Normally, the equations for neutron kinetics are obtained from the general space and time dependent multi-group diffusion equations for a finite reactor of non-uniform properties. The 3-D neutron kinetic calculations are often made with fast and thermal energy groups and six groups of delayed neutrons. To reduce the sophistication, a modified one group diffusion equation can be considered for fast calculations. In this, only the dynamics of the thermal energy group is simulated in detail.

For simplicity of calculations, most of the large system codes consider the point kinetics model assuming the overall buckling to be almost constant during the transient. The parameters used in these equations such as delayed neutron fraction, prompt neutron life time, delayed neutron precursor concentration, etc. are defined taking into account some phenomena formally neglected during the derivation of the point kinetics equation. An improvement to the point kinetics model is reflected in the one-dimensional kinetics equation that is also adopted in BWR calculations. In this, the radial buckling is assumed not to change during the transients. In point kinetics, reactivity coefficients for the void and Doppler effect are provided as inputs for the whole core. However, 1D and 3D neutron kinetics models facilitate direct evaluation of these feed back effects.

For the single-phase case, the natural circulation flowrate is found by integration of the loop momentum equation and coupling this to the energy equation. Thus, the primary temperature increase is given by the standard result,

(17)

where the fluid thermal expansion drives the convective flow around the loop. Also, from the heat balance across the HX, we have a second relation for the power, taking the primary to secondary temperature difference as close to that due to conduction across the tube wall only, with a correction coefficient for any surface, plugging, corrosion, and thermal resistance effects. To maximize the heat removal we assume the HX to not be limiting in capacity, and hence may take the core outlet temperature as saturated boiling, and the inlet temperature as close to the HX secondary temperature.

From the two expressions for the power, there is the following result for the maximum heat removal in a natural convection loop with onset of bulk boiling at the core exit is the limit,

(18)

The trends are somewhat counter intuitive for several reasons. The maximum heat removal is very sensitive to the HX design, relies on maximizing the primary to secondary temperature drop, and hence minimizing the core to HX elevation difference, and also maximizing the loop flow resistance.

The generic objective of the activity is to achieve an estimation of the maximum thermal power removable by NC in PWR systems. As already mentioned, no care is given to the thermohydraulics-neutronics interaction and, therefore to the actual possibility of generating the considered fission power. The analysis is carried out into two parts the former related to ITF, to give more realism to the calculated results, the latter related to a PWR. In all cases, the primary system pressure and the SG conditions (pressure, level and feedwater temperature) are kept constant at the nominal values if not differently specified. The main circulation pumps are at zero speed and the locked rotor hydraulic resistance of the impeller is taken into account. The core power and the feedwater flowrate levels are consistently modified.

In 1996, a new project for the investigation of passive safety injection systems of ALWR’s begun. The new project, entitled "Investigation of Passive Safety Injection Systems of Advanced Light Water Reactors", is a part of the INNO cluster of the European Commission Nuclear Fission Safety (NFS-2) Programme. The project received funding from the European Commission. The general objectives of the new project are:

• to provide new and independent information about passive safety injection system performance,

• to contribute to a public data base for the users and developers of thermal-hydraulic computer codes on the phenomenological behaviour of PSIS’s in LOCA conditions, and

• to identify the accuracy, uncertainties and limitations of thermal-hydraulic computer codes in the modelling of passive safety injection system behaviour.

Tests have been performed with a Plate Condenser, which was installed within the Condenser Tank, see Fig. 10. The atmosphere tested was pure steam, a mixture of steam and oxygen, a mixture of steam and helium and a mixture of steam, oxygen and helium.

In Fig. 11 the influence of the non-condensables on the transferred power is shown. Fig. 12 shows one test with steam and oxygen, showing an accumulation in the lower part of the tank.

The LWR transient analysis code, RELAP5, was developed at the Idaho National Engineering Laboratory (INEL) for the U. S. Nuclear Regulatory Commission (NRC). Code uses include analyses required to support rulemaking, licensing audit calculations, evaluation of accident mitigation strategies, evaluation of operator guidelines, and experiment planning analysis. RELAP5 has also been used as the basis for a nuclear plant analyser. Specific applications have included simulations of transients in LWR systems such as loss of coolant, anticipated transients without scram (ATWS), and operational transients such as loss of feedwater, loss of offsite power, station blackout, and turbine trip. RELAP5 is a highly generic code that, in addition to calculating the behaviour of a reactor coolant system during a transient, can be used for simulation of a wide variety of hydraulic and thermal transients in both nuclear and non-nuclear systems involving mixtures of steam, water, non-condensable, and solute.

The MOD3 version of RELAP5 has been developed jointly by the NRC and a consortium consisting of several countries and domestic organizations that were members of the International Code Assessment and Applications Program (ICAP) and its successor organization, the Code Applications and Maintenance Program (CAMP). Credit also needs to be given to various Department of Energy activities, including the INEL laboratory-directed discretionary funding program. The mission of the RELAP5/MOD3 development program was to develop a code version suitable for the analysis of all transients and postulated accidents in LWR systems, including both large — and small-break loss-of-coolant accidents (LOCAs) as well as the full range of operational transients.

The RELAP5/MOD3 code is based on a non-homogeneous and non-equilibrium model for the two-phase system that is solved by a fast, partially implicit numerical scheme to permit economical calculation of system transients. The objective of the RELAP5 development effort from the outset was to produce a code that included important first-order effects necessary for accurate prediction of system transients but that was sufficiently simple and cost effective so that parametric or sensitivity studies were possible.

The code includes many generic component models from which general systems can be simulated. The component models include pumps, valves, pipes, heat releasing or absorbing structures, reactor point kinetics, electric heaters, jet pumps, turbines, separators, accumulators, and control system components. In addition, special process models are included for effects such as form loss, flow at an abrupt area change, branching, choked flow, boron tracking, and non-condensable gas transport.

The system mathematical models are coupled by an efficient code structure. The code includes extensive input checking capability to help the user discover input errors and inconsistencies. Also included are free-format input, restart, renodalization, and variable output edit features. These user conveniences were developed in the recognition that generally the major cost associated with the use of a system transient code is in the engineering labour and time involved in accumulating system data and developing system models, while the computer cost associated with generation of the final result is usually small.

Main characteristics of the code are as follows:

— One-dimensional system thermal hydraulic code (enabling also pseudo two — or three­dimensional modelling);

— Basic two-phase model with 6 balance equations (mass, momentum and energy equation for each phase) with additional capability to model boron in liquid phase and non­condensable gases in vapour phase;

— Set of closure correlations and models (best-estimate) connected to the basic model by help of system of flow regimes and set of heat transfer modes;

— Point kinetic model;

— Other physical models (dynamic model of fuel-clad gap, radiation heat transfer, metal — water reaction, etc.);

— Special models for NPP components (pumps, various types of turbine, accumulator, pressurizer, etc.);

— Extensive system of trips and control components for modelling control system;

— Very flexible character of the code itself and of the input data.

In the event of failure of the active residual heat removal systems, four containment cooling condensers (CCC) are designed to remove residual heat from the containment to the dryer — separator storage pool located above the containment. The CCCs are actuated by rising temperatures in the containment. They use natural circulation both on the primary and on the secondary sides. The nominal heat transfer capacity of each condenser is 4 MW based on a containment pressure of 3 bar (absolute) and a cooling water temperature of 100°C. In a hypothetical core melt accident the thermal capacity could be 2 or 3 times higher, depending on the higher containment pressure and temperature. The containment cooling condenser has been experimentally tested at nearly original scale at the PANDA facility of the Paul Scherrer Institute in Switzerland.

The working principle of the CCC is shown in Figure 3. It comprises a simple heat exchanger mounted about 1 m above the water level of the core flooding pool. If the temperature in the drywell atmosphere increases over that in the dryer-separator storage pool, the water inside the heat exchanger tubes heats up. It flows to the outlet line due to the slope of the exchanger tubes. The outlet line ends at a higher elevational level than the inlet line, so the lifting forces are increased for the whole system. Depending on the heat transfer rate and cooling water temperature, secondary-side flow can be either single-phase, intermittent or two-phase.

Depending on the type of loss-of-coolant accident (LOCA) and on the time after onset of accident conditions, the medium on the heat exchanger primary side is either nitrogen, a
nitrogen-steam mixture or pure steam. In the hypothetical case of a core melt accident, a hydrogen-steam mixture would also be possible. Given nitrogen, steam and mixtures thereof, primary flow is downwards because the densities of pure gases and a nitrogen-steam mixture increase with decreasing temperature. This results in the expected downward flow. Condensed steam drops into the core flooding pool. However, the opposite is true for a hydrogen-steam mixture, as the density of this mixture decreases with decreasing temperature, resulting in an upward flow through the heat exchanger tube bundle. This does not pose any problem for the SWR 1000 because both directions of flow on the primary side are equivalent.

Finned-tube

cooler

Nevertheless, problems will arise if more hydrogen is generated than can be accommodated in the containment above the CCCs. The atmosphere in the containment is stratified with a high hydrogen content above the CCCs and a high steam content below. With increasing hydrogen mass the boundary between both stratified regions would drop lower and lower. After a certain time the CCCs would become ineffective because they would be surrounded by cooled hydrogen. To prevent this condition from occurring Siemens designed a hydrogen overflow line. The upper end of this line is located at a higher elevation than the CCCs and its lower end is higher than the lower end of the vent pipes. When the heat transfer capacity of a CCC deteriorates, the pressure in the containment increases until the hydrogen overflow pipe is empty of water (but the vent pipes still contain some water). This produces a forced-flow condition of cooled hydrogen-steam mixture from the drywell to the wetwell. The steam condenses in the wetwell pool and the hydrogen rises into the wetwell pool atmosphere. The hydrogen mass flow is self-controlled. The boundary of the stratified regions stabilizes at a position at which as much steam is condensed as is being generated.

In the case of a hypothetical core melt accident, there first occurs normal natural circulation in the drywell. Later, there is natural circulation with opposed flow directions on the primary side and with stratification between the cold hydrogen above and hot steam below. In the final phases there is a self-induced forced-flow of hydrogen from the drywell to the wetwell. To
simulate all these effects by way of computer code modeling would pose a considerable challenge. The experimental tests performed at the PANDA test facility were much easier, and demonstrated that the entire system is effective and stable, and the pressure differential of several kPa between the drywell and wetwell was generated without any problems.

Natural circulation concept is used in the AC600/1000 passive safety system design. The system is actuated by gravity, natural circulation or pressurized gas. Following accident, AC600/1000 is able to maintain core cooling and containment integrity without operator’s intervention. This is an important safety requirement in the AC600/1000 design.

Emergency residual heat removal system (ERHRS) is used to remove decay heat from reactor core following accident. Secondary side of SQ emergency feedwater tank and air cooler establish a natural circulation cycle. Air coolers are located in a chimney. The heat transfer area is about 750m for each air cooler. Emergency feedwater tank volume is 25m for each SG cycle.

Passive safety injection system of AC600/1000 (PSIS) is mainly used to mitigate the consequence of LOCA. The core cooling water is provided through use of the following four

water sources: core make-up water tank (volume is 2 or 3 x 40m ), accumulator (volume is 2 or 3 x 40m ), refueling water storage tank and containment sump. Low pressure safety injection subsystem uses two or three active pumps, design flow rate of each which is 142 kg/s.

The containment of AC600/1000 is a two-shell structure. Between inner steel shell and outer reinforced cylindrical concrete shell, there is a baffle to form an annular wind duct. Containment top water storage tank capacity can meet the requirement of 72 hours for steel shell cooling after large LOCA. In the top of containment, there is a cooling water distribution to make the heat removal more efficient and quick in the early phase of an accident. For long term cooling, peak pressure of the containment is not larger than 90% of containment design pressure. For severe accident, containment pressure is below failure pressure. Containment spray system is eliminated in the AC600/1000 design.

The second series of the EC funded project included five experiments for the investigation of the break location and CMT scaling (smaller CMT) on the PSIS behaviour. PACTEL operators made small changes to the CMT instrumentation to better detect CMT level and thermal stratification behaviour. The PSIS consisted of a Core Make-up Tank, which had connections to downcomer through an injection line and to one cold leg through a pressure balancing line. The break location and CMT scaling influences can be summarised as follows:

• Flow of cold water from downcomer to cold leg of the broken loop occurred in both hot and cold leg break experiments. The cold water did not, however, flow to the CMT through the PBL. So, no condensation problems occurred in the CMT due to flow of cold water to the tank.

• The saturated liquid layer in the CMT was very thin in the experiment where the break located in the cold leg close to downcomer. In the hot leg and cold leg inlet part break experiments, the saturated liquid layer was thick, due to flow of water to the tank during the injection phase. In all experiments, the thickness of thermally stratified region in the CMT increased during the experiments.

• The overall primary loop behaviour was similar in experiments with small and large CMT. As expected, the availability of larger amount of ECC water in the experiment with large CMT delayed the core heat-up.

The experiment series also investigated the CMT behaviour in a situation where the CMT is initially full of hot water. This may happen in the AP600 plant if the injection line check valve leaks. The programme also included an experiment without flow distributor (sparger) in the CMT. The results of these two experiments can be summarised as follows:

• Practically no recirculation flow occurred in the experiment where the CMT was initially full of hot water. The recirculation flow did not begun although there was a small initial density difference between the PSIS lines. This did not, however, cause any problems for the safety injection from the CMT.

• The fact that the CMT was full of cold water influenced the water distribution in the primary loop during the transient. This had affects on the water level formation in the vessel and the timing of the core heat-up.

• The experiment without sparger demonstrated the importance of the flow distributor on the CMT behaviour: the removal of the sparger led to rapid condensation in the CMT, which stopped safety injection from the tank. The PACTEL operators had to terminate the experiment due to condensation problems in the CMT.

• Condensation problems occurred in GDE-35 experiment when water flowing to the CMT broke saturated water layer in the CMT. This did not happen in the experiments with sparger, which distributed the incoming water horizontally into the tank.

E. F. Hicken

Institute for Safety Research and Reactor Technology, Forschungszentrum Julich, Germany

H. Jaegers

Institute for Safety Research and Reactor Technology, Forschungszentrum Julich, Germany

Abstract. During shutdown the decay heat in commercial Boiling Water Reactors is removed from the core region by active and redundant pump/heat exchanger-systems which are, in addition, supported by emergency power. To study the capability of the newly developed emergency condensers to remove energy produced within the core region to a large water pool outside the Reactor Pressure Vessel by natural convection, a related test in the NOKO facility as performed. The pressure vessel in the NOKO facility has been flooded above the inlet line to the emergency condenser and heated up to about 100°C. The natural circulation resulted in a cool down of the water within the pressure vessel. With two specially designed grids equipped with up to 12 thermocouples the temperature fields in two cross sections were measured; no plume-effect was identified. The vertical temperature profile was measured with thermocouples. The test showed that decay heat could be removed some time after scram to an outside pool by natural convection processes; the time after scram depends on the emergency condenser heat exchange area.

1. INTRODUCTION

It is well known that also after scram decay heat in the range of 30 to 40 MW for a LWR with 1000-1300 MW(e) is produced within the reactor core. To avoid an evaporation of fluid active and redundant heat removal systems are mandatory; usually diesel generators are installed, in addition, to maintain the heat removal capability also in case of the Loss-Of — Outside-Power (LOOP).

The SWR 1000 is equipped with emergency condensers for the removal of decay heat mainly for transients without loss of coolant; for loss-of-coolant accidents these condensers will assist the decay heat removal for some time.

These emergency condensers do not need any power and no valves are needed to start the operation; they are “passive” by definition. Therefore, a test should be performed to evaluate the capability of these condensers to remove heat produced in the core region to the outside pool by natural convection and, in addition, at ambient pressure, as it will be during long term shutdown.

Although the removal of the decay heat cannot be expected from the beginning it would be beneficial if the decay heat could be removed after some time. This capability would then be diverse to the active heat removal systems.