1.1. THE ROLE OF NATURAL CIRCULATION SYSTEMS

The role of the natural circulation systems should be considered in the general context of implementation of passive safety systems into new nuclear power plant designs. It may be noted that many systems and equipment proposed in future reactor concepts include natural circulation phenomena as the main mechanism that determines the passivity of the fulfilment of a designated function. Naturally driven systems have been used in the existing reactors both to remove the core heat under normal operation conditions including full power operation (like the Russian boiling water reactor VK-50 or the Dodewaard reactor in the Netherlands) and to fulfil some safety functions (like the passive part of the emergency core cooling system [2] in PWRs). So, there is a good technical basis and some operational experience to use passive systems in new reactor concepts.

With respect to plant safety, application of passive systems/components is intended to simplify the safety systems and to improve their reliability, to mitigate the effect of human errors and equipment failures, and to provide increased time to enable the operators to prevent or mitigate severe accidents. Natural circulation systems typically do not require or accommodate repair or maintenance work during power operation. Therefore, a reduced number of safety system trains may be needed to perform the designated safety function with the required reliability. An IAEA conference [1] included discussions on the safety of future plants, and noted that the use of passive safety systems is a desirable method of achieving simplification and increasing the reliability of the performance of essential safety functions, and should be used wherever appropriate.

However, natural circulation systems have their own advantages and drawbacks in comparison with forced flow systems, both in the areas of plant safety and plant economics. Passivity itself does not mean that a natural circulation system should be automatically considered as more reliable with regard to fulfilment of the designated safety function. Therefore a reasonable balance of traditional systems and new passive means is adopted in many future reactor concepts as a possible way to improve safety and public acceptability of nuclear power, and at the same time to keep nuclear power competitive with conventional power technologies. Many considerations govern this balance and define the final design decision, such as:

(a) Application of passive systems should reduce the number of components, and yield design simplification, so that the number and complexities of safety actions can be reduced;

(b) Passive means should be taken, to the extent possible, from similar ones having operational experience at power plants or elsewhere, so that the efforts needed to demonstrate the reliability and licensability are not too large;

(c) Passive systems actuation should be more reliable than for an active system providing the same function; otherwise the increase of the system reliability projected by implementation of the passive system may be lost;

(d) Passive systems should eliminate the need for short-term operator actions during accidents

(e) Passive systems should minimise dependence upon off-site power, moving parts, and control system actions for normal operation as well as during design basis and beyond design basis accidents;

(f) Passive systems should reduce the construction, operation and maintenance costs.

The main drawbacks of natural circulation systems include lower driving forces and less possibility to alter the course of an accident if something undesirable happens (i. e. less operational flexibility). In particular, in certain conditions where forceful or rapid actions are required, active systems may be more suitable to carry out certain safety functions. Also, load follow operation may be limited in the reactors based on natural circulation of the primary coolant. Therefore, in some new reactor designs originally designed for natural circulation, forced circulation flow (by pumps) has been introduced to allow for a better load follow capability and to increase the reactor rated power. Scaling for natural circulation systems is more difficult than for active systems. Therefore usage of experimental or operational data obtained for a system with a size that differs from that of the system being designed may not be appropriate. Due to low driving forces, the operation of a natural circulation system may be adversely affected by small variations in thermal-hydraulic conditions. The lower driving forces might also lead to quite large equipment, and this factor may reduce the cost savings projected from elimination or downsizing of active components. Besides, larger components may cause additional difficulties in seismic qualification on some plant sites.

The design decisions with regard to balancing active/passive features may also depend upon the functions assigned to the given system. In particular, the system having an important role in the mitigation of severe accident consequences that is located in a potentially contaminated area (e. g. the part of the containment cooling system which is located inside the containment) could be designed to be as passive as reasonably achievable. This is because of the difficulty or even impossibility of access to such areas and because passive components may not require maintenance even during long-term operation.

All the above aspects are being taken into account by nuclear power plant designers, and as a result, both novel and more or less proven passive systems and features are proposed in many new designs [2]. Some designs have only added a few passive components to the traditional systems. Some other designs make use of the passive systems/components and natural circulation phenomena for power production in normal operation, or to fulfil a number of safety functions, intended to prevent severe accidents and mitigate their consequences.

A high-pressure natural convection loop (CAPCN) was constructed and operated to produce data in order to verify the thermal hydraulic tools used to design the CAREM integral LWR design, mainly the dynamical response. This is accomplished by the validation of the calculation procedures and codes for the rig working in states that are very close to the operating states of CAREM reactor. CAPCN resembles CAREM in the primary loop and steam generators, while the secondary loop is designed just to produce adequate boundary conditions for the heat exchanger. Water enters the heated section from the lower plenum. The nuclear core is simulated by electric heaters. The heated water flows up through the riser to the upper plenum where a liquid-vapor interphase exists. The water exits this plenum through an outer volume in contact with the steam generator. The steam generator has two coils, once through, secondary inside. The sub-cooled water flows down through a downcomer or cold leg to the lower plenum. Natural circulation flow may be regulated by a valve in the cold leg and a bypass to the bottom of the riser. This rig was constructed according to ASME for the following primary parameters: 150 bar and 340°C. The primary loop may operate in saturated or sub-cooled regimes, with a heating power up to 300kW and different hydraulic resistance. The circuit configuration allows the study of stationary states similar to CAREM conditions of pressure, specific flow and enthalpy. Height was kept in a 1:1 scale.

Many experiments were performed in order to investigate the thermal-hydraulic response of the system in conditions similar to CAREM operational states. The influence of different parameters like vapor dome volume, hydraulic resistance and dome nitrogen pressure was studied. Perturbations in the thermal power, heat removal and pressure relief were applied. The dynamic responses at low pressure and temperatures, and with control feedback loops were also studied. It was observed that around the operating point self-pressurized natural circulation was very stable, even with important deviation on the relevant parameters.

The data obtained is being used to test numerical procedures and codes. A sensitive analysis was performed and most sensitive parameters were identified. A representative group of transients were selected, in order to check computer models. Simulation models are in current development against a reference transient, without adjustment. When this contrast is clear, models will be compared against the representative group of transients. The information on specific models should be incorporated into CAREM modelling.

It was demonstrated that the change from stratified conditions to natural circulation (and vice versa) is essential for many passive safety systems of the SWR 1000. It would be useful if these flow phenomena could be studied by means of computer code analyses. But in many cases, suitable computer codes are not yet available and development of such codes to serve as a basis for actual decision-making would be too time-intensive.

Designers of new nuclear power plant concepts normally cannot wait until suitable computer codes are ready in order to find solutions to problems. Basic decisions are often made using only a calculator and sound engineering judgement. At the moment it is still easier for the designer to prove its preliminary decisions by way of experimental testing rather than calculational analyses. One can only hope that such computer tools will be available prior to entering the plant design licensing period and that computed results are in good agreement with the results of experimental testing.

PCCAC-2D is a two dimension computer code which has be already used in the design of AC600/1000 passive containment cooling system. PCCAC-3D is a three dimension computer code which will be finished by the end of next September. PCCAC-2D, 3D can be used to predict the pressure and temperature of mixing gas inside the containment following the accident of primary pipe rupture or main steam line rupture. The heat removal characteristics from inside containment to atmosphere through the water film on the out surface of steel shell and natural circulation flow of air can be also simulated and calculated by PCCAC-2D or PCCAC-3D.

It follows from preliminary calculation results that the maximum pressure 0.37 Mpa of mixing gas in the containment will occur in 1123S after double ended rupture accident of cold leg pipe. The maximum temperature 138°C will occur in 14S after the accident.

FIG. 13. Flowrate distribution in containment

The maximum pressure 0.34MPa of mixing gas in the containment will occur in 16.8S after double ended rupture accident of hot leg pipe. The maximum temperature 143°C will occur in

6.85 after the accident. The maximum pressure 0.387MPa of mixing gas in the containment will occur in 395.3S after main steam line rupture accident. The maximum temperature 153°C will occur in 3S after the accident.

The flow fields of coolant in the containment indicate that main flow vortex will be established at 0.9S following accident, will be in equilibrium at 5.2S and will be destroyed at

24.85 because of the end of blow down. The temperature fields of coolant in the containment show that high temperature area is in around and above the break and low temperature area is in the bottom near steel shell at the beginning of the accident. Then the temperature fields become uniform after main flow vortex equilibrium. The natural circulation mass flowrate of air in the channel between steel shell and concrete shell of containment varies with time and reaches maximum 168 kg/s at 8000S. After 72 hours, the spray water will be ended and the natural circulation mass flowrate will reach 120 kg/s. It can be noted from the preliminary analysis and calculation that AC600 passive containment cooling system is able to remove decay heat of reactor from inside to outside of containment.

3.1.

Station blackout accident calculation for AC600/1000

ERHRAC is a special computer code for AC600/1000 design. ERHRAC can be used to calculate the natural circulation flowrates of three cycles (primary coolant cycle, secondary side cycle of SG and air flow cycle). The links connecting those three cycles are the steam generator and air cooler, establishing a tandem system.

In the unlikely event of a station black out accident, the flowrate through reactor core rapidly reduces. The changeover of coolant flow in primary coolant loop from the forced circulation
to the natural circulation is initiated automatically when main pump coast down is ended. In the secondary side loop, from steam generator the steam passes on to an air cooler, in which its heat is transferred to the air and steam is condensed to water. Then condensed water returns back to the steam generator by gravity to establish a natural circulation flow cycle.

Natural circulation flow rate is: 4% rated flowrate for each primary coolant loop and 3% rated flowrate for each secondary side loop of SG respectively. Air flowrate is about 290Kg/s for each air cooler.

Tre(S)

4.

Due to obvious deficiencies in the calculations with previous APROS versions a new solution model for the thermal stratification of APROS code was developed [23]. The old method used upwind solution for the enthalpy. Due to numeric diffusion the code lost information about the stratified layer. The new higher order numeric method uses information from three consecutive nodes to solve the transported liquid enthalpy. The new enthalpy solution contains a special weight function, which is calculated from liquid enthalpies of the three nodes. The experiments GDE-41 and GDE-43 were recalculated with the new model (Fig. 4). The model was also tested separately with a standalone PSIS [24]. The calculation results were good. The new model eliminated significantly the numerical diffusion and restricted the spreading of the thermal front.

FIG. 4. Temperature distribution in the GDE-43 experiment vs. APROS calculation with the old and new models at time 4000 s.

Lappeenranta University of Technology in co-operation with VTT Energy has performed totally 24 experiments to investigate Passive Safety Injection system with PACTEL facility during past eight years. The experiments in PACTEL have provided valuable information about PSIS behaviour in SBLOCA’s. The PSIS system worked as planned when the CMT was equipped with flow distributor (sparger). The main source of disturbances was condensation, which took place when stem met cold water. The experiment data has been and will be used to further develop Finnish thermal hydraulic analyses code APROS. The computer simulations have reproduced the measured transient behavior with good accuracy despite of smoothing of the thermal front generated by numerical diffusion. The code has been improved by developing a new solution model for the thermal stratification, which largely eliminated the effect of numerical diffusion.

[4]

In general, thermal hydraulic modelling of nuclear reactor systems is based on the one­dimensional approach. Thermal hydraulic modelling of the steady state, transient and stability behaviour of two-phase natural circulation systems is no exception to this general approach. System codes have reached a highly developed modelling status and a wide acceptance. They can reproduce accurately enough most of the existing safety related steady state and transient experiments, so far as the dominant physical mechanisms are known, as adequate models are included in the codes, and so far as the dominant phenomena are also understood by the code users. Thus, these codes are an excellent tool to analyse in a parametric way the dynamics and interrelation of a larger number of components in complex systems with different physics involved. However, the use of multi-dimensional modelling of two-phase natural convection in large vessels (including the calandria vessel of pressure tube type heavy water reactors), sumps or plenums of nuclear reactors is essential as the flow in these cases are multi­dimensional in nature.

Some of the most commonly adopted thermal hydraulic models applicable for the 1-D and multi-dimensional analyses are briefly described below:

In designing heat removal systems, we often rely on both single and two-phase natural circulation. The fluid heats and expands, and may boil and the two-phase flow generates an increased pressure drop. The inlet flow is derived from a pump, or a downcomer with a hydrostatic head, so that when there are many channels in parallel, as in a reactor, heat exchanger or condenser, the boundary condition is a constant pressure drop.

We can derive a formulation for the natural circulation flow, which includes the subcooled length and the effect of flow on the loss coefficients; is consistent with homogeneous stability analysis, and extends previous work. This result can then be coupled with the instability treatment to determine the natural-circulation limit on instability.

Consider any natural-circulation system as a loop or U-tube, where the driving force for the flow is due to the gravitational head difference between the water in the downcomer and that in the boiling channel and riser. In the steady state, which is sufficient for our discussion, this head difference is balanced by the pressure losses incurred by the flow and venting of the boiling two-phase flow, so in a natural circulation loop, the pressure drop boundary condition is zero around the loop. Thus, the net gravitational hydrostatic head is exactly balanced by the friction and form losses in the single and two-phase flow. Consider a uniformly heated channel of equivalent hydraulic diameter, D, with a subcooled inlet flow rate, G. To a good approximation, we take the flow as steady and incompressible. The momentum equation in the vertically inclined channel is expressed in the usual way in Equation (1).

og з(g2/ p) 3t 3z	f"(f+k )G 2/p )+ )p cosq
(1)
By integrating the momentum (Equation (1)) around the loop, we have the sum of the pressure drop components (Rohatgi and Duffey, 1994),
L-Zx tf [A; k, 2ZH Pt Д* J 21 Pt aJ s{zjPt + (g — zj)ps — ZxPt — t fe — iz — 4W} =0 Here we have where, by definition, the subcooled length, is for Xin< 0,
(2)
hfpXinAGL
(3)
Q
and, G, pme, ki, ke, Xin, Xe and Xa are mass flux, mean mixture density in the boiling section, mixture density at the exit, inlet and exit loss coefficients, inlet and exit qualities, and mean quality of the boiling region, respectively. We have defined mean channel and exit densities given by,
	Vg + Xa Vg Vp + Xe vp Vg + Xa Vg
Equation (2) is now cast in phase-change and subcooling number form (Rohatgi and Duffey, 1994). The resulting expression is, after much algebra and neglecting the small terms with the vapor to liquid density ratio:
where,
(4)
(4.1)
Friction Number Nfr
(4.2)
JL_ 2D
(4.3)

Phase Change (Zuber) Number N? — ———

AGhfgPg

and

Downcomer Number

(4.5)

Equation (4) is quadratic in Np and Ns and the roots provide the possible set of solutions for all liquid single and boiling two-phase flow. We can easily obtain the power required for given flow rate and core inlet subcooling from Equation (4), and vice versa. The Froude number will vary with the flow direction; the downcomer dimension, L*, is a design parameter; with the non-dimensional length being the ratio of the channel to downcomer heights L*, and typical values can be taken for the loss coefficients.

The natural circulation map describes the two allowed power and flow relationships for a given heated channel length to downcomer head ratio, L*, for given loss coefficients and Froude number. The two solutions correspond to single (liquid only) and two-phase flow for the same downcomer number at a given subcooling. The solutions of interest in a boiling loop are in the two-phase region. The intersection at zero subcooling number, for a large exit loss, is given very nearly by,

Np (Ns=0) = (2Nf /keNfrL*) -1 (5)

It is clear that the dependency between Np and Ns is nearly linear for a given set of loss coefficients and downcomer height, L*. Thus, the power to flow ratio is uniquely defined, and the curves describe the allowed flow states for a particular set of design values for the loss coefficients and relative downcomer height, L*. It is evident the flow can be bistable about the saturated line.

Since the flow is effectively bistable (double valued) for a given set of conditions, theoretically one can have single or two phase flow for the same pressure drop, an effect that is often observed during flow reversals and instabilities too. The effect of the relative driving head is shown by the following figure, where for a given set of loss coefficients the solutions to Equation (4) is shown for variations in L*, the relative heated length to downcomer height.

Now variations in downcomer height also can correspond to variations in total loop fluid inventory. The effect of changing total loop inventory on flowrate is determined by the shifts in the void in the boiling regions. So initially decreasing the inventory leads to an increasing flowrate, as the driving head increases since there is more boiling on the “hot” side (the riser or heated section). Eventually as the void increases, the flow reaches a maximum and then decreases as the driving head falls in the “cold leg” or downcomer.

Physically, the argument can be shown in more detail, and to illustrate the effects, Equation (2) can be solved assuming steady flow, taking a separator elevation equal to the downcomer height, and mean densities taken for the two-phase regions in the hot (riser, R) and cold (downcomer, D) regions. Lumping the loss terms together into an “effective” loss coefficient (Duffey and Sursock, 1987) it is possible to derive simple equations that describe the general behavior of the flow (and agree with the trends in the data). Thus, the loop flowrate is approximately given by:

where,

is the maximum possible flow if it were all liquid, and Ф is a two-phase friction multiplier, and I is the fractional liquid inventory in the loop, W1 the single-phase flowrate, VR is a measure of the hot side fractional volume to the total system volume, and Y is a simple interpolation function for a smooth flow transition. Since it is evident that the two-phase mass flow goes as (1-I) , we could maximize the flow in a design by adjusting the values of the losses, relative volumes and heated lengths.

It is well known that a two-phase natural circulation flow can exhibit instabilities over certain regimes (Gulshani et al, 1995) so we now examine that phenomenon and give some relevant analytical results.

The decreasing of NC driving forces, the small temperature difference across U-Tubes of steam generators, and the occurrence of the CounterCurrent Flow Limiting Phenomenon (CCFL) at the entrance of U-tubes are at the origin of wide system oscillations of core inlet flowrate, e. g. Ref. [10]. The phenomenon has been investigated in Refs. [5] and [11], based on a natural circulation experiment performed in Lobi facility, Ref. [12]. Evidence of the phenomenon has been found also in other facilities. At mass inventories of the primary system around 70% of the nominal value, the efficiency of the condensation heat transfer across U-Tubes causes the release of almost all core thermal power in the ascending side of U-Tubes. Liquid level builds up and is prevented to drain down by the steam-liquid mixture velocity at the tube entrance, i. e. the CCFL condition occurs. Therefore, liquid level rises in the U-Tubes till reaching the top. During this period, typical duration of the order of 10 s, flowrate at core inlet is close to zero and core boil-off occurs. Once the liquid level reaches the upper bend of U-Tubes, the siphon effect occurs and causes the emptying of the ascending side of U-Tubes and the re-establishment of core inlet flowrate. A new cycle starts. The phenomenon is made more complex by the interaction of the several thousands of U-Tubes that constitute a SG tube bundle. Different groups of tubes may stay at a different stage of the oscillation at the same time, also causing flow reversal in the tube bundle. Suitable core cooling still can be achieved in these conditions.

J. D. Jackson, J. Li

University of Manchester, United Kingdom

Abstract. A fundamental study is reported of heat transfer from a vertical heated tube to air which is induced naturally upwards through it by the action of buoyancy. Measurements of local heat transfer coefficient were made using a specially designed computer-controlled power supply and measurement system for conditions of uniform wall temperature and uniform wall heat flux. The effectiveness of heat transfer proved to be much lower than for conditions of forced convection. It was found that the results could be correlated satisfactorily when presented in terms of dimensionless parameters similar to those used for free convection heat transfer from vertical surfaces provided that the heat transfer coefficients were evaluated using local fluid bulk temperature calculated utilising the measured values of flow rate induced through the system. Additional experiments were performed with pumped flow. These covered the entire mixed convection region. It was found that the data for naturally-induced flow mapped onto the pumped flow data when presented in terms of Nusselt number ratio (mixed to forced) and buoyancy parameter. Computational simulations of the experiments were performed using an advanced computer code which incorporated a buoyancy-influenced, variable property, developing wall shear flow formulation and a low Reynolds number k~e turbulence model. These reproduced observed behaviour quite well.

1. INTRODUCTION

Convective heat transfer from the outside surface of the steel containment vessel of a pressurised water reactor to a flow of air induced upwards by buoyancy through the space between the vessel and an external concrete shell has been proposed as a passive method of removing heat from the containment following a severe accident. Whilst there is no doubt that conditions of turbulent flow could be produced by this means, it is probable that the effectiveness of the heat transfer process would be poor. In view of the limited flow rate likely to be achieved, the heat transfer process within the passage will be buoyancy-influenced as well as buoyancy-driven. The mechanism of heat transfer will therefore be mixed free and forced convection. Published work on mixed convection heat transfer in vertical passages has been reviewed in a number of papers (see, for instance, Reference [1]). Some surprising trends have been identified. In the case of upward flow in a heated vertical passage, buoyancy aids the motion but contrary to expectation the values of heat transfer coefficient achieved can be lower than for conditions of turbulent forced convection. This is because the turbulence in the boundary layer is modified through the action of buoyancy with the result that the flow takes on the characteristics of one at lower Reynolds number. Impairment of heat transfer builds up gradually as the buoyancy influence is caused to become stronger, either by increasing the heat loading or reducing the flow rate. Eventually a very sharp reduction in the effectiveness of heat transfer is found to occur. With further increase of buoyancy influence, the process of heat transfer recovers. In contrast, for downward flow in a heated tube buoyancy opposes the motion, turbulence is increased and heat transfer is enhanced. The physical mechanism by which turbulent heat transfer is modified through the action of buoyancy influences in vertical passages was first identified by Hall and Jackson [2]. The contrasting influences on heat transfer for upward and downward flow are well described by a simple semi-empirical model of mixed convection developed by Jackson and Hall [3]. Figure 1 shows the predictions of that model for the case of a uniformly heated tube. These are found to be in good agreement with observed behaviour (see Reference [1]).

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It is of interest to consider whether impairment of heat transfer would be encountered under the conditions likely to be achieved in a buoyancy-driven flow system of the kind which has been proposed for passively cooling a nuclear reactor containment vessel. In this connection, a further matter needs to be considered. Most of the experimental studies of mixed convection reported to date have been carried out with a thermal boundary condition of uniform wall heat flux. However, in the case of a severe accident in a pressurised water reactor, where steam is released from the core into a steel containment vessel and is condensing on its inside surface, the vessel will take up a uniform temperature. Since the nature of the thermal boundary condition could certainly affect the process of heat transfer to the air, there is a need to consider whether the behaviour with uniform wall temperature will be similar to that with uniform wall heat flux.

The study reported here using a non-uniformly heated test section which can operate at uniform temperature was undertaken to clarity these matters. This naturally-induced cooling experiment (NICE) formed part of the DABASCO project funded by the European Commission to provide an experimental data base for containment thermal hydraulic analysis (see Reference [4]).

The experimental facility used in the present study is shown in Figure 2. The test section, which was made from stainless steel tube of inside diameter 72.9 mm and wall thickness 1.63 mm, is suspended vertically in a space of height 15 m. It can be used for experiments with either naturally-induced flow or pumped flow.

In the case of naturally induced flow, the entry box is absent and air passes from the laboratory into the bellmouth intake and upwards through an unheated flow development section of length 3 m and a heated section above it of length 6 m before discharging at the top. Motion occurs simply as a result of heat being applied to the test section. The velocity at inlet is measured using a calibrated hot film probe mounted in the centre of the bellmouth intake.

In the case of pumped flow, air is supplied to the entry box at the bottom of the test section by a blower and then flows upwards through the test section. The mass flow rate can be measured using a metering nozzle installed at the exit. The hot film probe was calibrated against this flow metering nozzle. Wall static pressure tappings are provided at the top and bottom of the heated section. The pressure difference is measured using a high precision electronic micro­manometer.

Numerous thermocouples attached to the outside of the heated length of the test section enable the wall temperature distribution to be measured in detail. The heated length is well lagged on the outside with pre-formed thermal insulation of low thermal conductivity to minimise heat transfer to the surroundings. The small loses which do occur can be accurately accounted for using data from calibration experiments which were performed at the commissioning stage. Heat can be applied to the test section either uniformly or non-uniformly by means of 40 separate, individually-controlled heaters distributed along its length. These were made using proprietary heater cable which was wound tightly around the outside of the stainless steel tube. Electrical power is supplied to the heaters from the mains via variable auto-transformers through the specially designed computer-based power control and measurement system shown in Figure 3. This supplies power to each heater at a rate needed to maintain the tube at a specified temperature at that location and also enables the power to be measured. The power is controlled by allowing a proportion of the half cycles of the incoming AC supply to pass to the heater. For each of the 40 heaters there is a control signal generator and a zero-crossing solid state relay. The former generates a control signal which enables the latter to pass a programmable number of half cycles from the supply to the heater during a specified time interval of 0.16 s. The signal generators are connected to a computer via a PC interface. The computer can write a number to each of the 40 signal generators under the action of software. The power to each heater is controlled by these numbers. Knowing this number, the voltage of the incoming supply and the electrical resistance of the heater, the power generated can be calculated. In operation, the computer reads temperatures using the Intercole data acquisition system. It then compares each wall temperature with a pre-set value, calculates a new number using the PID technique and sends it to the corresponding signal generator. The power supply system can also be operated in such a manner that a specified distribution of heat input is applied to the test section. A Pentium PC connected to a 208 channel data acquisition system is used for signal collection and processing. The software package which is used to drive the data logger and the power control and measurement system was developed and tested ‘in house’ using Visual Basic under the Windows 95 environment.

Initially, commissioning tests were performed on the test facility using the pumped flow arrangement to demonstrate that everything was operating satisfactorily. Friction factors determined from pressure drop measurements made under conditions of isothermal flow through the test section were compared with values calculated using the well-established correlation equation of Blasius for fully developed turbulent flow in a smooth tube. The maximum discrepancy was less than 6%.

Local values of Nusselt number were determined at various locations along the tube from experiments performed under conditions of forced convection with uniform heating. In Figure 4(a) the results are compared with the distribution of Nusselt number calculated using the established empirical correlation equation of Petukhov et al [5]. As can be seen, the agreement is very satisfactory.

Results from a mixed convection experiment with uniform wall heat flux are shown in Figure 4(b). The Reynolds number was about 10,000. Under such conditions heat transfer was found to be strongly impaired due to the influence of buoyancy. It can be seen that the local values of Nusselt number lie well below the curve for forced convection calculated using the Petukhov equation. The observed behaviour is very similar to that found in an earlier study by Li [6] using a uniformly heated test section of similar dimensions (see Jackson and Li [7]).

In Fig. 1, some European test facilities capable to perform thermal-hydraulic tests with passive safety systems are shown. It is evident that the wide spread of power and pressures will allow the testing of the same component in several facilities of different size and thus increasing the confidence in the assessment of the effectiveness of this component. It has to be mentioned that the PANDA test facility has a larger volume than the other test facilities.

European test facilities to perform thermal-hydraulic tests with passive safety

In Fig. 2, the layout of the NOKO facility is shown [4]. In principle, the facility consists of

three loops:

1) The high pressure (<10MPa) "primary" loop with the electrically heated boiler (<4 MW), the recirculation pump, the steam-water separator, the pressure vessel and the test object (passive initiator, emergency condenser, building condenser or plate condenser).

2) The medium pressure (<1-1.5 MPa) "secondary" loop with the condenser tank with inlet and outlet pipes for the different fluids (water, steam, non-condensables).

3) The ambient pressure "auxiliary" loop with condensation tank and heat exchanger which is cooled by river water or via a cooling tower.

A commonly used instrumentation is installed.

2.4. The NOKO Data Bank

Data as well as drawings, calculations and reports are stored in a data bank; all information are available through the Internet — protected by a password. In Fig. 3 an overview is given.