Condensation is defined as the removal of heat from a system in such a manner that vapour is converted into liquid. This may happen when vapour is cooled sufficiently below the saturation temperature to induce the nucleation of vapour. This mode of heat transfer is often used in engineering because of possible high heat transfer coefficients. However, condensation heat transfer is degraded when non-condensable gases present in the condensing vapour. The presence of even a small amount of non-condensable gas in the condensing vapour has a profound influence on the resistance to heat transfer in the region of the liquid- vapour interface. The non-condensable gas is carried with the vapour towards the interface where it accumulates. The partial pressure of gas at the interface increases above that in the bulk of the mixture, producing a driving force for gas diffusion away from the surface. This motion is exactly counterbalanced by the motion of the gas-gas mixture towards the surface. Since the total pressure remains constant the partial pressure of gas at the interface is lower than that in the bulk mixture providing the driving force for gas diffusion towards the interface [1].

In-tube condensation of steam-air mixture is important for the passive thermal safety features of new advanced designs and therefore is subjected to experimental investigations at several research institutes. In addition to the studies of the University of California, Berkeley [2] and Massachusetts Institute of Technology [3], some detailed information and previous results of the experimental research being carried out at Middle East Technical University, Mechanical Engineering Department [4].

When the pure steam experiments are considered as the reference for comparison, a first indicator of the effect of air is a remarkable decrease in centreline and inner wall temperatures. Comparisons show that difference between saturation temperature, corresponding to the pure gas case, and measured centreline temperatures varies between 10 K and 50 K, depending on inlet air mass fraction. In other words, the temperature difference increases considerably as air mass fraction increases. It is found that there is a drastic decrease in the performance of the heat exchanger as the inlet air mass fraction increases. The inhibiting effect of air on condensation manifests itself as reduction in heat transfer coefficient. However, the inhibiting effect of air diminishes as system pressure and gas flow rate increase. The heat transfer coefficient can be based on either the measured centreline temperatures (Tc) or on the predicted one (Ts). The heat transfer coefficient considerably decreases when Ts is used since Ts is always greater than Tc. The ratio of the heat transfer coefficients computed from these two methods shows that increase in air mass fraction leads to larger deviation from the measured one calculated by Tc [4].

In the new advanced passive boiling water reactor design (SBWR and ESBWR), the main component of the passive containment cooling system (PCCS) is the isolation condenser (IC). The function of the IC is to provide the ultimate heat sink for the removal of the reactor coolant system sensible heat and core decay heat. In performing this function, the IC must have the capability to remove sufficient energy from the reactor containment to prevent it from exceeding its design pressure shortly following design basis events and to significantly reduce containment pressure in the longer run.

After a loss of coolant accident, the steam/air mixture from the reactor containment may flow to the IC which will then reject decay heat to a pool of water [5]. Similar advanced design features are also envisaged for AP-600. The researchers are focusing their attention on the AP — 600’s passive safety core cooling systems, whose major elements are a steel reactor containment shell surrounded by a concrete shield building, natural air conditioning between the containment and the shield building, and large volumes of gravity-fed water stored in tanks above the reactor itself. The passive systems performs the principle safety functions such as primary coolant inventory control, reactivity control and residual heat removal, but they rely on natural forces like condensation, evaporation and gravity rather than the mechanical equipment that is standard in traditional active designs.

For the supercritical fluid case, near and above the critical point at an absolute temperature, T, there is no distinguishable phase-change, but we still have a thermally expandable near-perfect gaseous fluid, with Д ~ 1/ T. Thus, to first order only, we have the supercritical flow rate wsc given by

Г 1 )1/3

WSC * W1F fp = W1Fgas

However, the thermal expansion coefficient is, in fact, non-linear with temperature changes near the critical point, as are many other properties, and the Boussinesq approximation is no longer a good approximation. The virial coefficients accommodate this deviation from perfect gas behaviour, but the properties are extremely non-linear near the critical point.

Therefore, for the higher-pressure case we may adopt a numerical analysis, based on iterative integration around the loop of the momentum equation (since mass is also conserved) for varying loop power inputs, using the thermophysical properties of the supercritical fluid as a function of actual thermodynamic state. Thus the general flow variation with major loop parameters (elevations, losses etc.) follows Equation (4) but with a non-linear expansion coefficient.

Taking the necessary boundary condition for parallel channels of constant pressure drop, differentiating the integral form of the mixture momentum equation, we may solve for the critical mass velocity, <G>, when the flow is unstable. The case and result for a heated supercritical flow is quite complex and requires numerical evaluation. For the adiabatic supercritical flow case, the analysis greatly simplifies. After making some algebraic manipulations, the criterion results in an unstable mass velocity given by:

G C m

< G >« C D ’ where Ц is the kinematic viscosity, De the equivalent diameter, and C the constant of

proportionality for the Reynolds number dependency of the friction factor. Thus, the dependency of the friction losses and the viscosity variation are quite important.

To investigate the feasibility of natural convection cooling for the primary circuit of a supercritical water-cooled reactor called, a simple steady-state, natural-circulation program was written, including the full physical variations of the thermophysical properties of supercritical water. With the initial and boundary conditions to the core, the operating pressure and temperature, the circuit resistance coefficients and the elevation difference between the core and the heat exchanger were specified. With an initially assumed flow the analysis iterated around the loop on flow to calculate the steady-state density and enthalpies in the circuit. To understand the parametric trends, many thousands of these calculations were done for different input conditions. The trends are shown in Figure 4 which confirms the expected that for a given inlet temperature the outlet temperature increases with increasing channel power, with decreasing elevation difference between core and heat exchanger, and with increasing circuit loss coefficient.

The effect of the large density and enthalpy changes around the critical temperature can be seen when the inlet temperature is a parameter. Below about 370oC, the outlet temperature/channel power surface is relatively flat (except for high loss coefficient combined with high power), whereas as soon as the inlet temperature exceeds 380oC the outlet temperature rises sharply regardless of the channel power and loss factors. If the fluid enters the core below the critical temperature, it is at relatively high density and low enthalpy, and exits above the critical temperature at low density and high enthalpy. The large density difference gives a large natural-convection driving force: the large enthalpy change allows a high channel power with relatively low flow and pressure drop.

To utilize the maximum design flexibility of elevation and loss coefficients within a maximum outlet temperature limit with a high-powered channel, it is necessary to keep the inlet temperature below the critical temperature and the outlet above the critical temperature. If the inlet temperature is allowed to rise above the critical temperature, the much-reduced density and enthalpy changes result in a very much higher outlet temperature.

A NC study was conducted utilizing a nodalisation for the PWR-1 in Table II, suitable for the Relap5/mod3.2.2, Ref. [14]. ‘Quasi’ steady-state thermalhydraulic NPP conditions are obtained at the end of transient calculations. The aim is to derive mutual relationships between significant NC parameters and to search for realistic boundary conditions allowing the maximum core power in NC. Use is also made of the NCFM derived above. Relevant results are given in Figs 7 and 8 and in Table VI.

TABLE VI. REMOVABLE POWER BY NATURAL CIRCULATION IN PWR-1

All the reported data relate to conditions where core power equals SG removed power. This is also valid when dryout situations occur and testifies of the small excursion of rod surface temperature. The excursion is actually limited to a few tens of Kelvin and is stable as a function of time. The main comments to the achieved results are:

• SPNC can be obtained up to about 20% core power, thus confirming the results related to ITF in Table V.

• TPNC allows removal of up to about 70% core power assuming nominal system conditions, again confirming the results related to ITF.

• Lowering SG pressure and increasing primary system pressure bring to increases in the NC thermal power removal capabilities. More than 90% core power can be removed in NC with SG pressure as low as 2.5 MPa.

• Dryout occurrences are undesirable. However, temperature excursions of rod surfaces are limited in space and do not affect the ‘stable and steady’ NC scenario.

• The NCFM obtained from the analysis of experimental NC scenarios at low core power values has been used as reference for high power NC scenarios. The information in Fig. 7, mainly the values of G/P and RM/V when dryout occur, shows that the NCFM (G/P versus RM/V) can be adopted also for high core power values. Dryout occurs with G/P close to unity (Kg/s/MW) or below this threshold. Lower values of G/P at dryout are experienced at higher core power.

• 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 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.

Tests have been performed with a passive flooding system. The objective was the experimental verification of the flooding of a pressure vessel with water from an outside water pool, using a special check valve.

3.4 Passive decay heat removal during shutdown

With a flooded pressure vessel it could be shown that it is possible to remove decay heat from the core region to an outside pool by natural convection, see [3].

FLUENT code has been developed by Fluent Inc.(USA). The FLUENT code is a state-of-the- art computer code for modelling incompressible and compressible fluid flow and heat transfer in complex geometries. FLUENT provides complete mesh flexibility, solving flows problems with unstructured meshes that can be generated about complex geometries with relative ease. Supported mesh types include 2D triangular/quadrilateral, 3D tetra — hedral/hexahedral/pyramid/wedge, and mixed (hybrid) meshes. FLUENT also allows to refine the grid structure, depending on the flow solution. This solution-adaptive grid capability is particularly useful for accurately predicting flow fields in regions with large gradients, such as free shear layers and boundary layers. In comparison to solutions on structured or block — structured grids, this feature significantly reduces the time required to generate a "good" grid. Solution-adaptive refinement makes it easier to perform grid refinement studies and reduces the computational effort required to achieve a desired level of accuracy; since mesh refinement is limited to those regions where greater mesh resolution is needed.

The FLUENT has the following modelling capabilities:

— Flows in 2D or 3D geometries using unstructured solution-adaptive triangular/tetrahedral, quadrilateral/hexahedral, or mixed (hybrid) grids that include prisms (wedges) or pyramids: (Both conformal and hanging-node meshes are acceptable.);

— Incompressible or compressible flows;

— Steady-state or transient analysis;

— Inviscid, laminar, and turbulent flows;

— Newtonian or non-Newtonian flow;

— Convective heat transfer, including natural or forced convection;

— Coupled conduction/convective heat transfer;

— Radiation heat transfer;

— Inertial (stationary) or non-inertial (rotating) reference frame models;

— Lagrangian trajectory calculations for a dispersed phase of parti cles/droplets/bubbles, including

— coupling with the continuous phase flow through porous media

— one-dimensional fan/heat-exchanger performance-models

— two-phase flows, including cavitation

— free-surface flows with complex surface shapes.

The four subsystems making up the passive core flooding systems connect the four core flooding pools with the headers of the four emergency condensers. These subsystems each consist of a connecting line and an integrated check valve. Normally the check valve should open when the RPV pressure drops below the pressure at the bottom of the flooding pool, and the water should flow from the flooding pool to the RPV by gravitational force alone. Under such conditions, however, only small forces are available to open the check valve. To eliminate this problem, a spring integrated into the check valve opens the valve at an RPV pressure some 2.5 bar higher than in the case of normal check valves.

When this low RPV pressure is reached, the RPV water level is lower than the inlet nozzle of the reflooding line. This means that the emergency condenser headers are filled with steam. When the spring-loaded check valve opens, steam flows from the header to the core flooding pool and is condensed there by means of a device similar to a quencher. At this point in the process the check valves serve the same function as supplementary relief valves and the pressure drop in the RPV accelerates after the check valves open. This steam blowing phase is the only phase in which natural circulation is used. The warm water of the core flooding pool flows to the water surface and the quencher is always surrounded by cold water.

Once the RPV pressure reaches the same value as that at the bottom of the core flooding pool, flow direction in the core flooding line reverses, and the water flows by gravitational force from the core flooding pool into the RPV irrespective of whether natural convection or stratification prevails in the pool or in the RPV.

The passive core flooding system was also tested at the emergency condenser test facility at Julich. Experimental results were in a good agreement with the related theoretical analyses.

The main purpose of the experimental study is to demonstrate capability of decay heat removal, to determine behavior of the system and components, to find start-up characteristics and procedures and to obtain a database for developing computer codes for AC600/1000 passive emergency residual heat removal system.

The total 166 sets of experimental data are identified. The test results illustrate that short disturbances of wind speed, power and valve opening have no significant impact on natural circulation flow, implying that the system seems to tolerate these disturbances. A correlation of two-phase, natural circulation flow rate is derived from constitutive equations by use of lumped system parameters were obtained. The empirical coefficients m and n were obtained by non-linear regression of 83 test data. Compared with 166 sets of the measured data, the deviation of 98.8% of the data points is within ±15%.

It is noted from transient research that AC600/1000 ERHR system is able to remove decay heat. Three start-up modes are available to trigger natural circulation flow. Computer code ERHRAC simulating natural circulation flow characteristics of ERHRS for AC600/1000 has been developed based on the test data.

The PACTEL experiment GDE-11 was calculated with APROS 2.11 codes. This experiment contained a rapid condensation and long lasting mixing period in the CMT. A steep thermally stratified layer was already formed before the break opening. All simulations used previously created and tested models for PACTEL, which were modified according to the needs of the PSIS. The results of the calculations have been presented in references [6, 12, 13, 14, 15, 16].

The APROS calculation used 5-equation model. The drift flux model was used to describe the velocity difference between phases. The CMT was modelled with five equal sized nodes. Sensitivity analyses with different number of nodes was not performed. With five nodes, it was not possible to model exactly the thermal stratification in the CMT. In the base case calculation, the top node liquid temperature was initiated to 140°C, which corresponds to the average temperature of the upper part of the CMT.

• CMT level did not drain until the top node liquid temperature had reached saturation. Draining stagnated to the level of the node boundary and continued after the liquid in the node next below was in saturation. During the stagnation the injection flow decreased dramatically and started to oscillate.

• This stepwise draining mode was observed also with the initial temperature of 50°C in the CMT top node.

• The CMT behaviour dominated the primary system behaviour.

The test was started by opening the valve in the outlet line, see Fig. 1.

The outlet line has an inner diameter of 65 mm. The velocity of the fluid entering the pressure vessel (inner diameter 448 mm) is about 0.2-0.3 m/s at the beginning. Analysing the thermocouple readings a mixing zone of about 70-100 mm in the lower part of the vessel can be evaluated; this is plausible. From the temperature distribution at the highest level (T2.02) it can be concluded that the length of the mixing zone did not change along the height of the vessel.

The two special grids have been installed to evaluate the possibilities of plumes. Fig. 5 and 6 show the temperature deviations from the mean temperature for the lower and upper grid, resp. With the exception of time period during which the mixing zone passes the grid the deviations are about < 0.2 K.

In Fig. 7 the temperature curves measured with the thermocouples are given. In Fig. 8 the time when the mixture zone passes a level is shown. Differentiating the curve (spline — approximation) given in Fig. 8 with time results in the velocity inside the pressure vessel, see Fig. 9. Multiplying the velocity with the density (p = 981 kg/m3 ) and the cross section area of the pressure vessel (0.158 m2) result in the mass flow M.

The vertical temperature profiles are given in Fig. 10.

FIG. 5. Temperature deviations from the FIG. 6. Temperature deviations from the mean temperature for the lower grid. mean temperature for the upper grid.

FIG. 7. Temperature profile in the pressure FIG. 8. Time to pass for the mixture zone vessel. through the pressure vessel.

The energy transferred from the vessel to the condenser part can be calculated from the mass flow M and the difference between the inlet and outlet enthalpies of the condenser

Li = M(hin-hout).

In Fig. 11 the transferred energy is shown. At the beginning of the natural circulation the transferred power is about 180 kW.