increase of the heat flux with time. The heat transfer coefficient in CESAR input deck couldn t be prescribed with time. On the other hand, the temperature could be prescribed with time. Consequently, the p% of heat flux through the wall element was calculated as follows: Q p% in = 100 Twall Tinitial T ) +. + % in p Tinitial p (11) hs 100 ( % The initial temperature of wall was set to 383 K. The heat flux through the wall increased slowly from 0% to 100%.

30 6. Calculation results of HERMES-HALF with heat in CESAR The objective of CESAR thermo-hydraulic simulations was to predict the natural circulation mass flow rate in HERMES-HALF geometry induced in ERVC loop due to heat transfer from reactor wall and vapour formation in riser part of the loop. In this section, the results of CESAR thermo-hydraulic calculations will be presented. The calculations of water circulation mass flow rate will be compared with CESAR hydraulic calculations with hydrogen injection (validated model in Section 5.) and experimental results with air injection. Consequently, a representativeness of experimental results with air injection will be discussed Water circulation mass flow rate for outlet area 0.15 m and m In this section, the calculation of water circulation flow rate with a change of heat flux through the wall from 0% to 100% of total value will be presented. The heat flux corresponding to 100% value is shown in MAAP curve in Figure 7 [7]. The CESAR calculations were performed for different inlet areas and the outlet areas 0.15 m and m. In the figures below, we can see the oscillations in circulation mass flow rate up to about 60% of heat flux. This circulation mass flow rate reflects the oscillations of void fraction. The CESAR calculation of void fraction in the cooling channel shows significant instabilities especially for lower heat flux values. The numerical oscillations in CESAR were also discussed in CESAR calculations presented by another authors [9]. In the model of HERMES-HALF with hydrogen injection presented in Chapter 5, the gas was injected into a lower part of the circuit. In the model of HERMES-HALF with heat, the first vapour bubbles were generated at the upper part of the cooling circuit (because of decrease of hydrostatic pressure). Consequently, because of vapour formation and decrease of density in the upper part of the circuit, the vapour bubbles were created also in the lower part of the circuit. The creation of vapour bubbles also in lower part of the circuit depends also on circulation mass flow rate. There exist certain geometrical arrangement (cross-section area of inlet and outlet to/from riser) when the vapour bubbles do not occur in the lower part of the circuit because of efficient cooling (high circulation mass flow rate). There is a suspicion that for some geometry, the injection of air/hydrogen into a lower part of the circuit can underestimate the prediction of circulation mass flow rate, as it would occur with heated wall. The geometries that give the circulation flow about 150 kg/s presented in Section 5. gave the circulation flow about 400 kg/s with heat and vapour formation. These calculations are not presented in this work and the effects that caused this huge difference should be studied in further work. Nowadays, the only known difference between hydrogen and heat calculation is the position distribution of void fraction in the cooling channel.

31 80. HERMES-HALF ASTEC Ver Circulation Mass Flow Rate[kg/s] ASTEC Heat (I.A:0.013m) ASTEC H (I.A:0.013m) Experiment Air (I.A:0.013m) Flow Rate alias Heat Flux [%] Figure 6: Circulation mass flow rate for outlet area 0.15 m and inlet area 0.013m 60. HERMES-HALF ASTEC Ver Circulation Mass Flow Rate[kg/s] ASTEC Heat (I.A:0.0088m) ASTEC H (I.A:0.0088m) Experiment Air (I.A:0.0088m) Flow Rate alias Heat Flux [%] Figure 7: Circulation mass flow rate for outlet area 0.15 m and inlet area m

32 30. HERMES-HALF ASTEC Ver Circulation Mass Flow Rate[kg/s] ASTEC Heat (I.A:0.0044m) ASTEC H (I.A:0.0044m) Experiment Air (I.A:0.0044m) Flow Rate alias Heat Flux [%] Figure 8: Circulation mass flow rate for outlet area 0.15 m and inlet area m 60. HERMES-HALF ASTEC Ver Circulation Mass Flow Rate[kg/s] ASTEC Heat (I.A:0.0088m) ASTEC H (I.A:0.0088m) Experiment Air (I.A:0.0088m) Flow Rate alias Heat flux [%] Figure 9: Circulation mass flow rate for outlet area m and inlet area m

33 30. HERMES-HALF ASTEC Ver Circulation Mass Flow Rate[kg/s] ASTEC Heat (I.A:0.0044m) ASTEC H (I.A:0.0044m) Experiment Air (I.A:0.0044m) Flow Rate alias Heat flux [%] Figure 30: Circulation mass flow rate for outlet area m and inlet area m 6.. Vessel failure and critical heat flux Failure of the lower head submerged into water may be caused due to boiling crisis on the external (i.e. cooled) RPV surface. It occurs, when the heat flux through RPV wall exceeds the critical heat flux. The result is that the flow regime transits suddenly from nucleate to film boiling, what results in significant decreasing of heat transfer coefficient and thus increasing of RPV wall temperature. As soon as there is subcooled or nucleate boiling on external RPV surface, the heat transfer coefficients here are sufficient for reliable cooling of RPV [1]. The results presented in a section above demonstrated that in most of cases, the experimental results with air injection corresponded to the circulation mass flow rate as would occur in a case of heat transfer and vapour formation. On the other hand, the vessel failure and values of critical heat flux can never be estimated from the experiment with air. In a case of in-vessel retention concept that will be implemented at APR1400 plant and other plants, the estimation of critical heat flux is very important. It is not possible to prove, only using hydraulic test that there will be no vessel failure and that the natural circulation is sufficiently high to prevent the boiling crisis on the outer RPV surface.

34 The CESAR calculations with heat generated in reactor wall could afford the ideas about the values of critical heat flux and possibilities of vessel failure. The calculations in this work were performed only for half-scaled geometry of APR1400 NPP. To predict real situation that would occur in a case of severe accident at APR1400 plant, it is recommended, in future, to perform the CESAR calculations for whole APR1400 geometry Equations for critical heat flux For the evaluation of the critical heat flux value, CESAR uses the physically based Zuber s correlation for pool boiling with the correction on liquid subcooling due to Ivey and Morris [11]: ϕ T CHF L < T = 0.149C sat C sub sub gσ ρ L ρ = 1 ( ρ ) G G 0.5 ρ G h LG (1) This correlation was developed for flow in homogeneously heated narrow vertical tube and does not take into account curvature and inclination of the heated surface. However, the CHF value depends on concrete geometrical configuration, which enables natural circulation of water in flooded cavity and thus effective heat removal from RPV wall to confinement atmosphere. This could significantly decrease or increase (e.g. in case of optimisation) the value of CHF compared to results using equation (1). The equation that could correctly represent the values of critical heat flux respecting all-important parameters for in-vessel retention was estimated from SULTAN experimental results [1]. However, the SULTAN correlation is a result of experiment with forced convection. The SULTAN correlation is giving heat flux (F) in MW/m, in term of cover pressure (P) in MPa, mass velocity (G) in kg/s/m, local thermodynamic quality (X), gap (E) in m and inclination (θ = sin (l), l inclination above horizontal): F = A0 + A1* X + A* X A0 = B0 + B1* E * G + B / P A1 = B7 * G + B8* E * G + A3* Θ + A4* Θ A = B9* E A3 = B10* G + B11* E * P + B1* X * G A4 = B13* P + B14* G + B15* X + B16* E G = LN ( G) + B3* G + B4 * E / P + B5* E / P + B6* P * G (13)

35 B0 = B4 = B8 = B1 = B16 =.636 B1 = B5 = B9 = B13 = B = B6 = B10 = B14 = B3 = B7 = B11 = B15 = The Zuber s correlation (1) for critical heat flux calculation is yet implemented into the CESAR source. There has been an effort to obtain the values of critical heat flux using SULTAN correlation (13). It can be expected that especially for heated surfaces inclined from vertical position, the critical heat flux predicted by SULTAN correlation is significantly smaller than the critical heat flux predicted by Zuber s correlation. Thus, the values of physical parameters needed to SULTAN correlation should be calculated correctly by CESAR up to the moment of critical heat flux. An automatic procedure using the SIGAL tool (written in Fortran language and used in ASTEC to manage the dynamic database of the code variables along the calculation) has been written to obtain the critical heat flux values using SULTAN correlation. Thanks to this procedure, the user can obtain the results of equation (13) in every moment of CESAR calculations. The results can be visualized and plotted. The quantity X for SULTAN correlation is calculated from the CESAR void fraction as follows: α X = 1 α (14) ρ Lu L α + ρ u 1 α V V where α is the void fraction, ρ L,V is the density of liquid or vapour and u L,V is the velocity of liquid or vapour in m/s CESAR calculations of critical heat flux in HERMES-HALF The CESAR calculations with heat were performed for different inlet/outlet geometries. The results were presented in Section As for the critical heat flux values, two different correlations were compared (Zuber s and SULTAN). On Figure 31 we can see the heat flux through the vessel wall (black color), the Zuber s prediction of critical heat flux and SULTAN correlation. It is possible to plot the Figure 31 for every tested geometry. For conclusion, the critical heat flux predicted by SULTAN or by Zuber s correlation was never reached. It is important to note, that the SULTAN correlation of critical heat flux depends on many parameters (see Section 6...1). The prediction of CHF obtained with this equation reflects all fluctuations of circulation mass flow rate. If we plot the values of critical heat flux

36 predicted by SULTAN e.g. close to wall RP5 with time, the curve of CHF would be oscillating. Moreover, the SULTAN correlation was derived from experiments with forced circulation and thus, it is not defined if counter current flow occurs (see LN in equation (13)). On Figure 31 we can see that the CHF by SULTAN is increasing with polar angle (inclination of RP1 is about 10 and RP8 is about 90 ). We can see that Zuber s correlation differs from SULTAN correlation especially for lower angle positions, because the Zuber s correlation was estimated for flow in a homogeneously heated narrow vertical tube. The SULTAN correlation is respecting the polar angle. Moreover, it is important to note that the bottom of the reactor vessel (angle 0 ) represents the flow along horizontal wall. Consequently, the Zuber s correlation especially in the lower part of the vessel seems to be optimistic... HERMES-HALF with HEAT ASTEC Ver RP8 Heat Flux[MW/m] RP Zuber Correlation Heat flux from wall SULTAN Correlation Wall[RP]: possiton along vessel wall from bottom to top Figure 31: Critical heat flux predicted by Zuber s and SULTAN correlation

37 7. Conclusion The in-vessel retention concept is implemented into a design of APR1400 NPP in Korea. The applicability of this concept is based on experimental program (HERMES) and computer codes calculations (RELAP) [6]. In this work, the simulations of HERMES-HALF experiment with code ASTEC (module CESAR) were presented. At first, the all-important characteristics of HERMES- HALF experiment were summarized. The HERMES-HALF experiment has been performed to measure the two-phase natural circulation mass flow rate through the annulus gap between the reactor vessel and the insulation material. The aim of the experiment was to observe and evaluate the two-phase natural circulation phenomena in a case of external reactor vessel cooling of APR1400 high power reactor. Since the heating method is very difficult and expensive for a large-scale and three-dimensional spherical test section, the non-heating method of an air injection was decided for HERMES-HALF experiment. However, the air injection capacity was available only up to 4 % of nominal rated value. The HERMES- HALF experiment represents an interesting hydraulic test for CESAR validation in IVR conditions because of annular space and curved channel geometry, what are characteristic features of ERVC loops. An input deck for the model of HERMES-HALF experiment was created for the code ASTEC (module CESAR). At the present time only hydrogen is available in CESAR as an incondensable gas. Therefore, the hydrogen injection into the cooling channel was used instead of air injection, used in experiment. However, if the volume mass flow rate of air and hydrogen is kept, the effect of air injection and hydrogen injection is similar. Even if the density of air and hydrogen is not the same, the difference is negligible compared to the density of water. The CESAR calculation could be compared with the HERMES-HALF experimental results only up to 4% of air injection flow rate. In this range, the CESAR calculations correspond with the experimental results. The maximum differences between the CESAR results and the experimental results are about 10-0%. The comparison of the ASTEC code simulation with the HERMES-HALF experimental results is an important step for the twophase natural circulation flow code s validation in the area of ERVC applications because of the specific geometry of this experiment. The second part of this work represents the thermal-hydraulic calculation as could occur in a case of heating experiment. Using the HERMES-HALF geometry, the heat flux through the wall was prescribed as a boundary condition. The calculated results showed the similar values of circulation mass flow rate for most of input/output geometries. Finally, the calculations of critical heat flux values were presented. The importance of calculations of critical heat flux values is because the failure of the lower head submerged into water may be caused due to boiling crisis on the external (i.e. cooled) RPV surface. From the thermo-hydraulic calculations, time and position of vessel failure can be revealed. However, the standard code s correlation for CHF estimation was derived for vertical bundle of fuel rods and thus is not conservative for estimation of heat flux from inclined curved surface