Analyses of Unprotected Transients in the Lead/Bismuth-Cooled Accelerator Driven System PDS-XADS

Transcription

1 Analyses of Unprotected Transients in the Lead/Bismuth-Cooled Accelerator Driven System PDS-XADS Tohru Suzuki, Xue-Nong Chen, Andrei Rineiski, and Werner Maschek Forschungszentrum Karlsruhe, Institute for Nuclear and Energy Technologies Postfach 3640, D Karlsruhe, Germany ABSTRACT Safety investigations for the lead/bismuth (Pb/Bi)-cooled experimental accelerator driven system (XADS) were performed with the computational code SIMMER-III. The code has been extended so as to describe the ADS configuration with heavy-metal coolant, sub-critical core and strong external neutron source. As transient scenarios, unprotected loss of flow (ULOF), unprotected subassembly blockage and unprotected transient over-power (UTOP) were simulated. The calculation results showed that the current Pb/Bi-cooled XADS design proposed in the framework program of the European Union has remarkable resistance against severe scenarios. I. INTRODUCTION Accelerator Driven Systems (ADSs), which combine a subcritical reactor with a high-energy proton accelerator and a spallation target, are under investigation for their transmutation potential of long-lived nuclear waste. Within the 5th framework program of the European Union, Preliminary Design Studies of an Experimental Accelerator Driven System (PDS-XADS) are performed [1]. This work consists of two options: heavy metal (lead/bismuth)-cooled and gas (helium)-cooled ADSs [2]. Besides the design of these devices, a major part of the work is devoted to safety investigations of this innovative system. The transient behavior of the ADSs under conditions leading up to accidents is investigated. For the safety investigation of the lead/bismuth (Pb/Bi)-cooled XADS, the authors have been using the computational code SIMMER-III which was originally developed for the safety analysis of sodium-cooled fast reactors [3, 4]. In order to adapt SIMMER-III to the Pb/Bi-cooled XADS analyses, the code has been extended so as to describe a subcritical core with a strong external neutron source [5, 6]. Moreover, additional improvement has been also done for the simulation of multi-phase flows in Pb/Bi-cooled ADSs, which differ significantly from sodium-cooled reactors. In the present study, some unprotected transients in the Pb/Bi-cooled XADS are analyzed with the extended SIMMER-III version. As severe transient scenarios, unprotected loss of flow (ULOF), unprotected subassembly blockage accidents and unprotected transient over-power (UTOP) are simulated. Within these safety investigations the whole spectrum from operational transients to accidents is analyzed. II. SAFETY INVESTIGATION FOR ADS Within the PDS-XADS program, a small scale Pb/Bicooled ADS using Mixed-Oxide (MOX) fuel is under investigation. The power level and the subcriticality level at the beginning of life (BOL) are 80 MWth and 0.970, respectively. Since the basic transient behavior of an ADS should be investigated at this stage, no dedicated transmuter fuel [7] with Minor Actinide (MA: Np, Am, Cm) is loaded in the present analyses. A specific property of the Pb/Bi coolant is its high boiling point (1943 K). This implies that during core disruption clad melting and pin disruption will occur before coolant boiling. II.A. Configuration of a Typical ADS The configuration of a typical ADS is illustrated in Fig. 1. As mentioned before, the ADS has a subcritical core. Protons from a linear accelerator (LINAC) are introduced into the target unit and hit the lead target. As the result of a spallation process, neutrons will be supplied to the core. The heat produced in the core will be carried away by the molten Pb/Bi coolant. The coolant flow is mainly driven by natural circulation, which is enhanced by a gas-lift pump with argon gas. The cross plane of the current XADS core is schematically depicted in Fig. 2. The center of the cross plane is occupied by a beam pipe and the target unit for the external neutron source. The core consists of 120 fuelassemblies, 162 dummy-assemblies and 12 outer absorber-assemblies. Each hexagonal fuel-assembly contains 90 fuel-pins filled with MOX fuel (PuO 2 and UO 2 ), which has a plutonium enrichment of atom% and a porosity of 5.0 %. It should be noted that recently in an updated XADS design a two enrichment-zone core has been developed. The volume fractions of fuel,

2 Argon circulation (gas-lift pump) Intact core Pin fuel Can wall Cladding Gas Plenum-gas leak through cracks (gas blowdown) Linear accelerato (LINAC) Local disruption Coolant Melting of cladding steel (bubbles and steel droplets) Heat exchanger Pb/Bi coolant Spread of disruption Fig. 3. DEC scenario in Pb/Bi-cooled ADS. Release of fuel particles Core Proton Target Neutron Fig. 1. Configuration of a typical ADS. Fuel assembly Target unit Dummy assembly Absorbers # For 2D circular-symmetric simulation; Assign to 4th ring Assign to 5th ring Assign to 6th ring Assign to 7th ring Assign to 8th ring Assign to 9th ring Fig. 2. Cross plane of current XADS. cladding/can-wall steel and coolant in the fuel-assemblies are 0.22, 0.13 and 0.58, respectively. II.B. DEC Scenario with Core Disruption in ADS For a complete safety analyses, Design Extension Conditions (DECs) covering severe accidents with the potential of fuel melting and pin disruption have been analyzed. The scenario of core disruption in the Pb/Bicooled ADS is quite different from that in sodium-cooled fast reactors. In the case of sodium-cooled reactors, coolant boiling will induce clad melting and pin disruption. In Pb/Bi-cooled ADSs, clad melting and fission-gas (helium) blowdown will precede any coolant boiling process. Hence, fuel relocation and multi-phase flow processes become different from those in sodiumcooled reactors. A potential DEC scenario with core disruption in the Pb/Bi-cooled ADS is illustrated in Fig. 3. If coast down of gas-lift pump or blockage in the coolant channel occurs, the heat cannot be removed by coolant flow. As a result, pin disruption in local scale might take place. In this local disruption, a plenum gas will leak through cracks in the cladding steel. This phenomenon is the so-called gas blowdown. When the temperature of the cladding reaches its melting point, bubbles and molten-steel droplets will be released into the coolant channel and go upward by buoyancy. After the disruption of cladding, fuel particles will be also released. Finally, disruptions will spread to the entire core region. We should notice that the behavior of bubbles, steel droplets and fuel particles would affect the possibility of a recriticality. III. COMPUTATIONAL CODE The computational code SIMMER-III used in the present study was developed by Japan Nuclear Cycle Development Institute (JNC) in collaboration with Forschungszentrum Karlsruhe (FZK), Commissariat à l'energie Atomique (CEA) and Institut de Radioprotection et de Sûreté Nucléaire (IRSN). Recently, SIMMER-III has been extended by FZK to describe the ADS specifics for its safety assessment [5, 6]. III.A. Outline of SIMMER-III SIMMER-III is a two-dimensional, three-velocityfield, multi-phase, multi-component, Eulerian fluid

3 dynamics code coupled with a structure model (fuel pins etc.) and space-, time- and energy-dependent neutron dynamic model [3, 4]. The overall fluid-dynamics solution algorithm is based on a time-factorization approach developed for the Advanced Fluid Dynamics Model (AFDM), in which intra-cell interfacial area source terms, heat and mass transfers, and the momentum exchange functions are determined separately from inter-cell fluid convection [8]. In addition, an analytical equation-of-state (EOS) model is introduced to close and complete the fluid-dynamics conservation equations [9]. The structure model represents the configuration and the time-dependent disintegration of fuel pins and subassembly can walls. A simple fuel-pin (SPIN) model, in which the fuel pellet is represented with surface- and interior- temperature nodes, is currently used. The breakup of structure components is based on thermal conditions and no mechanical failure is modeled. The advanced pin model (detailed fuel-pin model: DPIN2) has been successfully applied and tested for the CABRI program [10]. The modeling strongly reflects sodium coolant conditions and the model will be tested for heavyliquid-metal condition in the future. In neutronics, the transient neutron flux distribution is calculated with the improved quasi-static method [11]. For the space-dependent part, a TWODANT-based flux shape calculation scheme (transport theory) has been implemented [12]. III.B. Code Improvement for ADS Safety Analysis For ADS applications, an external space-, time- and energy-dependent neutron source has been implemented in the kinetics equations [7, 13]. For the calculation of the Pool average void fraction, α p [-] Experiment SIM-III (original) SIM-III (improved) Dimensionless gas flux, j g + [-] Fig. 4. Code improvement for Pb/Bi two-phase flow. different kinetic quantities, different weighting functions can be chosen. The source effectiveness indicating the position of the source relative to the bulk of fuel is transiently determined. SIMMER-III uses an elaborate scheme of EOS functions for fuel, steel, coolant, absorber and simulation materials. In the present analysis, EOS functions for Pb and Pb/Bi eutectic coolants have been newly prepared and validated [6]. The fluid-dynamics portion is also improved so as to estimate the gas-pb/bi two-phase flows as formed in the disrupted core of Pb/Bi-cooled ADS. In this improvement, the inter-phase drag between bubbles and molten Pb/Bi can be estimated according to bubble shape [14]. The result of the improvement is presented in Fig. 4. The figure suggests that the improved SIMMER-III has enough accuracy and reliability for the simulation of gas- Pb/Bi two-phase flows. In addition, the gas-blowout model originally developed for SAS4A code [15] is combined with SIMMER-III in order to treat the fission-gas blowdown process during pin disruptions. III.C. Geometric Model for PDS-XADS Analysis The geometric model of SIMMER-III for Pb/Bicooled XADS is illustrated in Fig. 5. The present simulation assumes a two-dimensional circular-symmetric calculation. The sizes of the radial fluid dynamics cells were determined so that the cross areas of hexagonal assemblies categorized with different colors in Fig. 2 would be equal to those of concentric cylindrical rings in the computational circular-symmetric geometry. The sizes of the axial fluid dynamics cells were set up so as to reproduce the actual XADS configuration as closely as possible. The neutronics cells for neutron flux calculation were set up to cover the entire core region. The sizes of the neutronics cells were smaller than 3cm. For these calculations, the SPIN model in SIMMER-III has been chosen, and the effects of the center hole and the porosity in the fuel pellet have also been taken into account. The adequacy of the present geometric model of SIMMER-III was validated through comparison with designed values such as total thermal-power, mean coolant velocity, dynamic pressure drop, and the temperatures of coolant, cladding, and fuel [16]. For the intact core in steady state, the difference between these designed parameters and calculated values is within 1 %. IV. CALCULATION RESULTS AND DISCUSSIONS In the present study, three types of severe transients were simulated in order to investigate the safety potential of the Pb/Bi-cooled XADS. They were unprotected loss of flow (ULOF), unprotected subassembly blockage case and unprotected transient over power (UTOP). Other

4 Target loop Coolant outlet 1.1 Beam pipe Target unit (2) Target unit (1) Target unit (3) Target unit (4) Target unit (5) 1st ~ 3rd rings; mm 4th ring; mm 5th ring; mm 6th ring; mm Grid plate Upper extension Upper plenum Active fuel Lower plenum Upper end-plug Lower end-plug 7th ring; mm 8th ring; mm Lower coolant path Thermal shield Dummy assemblies Thermal shield 9th ring; mm 1100 mm 1272 mm 715 mm Coolant inlet Fig. 5. Geometric model for SIMMER-III analysis. Normalized flow rate and power [-] Cladding temperature at innermost ring [K] Normalized flow rate Normalized power Fig. 6. Coolant flow rate and power for ULOF case Bottom of fuel assembly Midplane of fuel assembly Top of fuel assembly types of severe transients such as inadvertent beam trip, unprotected transient beam over currents and protected transient over power had already been performed in the our previous paper [16]. For the safety investigation of XADS, the word unprotected means no shutdown of the beam from LINAC during severe sequences. IV.A. Unprotected Loss of FLOW (ULOF) ULOF for the present XADS means complete loss of forced coolant circulation. This is a simulation for the coast down of the gas-lift pump. The transient of coolant flow rate and the normalized power are shown in Fig. 6. In the present ULOF calculation, the power generated in the core has been assumed to remain constant rather than to change according to the reactivity feedbacks. After the coast down of the gas-lift pump at 0 sec, the coolant velocity would initially decrease and approach a stable value after reaching a minimum value. This result means the remaining natural convection would form a new steady state. The cladding temperatures in the innermost Fig. 7. Cladding temperatures for ULOF case. fuel assembly are presented in Fig. 7. As shown in the figure, the highest cladding temperature at the top position of the core would never reach its melting point (1800 K) because of the remaining natural convection. Hence, no pin failure can be expected. In the present simulation, the effect of the Reactor Vessel Air Cooling System connected to the intermediate heat exchanger is not considered. Moreover, the associated proton-beam shutdown, which is actuated in the case of high lead/bismuth temperature at core outlet, is also omitted. However, the stable coolant flow rate obtained in the present analysis can give the cladding temperature under the physical barriers. This result supports the safety potential of the current XADS design. IV.B. Unprotected Subassembly Blockage Case The second simulation assumes an unexpected blockage in the innermost ring of fuel assemblies around

5 Highest cladding temperature [K] Pin failure No clad melting Blockage: A/A 0 =0.025 Without radial heat transfer With radial heat transfer Fig. 8. Cladding temperatures for blockage case. Normalized power [-] Fuel sweep-out 1.0 Reactivity Normalized power Fig. 9. Power and reactivity for pin failure case. Reactivity [pcm] the target unit. In this case, the flow area of the assembly ring was artificially reduced to 2.5 %. In order to estimate the effect of radial heat transfer through can-walls, two calculations were performed: with and without radial heat transfers. Fig. 8 shows the highest cladding-temperature in the innermost assembly ring with/without radial heat transfers. After the blockage at 0 sec, the cladding temperature would increase because of decreasing coolant flow rate in the corresponding assemblies. In the case where the radial heat transfer is taken into account, the cladding temperature stayed 100 K below the melting point. It can be seen that the cladding would be prevented from melting by the radial heat transfer through can walls. Since this calculation is performed in a two-dimensional model, heat loss in the azimuthal direction has not been taken into account. If only one subassembly takes the blockage in an actual three-dimensional geometry, we can expect a much lower cladding temperature. Moreover, in the actual PDS-XADS design, the heat in the hot fuel assembly can be removed by the bypass flow of coolant in the gap between adjacent fuel-assemblies. Although this effect is not taken into account in the present simulation, the bypass flow will also bring about lower fuel and cladding temperatures than the current calculation results. On the other hand, in the case without the radial heat transfer, the pin failure occurred at 31 sec. The pin failure and the fuel release into the coolant channel would lead to increases in power and reactivity. This unrealistic scenario is further investigated to test the code on transients going into melting and to assess the basic recriticality potential under melting conditions. The transients of power and reactivity for the pin failure case without radial heat transfer are presented in Fig. 9. After the pin failure at 31 sec, molten cladding droplets and released fuel particles were dispersed in the subassembly. As a result of pin-failure expansion in the vertical (32 sec) (88 sec) (98 sec) Cladding steel 0.01 m/s SA can-wall Volume fraction of released fuel particles Base vector for particle velocity Fig. 10. Fuel sweep-out for pin failure case. direction after 31 sec, the power and reactivity increased. However, the fuel sweep-out into the upper plenum region at 94 sec would bring reactivity reduction and no severe power excursion would occur. Fig. 10 shows the fuel particle distribution after the pin failure and the expanding damage in the innermost assembly. The figure also indicates that the fuel particle could be swept away from the core region and the reactivity would be reduced as a consequence. As an additional simulation in this framework, the upper core region was assumed to be entirely blocked with released steel particles. This blockage simulation was initiated from a condition in which the fuel pins in the upper 40 % part of the core were without cladding and the cladding steel was concentrated in upper region beforehand to form the blockage with maximum packing fraction. In order to consider the severest condition, heat and mass transfers from the core region to the target unit region were artificially suppressed. The transients of power and reactivity for this assumed entire-blockage case are shown in Fig. 11, and the fuel-particle

6 Normalized power [-] Maximum reactivity = pcm Normalized power Reactivity Outer can-wall breakup Fig. 11. Power and reactivity for upper blockage case Reactivity [pcm] Normalized power [-] Normalized power Reactivity Fig. 13. Power and reactivity for UTOP case. Reactivity [pcm] (0 sec) (1 sec) (30 sec) Fuel temperature at innermost ring [K] Bottom of fuel assembly 800 Midplane of fuel assembly Top of fuel assembly Fig. 14. Fuel temperatures for UTOP case. (38 sec) (44 sec) (49.3sec) Cladding steel 0.50 m/s SA can-wall Volume fraction of released fuel particles Base vector for particle velocity Fig. 12. Fuel sweep-out for upper blockage case. distribution process is presented in Fig. 12. After the start of calculation, the un-clad fuel pellets break up into fuel particles. The disruptions of the cladded part of the pins in the lower 60 % part of the core started at 28 sec. As a result, the power and reactivity have specifically increased. The maximum value of reactivity in the sequence was pcm as shown in Fig. 11. The power and reactivity would finally be decreased by fuel dispersions with molten-core pool convections, and would be reduced rapidly by a fuel sweep-out into the outer dummy-assembly region through broken can-wall at 49.3 sec. The calculational results for these unprotected blockage cases with significant or total core disruption suggest that the current XADS design has remarkable resistance against severe scenarios and energetic events even after massive pin-failure cases. IV.C. Unprotected Transient Over-Power (UTOP) A UTOP will be caused by a reactivity perturbation in the core. As no scram-rod system is provided in the current XADS design, the core must survive certain reactivity increases. The possible positive reactivity additions have been already analyzed and identified [1]. For the current design of Pb/Bi-cooled XADS, a positive reactivity addition of 1 $ with a rump rate of 2 $/s at the BOL condition is investigated in this paper.

7 The results of the normalized power and reactivity for UTOP with the condition above are displayed in Fig. 13, where the reactivity addition was installed at 10 sec. The increase of power after the reactivity addition is 10.2 %. Fig. 14 shows the transient of fuel temperature in the innermost subassembly ring. With the UTOP, the maximum increase of fuel temperature at the hottest cell is only 63 K as shown in the figure. Since MOX fuel is used in the present simulation, this fuel-temperature increase would induce the Doppler effect by which reactivity after 10 sec decreases as shown in Fig.13. V. CONCLUSIONS For the safety investigation of Pb/Bi-cooled XADS, unprotected severe transients were simulated with SIMMER-III code. The conclusions of the present study are summarized as follows: (1) Concerning the ULOF case, no pin failure would occur because of remaining natural convection. (2) In the subassembly blockage case with radial heat transfer, no clad melting must be expected. In the case without radial heat transfer, a pin failure occurred but the fuel sweep-out into the upper plenum region would bring a reactivity reduction and no power excursion would result. Finally, the simulations of a massive above-core steel blockage with progressive core disruption show that even under such assumptions no energetic recriticality takes place. (3) For the UTOP case, the calculation shows that the value of the considerable over power is far away from the value that leads to fuel melting. These simulation results suggest that the current PDS-XADS design has remarkable resistance against severe transient scenarios. ACKNOWLEDGMENTS The authors would like to acknowledge their sincere gratitude to D. I. Michael Flad for his technical assistance and fruitful suggestions concerning EOS preparations. This work has been partly funded by the EU Program PDS-XADS, Contract No: FIKW-CT REFERENCES 1. PDS-XADS Preliminary Design Studies of an Experimental Accelerator Driven System, EU Contract No.FIKW-CT (2001). 2. L. MANSANI, K. W. BURN, R. TINTI, B. GIRAUD, R. SUNDERLAND and J. CETNAR, Proposed Core Configuration for a Gas Cooled and a Lead- Bismuth Eutectic Cooled ADS System, ENC2002 Scientific Seminar, Lille, France (2002). 3. SA. KONDO, Y. TOBITA, K. MORITA and N. SHIRAKAWA, SIMMER-III: An Advanced Computer Program for LMFBR Severe Accident Analysis, Proc. International Conference on Design and Safety of Advanced Nuclear Power Plant (ANP 92), Vol. IV, pp , Tokyo, Japan (1992). 4. SA. KONDO, Y. TOBITA, K. MORITA, D. J. BREAR, K. KAMIYAMA, H. YAMANO, S. FUJITA, W. MASCHEK, E. A. FISCHER, E. KIEFHABER, G. BUCKEL, E. HESSELSCHWERDT, P. COSTE and S. PIGNY, Current Status and Validation of the SIMMER-III LMFR Safety Analysis Code, Proc. 7th International Conference on Nuclear Engineering (ICONE-7), No. 7249, Kyoto, Japan (1999). 5. W. MASCHEK, A. RINEISKI, K. MORITA, E. KIEFHABER, G. BUCKEL, M. FLAD, P. COSTE, S. PIGNY, G. RIMPAULT, J. LOUVET, T. CADIOU, SA. KONDO, Y. TOBITA, T. SUZUKI, H. YAMANO and S. FUJITA, SIMMER-III, A Code for Analyzing Transients and Accidents in ADS, Proc. International Topical Meeting on Nuclear Applications of Accelerator Technology (AccApp 00), Washington D. C., USA (2000). 6. W. MASCHEK, A. RINEISKI, E. KIEFHABER, G. BUCKEL, K. MORITA, M. FLAD, P. COSTE, S. PIGNY, J. LOUVET, T. CADIOU, G. RIMPAULT, SA. KONDO, Y. TOBITA and T. SUZUKI, SIMMER-III Code Development for Accelerator Driven System (ADS), Proc. 1st CAPRA Seminar, Windermere, UK (2000). 7. A. RINEISKI, W. MASCHEK and G. RIMPAULT, Performance of Neutron Kinetics Models for ADS Transient Analysis, Proc. International Topical Meeting on Nuclear Applications of Accelerator Technology (AccApp 01), Reno, Nevada, USA (2001). 8. W. R. BOHL, D. WILHELM, F. R. PARKER, J. BERTHIER, L. GOUTAGNY and H. NINOKATA, AFDM: An Advanced Fluid Dynamics Model. Scope, Approach and Summary, Los Alamos National Laboratory Report, LA MS (1990). 9. K. MORITA and E. A. FISCHER, Thermodynamic Properties and Equation of State for Fast Reactor Safety Analysis, Part I: Analytic Equation-of-state Model, Nuclear Engineering and Design, 183, pp (1998). 10. T. CADIOU, W. MASCHEK and A. RINEISKI, SIMMER-III: Applications to Reactor Accident Analysis, Proc. Joint IAEA/NEA Technical Meeting, Pisa, Italy (2002). 11. K. O. OTT and R. J. NEUHOLD, Nuclear Reactor Dynamics, ANS, La Grange Park, USA (1985). 12. G. BUCKEL, E. HESSELSCHWERDT, E. KIEFHABER, S. KLEINHEINS and W. MASCHEK,

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