Thermal hydraulic parametric investigation of decay heat removal from degraded core of a sodium cooled fast Breeder reactor

Nuclear Engineering and Design 313 (2017) 285–295

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Thermal hydraulic parametric investigation of decay heat removal from degraded core of a sodium cooled fast Breeder reactor Lokesh Verma a, Anil Kumar Sharma b,⇑, K. Velusamy b a b

Department of Physics and Astrophysics, University of Delhi, Delhi 110007, India Reactor Design Group, Indira Gandhi Centre for Atomic Research, HBNI, Kalpakkam, India

h i g h l i g h t s
Decay heat removal from degraded core of a typical SFR is highlighted.
Influence of number of DHXs in operation on PAHR is analyzed.
Investigations on structural integrity of the inner vessel and core catcher.
Feasibility study for retention of a part of debris in upper pool of SFR.

a r t i c l e

i n f o

Article history: Received 21 July 2016 Received in revised form 8 December 2016 Accepted 12 December 2016

Keywords: Core disruptive accident Post accident decay heat removal Natural convection Fast reactor Core catcher

a b s t r a c t Ensuring post accident decay heat removal with high degree of reliability following a Core Disruptive Accident (CDA) is very important in the design of sodium cooled fast reactors (SFR). In the recent past, a lot of research has been done towards the design of an in-vessel core catcher below the grid plate to prevent the core debris reaching the main vessel in a pool type SFR. However, during an energetic CDA, the entire core debris is unlikely to reach the core catcher. A significant part of the debris is likely to settle in core periphery between radial shielding subassemblies and the inner vessel. Failure of inner vessel due to the decay heat can lead to core debris reaching the main vessel and threatening its integrity. On the other hand, retention of a part of debris in core periphery can reduce the load on main core catcher. Towards achieving an optimum design of SFR and safety evaluation, it is essential to quantify the amount of heat generating core debris that can be retained safely within the primary vessel. This has been performed by a mathematical simulation comprising solution of 2-D transient form of the governing equations of turbulent sodium flow and heat transfer with Boussinesq approximations. The conjugate conduction-convection model adopted for this purpose is validated against in-house experimental data. Transient evolutions of natural convection in the pools and structural temperatures in critical components have been predicted. It is found that 50% of the core debris can be safely accommodated in the gap between radial shielding subassemblies and inner vessel without exceeding structural temperature limit. It is also established that a single plate core catcher can safely accommodate decay heat arising due to 70% of the core debris by establishing natural circulation in the lower sodium pool. The influence of heat removal rate by natural circulation on availability of the number of decay heat exchangers (DHX) dipped in the upper pool is also analyzed. It is seen that the temperatures in the inner vessel, source plate and the maximum debris temperature do not increase significantly even when the DHXs are deployed 5 h after the accident, demonstrating the benefit of large thermal inertia of the pool. Ó 2016 Elsevier B.V. All rights reserved.

1. Introduction A severe accident scenario like the Core Disruptive Accident (CDA) leading to whole core meltdown is a very rare incident

⇑ Corresponding author. E-mail address: [email protected] (A. Kumar Sharma). http://dx.doi.org/10.1016/j.nucengdes.2016.12.017 0029-5493/Ó 2016 Elsevier B.V. All rights reserved.

(frequency of occurrence <106/ry). A CDA can be initiated by two events in a pool type SFR (Chellapandi et al., 2013). The first event is ULOFA (Unprotected Loss of Flow Accident), which can hamper the structural integrity of the main vessel. The second event is UTOPA (Unprotected Transient Overpower Accident), which occurs due to uncontrolled withdrawal of control rods causing a surge in the reactivity. The comparison of the behavior of two core designs for the ASTRID reactor in case of severe accidents is

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done by Bertrand et al. (2016). First core design is a classical homogeneous core with large pin diameter. The second concept is an axially heterogeneous core. The comparison of the cores relies on two typical accident families: UTOP and ULOF. The results presented in this paper have contributed to retain the axially heterogeneous core for the studies currently carried-out for the ASTRID design. The core debris formed during a severe accident may cause the melting of other structures like grid plate during its passage. This debris would ultimately reach the main vessel and settle over it. It continues to generate heat due to the decay of fission products and may expose the main vessel to the ‘decay heat’ and hamper its structural integrity. A degraded core and the progression of a severe accident are depicted in Fig. 1. Passive safety with minimum human interference is one of the main features of Generation IV nuclear reactor designs. It is known that post shutdown heat removal is an important safety function which has to be accomplished with a good degree of certainty under accident scenarios. Raj et al. (2015) have detailed the necessary safety criteria for decay heat removal from a degraded core after a severe accident in sodium cooled fast reactors (SFR). Most of the fast reactors depend on passive systems to remove the decay heat after shutdown. This is done by the establishment of favorable natural convections possible in the sodium pools as well as in decay heat removal circuits. The natural convection removes the heat from the reactor pool and rejects it to the ambient air that acts as the final heat sink. It is important to carefully evaluate the development of natural circulation and the associated temperatures in various critical components during the accident phase. In the recent past, a lot of research work has been carried out in this area to establish effective natural convection path for safe removal of the decay heat from degraded core debris in SFR. A detailed computational fluid dynamic simulation of the passive decay heat removal in a SFR was carried out by Hung et al. (2011). They concluded that the passive reactor air cooling system can be an efficient system in removing decay heat through heat conduction, convection, and thermal radiation without the requirement of the operation of any active heat removal mechanisms. They also found that, by adding an extended part of the reactor liner and/or core periphery, a recirculation would occur above the partition in the upper sodium pool and it was preventing the downward flow from partition to

Fig. 1. Progression of a severe accident (Gnanadhas et al., 2011).

the core area and enhancing the convective heat transfer between the pools. A new passive residual heat removal system (PRHRS) is designed by Lv et al. (2016). In this system a cooling tank which works as an intermediate buffer device is used to transfer the core decay heat to the sodium to air heat exchanger (AHX), and then the heat is transferred to the atmosphere finally. Their study reveals that the Passive Residual Heat Removal System (PRHRS) could remove the decay heat from the primary loop effectively, and natural circulations can be established in the primary circuit and the PRHRS circuit. Hughes and Blandford (2016) carried out the experimental investigation of a directionally enhanced decay heat exchanger concept for high temperature direct reactor auxiliary cooling systems. They have optimized such a heat exchanger so that shell-side heat transfer is enhanced in one primary coolant flow direction and degraded in the opposite coolant flow direction, utilizing the fluidic diode concepts. To maintain the structural integrity of the main vessel during a CDA, the core debris needs to be prevented from directly falling on the main vessel. Various design concepts have been proposed and incorporated to take care of these issues. A core catcher plate has been placed below the grid plate to accommodate the core debris after breach in the grid plate. Two major types of core catcher concepts are prevalent in the nuclear reactors across the world. First is the in-vessel core catcher in which the corium is retained inside the main vessel and the other is the ex-vessel core catcher concept in which the corium is cooled outside the main vessel. For the pool type sodium cooled fast reactors, the integrity of the main vessel is a must and therefore, as stated by Jasmin Sudha et al. (2014), the in-vessel core catcher concept is preferable. In fact, the preference for an in-vessel core catcher rather than inevitable is more dictated by technological deficiencies in implementing ex-vessel core catcher designs and thus, ex-vessel retention strategy can be considered as a future design option, if such design maturity is demonstrated. External reactor vessel cooling coupled with in-vessel retention as a strategy to cool the molten core from outside the reactor vessel for Loviisa nuclear power plant in Finland was reported by Tuomisto and Theofanous (1994). An extensive study has been carried out on the different designs of core catchers in different nuclear power plants in the world. A brief review on different core catcher concepts used in SFRs across the world can be found in Jasmin Sudha et al. (2014). For the Indian Prototype Fast Breeder Reactor (PFBR), which uses mixed oxide fuel i.e. uranium plutonium mixed oxide (UO2-PuO2), an in-vessel core catcher was adopted (Chellapandi et al., 2013). Sharma et al. (2009) carried out a conjugate heat transfer study for the turbulent natural convection of sodium in a cylindrical enclosure with multiple internals heat sources to numerically evaluate feasibility of multi-tray core catcher concept. A multi-layer core catcher concept was proposed by Jasmin Sudha et al. (2014). It is clear from the above literature that active research is being done to prevent the corium from falling on the main vessel. In doing so, main attention is given to the design concept for the core catcher. However, during a CDA, some fraction of the core debris may disperse and settle on the various components in the upper pool of the reactor as discussed above. Therefore, the present work considers the situation, where it is assumed that some fraction of the core debris are relocated to a gap which is present between the core periphery and the inner vessel in the upper plenum and remaining fraction on core catcher in the lower plenum. The coupled heat transfer analysis is carried out for a case when the total heat load is distributed between the core catcher plate and the gap in consideration. The estimation of the amount of debris that can be accommodated safely in this gap and on the core catcher is also carried out. Moreover, the influence of DHXs on heat removal from the core debris is carried out. The transient forms of 2-D governing equations of turbulent natural convection are solved in entire

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qk2 e

lt cp : rT

domain including upper plenum and lower plenum of the reactor system.

leff ¼ l þ lt ; lt ¼ C l

2. Mathematical model

The following constants are used in k-e turbulence model (Henkes et al., 1991):

The mathematical description of the natural turbulent flow is given by the Reynolds Averaged Navier Stokes (RANS) equations, which also take into account the time averaged energy equation for describing the flow driven by buoyancy force. While defining the governing equations, the model assumed is two dimensional and axi-symmetric. The flow is considered to be viscous, Newtonian and incompressible and is assumed to follow Boussinesq’s approximation, i.e. density difference is solely important in producing the buoyancy. All the physical properties are assumed to be invariant. 2.1. Governing equations With the above assumptions, the basic governing equations that describe the unsteady turbulent fluid flow include the transient form of RANS equations for mass, momentum and energy conservation and the equations for turbulent kinetic energy (k) and its dissipation rate (e) are as follows: Continuity

1 @ðruÞ @ðv Þ þ ¼0 r @r @z

C l ¼ 0:09;

C e1 ¼ 1:44;

rK ¼ 1:0 and re ¼ 1:3

and keff ¼ K þ

C e2 ¼ 1:92;

rT ¼ 0:9;

The standard k-e model is adopted after a detailed sensitivity study employing other turbulent models which is described in Section 4.2. 2.2. Physical model, initial and boundary conditions The physical model considered in the study is depicted in Fig. 2 (a) and the enlarged view of the gap between the core periphery and inner vessel with the core catcher is shown in Fig. 2(b). The height and diameter of the main vessel are taken as 12.3 m and 12.6 m respectively, which are equivalent to the vessel dimensions of the 500 MWe Indian PFBR. The diameter of the core catcher

ð1Þ

Radial Momentum



 @ðquÞ @u @u @p 1 @ @u ¼ þ þq u þv r leff @t @r @z @r r @r @r
 @ @u u  leff 2 þ leff @z @z r

ð2Þ

Axial Momentum



 @ðqv Þ @v @v @p 1 @ @v þq u r leff þv ¼ þ @t @z r @r @r @z @r
 @ @v leff þ þ qg z bðT  T ref Þ @z @z

ð3Þ

Energy




 @ðqC p TÞ @T @T 1 @ @T @ @T þ qC p u þv ¼ rkeff þ keff þ q000 @t @r @z r @r @r @z @z ð4Þ Turbulent kinetic energy

  
@ðqkÞ @ @ 1 @ l @k þ ðqukÞ þ ðqv kÞ ¼ r lþ t @t @r @z r @r rk @r   
@ lt @k þ Pk þ Gk  qe lþ þ @z rk @z ð5Þ Dissipation rate of turbulent kinetic energy

  
@ðqeÞ @ @ 1 @ l @e þ ðqueÞ þ ðqv eÞ ¼ r lþ t @t @r @z r @r re @r   
@ lt @ e lþ þ @z re @z þ ½C e1 ðPk þ C e3 Gk Þ  C e2 e "
 #
2
2 2 u2 @u @v @u @ v ; P k ¼ lt 2 þ2 þ þ2 þ @z @r @z @r r

e k

Gk ¼ 

ð6Þ

lt @T gb ; rT @z

Fig. 2. (a) Physical model; (b) Enlarged view of the source location.

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Fig. 5. Variation of heat flux on the heat shield plate for 1800 s for different turbulence models.

Fig. 3. Computational grid used in the study.

Fig. 6. Variation of the temperature at the inner vessel, HSP and the maximum temperature at the debris bed for 10% at gap.

The roof slab and the outer surface of the main vessel are considered to be adiabatic. The initial temperature of sodium is taken to be at 673 K. No slip boundary condition is used for the solidfluid interface. The debris bed is modeled as a solid bed on the core catcher plate and in the gap between the core periphery and the inner vessel. It is assumed that the debris bed is a homogeneous mixture of fuel and steel along with sodium. The effective thermal conductivity of the debris bed is calculated using the correlation suggested by Kampf and Karsten (1970) given by:

  3 k ð1  /Þ 1  kNap  5 kb ¼ kNa 41  1 kp k 3 1  p þ ð1  /Þ kNa kNa 2

Fig. 4. Comparison of the obtained temperature along the central axis along with experimental data.

plate is considered as 6.5 m and its plate thickness is 20 mm. A single plate core catcher is adopted for the study. The gap between the core periphery and the inner vessel is 100 mm. The opening in the grid plate due to meltdown is assumed to be 2.2 m. The total volume of the core debris is 5.12 m3 (equivalent to typical 181 fuel sub-assemblies of 500 MWe SFR), which is distributed at different regions under consideration in the reactor.

ð7Þ

where U is the porosity of the bed (assumed to be 0.4), kNa is the thermal conductivity of sodium, kp is the effective thermal conducP ai and ai is tivity of the solid particles (fuel and steel) given by k1p ¼ ki mass fraction of component with conductivity ki. The specific heat P of the debris bed is determined by C pb ¼ ai C pi , where Cpi is the specific heat of the component i. 2.3. Description of the heat source The volumetric heat generation rate in the corium decreases exponentially with time. According to Sridharan (1989), the decay power is about 21% of the operating power 1 s after shutdown and

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eration by the debris bed is taken as 1000 s. Taking this into account, the decay power is modeled using the data available from literature. The equation used for decay power is

q000 ¼ 6e7  t 0:33 W=m3

ð8Þ

where q000 is the volumetric decay power and ‘t’ is the time (s) elapsed post accident. This volumetric heat generation rate as source is distributed at different locations considered in the study. 2.4. Modeling of the heat sink The decay heat is removed by a passive Safety and Decay Heat Removal System (SGDHRS) comprising 4 DHX with maximum heat removal capacity of 8 MW each at a pool temperature of 820 K. In the present study, DHX is modeled as a vertical annulus with a heat removal rate proportional to the upper pool temperature. The equation used for the volumetric decay heat removal is as follows. Fig. 7. Variation of the temperature at the inner vessel, HSP and the maximum temperature at the debris bed for 20% at gap.

q0 ¼ 22337  T  9e6 W=m3

ð9Þ

0

where q is the volumetric decay heat removal capacity and ‘T’ is the average upper pool temperature in K. 3. Numerical methodology A commercial CFD code is used to carry out the transient simulations. A through grid sensitivity study has been done. The computational grid used for calculation in the present study is shown in Fig. 3. Second order implicit scheme is used for time discretization. Body force weighted scheme is used for the discretization of the momentum equation whereas first order upwind scheme is used for the discretization of the convection terms in the conservation equations. Pressure-velocity coupling is resolved by using the PISO algorithm. The solutions are assumed to be fully converged when the residuals are of the order of 104 for all the governing equations except for energy equation, for which it is set at 108. 4. Validation Fig. 8. Variation of the temperature at the inner vessel, HSP and the maximum temperature at the debris bed for 30% at gap.

it decreases to 1.5% in 1 h. The relocation time of molten corium to reach the lower plenum is around 1000 s as estimated by Sudha and Velusamy (2014). Thus, the initial time for the decay heat gen-

4.1. Validation with experimental data To validate the numerical methodology, simulation of natural circulation in PATH experimental facility is carried out and the predicted results are compared with experimental data. PATH is a 1:4 water model of the PFBR and is dedicated for the post accident thermal-hydraulics experiments and their numerical validation (Gnanadhas et al., 2011). A heat load of 18 kW is provided at the source plate and the evolution of axial temperature distribution in the water is measured. The comparison of temperature profile along the central axis after 2 h obtained numerically and that with experiment is shown in Fig. 4. It can be seen that there is a good agreement between the experimental and numerical results obtained by computational model considered in the present study. 4.2. Sensitivity study for various turbulent models

Fig. 9. Variation of the temperature at the inner vessel, HSP and the maximum temperature at the debris bed for 50% at gap.

A sensitivity analysis is carried out to assess the applicability of practical, turbulence model for such kind of flow simulations. The target variable for the sensitivity study is the heat flux. The heat flux on the heat shield plate was monitored for the first 1800 s using five different turbulence models. The five different models considered are k-x standard, k-x SST, k-e standard, k-e RNG and k-e realizable. It was noticed that the two k-x models showed marked oscillations in the value of heat flux at the heat shield plate surface. Among the k-e models, the k-e realizable model showed a very poor and slow convergence. The k-e RNG and standard models

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Fig. 10. Stream functions at 1 h for 10% core debris in gap and 90% in core catcher plate.

Fig. 11. Isotherms at 1 h for 10% core debris in gap and 90% in core catcher plate.

showed smooth transitions in the value of heat flux but the standard k-e model converged much faster with a time step size of 0.5 s. On the other hand the RNG model required a much smaller time step of 0.2 s to converge. A plot of heat flux at the heat shield plate with time with the different turbulence models is shown in Fig. 5. Thus, based on the sensitivity studies employing various models, standard k-e turbulence model is found to be more stable and suitable model to carry out such simulations.

between the core periphery and the inner vessel) and on core catcher placed in the lower plenum. Since whole core meltdown scenario is considered, the entire core debris is relocated either in the gap mentioned or the core catcher plate placed below the grid plate and the final sink is DHX dipped in the upper pool.

5. Results and discussion

To analyze the effect of core debris settlement in upper and lower pools, different percentages of core debris are assumed to be settled in the gap between core periphery and inner vessel and the core catcher. The transient evolution of the temperature

The primary aim of this study is to analyze the effect of the settlement of the dispersed core debris in the upper pool (gap

5.1. Analysis to estimate safe fraction of core debris on structural integrity of components

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Fig. 12. Stream functions at 1 h for 50% core debris in gap and 50% in core catcher plate.

Fig. 13. Isotherms at 1 h for 50% core debris in gap and 50% in core catcher plate.

at the inner vessel, Heat shield Plate (HSP) and the maximum temperature at the debris bed in the lower plenum for four different debris distributions varied from 10%, 20%, 30% and 50% in the gap is shown in Figs. 6–9, respectively. It may be clarified that when 10% of the total core debris is settled in the gap and the remaining 90% of the debris is relocated to the HSP. It is observed that the average temperature of the inner vessel close to the solid debris bed reaches a maximum of 928 K in 25 min, which is well within the safe temperature limit of 1173 K (Fig. 6). However, the temperature on the shield plate in the lower plenum is found to reach a maximum value of 1219 K in 35 min, which remains

above the safe temperature limit of 1173 K for a duration of 50 min. Moreover, the maximum temperature in the debris bed in the lower plenum for 90% load is 1336 K which is also above the sodium boiling limit of 1213 K and it remains above the limit for 90 min. The stream function and isotherms at 1 h after accident are depicted in Figs. 10 and 11 respectively. It can be observed from these figures that a good natural circulation of sodium is established in the lower and side pools. The core debris heat generated in the lower plenum is transported to hot pool by natural convection of sodium through the gap created due to meltdown in the grid plate and the core. From the hot pool where the

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core debris is assumed to be settled in the gap, it is seen that the average temperature in the inner vessel reaches a maximum of 1036 K in 37 min, which is within the safe temperature limit. The temperature on the HSP which is found to reach a maximum of 1073 K in 30 min is also below the safe temperature limit (Fig. 8). The maximum temperature in the debris bed now is 1176 K, which is within the sodium boiling limit of 1213 K. Finally a case of 50% core debris settlement in the gap and the balance being relocated to the heat shield plate in the lower plenum is assumed. Fig. 9 shows the evolution of temperature in the inner vessel, HSP and the maximum debris bed temperature with time. It is clear from the figure that the temperature of the inner vessel increases to 1061 K in 40 min. The temperature in the HSP reaches 951 K in 25 min, and the maximum temperature in the debris bed is 998 K. All these temperatures are also within the acceptable threshold temperature limits. The streamlines for this case depicted in Fig. 12 show the establishment of natural circulation of sodium in the lower plenum is similar and the path for heat transfer from lower to upper plenum is through the gap created due to core meltdown. The isotherms represented in Fig. 13 show rise in the maximum temperature in the source region in the gap between the core periphery and the inner vessel. However, the temperature at the inner vessel is within safe temperature limits due to the continuous circulation in the side pool. Hence, it is found from the above observations that the structural integrity of the inner vessel placed in the upper pool is maintained even for 50% core debris settlement in the gap between the core periphery and the inner vessel. Further, a single core catcher system is found to be within the safe temperature limit for 70% core debris settlement on the core catcher and 30% in the upper pool. 5.2. Effect of DHX availability on heat transfer characteristics Post Accident Heat Removal (PAHR) analysis is carried out to investigate the effects of non-availability of decay heat exchangers on the structural temperature. The number of DHXs available for heat removal is varied from 4 to nil. Its influence on the heat transfer characteristics in the plenum is analyzed in detail. The heat source is assumed to be distributed 50% in upper plenum, i.e., gap between the core periphery and inner vessel and the remaining 50% in the lower plenum, on core catcher plate. Fig. 14(a)–(c) depict the evolution of temperature at three critical locations in

Fig. 14. Transient temperatures evolution at (a) inner vessel; (b) HSP surface and (c) maximum debris temperature for different number of DHXs in operation.

DHXs are dipped the heat is eventually dissipated to air via SGDHRS. For a case of 20% core debris settlement in the gap and remaining 80% being relocated on the HSP, it is observed that the average temperature in the inner vessel reaches a maximum of 1003 K in 33 min before it starts decreasing again (Fig. 7). The temperature on the HSP reaches a maximum of 1140 K in 30 min. These temperatures do not exceed the threshold temperature limit. However, the maximum temperature in the debris bed is 1250 K, which exceeds the sodium boiling limit for 40 min. When 30% of the

Fig. 15. Variation of pool temperature for different number of operational DHX for 5 h.

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Fig. 16. (a)–(e): Isotherms at 5 h representing temperature evolution with different number of operational DHXs (all scales in K).

the pool with different number of DHX in operation. When only 3 DHX are assumed to be in working condition, the temperature in the inner vessel is found to be 1064 K at 43 min after the event, the HSP surface temperature is 956 K at 25 min and the maximum temperature at the debris bed in the lower plenum is 1003 K at 23 min. This confirms that there is not significant increase in the temperature values at the different regions by non-availability of one DHX. When only two of the four DHX are available, the temperature in the inner vessel is seen to reach 1065 K in 45 min and the same in the HSP surface reaches 958 K in 25 min. The maximum temperature at the debris bed in the lower plenum reaches 1005 K in 23 min. In this case also, there is no significant increase in the temperature values at the different regions by non-availability of two DHX. Similarly, the temperatures at the inner vessel, core catcher and the maximum debris bed temperature are observed for the cases of availability of only one DHX and no DHX cases. It can be observed from these figures (Fig. 14(a)–(c)) that when no DHX are available, the temperatures at various regions start to increase after 1 h after accident. Hence, it is observed that the availability of the DHX does not play a significant role in heat removal in the initial hours of

PAHR, demonstrating adequate grace time for deployment of heat sink, thanks to the availability of large thermal inertia of the pool. 5.3. Effect of different DHXs on pool temperature It is clear from the previous section that the DHXs do not remove significant amount of heat during the initial hours of the accident. The decay heat from the degraded core debris increases the bulk temperature of the sodium present in the pool. Transient analyses are carried out to find the bulk sodium temperature rise which directly affects the structural temperature during PAHR posing threat to structural integrity. To investigate the decay heat of debris transferred to bulk sodium, the average pool temperature is obtained for the case with 50% source distribution to both the HSP and the gap. The effect of availability of different number of DHXs on the pool temperature is observed for 5 h after the accident. As the number of DHXs available is reduced from 4 to none, the average temperature of the pool is observed to rise with time. The variation in the pool temperature for different number of operational DHX for initial 5 h is shown in Fig. 15. It is found that the bulk sodium temperature initially remained nearby constant at 672 K for 75 min and thereafter reduced if all the four DHXs

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Fig. 17. (a)–(e): Stream functions at 5 h with different number of operational DHXs (all scales in kg/s).

are working satisfactorily during PAHR. For the case of three and two DHXs in operation, the maximum bulk sodium temperature rises to 692 K and 706 K and remains more or less constant after 170 min and 185 min of the event respectively. It is also clear from this figure that the availability of even one DHX in the pool can effectively remove the decay heat from the degraded core debris without significant increase in the structural material temperature during the initial hours. The study also revealed that if no DHX is available for initial 5 h of PAHR, there is not much concern of decay heat removal from the destroyed core and the heat of debris is completely transferred to increase the bulk sodium temperature in the pool. At the same time, the decay heat is also reduces with the associated change in pool temperature. The isotherms at 5 h after accident for the cases with different number of operational DHX are depicted in Fig. 16(a)–(e). It can be observed from the figures that the pool temperature increases as the number of DHXs in operation are reduced. The increase in pool temperature continues for a longer duration of time if the number of DHX in operation is less. The stream function for the different cases is depicted in Fig. 17(a)–(e). The strength of the natural circulation with different decay heat exchangers can be seen from these figures. It is clear from these figures that the stream function values are reduced to

half when no DHXs are in operation as compared to the case where all four DHXs are in functional mode. However, due to the excellent heat diffusion property of sodium, the heat is dissipated from the source to the bulk of the fluid which in turn reduces the source and structural temperature of the inner vessel.

6. Conclusion Thermal hydraulic investigations of post accident decay heat removal from a degraded core of a typical 500 MWe pool type sodium cooled fast reactor have been carried out. The main focus of the study is to determine the amount of heat generating core debris that can safely be accommodated in the radial gap between the core periphery and the inner vessel in upper plenum and on the core catcher plate in the lower plenum without exceeding safe structure temperature limits. The analyses also focused on the grace time available after CDA for deployment of decay heat removal system from the considerations of structural temperature in main vessel and bulk coolant temperature. Towards these the transient form of 2-D conservation equations of governing turbulent natural convection have been solved by a finite volume based

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solver. Validation of the mathematical model is done by in-house experimental data obtained in the PATH experimental facility. It is found that the temperature of the inner vessel does not exceed the safe temperature limit even when 50% of the debris settles in the upper plenum. Further, a single tray core catcher can safely accommodate decay heat due to 70% of the whole core debris by establishing natural circulation in the lower pool. The peak temperatures in the inner vessel, core catcher plate and the debris do not increase significantly even when the decay heat exchangers are deployed after 5 h. This is due the large thermal capacity of sodium in the primary system. These results provide significant data for possibility of retaining of core debris in upper and lower plenum, for post accident heat removal in the design of future SFRs. References Bertrand, F., Marie, N., Prulhière, G., Lecerf, J., Seiler, J.M., 2016. Comparison of the behavior of two core designs for ASTRID in case of severe accidents. Nucl. Eng. Des. 297, 327–342. Chellapandi, P., Srinivasan, G.S., Chetal, S.C., 2013. Primary containment capacity of Prototype Fast Breeder Reactor against core disruptive accident loading. Nucl. Eng. Des. 256, 178–187. Gnanadhas, L., Sharma, A.K., Malarvizhi, B., Murthy, S.S., Hemanth Rao, E., Kumaresan, M., Ramesh, S.S., Harvey, J., Nashine, B.K., Chellapandi, P., Chetal, S.C., 2011. PATH- an experimental facility for natural circulation heat transfer

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