Development of a Passive Reactor Shutdown Device to Prevent Core Disruptive Accidents in Fast Reactors: A Preliminary Study

Fast reactor cores are not in their most reactive configuration under normal operating conditions, and therefore, they have possibility to result in positive reactivity changes when assuming coolant boiling, cladding discharge, and fuel concentration [1]. Since the early stages of fast reactor development, this potential has highlighted the importance of safety evaluations focusing on core disruptive accidents (CDAs). Based on the lessons learned from the Fukushima nuclear accident, new safety standards for nuclear power plants (NPPs) (i.e., the New Regulatory Requirements) have been formulated in Japan to control the prevention and mitigation of severe accidents. Internationally, it has been suggested that safety designs should consider measures to prevent and mitigate the occurrence of severe accidents in design extension conditions (DEC), based on the defense-in-depth concept [2]. Therefore, it is essential to improve the safety of fast reactors against CDAs to commercialize fast reactors. If the frequency of core disruptive events can be significantly reduced and practically eliminated from safety design considerations, a fast reactor system with excellent social acceptability can be realized.

In addition to active reactor shutdown systems, passive safety systems working against anticipated transient without scram (ATWS), such as the self-actuated shutdown system (SASS) [3] and the gas expansion module (GEM) [4], have been developed, and investigated as countermeasures to prevent CDAs in fast reactors [5,6]. To prevent large releases of mechanical energy (i.e., energetics) during CDAs, various mitigation strategies, including suppressing positive reactivity via coolant voiding [7], and controlling core material relocation (i.e., controlled material relocation (CMR) [8]), have been proposed to avoid recriticality during CDAs [9]. Applying these measures to establish in-vessel retention (IVR) to stabilize and isolate the damaged core materials in the reactor vessel is an effective and reasonable approach to ensure the containment of radioactive materials.

In this study, a new subassembly-type passive reactor shutdown device that enhances the versatility and robustness of CDA prevention measures is developed, and a four-year plan (initiated in FY2019) to drastically improve the safety of sodium-cooled fast reactors (SFRs) is outlined in detail; Kyushu University, University of Fukui, Tokyo City University, Japan Atomic Energy Agency (JAEA), and Tokyo Institute of Technology are participating in this project. As a part of this project, the engineering feasibility of the device will be comprehensively analyzed in terms of the fuel materials, the pin structure that promotes fuel relocation during accidents, and the nuclear and thermal-hydraulic characteristics during device operation. This report outlines the project aims and stages presents preliminary results regarding the characteristics of the core in which the devices are implemented, and discusses the device response.

The proposed device (Fig. 1) is a subassembly of the pin bundle, which contains a fuel material that is solid under the core temperature conditions during normal reactor operation but liquefies when the core temperature rises during an ATWS accident. During such an accident, a large negative reactivity can be applied to the core by moving (relocating) the liquefied fuel to a region of low reactivity in the device pins by simple physics alone. By replacing some of the normal fuel subassemblies with these devices, the reactor can passively be brought to a subcritical state, which would prevent core damage and terminate the event before the normal driver fuel is damaged by the ATWS. Therefore, the proposed device represents a safety equipment component with advanced features that are innovative and novel in the following aspects.

• In the fourth layer of the defense-in-depth concept, the proposed device functions as a passive reactor shutdown system that is effective for both unprotected loss of flow (ULOF) and unprotected transient overpower (UTOP) accidents, which are typical events in fast reactor ATWS.

• When the core temperature rises during an accident, the proposed device's function is triggered immediately due to natural phenomena (e.g., liquefaction of low-melting fuel, gravity-induced drop of liquefied fuel), and it demonstrates high operational reliability in the case of a reactor shutdown.

• The proposed device can easily be introduced by replacing some of the existing fuel subassemblies in the reactor core, and it is highly compatible with the existing core design because it has reactor physics performance equivalent to that of other driver fuels during normal operation.

• Similar to an ordinary driver fuel, the proposed device can be replaced periodically, so there is no need for in-service inspections; moreover, the replaced device fuel can be recovered and recycled.

• After device operation, the fuel that moved within the pins remains in the pins, which facilitates the process of restarting the reactor by replacing the device subassemblies.

Table 1 compares the conventional passive shutdown system for fast reactors [6] with the subassembly-type device proposed in this study, which can be equipped in conjunction with the existing measures to prevent severe accidents. Therefore, this device can be used as an independent line of protection that exhibits diversity and robustness and can help prevent CDAs. In other words, this device is expected to dramatically improve reactor safety so that CDAs can confidently be regarded as extremely unlikely events.

To demonstrate the operating principle of the passive subassembly-type shutdown device, this study considers a device structure comprising hollow fuel pellets that melt at a designated temperature, as shown in Fig. 2. During normal operation, the device fuel is supported on top of the lower plenum region, and when the fuel liquefies following an increase in the core temperature during an accident, gravity pulls the liquid fuel to the lower plenum region. The fuel used in the device is a metallic alloy (e.g., Pu-U-Fe) or chloride (e.g., NaCl-PuCl₃-UCl₃) that has the characteristics of a fast reactor fuel and a relatively low melting point. By controlling the composition of the fuel, it is possible to obtain fuels that change from a fully solid state at the SFR inlet temperature during normal operation to a liquid or solid–liquid mixed phase during device operation.

The present work aims to (i) design the basic structure of the device and (ii) select a fuel material that has suitable properties under accident conditions. In addition, this project involves evaluating a series of operating characteristics (e.g., fuel melting and movement within the pin) under rising temperature conditions. The flow characteristics of the liquefied fuel in the device pin must be clarified to confirm that the device will operate under the expected accident conditions, i.e., that the device fuel will be relocated to the plenum region reliably. To maximize the operational reliability of the device, a method for ensuring the relocation of the liquefied fuel in the pin should also be investigated. Finally, basic experiments will be conducted using simulated materials to verify the operating principle of the device and to formulate a plan for future demonstrations. The reference device specifications were defined as follows:

• The melting point is set at 700 °C. Hollow fuel is used to facilitate the movement of the liquid fuel during device operation.

• The fuel shall be in contact with and adhere to the cladding; sodium bonding is not used. However, the effect of irradiation-induced swelling on the fuel properties will be considered. During normal operation, the cladding temperature of the device pins is below 560 °C because device fuel melting is prevented. Therefore, the corrosion effect due to the interaction between the fuel and cladding is considered to be small.

• Preheating pins containing metallic fuel in the upstream region of the core is introduced into the device subassembly. Heating the coolant flowing around the device fuel pin during normal operation promotes fuel liquefaction and fluidization during device operation and increases the temperature at the lower part of the device fuel section, preventing the falling molten fuel from solidifying.

• The device fuel support structure is made from a low-melting-point material to ensure the relocation of fluidized fuel to the lower plenum once the temperature rises during the accident. Alternatively, a fusible plug separating the lower plenum from the device fuel region can be introduced, in which case, the melting of the fusible plug facilitates fuel relocation from the pre-pressurized device fuel region.

• During normal operation, the core inlet temperature is set to about 400 °C, and the fuel temperature at the lower end of the device is set to 450 °C, which reflects the effect of the preheating pins.

• The device operating temperature during ATWS should be below the boiling point of sodium (i.e., 800–900 °C).

A 750-MW_el-class SFR core [10] was selected as the reference in this study, and mixed oxide (MOX) fuel core specifications were used to determine the core characteristics (e.g., power, reactivity coefficients) and to study the core performances under anticipated transient events (e.g., ULOF and UTOP events). In this scenario, the medium-sized MOX fuel (with high internal conversion) core specification values from the past design study [10,11]. This conceptual design embodies ambitious performance goals using a compact plant approach. Its main features are: (i) coolant inlet temperature = 395 °C, average outlet temperature = 550 °C; (ii) high linear heat rating for oxide fuel pellets aiming at a compact core size (core height = 100 cm); (iii) large bundle pitch (20.6 cm) with the higher fuel volume fraction; (iv) internal duct with rhombus space integrated in fuel subassemblies to remove fuel after (hypothetical) severe damage to fuel pins; (v) high burnup core with average burnup (∼150 MWd/kg); (vi) long-term operation cycle length = 832 days/cycle, and four-batch refueling in a uniformly dispersed manner. The reference core layout used in this study is presented on the left side of Fig. 3.

In this study, the fresh fuel compositions (for equilibrium cycles) and regional average compositions (atomic number densities) of the end of equilibrium cycle (EOEC) state were determined according to the core specifications. Nuclear characteristics were calculated using the conventional homogeneous models, which have previously been applied for fast reactor core analyses, using a 70-group cross section set for fast reactors (JFS-3-J4U1 [12,13]) combined with an effective cross section processing code (SLAROM [14]). Power distributions in the core (coolant orifice zone) were calculated using a three-dimensional model, and the spatial distributions of the reactivity coefficients (e.g., fuel Doppler, fuel mass, coolant density, and structural material density coefficients) were computed for the two-dimensional RZ-model comprising the lateral flow region and axial node using a diffusion theory code DIF3D [15]. Table 2 shows the core region-integrated reactivity coefficients, which are consistent with the typical values expected for large oxide-fueled cores. The high burnup design increased the sodium void reactivity to a larger positive value.

The diagram on the right side of Fig. 3 shows the tentative layout for the 16 devices that were used to estimate the effects of fuel relocation (from the core region to the bottom part of the device pin) in the device subassembly. Figure 4 shows the typical reactivity profile of the Pu-U-10at%Fe alloy fuel material in the axial direction within the core region. The axial distribution of reactivity values at each axial node position is the average distribution of the 16 device subassembly locations loaded, relative to the axial average of the reactivity values of the device fuel material in the core region. In this figure, the fuel material exhibits the maximum positive reactivity near the core's midplane; this peak in the reactivity profile reaches approximately 1.4, which corresponds to the square of the axial power peaking factor (1.2 for this core). Based on this trend in the axial reactivity, the device fuel materials were placed below the core's midplane in this study.

The solid fuel of the device is assumed to be a hollow metal (alloy-based) fuel slug in which a Pu-U-10at%Fe alloy adheres in an annular shape, with its inner surface along the cladding tube. The fuel relocation reactivity potentials were estimated by assuming the initial and final states and using static neutronic eigenvalue calculations for the core states based on the ratio of Pu to depleted uranium (DU) in the alloy. Table 3 shows the calculated reactivity of the device (relationship between the amount of relocated material in the device and the reactivity of the fuel materials) and the amount of heat generated in the core region of the device subassembly at different Pu:U ratios.

The changes in reactivity due to fuel material relocation were in the range of −3.5 $ to −6 $. Assuming that 50% of the fuel material was relocated out of core region in a short time, a reactivity insertion in the range of −1.7 $ to −3 $ can be expected. The relationship between the amount of relocated Pu and the change in reactivity is almost proportional.

To obtain a rough understanding of the transient dynamics of the core during device operation, the characteristics of the device-loaded core were evaluated when the fusible plug melted and the fluidized fuel relocated, assuming the candidate device specifications. Based on the power and reactivity characteristics of the core (discussed in Section 4.1), the flow network model transient analysis code-named ARGO [16] was used as a dynamic evaluation model to calculate the response characteristics of the device-loaded core during device operation in the initiating phases of ULOF and UTOP events. As the temperature of the device fuel rises and the solid phase fraction becomes smaller as liquefaction progresses, the molten fuel fluidizes. In this evaluation, it is assumed that the device fuel begins to flow with a large viscosity when it reaches its melting point.

The specifications of the target device-loaded core (750-MW_el-class homogeneous core) and device subassemblies are presented in Table 4. As shown in Fig. 3, 16 device subassemblies are distributed throughout the reactor core. The device fuel material is a Pu-U-Fe alloy with density = 14840 kg/m³, specific heat = 200 J/kg·K, latent heat of fusion = 61 kJ/kg, viscosity = 6.5 $×$ 10⁻³Pa·s, and thermal conductivity = 20 W/m·K. The axial length of the fuel region in the device fuel pin was 50 cm, as shown in Fig. 5. Two types of pins were used in the device fuel subassembly; specifically, 169 device fuel pins (diameter = 8.5 mm) and 127 preheating pins (diameter = 4.8 mm) per device subassembly. As discussed in Section 4.1, the negative reactivity required for reactor shutdown triggered by the device operation is more than 3 $ when all device fuels are relocated from the core. The thermal power of the device fuel per device subassembly was set to 2.9 MW (equivalent to 330 W/cm in linear power), and the thermal power of the preheating pins was set to half of that. As a result, the thermal power of the 16 device subassemblies (out of 286 total fuel subassemblies including the device subassemblies) accounts for about 4% of the reactor's total thermal power. The ratio of thermal power to coolant flowrate in the device subassemblies was assumed to be the same as that in the core fuel subassemblies.

The distributions of the fuel Doppler coefficients, fuel density coefficients, coolant density coefficients, and structural material density coefficients of the conventional homogeneous core evaluated in the previous section were used as the reactivity coefficients of the device-loaded core. The effective delayed neutron fraction was set as 3.7 × 10⁻³, and the sodium void reactivity was 7.3 $ in the core. In the case of ULOF, the primary pump coast-down flow halving time was set to 5 s. In addition, instantaneous adiabatic insulation was assumed after the initiation of the loss of flow in the steam generator, which is the heat removal source. In the case of UTOP, one control rod with a reactivity value of 90 ¢ was set to be withdrawn in 30 s, which is the maximum control rod withdrawal rate set as a typical TOP condition in the licensing safety analysis for the design basis event of the Japanese prototype reactor Monju. This relatively long reactivity insertion time was assumed in this work to conservatively model the fuel's rising temperature. It was also assumed that the rated power would be removed from the core after the start of UTOP.

Figure 6 shows the evolution of power, coolant flowrate, and P/F over time from the start of ULOF with no device operation. For about 14 s after the start of ULOF, the power is generally constant (remaining at the rated power), while the P/F increases linearly and the flowrate decreases with time. Since the void reactivity in this core is as large as about 7 $, which compensates for negative reactivity effects associated with the coolant temperature rise, there is almost no reduction in power. After 14 s, the coolant in some fuel subassemblies begins to boil, thus causing a power fluctuation from the insertion of positive reactivity; meanwhile, the P/F increases rapidly as the coolant continues boiling. Figure 7 shows the variation in the fuel center temperature, fuel surface temperature, and coolant temperature at the top (center of the core in the axial direction) and bottom of the device fuel over time after the start of ULOF. However, it is assumed here that the thermal conductivity of the device fuel is reduced to 1/3 because of swelling, and the latent heat of fusion is not taken into account. After about 11 s, the coolant temperature at the lower end of the device fuel and the fuel surface temperature rise to approximately 600 °C and 700 °C, respectively. Therefore, if the melting point of the device fuel is set to 700 °C, the fuel temperature will be below the melting point in all regions during rated operation, and the fuel temperature will exceed the melting point in all regions of the device fuel within 11 s after the start of ULOF.

On the basis of these results, it was assumed that the fusible plug melts 11 s after the start of ULOF, and the device operates when the liquefied fuel starts moving into the lower plenum, which inserts negative reactivity. During device operation, the fuel temperature at the lower end of the device fuel is considered to be above the melting point of 700 °C, and therefore, the central hole of the hollow pellet is not blocked by the falling molten fuel. The time required for the device fuel to move into the lower plenum is assumed to be 10.5 s, and a total of 3 $ of negative reactivity is assumed to be inserted over this time, as all 16 device subassemblies begin operating simultaneously. The device operation time was calculated by assuming that the upper 80% of the molten fuel flows down along the inner wall of the hollow fuel pellet due to gravity alone. Although it is estimated that the molten fuel actually moves to the lower plenum in a few seconds, the device operating time was conservatively estimated assuming that the enthalpy of the molten fuel decreases during the process of moving to the lower-temperature region and that its viscosity exceeds that of the liquid phase by more than 2000 times.

Figure 8 shows the variation in power, coolant flowrate, and P/F over time when the device is operated under the conditions described above. The power decreases when the device operation initiates negative reactivity insertion, and reactor shutdown is almost achieved in about 20 s. Figure 9 shows the variation in the net reactivity, including the breakdown into fuel Doppler, coolant density, cladding expansion, wrapper tube expansion, fuel expansion, and external reactivity (each of these values is relative to the onset of ULOF). In this case, the external reactivity corresponds to the negative reactivity insertion from the device operation. The net reactivity at the start of the device operation is about 0.02 $, but the device operation causes the net reactivity to become negative and the reactor power to decrease over about 12 s. In addition, no coolant boiling occurs during the 30 s transient event.

Figure 10 shows the variation in the P/F over time during UTOP without device operation. The reactor power increases about twofold after 30 s, and then the power decreases as the coolant inlet temperature increases because only the amount of heat equivalent to the rated power is removed. Figure 11 shows the variations in the fuel center temperature, fuel surface temperature, and coolant temperature at the top and bottom of the device fuel pins over time starting at the onset of UTOP. The fuel center temperature at the top of the device reaches about 700 °C about 30 s after the start of UTOP, and the coolant does not boil until about 40 s into the calculated transient event.

In the case of UTOP, the central temperature of the lower end of the device fuel almost exceeds the melting point of 700 °C after about 30 s, so the device is assumed to be operational 30 s after the start of UTOP. It is also assumed that a negative reactivity of 3 $ is inserted during the first 10.5 s in the case of UTOP. Figure 12 shows the variation in the P/F over time with device operation. The reactor power decreases during device operation; specifically, the relative power decreases from 2 times to 0.3 times in the 10 s following the onset of the device operation. Figure 13 shows the variation in the net reactivity over time, and the breakdown into the various contributing elements. Here, the external reactivity corresponds to the positive reactivity insertion from transient overpower (TOP) up to 30 s, and the negative reactivity insertion from the device operation after 30 s. After the device operation begins, the net reactivity is negative for a short time. As the temperature rises after the onset of UTOP, the fuel Doppler reactivity and the reactivities due to the expansion of the coolant, structural material, and fuel also become negative. However, as the core temperature decreases because of the device operation, these reactivities return to zero in about 40 s.

Overall, in the ULOF and UTOP initiating phases, i.e., when the device is activated and the 3 $ of negative reactivity is inserted over 10.5 s, stable core cooling is maintained before the coolant boils in the driver fuel subassembly, and both events are expected to be terminated. Figure 14 presents a map of the successful reactor shutdown processes induced by the proposed device operation; these results were obtained from survey evaluations using the device reactivity and reactivity insertion time as parameters. Even assuming that 1/3 to 1/2 of the devices loaded in the core do not operate, it is clear that there is enough margin to achieve reactor shutdown if the device operation time can be shortened to some extent.

This report describes a project aiming to develop a new subassembly-type passive reactor shutdown device to enhance the versatility and robustness of CDA prevention measures in SFRs. To obtain a rough understanding of the core transient behavior during the proposed device's operation, a preliminary evaluation of the core characteristics and device response was performed for the candidate device specifications. The results outlined conditions for the device operation during the initiating phases of ULOF and UTOP events to (i) promote stable cooling of the core prior to coolant boiling in the driver fuel subassemblies and (ii) terminate both events.

Special thanks to Dr. Yasushi Tsuboi of Toshiba Energy Systems & Solution Corp. We are deeply grateful to Dr. Hiroshi Endo for collaboration during the early stages of this work. We also thank Suzanne Adam, Ph.D., from Edanz for editing a draft of this paper.

• Ministry of Education, Culture, Sports, Science and Technology (MEXT) (Funder ID: 10.13039/501100001700).

The datasets generated and supporting the findings of this article are obtainable from the corresponding author upon reasonable request.

P/F =

power to coolant flow ratio relative to the start of a transient

Acronyms and Abbreviations

ASTRID =

Advanced Sodium Technological Reactor for Industrial Demonstration

ATWS =

anticipated transient without scram

ARGO =

plant dynamics analysis code

BN-800 =

800-MW_el sodium-cooled fast reactor at the Beloyarsk nuclear power station in Russia

CDA =

core disruptive accident

CMR =

controlled material relocation

DEC =

design extension condition

DIF3D =

code to solve one-, two-, and three-dimensional finite difference diffusion theory problems

DU =

depleted uranium

EFPD =

effective full power days

EOEC =

end of equilibrium cycle

GEM =

gas expansion module

IVR =

in-vessel retention

JAEA =

Japan Atomic Energy Agency

JENDL =

Japanese Evaluated Nuclear Data Library

JSFR =

Japan Sodium-cooled Fast Reactor

LIM =

lithium injection module

LOF =

loss of flow

MEXT =

Ministry of Education, Culture, Sports, Science, and Technology

MOEC =

middle of equilibrium cycle

MOX =

mixed oxide

NPP =

nuclear power plant

PGSFR =

Prototype Generation-IV Sodium-cooled Fast Reactor

PRISM =

Power Reactor Innovative Small Module

RAPID =

Refueling by All Pins Integrated Design

SASS =

self-actuated shutdown system

SFR =

sodium-cooled fast reactor

SLAROM =

code for cell homogenization calculations of fast reactors

TOP =

transient overpower

ULOF =

unprotected loss of flow

UTOP =

unprotected transient overpower