In most innovative reactor designs long-term retention of a totally molten core inside the RPV is foreseen by ex-vessel cooling, provided by flooding the reactor cavity from the outside [16]. In several concepts (e. g. AP600, WWER-640) the emergency water pool is used for this purpose. The emergency pool is a containment sump that covers the RPV and the lower part of the RCS. After depletion of the emergency water sources the level in the pool exceeds the elevation of the main coolant lines, thus allowing for a post accident recirculation through the vessel in a natural circulation mode.

In the SWR-1000 the reactor cavity is flooded from the core flooding pool when the water inventory of the RPV drops to 40% [15]. The water enters the space between the vessel wall and insulation through gaps at the control rod drive penetrations and heats up to the saturation state at the vessel wall. The steam leaves this gap through windows in the insulation and will be condensed at the BC. The condensate will refill the core flooding pool closing the NC loop. Retention of the core melt within the RPV means that there can be no steam explosion in the containment or corium/concrete reaction [17]. In some designs the reactor cavity is flooded via special gravity driven systems. Other concepts utilises accumulators with gas under pressure. Passive reactor cavity flooding for core retention by ex-vessel cooling is also adopted in integral reactor design (e. g. ABV-6 [18]).

In-vessel core retention systems (IVCRS) based on in-vessel cooling of the molten core materials can prevent the RPV from failure by melt-through utilising in-vessel flooding of the molten core materials [19]. Direct contact between corium and steel must be avoided and a cooling capability must be provided. TMI-2 experience showed, that a small gap between the corium (crust) and the vessel can be sufficient for corium cooling by evaporation of water. An IVCRS for in-vessel cooling can consist of a structure artificially providing such gap (in­vessel core catcher), in combination with a special gravity driven corium flooding system. As an example for an in-vessel core catcher a steel shell with some distance from the inner surface of the lower plenum is proposed. For heat removal a mass flow rate in the order of 10 kg/s is sufficient. Alternatively, IVCRS are proposed using sacrificial material having a large heat capacity inside the lower plenum.

(a) Natural circulation systems are now implemented or being considered in many future reactor designs. Natural circulation is applied both to ensure the sufficient coolant flow in the core (reactors based on primary coolant natural circulation during normal operation) and to fulfill or support the fundamental safety functions. It should be noted that reactors based on natural circulation during normal operation, e. g. the Dodewaard Reactor in the Netherlands and VK-50 in Russia operated for an extended period of time;

(b) There is a consensus that natural circulation systems can provide more reliable means for a number of safety functions (e. g. decay heat removal from the core, heat removal from the containment atmosphere) and can be cost effective as these systems are less vulnerable to failures and therefore, the number of parallel trains in one system can be reduced. Different designs use natural circulation in different ways. Common design concepts and their respective needs relevant to natural circulation should be identified;

(c) If natural circulation is used as an operational principle of design (full power to be removed from the core by natural circulation) the most important task seems to be preventing the core from two phase natural circulation instabilities induced by neutronic feedback;

(d) If natural circulation is used as a working principle for safety systems, 1D codes can cope with the phenomena of natural circulation in many cases with some limitations and some uncertainties. The main drawbacks of natural circulation systems include the lower driving forces and less possibility to alter the course of an accident if something unexpected happens. Due to low driving forces, the operation of the natural circulation systems may be adversely affected by small variations in thermal-hydraulic conditions. The lower driving forces might also lead to quite large equipment where the role of 3D phenomena is essentially increased;

(e) In cases where multi-dimensional phenomena are present, the safety relevance and need for a 3D approach should be investigated first. An example is the effect of thermal stratification on the building condenser, observed in PANDA;

(f) The effect of non-condensable gases on condensation has been known for many years, and design features to address the problem (e. g. steam jet air ejectors for steam turbine cycles) are provided. In order to cope with this problem in future natural circulation systems, special design features should be considered for condensers.

The SPOT is 16 air cooled heat exchangers-condensers (four ones for each steam generator), located outside around containment. The heat exchanger is canned in the box connected with exhaust stack above to ensure the air circulation. The design basis of the SPOT is the removal of 60 MW heat at the maximum design temperature of the external air (plus 50°C) taking into account failure of one train, so the power of one heat exchanger is 5MW. The air flowrate is controlled by partial opening of the air gates to ensure that the power of the heat exchanger should not be excessive in the minimum design temperature of external air (minus 40°C). Under standby conditions the air gates are closed, but the heat exchangers are in the hot state due to air leakage through air gates and heat transfer through the building constructions. These thermal losses are estimated in the design by value less than 0,1% of reactor rated power. The levers with load open the gates under de-energization of the electromagnets keeping the air gates in the closed position, and the air natural circulation through heat exchangers begins.

To confirm the design solutions mentioned above and the system characteristics, the extensive experimental investigations have been performed in OKB Gidropress. The investigations were carried out at the external air temperature from minus 19°C to plus 30°C and the pressure in the steam-condensate circuit from 0,5 MPa to 6,4 MPa. The start-up of the system from hot standby condition is performed, the heat losses are determined at the closed air gates, the system power against external air temperature and steam pressure is determined, the heat removal control is checked by the partial opening of the air gates. The investigations have given the convincing substantiation that the SPOT design functions are performed for WWER-1000/V-392 in the real conditions.

The test rig constructed in OKB Gidropress in 1991, includes the full scale heat exchanger of the SPOT system with representation of air and steam-condensate circuits as shown in Figure

4. The heat exchanger consists of about 200 flat spirals having small slope in relation to horizontal. The tube bank about 2 m high and total surface more than 300 m has two inlet and two outlet collectors. On the sections with cross-sectional air flow, the tubes have ribs for intensification of the heat transfer to the air.

The tube bank is canned in heat-transfer box of square cross-section, which forms the heat exchanger air flow circuit. The lower and upper part of this box are connected with horizontal boxes for air inlet and outlet accordingly. The box upper outlet is connected to exhaust stack. The air gates are installed before of the heat exchanger and after one in pairs in one plane, each of these overlaps half of the cross-section of the circuit. The gate is the flat plate rotated about one’s own axis. The lever with load which is retained by the electromagnet, when the gate is closed, is attached to the axis. The horizontal inlet box, the box of the heat-transfer bank, the horizontal outlet box and exhaust stack together create the circuit for air natural circulation. In this circuit (see Figure 6) the elevations of its components (including the height from inlet box axis up to top of the exhaust stack) and cross sections of all air circuit elements per one real heat exchanger are represented in the full scale. Whole air circuit, starting from the heat exchanger, has been covered by thermal insulation and sheathed by aluminium sheets to create adequate conditions concerning to the heat losses of the real SPOT system.

During the tests, the parameters necessary both for the confirmation of the SPOT design operation and for the development and validation of calculation models for the whole system and its separate elements have been measured. The set of the measured parameters includes:

• flow rate, pressure and temperature of the superheated steam,

• feedwater flow rate,

• flow rate and temperature of the condensate,

• flow rate, pressure and temperature of the saturated steam supplied to the heat exchanger,

• pressure and temperature of the atmospheric air,

• pressure difference along the heat exchanger in the steam-condensate circuit,

• air temperature in the air circuit elements,

• temperature of the heat-transfer tube surface,

• temperature of the thermal insulation.

These measurements have been shown in the control room of the test rig and the registration of the parameters have been realized by the data acquisition system IMPACT 3590 with subsequent treatment by a personal computer.

The investigations to define the heat losses in hot standby condition with the closed air gates have been performed from 07.04.92 to 16.02.93 (totally 30 experiments). In experiments, the steam pressure before the heat exchanger accounted from 0,6 MPa to 6,39 MPa, the steam temperature before the heat exchanger accounted from 159°C-280°C and external air temperature accounted from minus 8°C to plus 18°C. It was found that the direct recalculation of the test results to define the system heat losses results in a larger losses than it is anticipated by the SPOT design (less than 0,1% of the reactor rated power). It was also determined that temperature of the external surface of the heat exchanger thermal insulation exceeds significantly the design value 45°C. The reasons of these differences for the test rig have been found and the recommendations have been developed for the design modification of the SPOT system thermal insulation for WWER-1000/V-392.

To define the dependence of the thermal power removed by the heat exchanger against the external air temperature and steam pressure, the tests have been performed under stationary conditions in the range of external air temperature from -19°C to +30°C and steam pressure at the heat exchanger inlet from 0,54 MPa to 6,39 MPa. During the period from 27.03.92 to 15.02.93, thirty experiments have been performed under positive external air temperature and ten experiments under negative one; the air gates were open partially in eleven experiments to evaluate possibility of the SPOT power control. It was determined that the heat exchanger removes from 5 MW to 7.4 MW at the pressure 6,3 MPa and the external air temperature from plus 30°C to minus 19°C. It was confirmed that the heat exchanger design geometrical characteristics ensure necessary power of the heat exchanger at the maximum design temperature 50°C and fully open air gates. For evaluation of the removed power control by means of partial closing of the lower air gates, two series of the experiments have been performed at the external air temperature plus 18-19°C and minus 11-16°C and steam pressure 6,3 MPa. The gates have been deviated on 10, 20 and 30 degrees in relation to horizontal. It was determined that the design configuration of the air gates ensures the power control of the heat exchanger under turn angles 0-40 degrees, after that the regulating capacity is lost. For example, the heat exchanger power at the turn angle of the gates 10 degrees is approximately 2,5 times less than under fully open gates, and at the angle 30 degrees the power is only approximately 15% less than the power at fully open gates.

To investigate the influence of the non-condesable gas on the heat exchanger operation, two experiments have been performed with supply of the different nitrogen quantity to the steam generator model. The gas quantity in these experiments was approximately 5,5 and 1,3 times more than quantity which corresponds to the real conditions of the reactor plant operation. A number of effects related to the nitrogen supply have been noted. In particular, the condensate temperature is reduced concerning saturation temperature due to decreasing of the partial steam pressure. The nitrogen is distributed on the collectors unequally, according to their location concerning steam pipeline. Discharge of a part of nitrogen with condensate flow was observed. The experiments have been performed with putting the heat exchanger into operation at the fresh steam supply and at the different quantity of nitrogen injected. It was determined that in all cases the heat exchanger power is stabilized in 40-50 seconds, and the value of the stationary power does not depend on quantity of the nitrogen injected and is only determined by the external air temperature. The tests performed have shown that the nitrogen is discharged gradually with the condensate. So, at the injection of the nitrogen quantity, which exceeds possible injection in real conditions about 5,5 times, the nitrogen has practically discharged completely from the heat exchanger in 10 minutes after the air gates opening.

The investigation of the stability of the SPOT heat exchangers parallel operation for the different loops has been performed on TDU-1 facility in one of the institutes of the Science Academy of Belorussia. The experiments have been performed on the two-loop three-circuit test rig by 1 MW power. The steam generator and the SPOT air heat exchanger-condenser were modelled in each loop. It has allowed to model operation of two train of the SPOT in parallel at the different loading of the heat exchangers. The experiments have been performed at the air temperature from plus 5°C to plus 31 °C and at the steam pressure from 0,6 MPa to 5,4 MPa. The experiments have demonstrated steady operation of the heat exchangers, the variations of the condensate and heat transfer tubes temperatures were not noted. In the same institute the experiments have been performed on the SPOT-2 facility, which models the circulation circuit WWER-1000 and the SPOT circuit in scale 1:5500 in relation to the power and ensures the hydraulic similarity and real difference of the equipment elevations. These experiments have confirmed the possibility of the long-time passive heat removal in case of the main circulation pipeline rupture and station blackout.

The containment model have been developed and constructed in scale 1:80 for the experimental investigation of the possible influence of the wind on the SPOT effectiveness. The investigations have been performed for the wind speed from 0 to 90 m/s (from the calm atmosphere to hurricane) at the wind direction from 0 to 360 degrees in relation to the reactor building axes. These experiments have shown the absence of the circulation reversal in the exhaust air stacks and have confirmed the design solution correctness, at which the SPOT trains are connected by the common inlet collector and the common outlet collector with the deflectors.

5. CONCLUSION

Broad objectives for advanced nuclear power plants have been documented [7] by the International Atomic Energy Agency. With regard to the safety enhancements, this document states that the plant design should seek to take the maximum, feasible advantage of inherent safety features, and efforts should be made to utilize passive safety systems to the extent that they can be shown as reliable and cost effective as active systems for the same function. Following these recommendations, a reasonable balance of active and passive systems based on the weighing of their advantages and disadvantages with regard to the designated functions, overall plant safety, and construction and operation costs has been found in WWER-1000/V-392 and WWER-640/V-407 designs.

A number of the relatively innovative passive safety means are used in new Russian plant designs with V-392 and V-407 reactors to fulfil the fundamental safety functions, such as reactivity control, fuel cooling and radioactivity confinement. Their implementation allowed to significantly increase the power plant safety in terms of the expected severe core damage and excessive radioactivity release frequencies. For example, the estimated core melt frequency of WWER-1000/V-392 is three orders of magnitude less than for the operating power units with WWER-1000/V-320 reactor.

As the sufficient operational experience for some passive systems and components is absent, the extensive experimental investigation and tests have been carried out or planned to prove the functioning of these systems under plant conditions. In particular, the experiments are already performed for residual heat removal system (SPOT), quick boron supply system, the system to keep the rarefied atmosphere in the containment wall’s gap, emergency core cooling system, and some others.

The experimental investigations and tests performed have confirmed the design functioning of the passive safety means proposed and allowed to optimize their general configuration and initiating signals. These investigations have also created the necessary experimental data base for the modeling of the passive safety means by the system thermohydraulic codes. Further investigations are being planned for additional verification of the passive safety systems and for the optimization of their design.

Two-phase natural circulation is the normal mode of coolant circulation in many advanced designs of BWRs. Typical examples are the simplified boiling water reactor (SBWR) and the advanced heavy water reactor (AHWR). Two-phase natural circulation is also the predicted mode of coolant circulation in most of the current designs of water cooled reactors during small break LOCA. Hence two configurations of loops as shown in Fig 2, one with a heat exchanger and another with a steam-water separator, are of interest to nuclear industry. For the development of the scaling relationships, only uniform diameter loops with adiabatic pipes operating without any inlet subcooling are considered here. The scaling laws for two — phase natural circulation can be obtained from the conservation equations of mass, momentum and energy applicable for homogeneous equilibrium flow.

dp 1 dW

dt A ds

d(ph) 1 d(Wh) _ 4q dt A ds D

H H

d(ph) 1 d(Wh) _ 0 dt A ds

d(rh) + ^ 0Wh) = _ 4U(Tsat — Ts)

dt A ds D

Ws, hs

Steam

FIG. 2a (left) Uniform diameter two phase loop with condenser. FIG. 2b (right) Uniform diameter two phase long with separator.

For the loop with the steam separator, instead of Eq. (23c) a point model of mass and energy equations is used, which assumes complete separation and thermal mixing in the separator. The feed water flow rate is assumed to be controlled to match the steam production rate. Integrating the momentum equation over the loop we get,

These equations are nondimensionalised using the following substitutions

Where tr=VtpL/Wss and Apss (always taken as positive) and Ahss are respectively the steady state density and enthalpy differences across the heated section. The nondimensional equations obtained are

with (fLO)2=1 for gingle-phage region and li = Li/Lt. In writing the above equation, the local preggure loggeg were congidered to be negligible.

4.1. Steady state flow

The gteady gtate governing equationg (all temporal derivativeg = 0 and wgg = 1) are da

dS

The LHS of the momentum equation becomeg zero for the uniform diameter loop ghown in Fig. 2a. For non-uniform diameter loopg with a tall riger (H in Fig. 2b) itg value becomeg negligible compared to the friction preggure drop. Hence,

% jP’dz = JNl RegA 2Rebg

For a loop with horizontal heater and cooler, it can be ghown that the <j" p * dZ = 1. Hence,

Where C=(2/p)r and r=1/(2-b). For laminar flow (b=1, p=64), C=0.03125 and r=1 and for turbulent flow (b=0.25 and p=0.316), C=2.87 and r=0.5714. The game relationghip can be obtained for the loop with vertical heater and cooler if we uge the elevation difference between the thermal centreg of the cooler and heater in place of the loop height in Grm. In

case, the value of the LHS in equation (33) is significant, then an explicit equation for Ress cannot be obtained. Instead, a polynomial in Ress can be obtained as

Where xex is the heater exit quality. Since xex=Q/Wsshfg without inlet subcooling, Eq. (36) can be rewritten as

For laminar flow, b=1, an explicit expression for Ress can be obtained. For turbulent flow, Ress can be numerically calculated from the above polynomial.

It can be shown that the Apss used in the Grm can be estimated as aexpfg+ (pL — pin). For no inlet subcooling, the density at the inlet is pL, then Apss=aexpfg where aex is the void fraction at the exit of the heater. If we use the homogeneous model for the evaluation of the heater exit void fraction, then aex=pL/[(pL-pG)+WsshfgpG/Q]. For the evaluation of the (fLO) , several homogeneous models are reported in the literature. If we use Owens (1961) model, then (fLO) =pL/ptp=1/[1-aex(1-pG/pL)]. Thus knowing the heater exit void fraction both Grm and NG can be evaluated. Although the homogeneous flow assumption was used in the derivation of Eq. (35), the difference in the velocities of the two-phases can be accounted by selecting appropriate models for aex and (fLO) .

Conclusions on the conditions in the reactor sump can in principal be drawn from the experi­mental results by two means, by applying scaling laws, and by applying CFD methods. Scaling laws were applied by Knebel & Muller (1997) to determine from the 2d experiments that in the first days the temperatures in the water above the core melt will be large enough to achieve boiling. As a consequence, the investigations in SUCOS are only applicable for the long term cooling after about ten days. A condition for applying scaling laws is, that the physical trans­port phenomena are of similar relevance in the model experiment and in the reactor sump. The analyses of SUCOS-2D in Carteciano et al. (1999) have shown that specific 3d effects, like the
feed water pipes to the coolers, the bolts, and screws going through the fluid domain strongly affect the experimental results. In the small cross section of the SUCOS-2D slab geometry these structures become more important for the heat transfer and for the flow resistance than in the huge reactor sump. Thus, the results of this experiment cannot be used for scaling up to predict accurately the temperatures in the reactor sump. In contrast, the first numerical cases, which did not record those structures, are more feasible as a basis for this extrapolation. The experimental results of SUCOS-3D could be a better basis because there the cross sections oc­cupied by those structures are of relatively less importance, but there the above mentioned strange flow distribution was found which is specific for this experiment and not for the reactor sump.

Scaling up by CFD-tools is also not free of serious problems in this context. On one hand the SUCOS tests could successfully be interpreted with the FLUTAN code. Such CFD codes have the flexibility to tackle all the experiment or reactor specific structures. On the other hand the model experiments showed laminar flows, whereas the reactor sump will have turbulent flow conditions (Knebel & Muller 1997). Hence additional investigations are necessary by experi­ments of similar flow types to validate the turbulence models and boundary conditions used with the turbulence models. Purely buoyant flows are currently a challenge for any turbulence model, see e. g. Hanjalic (1994, 1999). Also the TMBF, which is explicitly developed for buoy­ant flows, has up to now only been validated for two-dimensional forced, mixed, and natural convection (Carteciano et al. 1997, 1999b). An additional feature, which is inherent to most purely buoyant flows, is its local time dependence. The cold plumes plunging down through the chimney are a low frequent phenomenon which was also not completely filtered out in the experiment by time-averaging over two minutes. This causes the wall heat flux on the copper plate to change in time, Fig. 4. And also the hot plumes rising from the surface of the copper plate are no stationary phenomenon. The surface temperature on the copper plate, Fig. 10, is very straggly as it is typically found in Rayleigh-Benard convection in which the hot plumes rise mainly from the knots of those straggly structures (Worner & Grotzbach 1997).

As a consequence of the 3d and time-dependent nature of buoyant flows and of the current status of development of the standard turbulence models, there is currently no way to achieve reliable results by standard models on the temperature fields in the reactor sump. The only way which is nowadays often considered to give a better solution for 3d time-dependent flows is to apply Large Eddy Simulation methods (LES). Indeed, there exist already several applications of LES to reactor typical flows; for an overview see Grotzbach & Worner (1999). These show the tremendous potential of the LES method and that its possibilities are going far beyond those of standard Reynolds averaged turbulence models. The main problems which need to be solved for LES methods in this context are e. g. the development of more universal subgrid scale models, of boundary conditions for buoyant flows, and of numerical methods in commer­cial codes that fulfill LES requirements.

5. CONCLUSIONS

The former numerical interpretation of SUCOS-2D experiments with the FLUTAN computer code showed that good agreement between experiment and calculation can be achieved when local thermal disturbances like the feed water pipes to the coolers are recorded in the simula­tion. Here, the single phase natural convection experiment SUCOS-3D is interpreted. The nu­merical results confirm, natural convection in large pools, in which pressure drops are negligi­ble, is very sensitive against small disturbances. Here, some other discrepancies are found by analyzing the experimental results from SUCOS-3D: In these experiments the maximum tem­perature did not occur far above the horizontal coolers as in SUCOS-2D, but below the tilted roof, and the heat fluxes over the horizontal and vertical coolers were not homogeneously dis­tributed. Comparable pool temperatures could only be achieved numerically by using the measured heat fluxes at the coolers instead of temperature boundary conditions, but the calcu­lated flow pattern was still different. One explanation for this strange result is supported by the experimental data, this is to postulate that one of the horizontal coolers was slightly tilted against the main flow direction. Additional numerical investigations indeed show that a slope of only one percent would explain the experimental flow field. From this problem of the ex­periment one can learn how to improve this sump cooling concept: Foreseeing a small slope of the horizontal coolers downwards in the expected flow direction would stabilize the flow and would drastically increase the efficiency of the horizontal coolers.

Based on the performed numerical investigations it can be concluded that a transformation of the experimental results from SUCOS-2D and -3D to reactor sump conditions by means of scaling laws is questionable. Such transformation will only be possible by applying well vali­dated CFD-Codes and experienced code users with a sound physical and engineering back­ground. Current standard turbulence models form the working basis for engineers, but they can only be used for approximate predictions, because the statistical models fail for buoyant flows which are locally time-dependent. The only more reliable solution could come from adequate Large Eddy Simulation methods and LES-suitable codes for complex geometries.

[1] .

From the modelling point of view, two types of instabilities are important. These are the static and dynamic instability. Static instability can be fully described by the steady state governing equations. Examples of this type of instability are Ledinegg (flow excursion), flow pattern transition instability, flashing instability, etc. For the dynamic instability, the feedback effects are important and hence it requires the solution of the time dependent conservation equations of mass, momentum and energy. Typical example is the density wave oscillations observed in two-phase systems. Often, thermalhydraulic systems exhibit compound instability that is usually a combination of two mechanisms. Here, the instability may begin with one of the known mechanism for static instability that may trigger some secondary mechanisms resulting in an oscillatory behaviour. Typical examples are parallel channel instability and coupled neutronic-thermalhydraulic instability of BWRs.

In generally, a two-phase natural circulation system can be designed to avoid all the different types of instability discussed above. This is done by carrying out a stability analysis, which is usually phenomena-specific and therefore cannot be described in detail in a brief note as the present one. However, the objectives of most stability analyses can be listed as:

(a) To generate stability maps which delineate the stable and unstable zones of operation;

(b) To specify adequate stability margin for design purposes; and

(c) To predict the nature of the unstable oscillatory behaviour of flow, power and temperature (in time domain) if the system passes through an unstable operating zone during power/pressure raising.

Most static instability analyses are carried out with the prime objective of generating a stability map. For the dynamic analysis, however, all the three objectives become important. Both the linear and nonlinear methods are used for the analysis of this type of instability. Mathematically, the dynamic behaviour of any system can be considered to be linear for small perturbations around the steady state operating condition. An analytical solution is sought for the perturbed equations to obtain the characteristic equation for the stability of the system. The characteristic equation is solved numerically to generate the stability map (locus of all neutrally stable points) of the system. The computer codes based on the linear stability method (e. g. RAMONA, NUFREQ, etc.) are very useful to predict the stability maps without consuming much computer time. However, it is beyond its capability to predict the non-linear effects that come into play after the neutral threshold (to achieve the third objective). Hence, direct numerical solution of the non-linear governing equations is carried out using finite difference methods. An additional objective of the non-linear computer codes is to provide an understanding of the basic physical mechanisms involved in BWR behaviour beyond neutral stability. In principle, system codes like RELAP5, ATHLET, CATHARE, etc. can be used for this purpose, although this is not the current practice (generally). If the system codes are used, this will require the use of qualified fine nodalisation as those used for the normal transient analysis (normally a coarse nodalisation) may not even predict the existence of instability (as the numerical solution technique adopted in these codes are very robust which introduce numerical damping through artificial/numerical diffusion).

For simulating coupled neutronic-thermalhydraulic instability, usually a point kinetics model is used in addition to a fuel heat transfer model (which basically solves the conduction equation in the fuel, fuel-clad gap and clad to obtain the feedback effects of fuel). Point kinetics equations are first order ordinary differential equations that can be easily integrated by using a conventional numerical technique. However, the problem is the introduction of feedback constants in them, which are spatially averaged over the whole core. So uncertainty is therefore introduced at this level in addition to the inaccuracy in neglecting the changes in core power distribution. Similarly, the way in which the neutronics and the thermal hydraulics interact during the calculation has an important effect on the capability to simulate multi­dimensional core dynamics. Grouping of core channels in thermal hydraulic regions having similar geometric and operating parameters can have strong influence on the out-of-phase mode oscillations.

The other common design is when there is two-phase boiling flow with direct steam to the turbine. When steam is fed directly to the turbine, again there is a direct relation of the maximum power output to the secondary (turbine stop valve) pressure. The maximum heat removal is due to the circulation of two-phase mixture when the downcomer is liquid and the core is two-phase and heat removal is then totally evaporative

The maximum power output in a natural circulation boiling system without a HX is not derived on the basis when the natural circulation driving head is equal to the two-phase losses. Instead, as we have noted above, the ultimate or maximum power output is set by the onset of flow instability and hence subsequent CHF.

The form of the natural circulation line has been found above to be Np/Ns~ constant for a given downcomer head to core height ratio, L*. The limiting maximum power solutions for the unstable case are, from Equation (12) and (13) and the data comparisons in Figure 3,

Np/N~ 3 (19)

with a residual dependency on the loss coefficients. When there is a natural circulation loop, then the intersection of the stability region with the natural circulation flow is very nearly, for typical design values, when,

Np/Ns~ 2 (20)

By comparing a wide range of parallel or multichannel instability data at high pressure (5MPa) on a Np versus Ns plot, the data do indeed group around a line given by N*= 3+10/Ns (Rohatgi and Duffey, 1994).

Now the maximum two-phase flow in the whole natural circulation loop at intermediate inventories, is given by Equation 15:

1/3

Combining the equations for the maximum flowrate at the instability limiting case, we find the hypothetical maximum core power at this maximum flowrate to be:

This result clearly shows to optimize the design power output, the minimum loss coefficients, and the maximum elevation (driving) head and flow area should be obtained.

The maximum unstable channel power is less than that ultimately obtainable from the boiling flow in a channel which is given by:

which is typically of order 10MW(t). It is therefore important to compare the various limits to see which may be the design constraint.

Results are related to experimental facilities 2, 3, 5 and 6 listed in Table I and are summarized in Table V, ref. [13]. Calculations related to PWR-1 of Table II are also discussed in Ref. [13]. These are not considered hereafter owing to a large ‘mass error’ that characterizes the achieved results as documented in the same Ref. [13]. Qualified nodalisations of the ITF, suitable for the Relap5/mod 2 code, were used. The qualification came from the simulation of the NC tests performed in the considered ITF, e. g. ref. [12], and from the demonstration that calculated results adequately reproduce the available experimental values.

The main results of the study can be summarized as follows, ref. [13]:

• A uniform increase of NC flowrate with core power is calculated, until core power achieves values around 40% of the nominal value. Further increases of core power do not cause proportional increases in core flow.

• Oscillatory flows are calculated for core power larger than 40% in Bethsy and Lobi.

• The primary mass inventory decrease occurs via the pressurizer relief valve that is assumed to open and to close in order to keep constant the system pressure.

• PWR cores, in the actual configuration can operate in NC conditions with power up to about 15% the nominal value.

• The largest facilities are designed to operate at low core power (ITF design finalized to the simulation of small break LOCA). This may explain the small value, in terms of % core power, at which dryout occurs.

• Neglecting the Lstf case, up to 70% core power can be removed by NC before experiencing dryout. This can be assumed as the thermalhydraulic limit for system (not any more PWR) operation in NC.

The main problem in the first two test series in PACTEL was the rapid condensation in the CMT, which temporarily stopped the ECC water injection. The condensation took place when the saturated water layer in the CMT broke down, and the steam got into contact with cold ECC water. This happened when water flowed to the CMT after the level in the tank has already dropped. In the tests the water level in the CMT started to drop almost immediately after the opening of the break and the period of single phase natural circulation through the cold leg pressure balancing line to the CMT was very short. Due to this the saturated water layer in the CMT separating cold water from steam remained thin. For the third series, a flow distributor called sparger has been added to the CMT. The purpose of the sparger is to diminish possibilities of rapid condensation in the CMT. Further, the experiment procedure included filling of the PBL with hot water before the break opening. The CMT and the IL was full of cold water. All the earlier experiments begun with cold water in the CMT, PBL and IL.

The main objective of the third series was to investigate the influences of break size and the removal of the pressurizer PBL on the CMT behaviour in cold leg SBLOCA’s. The tests run with four different break sizes (from 1 to 5 mm in diameter) including reproducibility studies. Like in the earlier tests the main objective of the tests was to simulate the PSIS behaviour and not to try to simulate any proposed ALWR reactor concept in particular.

In all experiments the CMT ran as planned. There were no problems with rapid condensation in the CMT, such as was seen in the earlier passive safety injection experiments in PACTEL. The main reason was the new CMT arrangement, with a flow distributor (sparger) installed to the CMT. The sparger spread incoming flow to the CMT horizontally, and the breakdown of the saturated water layer due to incoming water was not possible. The hot liquid layer between the steam and cold water in the CMT remained stable, even in the experiments where the hot liquid layer was less than 5 cm thick.

Passive Initiators (PI) are small heat exchangers outside the Reactor Pressure Vessel (RPV); one line of this heat exchanger is connected to the RPV well above the normal water level and the other line well below the water level. Both lines are always open to the RPV. The secondary side of the Passive Initiators is filled with water. In case of a substantial water level decrease in the RPV steam will condense inside the PI and heat up the secondary side resulting in a pressure build-up. This can be used for control purposes. In Fig. 13a the four designs tested are schematically shown. In Fig. 13b the pressure build-up on the secondary side and the outflow is given for water and ethylacetate (ESEE), which has a lower boiling point than water.

a)

FIG: 13. Designs (a) and experimental results (b) of the different passive initiators.