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Conceptual design and analysis of a semi-passive containment cooling system for a large concrete containment
Nuclear Engineering and Design 199 (2000) 227 – 242 www.elsevier.com/locate/nucengdes
Conceptual design and analysis of a semi-passive containment cooling system for a large concrete containment C.S. Byun a, D.W. Jerng a, N.E. Todreas b,*, M.J. Driscoll b a
b
Korea Electric Power Research Institute, 103 -16, Munji-Dong, Yusung-Gu, Taejon, South Korea Department of Nuclear Engineering, Massachusetts Institute of Technology, Rm 24 – 243, 77 Massachusetts A6., Cambridge, MA 02139 -4307, USA Received 5 August 1999; received in revised form 27 September 1999; accepted 9 December 1999
Abstract An internal evaporator-only (IEO) concept has been developed as a semi-passive containment cooling system for a large dry concrete containment. The function of this system is to keep the containment integrity by maintaining the internal pressure not to exceed ultimate design pressure, i.e. 0.83 MPa (120 psia) in the absence of any other containment cooling following a severe accident, which postulates core damage and hydrogen combustion. The ability of the concept to protect the containment was evaluated for the design basis accident (DBA) large break loss of coolant accident (LB LOCA) and severe accident scenarios (LB LOCA without Emergency Core Cooling System (ECCS) and containment spray flow, 100% zirconium oxidation and complete hydrogen combustion). All were modeled using the GOTHIC computer code. It was concluded that a practical system requiring four IEO loops could be utilized to meet design criteria for severe accident scenarios. © 2000 Elsevier Science S.A. All rights reserved.
1. Introduction In general, passive type reactors which use natural driving forces such as gravity and buoyancy force ensure a high level of safety and could be cost-effective compared to active type reactors which depend on electric power and operator actions for ultimate safety, because operator acAbbre6iations: ECCS: Emergency Core Cooling System; ENEL: Ente Nazionale per L’Energia Elettrica; SBWR: Simplified Boiling Water Reactor; RCS: Reactor Coolant System * Corresponding author. Tel.: + 1-617-2535296; fax: +1617-2588863. E-mail address: [email protected] (N.E. Todreas)
tions are minimized and systems can be considerably simplified. Accordingly application of passive safety concepts are being studied worldwide even for active type reactors and the study introduced here is one of these efforts. Containment cooling is essential to ensure the safety of fission power plants since it maintains the integrity of the containment and thus the radioactivity is confined. Conventional nuclear power plants, which are of the active type use active cooling systems to cool the internals of the containment and depend on pumps and heat exchangers to recirculate the spray water. To achieve the ultimate safety of nuclear power
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plants, however, concepts for passive cooling of containments have been developed due to the aforementioned reasons. The applicable candidates for a passive containment cooling system (PCCS) depend on the containment type. For a containment with steel walls such as AP600, the method (IAEA, 1993) of spraying water on the outer surface of the containment can be applied. In this case, the steel wall provides the heat transfer area and the main heat transfer mechanism is the natural convection inside the containment and evaporative plus air cooling on the steel surface of the containment. However, the concrete containment type needs a passive means to extract the energy from the containment atmosphere despite the high thermal resistance concrete walls. Various passive containment cooling systems have been proposed (and some are currently under evaluation): such as heat pipes (Ahmad et al., 1983), the isolation condensers of SBWR and the thermosyphon loop proposed by ENEL (de Cachard et al., 1997). The heat pipe type requires considerable space in the upper dome in order to cope with a design basis accident (DBA) because of a small heat removal capability per single heat pipe (: 0.25 MW) so it is not appropriate for the large rating pressurized water reactor (PWR). In addition, the leakage probability and the integrity of upper dome may be a concern due to the presence of many penetrations. The isolation condenser was applied to a boiling water reactor’s post-loss of coolant accident (LOCA) requirements, which are inherently different from the PWR considered here. The ENEL approach has a water pool as the cooling water source outside the upper dome and two heat exchangers, one inside the containment and another inside the pool. Except for its installed location, this is somewhat similar to the closed thermosyphon loop (CTSL) concept proposed by MIT as a PCCS for a large rating PWR in which the evaporated steam is condensed in a heat exchanger submerged in the condensing pool (Leiendecker et al., 1997, 1998). In this work, a PCCS concept was developed for application to a dry concrete containment of a large capacity PWR which consists of an internal evaporator and external pool with connecting
pipes, hereafter designated as the internal evaporator-only (IEO) concept. The internal evaporator is located at the inside of the containment and absorbs heat. The water supply pool is located outside the containment and elevated above the internal evaporator. Therefore, natural circulation between the pool and the internal evaporator can be developed. The IEO concept can reduce the internal thermal resistance compared to the CTSL concept. In addition, the problem with non condensable gas buildup inside the closed loop system, which requires a vacuum pump system to periodically evacuate the loop vapor space, can be eliminated. For this new IEO concept, performance analyses were performed with a distributed parameter model using the GOTHIC code, under various situations such as hydrogen burning and aerosol deposition on the heat transfer surface. The primary objective was to evaluate the potential of the IEO system to mitigate containment pressure increases under severe accident conditions. In order to evaluate the performance of the IEO concept, the Korean next generation reactor (KNGR) was selected as a reference reactor, which is an evolutionary type large capacity PWR with a 4000-MWth system power and a free containment volume of 90 400 m3.
2. System overview The IEO concept requires the evaporator be placed inside the containment and a condensing pool be placed outside containment. Therefore, some inherent characteristics of the KNGR General Arrangement (GA) must be considered in order to adopt the IEO concept within the KNGR. As seen in Fig. 1, the KNGR has two pools, which can be used as the cooling water source at the top of the auxiliary building at the elevation of 53.3 m outside the containment, which are furnished for the passive secondary cooling system (PSCS). In addition, the KNGR has an annular gap of 30 cm width around the containment wall, so that mixing between the upper containment and the lower containment can be enhanced.
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Based on the above considerations, the IEO concept for the KNGR is proposed as shown in Fig. 2. As seen in Fig. 2, the IEO concept has a missile shield, which can protect the evaporator tubes from laterally directed missiles to reduce the containment bypass probability. In addition the isolation capability is increased by applying a double-series valve arrangement on both sides of the containment to meet containment isolation criteria. Furthermore a radiation monitor may be installed in the steam line in order to detect radiation release to atmosphere via a ruptured tube during operation. As seen in Fig. 2, one IEO unit installed above the operating floor (elevation 47.6 m) has five main components; evaporator tubes, separator, mixing plenum, loop piping and condensing pool. To provide redundancy, two
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IEO units will be installed in each PSCS pool. Since two PSCS pools are provided in the KNGR, a total of four IEO units will be considered as the reference number for performance analysis. The evaporator tubes are empty during normal operation and they start to remove heat from the containment atmosphere after filling with cold coolant from the outside PSCS pool installed at 53.3 m. After the IEO is filled with cold water and until the water level is equalized in the pool and the IEO, the water in the IEO is heated up through condensation heat transfer on the outside of the evaporator tubes. Boiling starts after the local internal temperature of the water in the tubes rises beyond the saturation temperature at the local pressure.
Fig. 1. Vertical view of KNGR.
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the recirculation line takes place.
2.1. Design parameters for the IEO
Fig. 2. Schematic of IEO concept. Table 1 Design parameters for single smooth tube Parameter
Value
Tube (SS) Tube length (m) Inner tube diameter (m) Outer tube diameter (m) Tube thickness (mm) Conductivity (W/m-K) Density (kg/m3) Specific heat (kJ/kg-K)
2 0.04 0.042 1 20 7820 0.46
There are three pipes in one IEO unit: one steam line, one feedwater line and a recirculation line. The feedwater line and the steam line have double series isolation valves which are powered from the Class 1E DC. The recirculation flow through the recirculation line is generated by the buoyancy force due to the density difference between two-phase liquid in the evaporator tubes and the all-liquid water in the recirculation line connecting the separator and inlet plenum. In the lower mixing plenum, mixing between cold water from the feedwater line and saturated water from
The evaporator consists of 500 stainless steel tubes. Each tube is 2 m long and has an outer radius of 2.1 cm and a wall thickness of 1 mm. Tables 1 and 2 show design parameters for the single tube, radial fins and tube bundle. The total outer heat transfer area of the smooth tubes is 131.9 m2 and that of the inner surface is 125.6 m2. Introducing extended surfaces of fins with high conductivity, the number of the tubes can be optimized in order to meet the performance target. The wall of an evaporator tube acts as a thermal resistance to the heat flow. If the tube wall is too thick, the temperature drop becomes significant and the heat transport capability will be reduced. Therefore, the tube wall is usually kept as thin as possible. A very thin-walled evaporator tube, however, cannot be used for a nuclear reactor application because it may fail structurally. The pressure inside the containment building may build up to 0.83 MPa within a short period following a core melt accident. Under these cirTable 2 Design parameters of fins and tube bundle Parameter
Value
Fin (copper) Outer fin diameter(m) Fin thickness(mm) Gap between fins(mm) Fin spacing (fins/m) Conductivity(W/m-K)
0.1 1 4 200 400
Tube number
500
Area Smooth tubes Total outer heat transfer area for smooth tube (m2) Total inner heat transfer area for smooth tube (m2) Finned tubes Total fin area Total smooth region area Total outer heat transfer area for smooth tube (m2)
131.9 125.6
2711.7 79.2 2790.9
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cumstances, the evaporator tubes inside the containment are subjected to 0.83 MPa external pressure. Therefore, the evaporator tubes must be strong enough to withstand this external pressure as well as any kind of rapid pressure surge that may occur inside the containment: for example hydrogen detonation or molten core-water interaction. The required thickness of the evaporator tube wall and the maximum allowable pressure on the tube wall were determined by meeting pressure vessel requirements of the ASME Boiler and Pressure Vessel Code. The codes for designing shells and tubes under external pressure were used. Requirements for the design are given in Section VIII, UG28, UB98 of the ASME Boiler and Pressure Vessel Code (ASME, 1971). The material used for the evaporator tube is stainless steel SA179 (ASME, 1971) and the external design pressure is assumed to be 0.83 MPa. A tube wall thickness of 0.75 mm is needed. On the other hand, a wall thickness of 1 mm will allow the tube wall to be subjected to an external pressure up to 2.24 MPa, which is almost 2.7 times the design pressure and will give adequate corrosion allowance if one considers that corrosion of the evaporator tube should be negligible because the evaporator tube will always be dry during normal operation. In order to enhance the condensation heat transfer outside the evaporator tube, the outer surface of the evaporator tube will have fins to provide an extended heat transfer area. In this study, radial and axial fin types are considered. A fin enhancement factor of 4 is used, which is defined by Leiendecker et al. (1997) as the ratio of the heat transfer rate for the finned tube to that for the smooth tube. This fin enhancement factor of 4 is selected conservatively allowing for the gap contact conductance between the fin and body of the tube, surface corrosion and fouling layers. The separator in the IEO serves two functions. One is to separate the liquid entrained by steam flow. The second function of the separator is to supply the recirculation head, acting as a water drum. Unlike the closed thermosyphon loop concept, there is no need for a highly efficient separator in the IEO concept since a moderate amount
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of liquid entrainment in the vented steam is tolerable. A gravity steam–water separator consisting of a vertically oriented cylindrical tank is adopted in this project. The separation effect can be determined by calculating the superficial velocity of steam at the two-phase interface. Hence an effective separation dictates a separator with a long water column. Since the equalizing water level in the separator is initially 9.1 m, the superficial vapor velocity at the interface is about 0.3 m/s, which is sufficiently low to ensure effective separation of the entrained liquid in the evaporator tubes. On the other hand, to reduce the water inventory in the evaporator unit, displacer tubes 8 m tall may be installed. In this case the equalizing water level increases somewhat compared to the separator without displacer tubes, i.e. from 9.1 to 9.46 m. In this manner, the heat-up time of the IEO unit to reach the saturation state will be reduced due to a considerable reduction of the water inventory from 30 475 kg for the IEO without displacer tubes to 13 261 kg for the IEO with displacer tubes. In the mixing plenum, the hot recirculation water and the cold feed water are mixed. In order to establish the degree of mixing, a detailed analysis will be eventually required, but in this report it is assumed that complete mixing occurs in the mixing plenum. Two PSCS pools, which are used as a coolant source for the PCCS were originally provided for the secondary cooling system and play the role of a cooling water source for the PSCS. Originally, tank capacity was determined based on the mission time of the PSCS, which is defined by the time duration for cooling down the RCS from the initial condition at PSCS actuation to the shutdown cooling entry condition for all design basis accidents. For this function, PSCS tank capacity was determined as 1211.3 m3 based on 8 h of operation after reactor trip without the need for refill, referring to the auxiliary feedwater system cooldown time of 6–8 h for cooling down the RCS to the shutdown cooling entry condition. However, in fact, the PSCS tank level will be controlled by the water supply system operated by a level controller installed in the PSCS tank.
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sub-compartments were modeled as lumped parameters. Almost all of the thermal conductors are internal thermal conductors except the dome and the cylinder wall of the upper containment. The cylinder wall was modeled as the ‘wall’ and the dome was modeled as the ‘ceiling’ of the upper containment. By this modeling, GOTHIC automatically calculates the distributed surface area of the conductor ‘wall’ in order to fit the size of a sub-cell.
3.1. IEO modeling The conductor model can be adopted as an appropriate model for the IEO. A thermal conductor model requires three modeling parameters: heat transfer area and wall thickness, condensation heat transfer coefficient on the containment side and the heat transfer option on the inner surface of the IEO. Fig. 3. Plan view of the noding scheme for the distributed parameter model of the GOTHIC code.
3. Modeling The lumped parameter model (LPM) devised for GOTHIC has 23 control volumes, 50 flowpaths and 53 thermal conductors. As components, there are two spray pumps and five valves. This LPM is used for the sensitivity analysis objective only and most of performance analysis for the IEO is performed using a distributed parameter model (DPM). A DPM was devised based on refinement of the lumped parameter model. A distributed parameter model can provide a detailed representation of the containment phenomena because the conservation equations within the subdivided volume are solved for the 3-dimensional distributions of velocity, temperature and pressure which represent approximations to actual conditions subject to the density of the noding scheme. In the DPM for KNGR, there are three subdivided control volumes: an upper containment and two outside pools (PSCS Tanks). Figs. 3 and 5 show the upper containment modeled by 105 cells (3x × 5y × 7z). As seen in Fig. 4, the other 21 cells including equipment rooms and steam generator
Fig. 4. Vertical view (B-B) of the noding scheme for distributed parameter model of the GOTHIC code.
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Unlike the closed loop concept, the IEO has no internal structure between the containment atmosphere and the external coolant on the atmospheric side, i.e. the evaporator tube wall is the sole thermal barrier. Hence, the thickness of the evaporator tube wall can be used as the equivalent wall thickness. For the heat transfer area of the equivalent thermal conductor, the inner surface area or outer surface area of the smooth tube can be used. In this study, the inner surface area of the smooth tube is used for the equivalent thermal conductor of the IEO. Hence the heat transfer area used per one IEO unit is the inner surface area of 125.6 m2 of the smooth evaporator tube.
3.1.1. Condensation heat transfer option on the outer surface of the IEO model The condensation heat transfer coefficient on the outer surface of the IEO model is modeled by the reduced Uchida condensation heat transfer coefficient, hred, which is obtained from a best estimate condensation correlation and two multipliers (l, N) to represent the effect of the extended area due to the fins. hred = houtl N(Dout/Din)
(1)
where hout is the nominal value of the condensation heat transfer coefficient, for which the Uchida condensation correlation calculated by GOTHIC is used, l is the ‘fin enhancement factor’ and N is a multiplier (based on the best estimate results obtained through experiment) on the condensation heat transfer coefficient. The term, Dout/ Din is the ratio of tube diameters, which is introduced because the heat transfer area of the equivalent thermal conductor is taken as the inner surface area of the IEO. The best estimate results, which are obtained by experiment, for the condensation heat transfer coefficient, give the multiplier, N, to apply to the nominal condensation heat transfer coefficient value (Liu et al., 1999). Based on this experiment for steam/air, Liu et al. (1999) recommends using the diffusion layer model (DLM) with suction in the containment analysis, which is an option not yet available in GOTHIC. However Liu et al. (1999) also recommends that the 2.2 times Uchida
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Option can be used for engineering design and analysis based on its simplicity and ease of application based on the good agreement of this option with experimental data and DLM predictions. The IEO tube bundles are envisioned to be constructed along the containment walls with integral air baffles to ensure a unidirectional flow field through the bundle. The hot, steam rich mixture will be drawn from higher elevations and directed through the IEO, with the cool, air rich mixture being discharged at lower elevations. This design isolates the flow field in the IEO tube bundle from possible disruption due to unpredictable containment mixing and makes the IEO concept more amenable to analysis. Based on the above consideration, a two times Uchida Option is used for the performance analysis of the IEO in the DBA LOCA scenario, to provide a conservative allowance for margin compared to the experimental results. From the runs to obtain the heat transfer coefficient in the presence of air/helium non condensables, Liu et al. (1999) also suggests utilization of the Uchida air-only correlation with a reduction factor of 20% to be conservative in the air/helium (simulating hydrogen) case as long as the total pressure is B 4.5 atm and there is B 30% hydrogen in the air–hydrogen mixture. Here the one times Uchida Option will be used for modeling the IEO, again with a large conservative allowance. The fin enhancement factor accounts for the effect of the extended surface for the finned tube on the heat transfer rate. The fin enhancement factor can be obtained as the ratio of the heat transfer rate for the finned tube to that for the smooth tube. Here, we will apply a reduction factor defined by the ratio of the fin enhancement factor (again with a conservative allowance) to that obtained by theoretical analysis for the same tube. For this introduced reduction factor, a value of :0.5 appears appropriate given our current level of understanding. Hence the fin enhancement factor of 4, reduced from the prediction of 9.5 for the proposed radial finned tube, will be used for the GOTHIC analysis (Leiendecker et al., 1997).
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3.1.2. Heat transfer option on the inner surface of the IEO The specific temperature option is applied to the inner surface of the IEO model because the bulk temperature of the fluid inside the evaporator tube is at its saturation temperature. Hence the inner surface temperature changes only slightly with axial position. The inner surface temperature of the evaporator tube may be characterized in three time phases. During Phase I, the evaporator tube heat-up process, the inner wall temperature will monotonically increase from the initial temperature (30°C) of the system to some temperature, which is determined by the containment and saturation temperatures at that time. During Phase II, the separator heat-up process, because the saturation temperature which corresponds to the initial water level is not changed, the inner wall temperature can be considered as constant if the containment temperature is fixed. Saturation temperature and the inner wall temperature can be obtained by considering the design temperature of 130°C for DBAs (note the containment temperature is roughly constant at 130°C during the blowdown phase). The saturation temperature is about 120°C and the inner wall temperature is 125°C at the containment temperature of 130°C. We can choose two temperature options for the inner surface temperature; the constant temperature option of 125°C and the variable temperature option obtained from the above considerations. Of these two options, the constant temperature option of 125°C will give the most conservative results in the GOTHIC code because GOTHIC will calculate the heat removal rate conservatively for Phase I and Phase III. Hence this constant temperature option will be used for the GOTHIC code analysis. 3.2. Hydrogen and aerosol layer modeling GOTHIC includes two hydrogen burn models: discrete burn model and continuous burn model (George et al., 1997). In this analysis, the complete burning of hydrogen using the continuous
burn model is assumed. It should be noted that the complete combustion assumption is unrealistically conservative because the inerting effect by steam and the limit of oxygen mole fraction for hydrogen burning are not considered. However this approach will give bounding results. To obtain a first estimate of the effect of the aerosol layer on the IEO performance, a uniform layer of aerosols is simply assumed to cover the IEO outer surface. The surface area is 125.6 m2 per IEO unit, which was used as the reference area of the IEO model in the GOTHIC code. The mass of aerosols is assumed as 50 kg, which is predicted for a in the typical 1000 MWe PWR with a large dry containment of about 105 m3 volume. The thermophysical properties of aerosol were assumed to be those of typical concrete which has a density of 2240 kg/m3 and a thermal conductivity of 1.13 W/mK. For this conservative case, the thickness of the aerosol layer becomes 0.044 mm, which is selected for analysis of the aerosol effect.
3.3. Lumped parameter modeling 6ersus distributed parameter modeling The comparison of the LPM versus the DPM presented below is sensitive to the mixing processes modeled in GOTHIC. In this comparison the IEO design and location is established to ensure a unidirectional flow downward through the bundle as described in Section 3.1.1. Hence while the comparison between models made in this section is applicable to this IEO design since the dominant mechanisms creating this flow field are mechanistically modeled, this comparison between model performance is necessarily applicable only to other dry containment configurations where these same mechanisms are dominant. Fig. 5 shows the difference of the pressure predicted by the GOTHIC 6.0a code between LPM and DPM. The following factors are employed: a fin enhancement factor of 4 for the finned tube model, the specified temperature option of 125°C inside the tube and the two times Uchida Option for the containment side.
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Fig. 5. Comparison of the pressure predicted by LPM and DPM of GOTHIC 6.0a for four IEOs (fin enhancements factor: 4; Specified temperature option for the pool side: 125°C; two times Uchida Option for the containment side). (1), LPM; (2), DPM; Elevation; 53.3 m, Nominal transient: DPM; Elevation; 88.2 m.
Fig. 6. Comparison of the performance of four IEOs with that of four CTSLs for DBA LB LOCA scenario (fin enhancement factor: 4; elevation; 88.2 m). (1), Four CTSL; turbulent convection option for the pool side; one time Uchida Option for the containment side; DPM for GOTHIC 3.4e; (2), Four CTSL; turbulent convection option for the pool side; two times Uchida Option for the containment side; DPM for GOTHIC 3.4e; (3), Four IEO; turbulent convection option for the pool side; one times Uchida Option for the containment side; DPM for GOTHIC 3.4e; (4), Four IEO; turbulent convection option for the pool side; two times Uchida Option for the containment side; DPM for GOTHIC 3.4e. Nominal transient: four IEO; specified temperature option for the pool side: 125°C; two times Uchida Option for the containment side; DPM for GOTHIC 6.0a.
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Two pressure curves for the DPM are plotted in this figure. One curve (with square markers) presents the pressure when four IEOs are assumed to be installed at an elevation of 88.2 m. The other curve (with triangular markers) shows the pressure trend when four IEOs are installed at 53.3 m. As seen in the figure, the pressure obtained using the LPM is higher than that from the DPM during the initial 300 s. This phenomenon can be explained by most of the thermal conductor being distributed in the lower elevation of the upper containment and the atmosphere being forced to be homogeneous and so there is less cooling due to expansion allowed in the LPM run. Hence the pressure for the DPM run is lower than that for the LPM run. After 300 s, however, the pressure for the DPM run is higher than that for the LPM run. This can also be explained by stratification of the steam. The ratio of the air density to the steam density at high level is less than that in the lower level, i.e. some stratification of steam occurs. Therefore the steam removal rate throughout the entire volume will decrease because of low steam density in the lower level where most thermal conductors are located. The above findings differ from previous work which predicted the pressure obtained from the DPM run to be always lower than that from the LPM run (Kim et al., 1998). Based on the same reason, we can explain the sensitivity of IEO performance to the installed location of the IEO. After 300 s, because the steam density in the high level is more than that in the low level, the performance of the IEO which is installed in a high position is better than that of an IEO which is located in the low position. However, the difference in peak pressure is about 13.1 kPa (549.5 kPa vs 562.6 kPa), which is not especially significant. In addition, natural convection circulation is established adequate to permit location of IEO units low in containment, such that the PSCS pools can serve as a source of water or heat sink. All results reported in the next section were obtained using the distributed parameter model.
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four IEOs is much better than that of four TSLs due to the reduced thermal resistance of the IEO compared to the TSL. In addition the two times Uchida Option effect for the IEO is larger than that for the TSL concept. Note, however, that pressure is not reduced by a factor of 2 in 24 h (86 400 s).
Fig. 7. Pressure prediction as a function of the number of IEO units used (fin enhancement factor: 4; two times Uchida Option for the containment side; DPM for GOTHIC 6.0a; Elevation; 88.2 m). (1), Four IEO; specified temperature option for the pool side: 105°C; (2), Six IEO; specified temperature option for the pool side: 105°C; (3), Ten IEO; specified temperature option for the pool side: 105°C. Nominal transient: four IEO; specified temperature option for the pool side: 125°C.
4. Performance analysis of the IEO
4.1. Large break loss of coolant accident (LB LOCA) as a DBA 4.1.1. Comparison of the IEO performance with the closed thermosyphon The performance of four IEOs compared with that of four closed thermosyphon loops is shown in Fig. 6 for the LB LOCA DBA case. For the analysis of the thermosyphon loop (TSL) the equivalent thermal conductor model with the modeling parameters in Leiendecker’s report are used except that the surface area for the single TSL is revised to be 125.7 m2 (based on the outer surface area). This value is then multiplied by a fin enhancement factor of 4. The heat transfer coefficient option of natural convection for the pool side and the one times or two times Uchida Options were used for the containment side. This figure shows that the performance of
4.1.2. IEO performance as a function of IEO units Fig. 7 shows the pressure trend as the number of IEO units is varied. In this analysis the specified temperature of 105°C is assumed, i.e. the water level in the separator is assumed to be as low as possible (e.g. by introducing a level controller). As seen in this figure, if we use ten IEOs, GOTHIC predicts the peak pressure of 0.42 MPa, so we can likely meet the design peak pressure of 0.41 MPa. However it is inherently difficult to meet the second performance criterion which is to reduce the containment peak pressure by half after 24 h of an accident because the temperature difference between the containment and the IEO wall is low in the long term period. From this figure, we can conclude that if the water level and hence the specified temperature is high, as in the present analysis, a large number of IEO units (beyond ten IEO units) may be required to meet the design criteria (peak pressure of 0.41 MPa). Therefore in order to improve the IEO performance, the water level in the separator must be kept as low as possible and the IEO should be installed in the highest position practicable. However the latter option is restricted by the location of the cooling water source, i.e. the water pool. As mentioned in Section 3, KNGR has the PSCS pool fixed at an elevation of 53.3 m. Therefore this approach to increase the performance of the IEO is not acceptable in the present stage of KNGR design. The former approach may be possible if we can introduce a reliable level controller such as a diaphragm valve which is operated by the difference of the water head between the IEO and the water pool, or a float type controller.
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4.2. LB LOCA as a se6ere accident 4.2.1. Hydrogen burning analysis Deflagrations can occur due to hot particles and autoigniting plumes if the atmosphere contains sufficient hydrogen without dense steam. These flammability conditions are typically determined by a hydrogen – air – steam flammability chart (Kumar, 1985; Marshal, 1986). These flammability charts, which are applicable when the atmosphere is about 375 K, shows that the hydrogen concentration must exceed the lower flammability limit (LFL) for a flame to propagate through the mixture for a fixed concentration of steam. Also there is a critical steam concentration of about 50 – 65% beyond which all concentrations of hydrogen are inerted. In this assessment, however, the potential temperature dependence of deflagrations must be recognized. Stamps and Berman (1991) have reviewed the temperature dependence of the lower flammability limit and found that the lower flammability limit decreases about 0.5 vol.%/100 K for upward propagating flames and about 1.0 vol.%/
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100 K for downward propagating flames. The data base for this temperature dependence goes up to about 800 K for dry systems and about 473 K for steam concentrations reaching about 60%. From this review and Kumar’s experimental data, the LFLs for upward and downward propagating flames are given by XH2(UP, XS, T) = 0.37+ 0.2381 XS − 5E(− 5)(T − 373)
(2)
XH2(Down, XS, T) = 0.075+ 0.2381 XS − 1.0135E(− 5)(T −373) (3) On the other hand, examination of the flammability chart (Kumar, 1985) shows that there is a steam concentration beyond which hydrogen concentration is always inerted, regardless of hydrogen concentration. The inerting limit is given by Xstm(inert)= 0.63+ 3E(−4)(T − 373)
(4)
where the temperature derivative, dXstm/dt 3 vol.%/100 K was measured by Kumar over the 373–473 K temperature range.
Fig. 8. Gas concentrations predicted by GOTHIC 6.0a assuming 55% zirconium oxidation.
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Fig. 9. Performance of IEOs for LB LOCA severe accident (fin enhancement factor: 4; specified temperature option for the pool side: 125°C; one time Uchida Option for the containment side; DPM for GOTHIC 6.0a; Elevation; 88.2 m). Base case: No IEO; 100% Zr oxidation; no hydrogen burn; no aerosol deposition. (1), Four IEO; 100% Zr oxidation; no hydrogen burn; aerosol deposition; (2), Four IEO; 100% Zr oxidation; complete hydrogen burn; aerosol deposition; (3), Four IEO; 100% Zr oxidation; no hydrogen burn; no aerosol deposition.
Using this condition we can assess the flammability of hydrogen. As seen in Fig. 8 assuming realistic oxidation of zirconium (55%), hydrogen concentration is about 0.028–0.03 and the steam concentration has the range of 0.64– 0.66 during hydrogen release. On the other hand, the containment temperature is 204°C (477 K) at the peak. Therefore for this temperature and the steam concentration of 0.64 (conservative selection), the LFLs for upward and downward propagation and the inerting limit are given by XH2(UP, 0.64, 477 K) =0.517
(5)
XH2(Down, 0.64, 477 K) =0.217
(6)
Xstm(inert) =0.66
(7)
This evaluation shows that hydrogen cannot be inerted by the steam concentration because the inerting limit is higher than the steam concentration of 64%. However, it is apparent that there is no deflagration because the predicted hydrogen concentration is lower than the lower flammability limits.
Also, hydrogen reacts with oxygen at all temperatures, although the reaction is strongly temperature dependent. Since this type of reaction occurs uniformly through a constant temperature mixture, we call it volumetric combustion. Autoignition can occur if the uniform gas temperature exceeds a critical value. An examination of autoignition criteria for the wellmixed, uniformly heated mixture is summarized by Stamps and Berman (1991). According to Stamps and Berman (1991), the autoignition temperature is over 520°C under the 0% steam and 30% hydrogen mixture condition and this temperature limit increases as steam mole fraction increases. Therefore, there is no possibility of volumetric combustion of hydrogen for our cases, because the hydrogen and steam concentration are calculated to be about 3 and 64%, respectively.
4.2.2. Sensiti6ity analysis of the IEO performance Fig. 9 shows the performance for a LB LOCA severe accident. Here, four IEOs with a fin enhancement factor of 4 are used. The
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Uchida correlation with no modification factor for the outer surface of the tube and the fixed temperature of 125°C at the inside of the tube are used. In addition an aerosol layer of 0.004 mm is assumed to be formed at the outside of the tubes as mentioned in Section 3. Hydrogen is assumed to be produced through 100% oxidation of zirconium in the active core, but its burning is not assumed. As seen in Fig. 9, four IEOs are found enough to easily meet the design criteria of 0.83 MPa peak pressure if we do not include the hydrogen burning effect. In addition Fig. 9 also shows that the long-term pressure can be maintained close to 0.41 MPa. Therefore, a set of four IEOs can meet the design criteria to maintain the containment integrity during severe accidents. In addition, there is a sufficient margin between the design
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criteria and performance results so that it may be possible to reduce the number of IEO units or the size of an IEO system. In Fig. 9, the performance degradation due to the aerosol layer is quite small. In fact, the aerosol layer thickness was calculated as 2 mm if we include fins in the total heat transfer area. Therefore, if we use 2 mm thickness instead of 0.044 mm and more realistic thermophysical data for the aerosol layer than those of concrete, the performance degradation of the IEO due to the aerosol layer will be practically negligible. Fig. 9 also shows the containment pressure prediction when hydrogen is totally and rapidly burned during the hydrogen release period. In this case, the complete burning limits were used. This approach for hydrogen burning is unrealistically conservative since:
Fig. 10. IEO performance as a function of IEO number used for LB LOCA severe accident (fin enhancement factor: 4; specified temperature option for the pool side: 125°C; one times Uchida Option for the containment side; DPM for GOTHIC 6.0a; Elevation; 88.2 m). Base case: no IEO; 100% Zr oxidation; no hydrogen burn; no aerosol deposition. (1), Two IEO; 100% Zr oxidation; no hydrogen burn; aerosol deposition; (2), Four IEO; 100% Zr oxidation; no hydrogen burn; aerosol deposition.
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1. The zirconium oxidation during the melt progression does not exceed 60% (NUREG-1150, 1987), or 50% (Yang et al., 1991). 2. If we use the realistic zirconium oxidation of 55%, there is no possibility of deflagration and autoignition as discussed in Section 4.2.1. 3. If we consider a hydrogen mitigation system such as an igniter and/or hydrogen recombiner, the global burning of hydrogen can hardly occur. Nevertheless, the peak pressure even with a complete burning condition was found to be about 0.76 MPa as seen in Fig. 9. This means that the large dry containment like KNGR can meet the design criteria of 0.83 MPa even under the most conservative assumptions for zirconium oxidation and hydrogen burning.
4.2.3. IEO performance 6ersus number of IEO units in ser6ice Fig. 10 shows the IEO performance as a function of the number of IEO units used. As seen in this figure, there is no significant difference of the pressure during the hydrogen release period as the number of IEO units changes because of small condensation heat transfer coefficients during this period. However, Fig. 10 shows that the long-term pressure can be considerably reduced even with only two IEO units in service. Therefore, it could be possible to meet the design criteria for the severe accident of 0.83 MPa for the long-term period (72 h) even with two IEO units. Also, the peak pressure is lower than the design pressure even under the most conservative hydrogen generation and burning situation as mentioned before.
5. Conclusions In this paper, the IEO concept for the PCCS of a large dry concrete containment was proposed, in which the in-pool condenser is eliminated compared to the closed thermosyphon loop proposed in previous work (Leiendecker et al., 1997). By eliminating the in-pool condenser of a CTSL it is predicted that the thermal resis-
tance can be roughly cut in half in the IEO concept. Furthermore, cost and space requirements should be similarly reduced and problems due to non condensable gas built-up inside the heat transport system can be avoided. In the performance analysis using the GOTHIC code for a LB LOCA as a DBA, the pressure obtained using a LPM is about 13.8 kPa higher than that from DPM during the initial 300 s since the atmosphere is forced to be homogeneous and so there is less cooling due to artificial mixing in the LPM. After 300 s, the pressure from the DPM is about 13.8 kPa higher than that from the LPM because the ratio of air to steam densities at higher level is lower than that in the lower level, i.e. a stratification of steam occurs and so the steam removal rate throughout the entire volume will decrease because of low steam density at the lower level where most thermal conductors are located. In addition, if we use six IEOs assuming that the separator water level is sufficiently low (less than 50 cm), we can likely meet the DBA design peak pressure of 0.41 MPa (g). However, it is inherently difficult to meet another design criterion (to reach half of peak design pressure within 24 h) because the temperature difference between the containment and the IEO wall is low in the long term. For the severe accident scenario, it is apparent that there is no deflagration because the predicted hydrogen concentration is lower than the lower flammability limits and autoignition point of hydrogen. Even with two IEO units, it is possible to meet severe accident design criteria (less than 0.83 MPa) during the long-term period with enough margin. In addition the peak pressure is just 0.76 MPa even assuming 100% zirconium oxidation and the complete burning of hydrogen. Calculation also shows that the fouling effect by aerosols on IEO performance is negligible. Based on the above findings from the performance analyses, we conclude that the IEO has more merit for mitigation of severe accidents than design basis accidents.
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We suggest further development of the IEO modeling for the results to be more realistic. As discussed in Section 3, the fixed temperature option of 125°C at the inside of the internal evaporator was used for the IEO modeling in the GOTHIC code to avoid limitations in the code calculation speed due to Courant’s time limit as discussed in Section 3 and it needs to be improved. The geometrical dimensions of the finned tubes used in the MIT experiments by Liu et al. (1999) were different from the configurations used for the IEO performance analyses. Therefore additional experiments with finned tubes may be required. Any new experiments should focus on parameters impacting IEO performance, for example, the fin effectiveness as a function of geometry and the presence of non condensables, tube bundle effects and tilted tube effects. Eventually a full-scale test of a prototype unit to quantify bundle effects may be needed to qualify the IEO system for service. Furthermore we need to apply more realistic hydrogen production and burning processes to the IEO performance analysis. In this study, we assumed that most of the hydrogen was released during the core melt progression. However, hydrogen can be produced by the interaction between the hot core melt and water in the cavity and also the interaction between the debris and the concrete. With regard to hydrogen burning, hydrogen mitigation systems such as igniters and recombiners should be modeled appropriately in future work.
Appendix A. Nomenclature Din Inner diameter of the evaporator tube Dout Outer diameter of the evaporator tube hred Reduced condensation heat transfer coefficient hout Nominal condensation heat transfer coefficient N Multiplier on the condensation heat transfer coefficient based on the best estimate results obtained through experiment
T X H2 Xstm l
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Temperature of the containment atmosphere Hydrogen mole fraction Steam mole fraction Fin enhancement factor defined as the ratio of heat transfer rate for finned tube to that for the smooth tube
Acknowledgements This study was sponsored by Korea electric power corporation under the Collaborative Program in Power Engineering with Massachusetts Institute of Technology. The lead author wishes to express his appreciation for the opportunity to carry out this project provided by KEPRI. The authors appreciate the help and advice of Young Sang Choi, Brett Mattingly, Haiyang Liu, Yong Hak Kim and Markus Leiendecker. In particular Markus Leiendecker’s previous work was of considerable help in preparation of the numerical analysis code by the lead author and Young Sang Choi motivated the authors to focus on the severe accident analysis. References Ahmad, A. et al., 1983. PWR Severe Accident Delineation and Assessment, NUREG/CR-2666, UCLA-ENG-8284. ASME, 1971. ASME Boiler and Pressure Vessel Code, Section VIII, Div.1. American Society of Mechanical Engineering, USA. de Cachard, F.D. et al., 1997. Thermal-Hydraulic Modeling of Finned Tube Containment Condensers, Annual Report 1997, Annex IV, PSI Nuclear Energy and Safety Research, pp. 49 – 58. George, T.L. et al., 1997. GOTHIC Containment Analysis Package-Technical Manual, Ver. 6.0a, EPRI. IAEA, 1993. Status of Advanced Containment Systems for Next Generation Water Reactors. IAEA-TECDOC, Vienna, p. 752. Kim, Y.H., Todreas, N.E., Driscoll, M.J., 1998. Distributed Parameter Modeling of the KNGR Containment using GOTHIC, MIT-ANP-TR-059. Kumar, R.K., 1985. Flammability limits of hydrogen-oxygendiluent mixtures. J. Fire Sci. 3 (4), pp. 245 – 262. Leiendecker, M., Todreas, N.E., Driscoll, M.J., Hurtado, A., 1997. Design and Numerical Simulation of a Two-Phase Thermosyphon Loop as a Passive Containment Cooling System for PWRs, Vol. I and II, MIT-ANP-TR-053, Rev. 1.
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Leiendecker, M., Todreas, N.E., Driscoll, M.J., 1998. Thermosyphon loops for concrete containment cooling. Trans. Am. Nucl. Soc. 78, 263. Liu, H., Todreas, N.E., Driscoll, M.J., 1999. An Experimental investigation of a passive cooling unit for nuclear plant containment. Nucl. Eng. Design, 199, 243. Marshal, B.W., 1986. Hydrogen: air: steam flammability limits and combustion characteristics in the FITS vessel, Rep. NUREG/CR-3468, SAND84-0383.
NUREG-1150, 1987. Reactor Risk Reference Document: Appendices J – O: Draft for Comment, USNRC. Stamps, D.W., Berman, M., 1991. High-temperature hydrogen combustion in reactor safety applications. Nucl. Sci. Eng. 109, 39. Yang, J.W. et al., 1991. Hydrogen Combustion Control and Value-Impact Analysis for PWR Dry Containments, NUREG/CR-5662, USNRC.
Conceptual design and analysis of a semi-passive containment cooling system for a large concrete containment C.S. Byun a, D.W. Jerng a, N.E. Todreas b,*, M.J. Driscoll b a
b
Korea Electric Power Research Institute, 103 -16, Munji-Dong, Yusung-Gu, Taejon, South Korea Department of Nuclear Engineering, Massachusetts Institute of Technology, Rm 24 – 243, 77 Massachusetts A6., Cambridge, MA 02139 -4307, USA Received 5 August 1999; received in revised form 27 September 1999; accepted 9 December 1999
Abstract An internal evaporator-only (IEO) concept has been developed as a semi-passive containment cooling system for a large dry concrete containment. The function of this system is to keep the containment integrity by maintaining the internal pressure not to exceed ultimate design pressure, i.e. 0.83 MPa (120 psia) in the absence of any other containment cooling following a severe accident, which postulates core damage and hydrogen combustion. The ability of the concept to protect the containment was evaluated for the design basis accident (DBA) large break loss of coolant accident (LB LOCA) and severe accident scenarios (LB LOCA without Emergency Core Cooling System (ECCS) and containment spray flow, 100% zirconium oxidation and complete hydrogen combustion). All were modeled using the GOTHIC computer code. It was concluded that a practical system requiring four IEO loops could be utilized to meet design criteria for severe accident scenarios. © 2000 Elsevier Science S.A. All rights reserved.
1. Introduction In general, passive type reactors which use natural driving forces such as gravity and buoyancy force ensure a high level of safety and could be cost-effective compared to active type reactors which depend on electric power and operator actions for ultimate safety, because operator acAbbre6iations: ECCS: Emergency Core Cooling System; ENEL: Ente Nazionale per L’Energia Elettrica; SBWR: Simplified Boiling Water Reactor; RCS: Reactor Coolant System * Corresponding author. Tel.: + 1-617-2535296; fax: +1617-2588863. E-mail address: [email protected] (N.E. Todreas)
tions are minimized and systems can be considerably simplified. Accordingly application of passive safety concepts are being studied worldwide even for active type reactors and the study introduced here is one of these efforts. Containment cooling is essential to ensure the safety of fission power plants since it maintains the integrity of the containment and thus the radioactivity is confined. Conventional nuclear power plants, which are of the active type use active cooling systems to cool the internals of the containment and depend on pumps and heat exchangers to recirculate the spray water. To achieve the ultimate safety of nuclear power
0029-5493/00/$ - see front matter © 2000 Elsevier Science S.A. All rights reserved. PII: S 0 0 2 9 - 5 4 9 3 ( 0 0 ) 0 0 2 2 8 - 4
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plants, however, concepts for passive cooling of containments have been developed due to the aforementioned reasons. The applicable candidates for a passive containment cooling system (PCCS) depend on the containment type. For a containment with steel walls such as AP600, the method (IAEA, 1993) of spraying water on the outer surface of the containment can be applied. In this case, the steel wall provides the heat transfer area and the main heat transfer mechanism is the natural convection inside the containment and evaporative plus air cooling on the steel surface of the containment. However, the concrete containment type needs a passive means to extract the energy from the containment atmosphere despite the high thermal resistance concrete walls. Various passive containment cooling systems have been proposed (and some are currently under evaluation): such as heat pipes (Ahmad et al., 1983), the isolation condensers of SBWR and the thermosyphon loop proposed by ENEL (de Cachard et al., 1997). The heat pipe type requires considerable space in the upper dome in order to cope with a design basis accident (DBA) because of a small heat removal capability per single heat pipe (: 0.25 MW) so it is not appropriate for the large rating pressurized water reactor (PWR). In addition, the leakage probability and the integrity of upper dome may be a concern due to the presence of many penetrations. The isolation condenser was applied to a boiling water reactor’s post-loss of coolant accident (LOCA) requirements, which are inherently different from the PWR considered here. The ENEL approach has a water pool as the cooling water source outside the upper dome and two heat exchangers, one inside the containment and another inside the pool. Except for its installed location, this is somewhat similar to the closed thermosyphon loop (CTSL) concept proposed by MIT as a PCCS for a large rating PWR in which the evaporated steam is condensed in a heat exchanger submerged in the condensing pool (Leiendecker et al., 1997, 1998). In this work, a PCCS concept was developed for application to a dry concrete containment of a large capacity PWR which consists of an internal evaporator and external pool with connecting
pipes, hereafter designated as the internal evaporator-only (IEO) concept. The internal evaporator is located at the inside of the containment and absorbs heat. The water supply pool is located outside the containment and elevated above the internal evaporator. Therefore, natural circulation between the pool and the internal evaporator can be developed. The IEO concept can reduce the internal thermal resistance compared to the CTSL concept. In addition, the problem with non condensable gas buildup inside the closed loop system, which requires a vacuum pump system to periodically evacuate the loop vapor space, can be eliminated. For this new IEO concept, performance analyses were performed with a distributed parameter model using the GOTHIC code, under various situations such as hydrogen burning and aerosol deposition on the heat transfer surface. The primary objective was to evaluate the potential of the IEO system to mitigate containment pressure increases under severe accident conditions. In order to evaluate the performance of the IEO concept, the Korean next generation reactor (KNGR) was selected as a reference reactor, which is an evolutionary type large capacity PWR with a 4000-MWth system power and a free containment volume of 90 400 m3.
2. System overview The IEO concept requires the evaporator be placed inside the containment and a condensing pool be placed outside containment. Therefore, some inherent characteristics of the KNGR General Arrangement (GA) must be considered in order to adopt the IEO concept within the KNGR. As seen in Fig. 1, the KNGR has two pools, which can be used as the cooling water source at the top of the auxiliary building at the elevation of 53.3 m outside the containment, which are furnished for the passive secondary cooling system (PSCS). In addition, the KNGR has an annular gap of 30 cm width around the containment wall, so that mixing between the upper containment and the lower containment can be enhanced.
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Based on the above considerations, the IEO concept for the KNGR is proposed as shown in Fig. 2. As seen in Fig. 2, the IEO concept has a missile shield, which can protect the evaporator tubes from laterally directed missiles to reduce the containment bypass probability. In addition the isolation capability is increased by applying a double-series valve arrangement on both sides of the containment to meet containment isolation criteria. Furthermore a radiation monitor may be installed in the steam line in order to detect radiation release to atmosphere via a ruptured tube during operation. As seen in Fig. 2, one IEO unit installed above the operating floor (elevation 47.6 m) has five main components; evaporator tubes, separator, mixing plenum, loop piping and condensing pool. To provide redundancy, two
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IEO units will be installed in each PSCS pool. Since two PSCS pools are provided in the KNGR, a total of four IEO units will be considered as the reference number for performance analysis. The evaporator tubes are empty during normal operation and they start to remove heat from the containment atmosphere after filling with cold coolant from the outside PSCS pool installed at 53.3 m. After the IEO is filled with cold water and until the water level is equalized in the pool and the IEO, the water in the IEO is heated up through condensation heat transfer on the outside of the evaporator tubes. Boiling starts after the local internal temperature of the water in the tubes rises beyond the saturation temperature at the local pressure.
Fig. 1. Vertical view of KNGR.
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the recirculation line takes place.
2.1. Design parameters for the IEO
Fig. 2. Schematic of IEO concept. Table 1 Design parameters for single smooth tube Parameter
Value
Tube (SS) Tube length (m) Inner tube diameter (m) Outer tube diameter (m) Tube thickness (mm) Conductivity (W/m-K) Density (kg/m3) Specific heat (kJ/kg-K)
2 0.04 0.042 1 20 7820 0.46
There are three pipes in one IEO unit: one steam line, one feedwater line and a recirculation line. The feedwater line and the steam line have double series isolation valves which are powered from the Class 1E DC. The recirculation flow through the recirculation line is generated by the buoyancy force due to the density difference between two-phase liquid in the evaporator tubes and the all-liquid water in the recirculation line connecting the separator and inlet plenum. In the lower mixing plenum, mixing between cold water from the feedwater line and saturated water from
The evaporator consists of 500 stainless steel tubes. Each tube is 2 m long and has an outer radius of 2.1 cm and a wall thickness of 1 mm. Tables 1 and 2 show design parameters for the single tube, radial fins and tube bundle. The total outer heat transfer area of the smooth tubes is 131.9 m2 and that of the inner surface is 125.6 m2. Introducing extended surfaces of fins with high conductivity, the number of the tubes can be optimized in order to meet the performance target. The wall of an evaporator tube acts as a thermal resistance to the heat flow. If the tube wall is too thick, the temperature drop becomes significant and the heat transport capability will be reduced. Therefore, the tube wall is usually kept as thin as possible. A very thin-walled evaporator tube, however, cannot be used for a nuclear reactor application because it may fail structurally. The pressure inside the containment building may build up to 0.83 MPa within a short period following a core melt accident. Under these cirTable 2 Design parameters of fins and tube bundle Parameter
Value
Fin (copper) Outer fin diameter(m) Fin thickness(mm) Gap between fins(mm) Fin spacing (fins/m) Conductivity(W/m-K)
0.1 1 4 200 400
Tube number
500
Area Smooth tubes Total outer heat transfer area for smooth tube (m2) Total inner heat transfer area for smooth tube (m2) Finned tubes Total fin area Total smooth region area Total outer heat transfer area for smooth tube (m2)
131.9 125.6
2711.7 79.2 2790.9
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cumstances, the evaporator tubes inside the containment are subjected to 0.83 MPa external pressure. Therefore, the evaporator tubes must be strong enough to withstand this external pressure as well as any kind of rapid pressure surge that may occur inside the containment: for example hydrogen detonation or molten core-water interaction. The required thickness of the evaporator tube wall and the maximum allowable pressure on the tube wall were determined by meeting pressure vessel requirements of the ASME Boiler and Pressure Vessel Code. The codes for designing shells and tubes under external pressure were used. Requirements for the design are given in Section VIII, UG28, UB98 of the ASME Boiler and Pressure Vessel Code (ASME, 1971). The material used for the evaporator tube is stainless steel SA179 (ASME, 1971) and the external design pressure is assumed to be 0.83 MPa. A tube wall thickness of 0.75 mm is needed. On the other hand, a wall thickness of 1 mm will allow the tube wall to be subjected to an external pressure up to 2.24 MPa, which is almost 2.7 times the design pressure and will give adequate corrosion allowance if one considers that corrosion of the evaporator tube should be negligible because the evaporator tube will always be dry during normal operation. In order to enhance the condensation heat transfer outside the evaporator tube, the outer surface of the evaporator tube will have fins to provide an extended heat transfer area. In this study, radial and axial fin types are considered. A fin enhancement factor of 4 is used, which is defined by Leiendecker et al. (1997) as the ratio of the heat transfer rate for the finned tube to that for the smooth tube. This fin enhancement factor of 4 is selected conservatively allowing for the gap contact conductance between the fin and body of the tube, surface corrosion and fouling layers. The separator in the IEO serves two functions. One is to separate the liquid entrained by steam flow. The second function of the separator is to supply the recirculation head, acting as a water drum. Unlike the closed thermosyphon loop concept, there is no need for a highly efficient separator in the IEO concept since a moderate amount
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of liquid entrainment in the vented steam is tolerable. A gravity steam–water separator consisting of a vertically oriented cylindrical tank is adopted in this project. The separation effect can be determined by calculating the superficial velocity of steam at the two-phase interface. Hence an effective separation dictates a separator with a long water column. Since the equalizing water level in the separator is initially 9.1 m, the superficial vapor velocity at the interface is about 0.3 m/s, which is sufficiently low to ensure effective separation of the entrained liquid in the evaporator tubes. On the other hand, to reduce the water inventory in the evaporator unit, displacer tubes 8 m tall may be installed. In this case the equalizing water level increases somewhat compared to the separator without displacer tubes, i.e. from 9.1 to 9.46 m. In this manner, the heat-up time of the IEO unit to reach the saturation state will be reduced due to a considerable reduction of the water inventory from 30 475 kg for the IEO without displacer tubes to 13 261 kg for the IEO with displacer tubes. In the mixing plenum, the hot recirculation water and the cold feed water are mixed. In order to establish the degree of mixing, a detailed analysis will be eventually required, but in this report it is assumed that complete mixing occurs in the mixing plenum. Two PSCS pools, which are used as a coolant source for the PCCS were originally provided for the secondary cooling system and play the role of a cooling water source for the PSCS. Originally, tank capacity was determined based on the mission time of the PSCS, which is defined by the time duration for cooling down the RCS from the initial condition at PSCS actuation to the shutdown cooling entry condition for all design basis accidents. For this function, PSCS tank capacity was determined as 1211.3 m3 based on 8 h of operation after reactor trip without the need for refill, referring to the auxiliary feedwater system cooldown time of 6–8 h for cooling down the RCS to the shutdown cooling entry condition. However, in fact, the PSCS tank level will be controlled by the water supply system operated by a level controller installed in the PSCS tank.
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sub-compartments were modeled as lumped parameters. Almost all of the thermal conductors are internal thermal conductors except the dome and the cylinder wall of the upper containment. The cylinder wall was modeled as the ‘wall’ and the dome was modeled as the ‘ceiling’ of the upper containment. By this modeling, GOTHIC automatically calculates the distributed surface area of the conductor ‘wall’ in order to fit the size of a sub-cell.
3.1. IEO modeling The conductor model can be adopted as an appropriate model for the IEO. A thermal conductor model requires three modeling parameters: heat transfer area and wall thickness, condensation heat transfer coefficient on the containment side and the heat transfer option on the inner surface of the IEO. Fig. 3. Plan view of the noding scheme for the distributed parameter model of the GOTHIC code.
3. Modeling The lumped parameter model (LPM) devised for GOTHIC has 23 control volumes, 50 flowpaths and 53 thermal conductors. As components, there are two spray pumps and five valves. This LPM is used for the sensitivity analysis objective only and most of performance analysis for the IEO is performed using a distributed parameter model (DPM). A DPM was devised based on refinement of the lumped parameter model. A distributed parameter model can provide a detailed representation of the containment phenomena because the conservation equations within the subdivided volume are solved for the 3-dimensional distributions of velocity, temperature and pressure which represent approximations to actual conditions subject to the density of the noding scheme. In the DPM for KNGR, there are three subdivided control volumes: an upper containment and two outside pools (PSCS Tanks). Figs. 3 and 5 show the upper containment modeled by 105 cells (3x × 5y × 7z). As seen in Fig. 4, the other 21 cells including equipment rooms and steam generator
Fig. 4. Vertical view (B-B) of the noding scheme for distributed parameter model of the GOTHIC code.
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Unlike the closed loop concept, the IEO has no internal structure between the containment atmosphere and the external coolant on the atmospheric side, i.e. the evaporator tube wall is the sole thermal barrier. Hence, the thickness of the evaporator tube wall can be used as the equivalent wall thickness. For the heat transfer area of the equivalent thermal conductor, the inner surface area or outer surface area of the smooth tube can be used. In this study, the inner surface area of the smooth tube is used for the equivalent thermal conductor of the IEO. Hence the heat transfer area used per one IEO unit is the inner surface area of 125.6 m2 of the smooth evaporator tube.
3.1.1. Condensation heat transfer option on the outer surface of the IEO model The condensation heat transfer coefficient on the outer surface of the IEO model is modeled by the reduced Uchida condensation heat transfer coefficient, hred, which is obtained from a best estimate condensation correlation and two multipliers (l, N) to represent the effect of the extended area due to the fins. hred = houtl N(Dout/Din)
(1)
where hout is the nominal value of the condensation heat transfer coefficient, for which the Uchida condensation correlation calculated by GOTHIC is used, l is the ‘fin enhancement factor’ and N is a multiplier (based on the best estimate results obtained through experiment) on the condensation heat transfer coefficient. The term, Dout/ Din is the ratio of tube diameters, which is introduced because the heat transfer area of the equivalent thermal conductor is taken as the inner surface area of the IEO. The best estimate results, which are obtained by experiment, for the condensation heat transfer coefficient, give the multiplier, N, to apply to the nominal condensation heat transfer coefficient value (Liu et al., 1999). Based on this experiment for steam/air, Liu et al. (1999) recommends using the diffusion layer model (DLM) with suction in the containment analysis, which is an option not yet available in GOTHIC. However Liu et al. (1999) also recommends that the 2.2 times Uchida
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Option can be used for engineering design and analysis based on its simplicity and ease of application based on the good agreement of this option with experimental data and DLM predictions. The IEO tube bundles are envisioned to be constructed along the containment walls with integral air baffles to ensure a unidirectional flow field through the bundle. The hot, steam rich mixture will be drawn from higher elevations and directed through the IEO, with the cool, air rich mixture being discharged at lower elevations. This design isolates the flow field in the IEO tube bundle from possible disruption due to unpredictable containment mixing and makes the IEO concept more amenable to analysis. Based on the above consideration, a two times Uchida Option is used for the performance analysis of the IEO in the DBA LOCA scenario, to provide a conservative allowance for margin compared to the experimental results. From the runs to obtain the heat transfer coefficient in the presence of air/helium non condensables, Liu et al. (1999) also suggests utilization of the Uchida air-only correlation with a reduction factor of 20% to be conservative in the air/helium (simulating hydrogen) case as long as the total pressure is B 4.5 atm and there is B 30% hydrogen in the air–hydrogen mixture. Here the one times Uchida Option will be used for modeling the IEO, again with a large conservative allowance. The fin enhancement factor accounts for the effect of the extended surface for the finned tube on the heat transfer rate. The fin enhancement factor can be obtained as the ratio of the heat transfer rate for the finned tube to that for the smooth tube. Here, we will apply a reduction factor defined by the ratio of the fin enhancement factor (again with a conservative allowance) to that obtained by theoretical analysis for the same tube. For this introduced reduction factor, a value of :0.5 appears appropriate given our current level of understanding. Hence the fin enhancement factor of 4, reduced from the prediction of 9.5 for the proposed radial finned tube, will be used for the GOTHIC analysis (Leiendecker et al., 1997).
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3.1.2. Heat transfer option on the inner surface of the IEO The specific temperature option is applied to the inner surface of the IEO model because the bulk temperature of the fluid inside the evaporator tube is at its saturation temperature. Hence the inner surface temperature changes only slightly with axial position. The inner surface temperature of the evaporator tube may be characterized in three time phases. During Phase I, the evaporator tube heat-up process, the inner wall temperature will monotonically increase from the initial temperature (30°C) of the system to some temperature, which is determined by the containment and saturation temperatures at that time. During Phase II, the separator heat-up process, because the saturation temperature which corresponds to the initial water level is not changed, the inner wall temperature can be considered as constant if the containment temperature is fixed. Saturation temperature and the inner wall temperature can be obtained by considering the design temperature of 130°C for DBAs (note the containment temperature is roughly constant at 130°C during the blowdown phase). The saturation temperature is about 120°C and the inner wall temperature is 125°C at the containment temperature of 130°C. We can choose two temperature options for the inner surface temperature; the constant temperature option of 125°C and the variable temperature option obtained from the above considerations. Of these two options, the constant temperature option of 125°C will give the most conservative results in the GOTHIC code because GOTHIC will calculate the heat removal rate conservatively for Phase I and Phase III. Hence this constant temperature option will be used for the GOTHIC code analysis. 3.2. Hydrogen and aerosol layer modeling GOTHIC includes two hydrogen burn models: discrete burn model and continuous burn model (George et al., 1997). In this analysis, the complete burning of hydrogen using the continuous
burn model is assumed. It should be noted that the complete combustion assumption is unrealistically conservative because the inerting effect by steam and the limit of oxygen mole fraction for hydrogen burning are not considered. However this approach will give bounding results. To obtain a first estimate of the effect of the aerosol layer on the IEO performance, a uniform layer of aerosols is simply assumed to cover the IEO outer surface. The surface area is 125.6 m2 per IEO unit, which was used as the reference area of the IEO model in the GOTHIC code. The mass of aerosols is assumed as 50 kg, which is predicted for a in the typical 1000 MWe PWR with a large dry containment of about 105 m3 volume. The thermophysical properties of aerosol were assumed to be those of typical concrete which has a density of 2240 kg/m3 and a thermal conductivity of 1.13 W/mK. For this conservative case, the thickness of the aerosol layer becomes 0.044 mm, which is selected for analysis of the aerosol effect.
3.3. Lumped parameter modeling 6ersus distributed parameter modeling The comparison of the LPM versus the DPM presented below is sensitive to the mixing processes modeled in GOTHIC. In this comparison the IEO design and location is established to ensure a unidirectional flow downward through the bundle as described in Section 3.1.1. Hence while the comparison between models made in this section is applicable to this IEO design since the dominant mechanisms creating this flow field are mechanistically modeled, this comparison between model performance is necessarily applicable only to other dry containment configurations where these same mechanisms are dominant. Fig. 5 shows the difference of the pressure predicted by the GOTHIC 6.0a code between LPM and DPM. The following factors are employed: a fin enhancement factor of 4 for the finned tube model, the specified temperature option of 125°C inside the tube and the two times Uchida Option for the containment side.
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Fig. 5. Comparison of the pressure predicted by LPM and DPM of GOTHIC 6.0a for four IEOs (fin enhancements factor: 4; Specified temperature option for the pool side: 125°C; two times Uchida Option for the containment side). (1), LPM; (2), DPM; Elevation; 53.3 m, Nominal transient: DPM; Elevation; 88.2 m.
Fig. 6. Comparison of the performance of four IEOs with that of four CTSLs for DBA LB LOCA scenario (fin enhancement factor: 4; elevation; 88.2 m). (1), Four CTSL; turbulent convection option for the pool side; one time Uchida Option for the containment side; DPM for GOTHIC 3.4e; (2), Four CTSL; turbulent convection option for the pool side; two times Uchida Option for the containment side; DPM for GOTHIC 3.4e; (3), Four IEO; turbulent convection option for the pool side; one times Uchida Option for the containment side; DPM for GOTHIC 3.4e; (4), Four IEO; turbulent convection option for the pool side; two times Uchida Option for the containment side; DPM for GOTHIC 3.4e. Nominal transient: four IEO; specified temperature option for the pool side: 125°C; two times Uchida Option for the containment side; DPM for GOTHIC 6.0a.
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Two pressure curves for the DPM are plotted in this figure. One curve (with square markers) presents the pressure when four IEOs are assumed to be installed at an elevation of 88.2 m. The other curve (with triangular markers) shows the pressure trend when four IEOs are installed at 53.3 m. As seen in the figure, the pressure obtained using the LPM is higher than that from the DPM during the initial 300 s. This phenomenon can be explained by most of the thermal conductor being distributed in the lower elevation of the upper containment and the atmosphere being forced to be homogeneous and so there is less cooling due to expansion allowed in the LPM run. Hence the pressure for the DPM run is lower than that for the LPM run. After 300 s, however, the pressure for the DPM run is higher than that for the LPM run. This can also be explained by stratification of the steam. The ratio of the air density to the steam density at high level is less than that in the lower level, i.e. some stratification of steam occurs. Therefore the steam removal rate throughout the entire volume will decrease because of low steam density in the lower level where most thermal conductors are located. The above findings differ from previous work which predicted the pressure obtained from the DPM run to be always lower than that from the LPM run (Kim et al., 1998). Based on the same reason, we can explain the sensitivity of IEO performance to the installed location of the IEO. After 300 s, because the steam density in the high level is more than that in the low level, the performance of the IEO which is installed in a high position is better than that of an IEO which is located in the low position. However, the difference in peak pressure is about 13.1 kPa (549.5 kPa vs 562.6 kPa), which is not especially significant. In addition, natural convection circulation is established adequate to permit location of IEO units low in containment, such that the PSCS pools can serve as a source of water or heat sink. All results reported in the next section were obtained using the distributed parameter model.
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four IEOs is much better than that of four TSLs due to the reduced thermal resistance of the IEO compared to the TSL. In addition the two times Uchida Option effect for the IEO is larger than that for the TSL concept. Note, however, that pressure is not reduced by a factor of 2 in 24 h (86 400 s).
Fig. 7. Pressure prediction as a function of the number of IEO units used (fin enhancement factor: 4; two times Uchida Option for the containment side; DPM for GOTHIC 6.0a; Elevation; 88.2 m). (1), Four IEO; specified temperature option for the pool side: 105°C; (2), Six IEO; specified temperature option for the pool side: 105°C; (3), Ten IEO; specified temperature option for the pool side: 105°C. Nominal transient: four IEO; specified temperature option for the pool side: 125°C.
4. Performance analysis of the IEO
4.1. Large break loss of coolant accident (LB LOCA) as a DBA 4.1.1. Comparison of the IEO performance with the closed thermosyphon The performance of four IEOs compared with that of four closed thermosyphon loops is shown in Fig. 6 for the LB LOCA DBA case. For the analysis of the thermosyphon loop (TSL) the equivalent thermal conductor model with the modeling parameters in Leiendecker’s report are used except that the surface area for the single TSL is revised to be 125.7 m2 (based on the outer surface area). This value is then multiplied by a fin enhancement factor of 4. The heat transfer coefficient option of natural convection for the pool side and the one times or two times Uchida Options were used for the containment side. This figure shows that the performance of
4.1.2. IEO performance as a function of IEO units Fig. 7 shows the pressure trend as the number of IEO units is varied. In this analysis the specified temperature of 105°C is assumed, i.e. the water level in the separator is assumed to be as low as possible (e.g. by introducing a level controller). As seen in this figure, if we use ten IEOs, GOTHIC predicts the peak pressure of 0.42 MPa, so we can likely meet the design peak pressure of 0.41 MPa. However it is inherently difficult to meet the second performance criterion which is to reduce the containment peak pressure by half after 24 h of an accident because the temperature difference between the containment and the IEO wall is low in the long term period. From this figure, we can conclude that if the water level and hence the specified temperature is high, as in the present analysis, a large number of IEO units (beyond ten IEO units) may be required to meet the design criteria (peak pressure of 0.41 MPa). Therefore in order to improve the IEO performance, the water level in the separator must be kept as low as possible and the IEO should be installed in the highest position practicable. However the latter option is restricted by the location of the cooling water source, i.e. the water pool. As mentioned in Section 3, KNGR has the PSCS pool fixed at an elevation of 53.3 m. Therefore this approach to increase the performance of the IEO is not acceptable in the present stage of KNGR design. The former approach may be possible if we can introduce a reliable level controller such as a diaphragm valve which is operated by the difference of the water head between the IEO and the water pool, or a float type controller.
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4.2. LB LOCA as a se6ere accident 4.2.1. Hydrogen burning analysis Deflagrations can occur due to hot particles and autoigniting plumes if the atmosphere contains sufficient hydrogen without dense steam. These flammability conditions are typically determined by a hydrogen – air – steam flammability chart (Kumar, 1985; Marshal, 1986). These flammability charts, which are applicable when the atmosphere is about 375 K, shows that the hydrogen concentration must exceed the lower flammability limit (LFL) for a flame to propagate through the mixture for a fixed concentration of steam. Also there is a critical steam concentration of about 50 – 65% beyond which all concentrations of hydrogen are inerted. In this assessment, however, the potential temperature dependence of deflagrations must be recognized. Stamps and Berman (1991) have reviewed the temperature dependence of the lower flammability limit and found that the lower flammability limit decreases about 0.5 vol.%/100 K for upward propagating flames and about 1.0 vol.%/
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100 K for downward propagating flames. The data base for this temperature dependence goes up to about 800 K for dry systems and about 473 K for steam concentrations reaching about 60%. From this review and Kumar’s experimental data, the LFLs for upward and downward propagating flames are given by XH2(UP, XS, T) = 0.37+ 0.2381 XS − 5E(− 5)(T − 373)
(2)
XH2(Down, XS, T) = 0.075+ 0.2381 XS − 1.0135E(− 5)(T −373) (3) On the other hand, examination of the flammability chart (Kumar, 1985) shows that there is a steam concentration beyond which hydrogen concentration is always inerted, regardless of hydrogen concentration. The inerting limit is given by Xstm(inert)= 0.63+ 3E(−4)(T − 373)
(4)
where the temperature derivative, dXstm/dt 3 vol.%/100 K was measured by Kumar over the 373–473 K temperature range.
Fig. 8. Gas concentrations predicted by GOTHIC 6.0a assuming 55% zirconium oxidation.
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Fig. 9. Performance of IEOs for LB LOCA severe accident (fin enhancement factor: 4; specified temperature option for the pool side: 125°C; one time Uchida Option for the containment side; DPM for GOTHIC 6.0a; Elevation; 88.2 m). Base case: No IEO; 100% Zr oxidation; no hydrogen burn; no aerosol deposition. (1), Four IEO; 100% Zr oxidation; no hydrogen burn; aerosol deposition; (2), Four IEO; 100% Zr oxidation; complete hydrogen burn; aerosol deposition; (3), Four IEO; 100% Zr oxidation; no hydrogen burn; no aerosol deposition.
Using this condition we can assess the flammability of hydrogen. As seen in Fig. 8 assuming realistic oxidation of zirconium (55%), hydrogen concentration is about 0.028–0.03 and the steam concentration has the range of 0.64– 0.66 during hydrogen release. On the other hand, the containment temperature is 204°C (477 K) at the peak. Therefore for this temperature and the steam concentration of 0.64 (conservative selection), the LFLs for upward and downward propagation and the inerting limit are given by XH2(UP, 0.64, 477 K) =0.517
(5)
XH2(Down, 0.64, 477 K) =0.217
(6)
Xstm(inert) =0.66
(7)
This evaluation shows that hydrogen cannot be inerted by the steam concentration because the inerting limit is higher than the steam concentration of 64%. However, it is apparent that there is no deflagration because the predicted hydrogen concentration is lower than the lower flammability limits.
Also, hydrogen reacts with oxygen at all temperatures, although the reaction is strongly temperature dependent. Since this type of reaction occurs uniformly through a constant temperature mixture, we call it volumetric combustion. Autoignition can occur if the uniform gas temperature exceeds a critical value. An examination of autoignition criteria for the wellmixed, uniformly heated mixture is summarized by Stamps and Berman (1991). According to Stamps and Berman (1991), the autoignition temperature is over 520°C under the 0% steam and 30% hydrogen mixture condition and this temperature limit increases as steam mole fraction increases. Therefore, there is no possibility of volumetric combustion of hydrogen for our cases, because the hydrogen and steam concentration are calculated to be about 3 and 64%, respectively.
4.2.2. Sensiti6ity analysis of the IEO performance Fig. 9 shows the performance for a LB LOCA severe accident. Here, four IEOs with a fin enhancement factor of 4 are used. The
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Uchida correlation with no modification factor for the outer surface of the tube and the fixed temperature of 125°C at the inside of the tube are used. In addition an aerosol layer of 0.004 mm is assumed to be formed at the outside of the tubes as mentioned in Section 3. Hydrogen is assumed to be produced through 100% oxidation of zirconium in the active core, but its burning is not assumed. As seen in Fig. 9, four IEOs are found enough to easily meet the design criteria of 0.83 MPa peak pressure if we do not include the hydrogen burning effect. In addition Fig. 9 also shows that the long-term pressure can be maintained close to 0.41 MPa. Therefore, a set of four IEOs can meet the design criteria to maintain the containment integrity during severe accidents. In addition, there is a sufficient margin between the design
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criteria and performance results so that it may be possible to reduce the number of IEO units or the size of an IEO system. In Fig. 9, the performance degradation due to the aerosol layer is quite small. In fact, the aerosol layer thickness was calculated as 2 mm if we include fins in the total heat transfer area. Therefore, if we use 2 mm thickness instead of 0.044 mm and more realistic thermophysical data for the aerosol layer than those of concrete, the performance degradation of the IEO due to the aerosol layer will be practically negligible. Fig. 9 also shows the containment pressure prediction when hydrogen is totally and rapidly burned during the hydrogen release period. In this case, the complete burning limits were used. This approach for hydrogen burning is unrealistically conservative since:
Fig. 10. IEO performance as a function of IEO number used for LB LOCA severe accident (fin enhancement factor: 4; specified temperature option for the pool side: 125°C; one times Uchida Option for the containment side; DPM for GOTHIC 6.0a; Elevation; 88.2 m). Base case: no IEO; 100% Zr oxidation; no hydrogen burn; no aerosol deposition. (1), Two IEO; 100% Zr oxidation; no hydrogen burn; aerosol deposition; (2), Four IEO; 100% Zr oxidation; no hydrogen burn; aerosol deposition.
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1. The zirconium oxidation during the melt progression does not exceed 60% (NUREG-1150, 1987), or 50% (Yang et al., 1991). 2. If we use the realistic zirconium oxidation of 55%, there is no possibility of deflagration and autoignition as discussed in Section 4.2.1. 3. If we consider a hydrogen mitigation system such as an igniter and/or hydrogen recombiner, the global burning of hydrogen can hardly occur. Nevertheless, the peak pressure even with a complete burning condition was found to be about 0.76 MPa as seen in Fig. 9. This means that the large dry containment like KNGR can meet the design criteria of 0.83 MPa even under the most conservative assumptions for zirconium oxidation and hydrogen burning.
4.2.3. IEO performance 6ersus number of IEO units in ser6ice Fig. 10 shows the IEO performance as a function of the number of IEO units used. As seen in this figure, there is no significant difference of the pressure during the hydrogen release period as the number of IEO units changes because of small condensation heat transfer coefficients during this period. However, Fig. 10 shows that the long-term pressure can be considerably reduced even with only two IEO units in service. Therefore, it could be possible to meet the design criteria for the severe accident of 0.83 MPa for the long-term period (72 h) even with two IEO units. Also, the peak pressure is lower than the design pressure even under the most conservative hydrogen generation and burning situation as mentioned before.
5. Conclusions In this paper, the IEO concept for the PCCS of a large dry concrete containment was proposed, in which the in-pool condenser is eliminated compared to the closed thermosyphon loop proposed in previous work (Leiendecker et al., 1997). By eliminating the in-pool condenser of a CTSL it is predicted that the thermal resis-
tance can be roughly cut in half in the IEO concept. Furthermore, cost and space requirements should be similarly reduced and problems due to non condensable gas built-up inside the heat transport system can be avoided. In the performance analysis using the GOTHIC code for a LB LOCA as a DBA, the pressure obtained using a LPM is about 13.8 kPa higher than that from DPM during the initial 300 s since the atmosphere is forced to be homogeneous and so there is less cooling due to artificial mixing in the LPM. After 300 s, the pressure from the DPM is about 13.8 kPa higher than that from the LPM because the ratio of air to steam densities at higher level is lower than that in the lower level, i.e. a stratification of steam occurs and so the steam removal rate throughout the entire volume will decrease because of low steam density at the lower level where most thermal conductors are located. In addition, if we use six IEOs assuming that the separator water level is sufficiently low (less than 50 cm), we can likely meet the DBA design peak pressure of 0.41 MPa (g). However, it is inherently difficult to meet another design criterion (to reach half of peak design pressure within 24 h) because the temperature difference between the containment and the IEO wall is low in the long term. For the severe accident scenario, it is apparent that there is no deflagration because the predicted hydrogen concentration is lower than the lower flammability limits and autoignition point of hydrogen. Even with two IEO units, it is possible to meet severe accident design criteria (less than 0.83 MPa) during the long-term period with enough margin. In addition the peak pressure is just 0.76 MPa even assuming 100% zirconium oxidation and the complete burning of hydrogen. Calculation also shows that the fouling effect by aerosols on IEO performance is negligible. Based on the above findings from the performance analyses, we conclude that the IEO has more merit for mitigation of severe accidents than design basis accidents.
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We suggest further development of the IEO modeling for the results to be more realistic. As discussed in Section 3, the fixed temperature option of 125°C at the inside of the internal evaporator was used for the IEO modeling in the GOTHIC code to avoid limitations in the code calculation speed due to Courant’s time limit as discussed in Section 3 and it needs to be improved. The geometrical dimensions of the finned tubes used in the MIT experiments by Liu et al. (1999) were different from the configurations used for the IEO performance analyses. Therefore additional experiments with finned tubes may be required. Any new experiments should focus on parameters impacting IEO performance, for example, the fin effectiveness as a function of geometry and the presence of non condensables, tube bundle effects and tilted tube effects. Eventually a full-scale test of a prototype unit to quantify bundle effects may be needed to qualify the IEO system for service. Furthermore we need to apply more realistic hydrogen production and burning processes to the IEO performance analysis. In this study, we assumed that most of the hydrogen was released during the core melt progression. However, hydrogen can be produced by the interaction between the hot core melt and water in the cavity and also the interaction between the debris and the concrete. With regard to hydrogen burning, hydrogen mitigation systems such as igniters and recombiners should be modeled appropriately in future work.
Appendix A. Nomenclature Din Inner diameter of the evaporator tube Dout Outer diameter of the evaporator tube hred Reduced condensation heat transfer coefficient hout Nominal condensation heat transfer coefficient N Multiplier on the condensation heat transfer coefficient based on the best estimate results obtained through experiment
T X H2 Xstm l
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Temperature of the containment atmosphere Hydrogen mole fraction Steam mole fraction Fin enhancement factor defined as the ratio of heat transfer rate for finned tube to that for the smooth tube
Acknowledgements This study was sponsored by Korea electric power corporation under the Collaborative Program in Power Engineering with Massachusetts Institute of Technology. The lead author wishes to express his appreciation for the opportunity to carry out this project provided by KEPRI. The authors appreciate the help and advice of Young Sang Choi, Brett Mattingly, Haiyang Liu, Yong Hak Kim and Markus Leiendecker. In particular Markus Leiendecker’s previous work was of considerable help in preparation of the numerical analysis code by the lead author and Young Sang Choi motivated the authors to focus on the severe accident analysis. References Ahmad, A. et al., 1983. PWR Severe Accident Delineation and Assessment, NUREG/CR-2666, UCLA-ENG-8284. ASME, 1971. ASME Boiler and Pressure Vessel Code, Section VIII, Div.1. American Society of Mechanical Engineering, USA. de Cachard, F.D. et al., 1997. Thermal-Hydraulic Modeling of Finned Tube Containment Condensers, Annual Report 1997, Annex IV, PSI Nuclear Energy and Safety Research, pp. 49 – 58. George, T.L. et al., 1997. GOTHIC Containment Analysis Package-Technical Manual, Ver. 6.0a, EPRI. IAEA, 1993. Status of Advanced Containment Systems for Next Generation Water Reactors. IAEA-TECDOC, Vienna, p. 752. Kim, Y.H., Todreas, N.E., Driscoll, M.J., 1998. Distributed Parameter Modeling of the KNGR Containment using GOTHIC, MIT-ANP-TR-059. Kumar, R.K., 1985. Flammability limits of hydrogen-oxygendiluent mixtures. J. Fire Sci. 3 (4), pp. 245 – 262. Leiendecker, M., Todreas, N.E., Driscoll, M.J., Hurtado, A., 1997. Design and Numerical Simulation of a Two-Phase Thermosyphon Loop as a Passive Containment Cooling System for PWRs, Vol. I and II, MIT-ANP-TR-053, Rev. 1.
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