Nuclear and thermo-physical data/correlations

CIAE used the neutronics data supplied by Argonne and the SAS4A/SASSYS-1 thermo-physical data correlations.

7.1.3. Thermal hydraulics methods and models 7.1.3.1. Code(s) used

The code used was SAS4A/SASSYS-1, version 3.0. SAS4A/SASSYS-1 is designed to perform deterministic analysis of severe accidents in liquid metal cooled reactors (LMRs).

Detailed, mechanistic models of steady state and transient thermal, hydraulic, neutronic and mechanical phenomena are employed to describe the response of the reactor core and its coolant, fuel elements and structural members to accident conditions caused by loss of coolant flow, loss of heat rejection or reactivity insertion.

7.1.3.2. Basic method

In space, each SAS4A/SASSYS-1 channel represents one or more subassemblies with either a single pin model or a multiple pin model. Many channels are employed for a whole core representation. Heat transfer in each pin is modelled with a two dimensional (r/z) heat conduction equation. Single-phase coolant thermal hydraulics is simulated with a unique, one dimensional (axial) liquid metal flow model. Reactivity feedbacks from fuel heating (axial expansion and Doppler), coolant temperature changes and fuel and cladding temperature changes are tracked with first-order perturbation theory. Reactivity effects from reactor structural temperature changes yielding radial core expansion are modelled. Changes in-reactor power level are computed with point-kinetics.

7.1.3.3. Model

For SHRT-17, the average inlet temperature, average outlet temperature, core outlet pressure and steady state coolant flow rate per pin were used in each channel as the core boundary conditions. For SHRT-45R, the reactivity feedbacks for Doppler, axial expansion, radial expansion and coolant density were calculated and point-kinetics was used.

The basic SAS4A/SASSYS-1 modelling options provide for only a single inlet plenum, and so the high pressure and low pressure inlet plena were not modelled separately. Instead, the normalized driving pressure was used as a primary circuit boundary condition for the reactor core.

7.1.4. Blind results 7.1.4.1. SHRT-17

The calculated peak fuel, cladding and coolant temperatures during the SHRT-17 test are illustrated in FIG. 32. The highest fuel, cladding and coolant temperatures are around 900K and appear at around 53s. The sodium saturation temperature is around 1088K, so there is more than a 100K margin to boiling for the coolant. The solidus temperature of the U-5Fs alloy is more than 1200K, which means there is ample safety margin for the fuel during this transient. The simulation results demonstrate well the inherent safety characteristics of the EBR-II reactor during the protected loss of flow transient.

FIG. 32. Peak in-core temperatures during SHRT-17, phase 1.

In order to measure a detailed temperature distribution at different elevations and in subchannels of subassemblies, instrumented subassemblies with thermocouples were placed in the core during the SHRT-17 test. Simulation results for temperatures at different elevations in XX09 are illustrated in FIG. 33. Subchannel modelling was not used, so no radial temperature variation was calculated.

7.1.4.2. SHRT-45R

CIAE did not generate any blind results for SHRT-45R.

7.1.5. Final results, data comparisons 7.1.5.1. SHRT-17

The calculated peak fuel, cladding and coolant temperatures during the SHRT-17 test are illustrated in FIG. 34. The highest fuel, cladding and coolant temperatures are around 822K and the highest temperatures appear at around 70s. The sodium saturation temperature is around 1088K, so there is more than a 100K margin to boiling for the coolant. The solidus temperature of the U-5Fs alloy is more than 1200K, which means there is ample safety margin for the fuel during this transient. The simulation results demonstrate well the inherent safety characteristics of the EBR-II reactor during this protected loss of flow transient.

FIG. 33. XX09 temperatures during SHRT-17, phase 1.

FIG. 34. Peak in-core temperatures during SHRT-17, phase 2.

The calculated core inlet and outlet temperatures during the transient are illustrated in FIG.

35. The experimental data are shown by the dashed lines, and the calculation results are plotted using solid lines. Both experimental and calculated inlet temperatures of the core remain almost invariant and fall on top of each other, while the outlet temperature of the core changes with the transient. The highest outlet temperature occurs at around 100s, and the highest outlet temperature is about 780K. From FIG. 35, it can be seen that the calculated outlet temperature has the general same trend as the experimental data but predicts a much higher peak temperature and faster decline than the experimental data.

Simulation results of temperatures at different elevations in XX09 are illustrated in FIG. 36.

The experimental data are plotted using dashed lines, and the calculation results are plotted using solid lines, as indicated in the figure. Only radially averaged temperature results are available, since subchannel modelling was not used. From FIG. 36, it can be seen that the calculation results follow curves similar in shape data to the experimental data but overshoot the actual temperatures.

FIG. 35. Core inlet and outlet temperature during SHRT-17, phase 2.

Simulation results of temperatures at different elevations in XX10 are illustrated in FIG. 37.

Again, the experimental data are plotted using dashed lines, and the calculation results are plotted using solid lines. From FIG. 37, it is clear that the calculation results do not match the experimental data very well, and this is a problem that remains unsolved.

FIG. 36. XX09 temperatures during SHRT-17, phase 2.

FIG. 37. XX10 temperatures during SHRT-17, phase 2.

7.1.5.2. SHRT-45R

The calculated peak fuel, cladding and coolant temperatures during the SHRT-45R test are illustrated in FIG. 38. The highest fuel, cladding and coolant temperatures are around 1050K, and the highest temperatures appear at around 60s. Sodium saturation temperature is around 1200K, so there is more than 100K margin to boiling. The solidus temperature of the U-5Fs alloy is more than 1200K, which means there is significant safety margin for the fuel during this transient. The simulation results demonstrate well the inherent safety characteristics of the EBR-II reactor during an unprotected loss of flow transient.

The calculated core total power, fission power and decay power are illustrated in FIG. 39. The experimental data for fission power are shown by the dashed line, and the simulation results are are plotted with solid lines. As can be seen in FIG. 39, after the station blackout occurs, the power of the reactor decreases quickly and eventually converges to a low power level. The calculated fission power compares very well to the recorded fission power.

The calculated and recorded core outlet temperatures during the transient are illustrated in FIG. 40. The outlet temperature of the core changes with the transient. The highest outlet temperature occurs at around 60s, and the highest predicted outlet temperature is about 980K.

From FIG. 40, it can be seen that the calculated results have the same overall trend as the experimental data but overpredict the outlet temperature, especially near the temperature peak.

FIG. 38. Peak in-core temperatures during SHRT-45R, phase 2.

The reactivity changes during the transient are illustrated in FIG. 41. Only the Doppler reactivity, axial expansion reactivity, radial expansion reactivity and coolant density reactivity were considered. From FIG. 41, it is clear that the largest reactivity is the sodium void reactivity.

FIG. 39. Power during SHRT-45R test, phase 2.

FIG. 40. Core outlet temperature during SHRT-45R test, phase 2.

FIG. 41. Reactivity during SHRT-45R test, phase 2.

Simulation results of average temperatures at different elevations in XX09 are illustrated in FIG. 42. It can be seen that the calculation results have the same general trend as the experimental data, with the peak temperature at each elevation close to the experimental data.

However the calculated temperatures decrease more quickly than do the recorded data.

Simulation results of average temperatures at different elevations in XX10 are illustrated in FIG. 43. As for XX09, the plots show that the calculation results follow shapes similar to the

experimental data, but in this case, the temperatures at each elevation throughout the transient are significantly lower than the experimental data. The reason for this discrepancy has not yet been identified.

FIG. 42. SHRT-45R temperatures at different elevations in XX09, phase 2.

FIG. 43. SHRT-45R temperatures at different elevations in XX10, phase 2.

7.1.5.3. Model improvements

For the final Phase 2 model of SHRT-17, the channel partition and the geometry model were changed appreciably from the Phase 1 model. Also, for each channel, the axial segments partitioning, flow path area and structure and reflector data were modified appreciably. These modifications made little difference to the results. The greatest change was in simulating a natural circulation flowrate of about 3%, determined by experience and the experiment results. Applying this flowrate produced core temperature results that confirmed the core model.

7.2. NORTH CHINA ELECTRIC POWER UNIVERSITY (CHINA) 7.2.1. Geometry/discretization

The main components in this simulation include the reactor core, IHX, sodium pool and pump. The boundary conditions for the secondary side of the IHX are the coolant inlet temperature and the mass flow rate.

All rows of the core subassemblies were divided into 9 channels, representing the driver fuel, stainless steel reflector, control rods and blanket regions. Each channel was divided into 26 axial slices with 4 radial nodes in the fuel and 1 node each in the gap, cladding, coolant and structure. Each channel had a separate power fraction and flow fraction distribution. All channels were treated using similar thermal hydraulic models. Coolant flow in each channel was parallel to the channel axis, without cross-flow between adjacent channels. At the same time, axial heat conduction in the coolant and fuel was neglected.

The nodalization of the intermediate heat exchanger (IHX) is shown in FIG. 44. All the heat transfer tubes were simplified to one representative tube. There are two basic assumptions: 1) ideal mixing of coolant at the inlet and outlet and 2) fully developed convective heat transfer.

There were four kinds of radial nodes: secondary coolant, tube, primary coolant and shell wall. The nodes in the coolant and structure were placed in a staggered manner.

FIG. 44. Nodal diagram for IHX.

7.2.2. Nuclear and thermo-physical data/correlations

Total fission power was calculated based on a neutron point-kinetics model containing six groups of delayed neutrons. The time-dependent portion of the decay heat contribution was handled by a tabular look-up of data supplied in the input file. Thus, paired points of time vs.

decay heat fraction were supplied by the user.

Contributions to reactivity feedback effects consisted of the Doppler effect, sodium density, fuel axial expansion and structural expansion bowing. All reactivity feedbacks were calculated based on one control volume for each axial channel.

The basic correlations and laws to calculate the thermo-physical properties, friction factor and heat transfer coefficients are described below. Thermal properties such as thermal conductivity, heat capacity, enthalpy, density, and viscosity for sodium are given in [28].

Friction factors were assumed to be a function of Reynolds number and roughness. Friction factor for laminar flow was calculated as f=64/Re. For turbulent flow, the friction factor was determined by the Moody diagram.

The heat transfer coefficient for liquid sodium flow through tube bundles was determined by modifying the Schad correlation [70].

For sodium flow through the piping, the heat transfer coefficient was determined by the Graber-Rieger correlation ([71], [72], [73], [74]).

7.2.3. Thermal hydraulics methods and models 7.2.3.1. Code(s) used

For the SAC-CFR code, the primary options and assumptions used for the EBR-II analysis included:

(a) Liquid sodium was regarded as an incompressible fluid;

(b) The assumptions for the IHX model were mentioned in the last section;

(c) Fractional power and flow deposited in each channel were specified through input and did not change during the transient simulation;

(d) The radial node mesh in the fuel pin used equal radius increments;

(e) A three-dimensional sodium pool model was adopted;

(f) The pressure drop was specified through input. After the steady state initialization, a loss coefficient was calculated. Then the loss coefficient remained constant during the transient simulation.

7.2.3.2. Basic method

The SAC-CFR code is divided into three major subroutine categories: read subroutine, steady state calculation and transient calculation. The basic call relationship among the subroutines is shown in FIG. 45.

Since the thermal properties are independent of pressure, the energy equation and momentum equations for the sodium coolant were decoupled to simplify the calculation. The overall solution logic for the transient simulation is as follows:

(a) HYDDRV—primary and intermediate loop hydraulic calculation to calculate the mass flow rate in the loop;

(b) POWDRV—fission power generation calculation, including the reactivity feedback;

(c) COLDRV—coolant heat transfer calculation in the core;

(d) FUELDRV—calculation of heat transfer in the fuel, including the heat transfer through the cladding;

(e) TLPDRV—loop energy calculation for primary and intermediate loops, including the IHX;

(f) PCSPPS—plant control system and protection system calculation.

SAC-CFR

FIG. 45. Call relationship between SAC-CFR subroutines.

In addition, a three-dimensional sodium pool model was used. The simplifications made to the pool model can be summarized as:

(a) The pool was modelled to 360° circumferentially using a cylindrical coordinate system.

The intermediate heat exchanger and pumps occupied one or more control volumes according to their actual size;

(b) The fuel handling structure was neglected;

(c) The openings on the primary side of the IHX were inlet boundaries for the pool model, with the rest of the IHX treated as a solid structure;

(d) The solid structures of the IHX and pumps were modelled with the porous medium method.

7.2.3.3. Model

EBR-II is modelled through the input file. The input file consisted of five parts: main vessel model, primary and intermediate loop model, property data, steady state initialization data and transient input. After reading the geometry data and steady state initialization data, the plant balance state was determined by steady state initialization. Then the transient simulation was started according to the trigger point or boundary conditions specified in the transient input file.

A neutron point-kinetics model was used to calculate the fission power. A single channel model was employed to simulate the thermal hydraulic response in the core. Flow and heat transfer in pipes and the heat exchanger were assumed to be one dimensional. The sodium pool was analyzed with a three-dimensional model.