FUNDAMENTAL SAFETY OVERVIEW VOLUME 2: DESIGN AND SAFETY CHAPTER S: RISK REDUCTION CATEGORIES

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1 PAGE : 1 / HYDROGEN MITIGATION ASSESSMENT Aspects of the Two-room Containment Due to its significance for hydrogen control, the process of transferring from a two-zone in a onezone containment are briefly reviewed in the present section. The general concept of hydrogen control and monitoring, and associated technical solutions is given in Chapter F.2.4. The technical approach being considered for transforming the containment from two-zone to onezone is based on three types of device: - Rupture foils, - Convection foils with integrated rupture foils - Mixing dampers similar to those used in the ventilation systems. Rupture foils open passively under the action of pressure differences. No missiles are generated by their opening. The foils open at a pressure difference of 50 mbar with a manufacturing tolerance of 30%. The 50 mbar difference is chosen to preserve some margin for pressure fluctuations caused by the ventilation system. The rupture foils are located above the steam generators and cover a large surface area. Convection foils open passively when the gas temperature just below the foil reaches 85 C. They open also at a pressure difference of 50 mbar ± 30 % due to the integrated rupture foils. The convection foils are located above the SG compartment together with the rupture foils and cover a large surface area. The mixing dampers open on a pressure difference of 30 mbar, an absolute pressure of 1.2 bar or in case of a power loss. Also a manual actuation by the operator is possible. The mixing dampers are installed in the wall between IRWST and annulus and cover an appropriate surface area. The doors between the accessible and non-accessible zones of the containment may open in the course of the accident. In that case they would also contribute to the total flow cross section between accessible and inaccessible compartments. Normal doors are modelled to open at a pressure difference of 100 mbar, radiation protection doors at a pressure difference of 500 mbar Justification of hydrogen control concept The ultimate objective of hydrogen control is to prevent the containment integrity from being endangered by the hydrogen combustion phenomena. The key parameter showing the efficiency of the systems implemented to achieve this goal is the pressure load on the containment wall due to these phenomena. Assessment of the pressure load of the containment is performed in 5 steps:

2 PAGE : 2 / 12 a) Calculation of the hydrogen production and the mass and energy release from the RCS into the containment with MAAP for scenarios relevant to the hydrogen risk. The selection of representative scenarios and bounding scenarios is based on a large number of MAAP calculations as described in Chapter S The hydrogen and steam generated during the ex-vessel MCCI phase is calculated with the COSACO code, while steam generation due to corium flooding and quenching is predicted with a phenomenological model (see Chapter S.2.2.4). b) The gas and temperature distributions within the containment for the selected scenarios are calculated with the CFD code GASFLOW. These calculations take into account all the measures foreseen for hydrogen control including recombiners and the measures for containment atmosphere mixing. The calculation is limited to the period of time during which a risk of hydrogen combustion exists, generally from the start of hydrogen release until a homogeneous distribution is achieved and, again, during the early ex-vessel MCCI phase. CFD calculations are used to assess the risk of fast combustion and DDT (see within this Sub-chapter) and the performance of the recombiner units (see within this Sub-chapter). c) Assessment of the hydrogen risk is based firstly on the AICC (Adiabatic Isochoric Complete Combustion) pressure, and secondly on the possibility of flame acceleration and DDT (deflagration detonation transition). The two latter risks are evaluated using the calculated gas concentration and temperature distribution together with criteria derived from tests. The test criteria provide a link between hydrogen related analytical work and experimental findings. A number of criteria deal with the non-occurrence of flame acceleration and hence the risk of fast deflagration. The most important of these criteria is the sigma criterion, which compares the rate of expansion (gas density before non-isochoric combustion divided by the density after this same combustion) with a limiting value derived from a range of experiments run primarily on a tube-type geometry. The sigma criterion may be consolidated by also considering the relationship between the mixture quality and the distance necessary for flame acceleration (run-up distance). Recently, corrective factors have become available to account for radial venting; these factors are important when experiments performed in closed tubes are extrapolated to containment-type geometries. Provided that fast deflagration cannot be excluded, another criterion known as 7 lambda assesses the likelihood of DDT by comparing the gas mixture quality, in terms of the detonation cell width, with the characteristic length of the hydrogen containing cloud, a compartment or of a group of compartments in the containment. The detonation cell width may be determined by interpolating cell lengths derived from tests for different values for the hydrogen and steam concentration, temperature and pressure. d) The hydrogen combustion mode is determined from the distribution of the gases and temperatures on the basis of assuming that ignition may take place at any moment and at any location if the gaseous mixture is inflammable. With this deterministic approach the probability of ignition is ignored and it is conservatively assumed that the ignition occurs just after the maximum accumulation of hydrogen has been reached. In addition, recombination is taken into account. The explicit calculation of hydrogen combustion is necessary to assess temperature changes, particularly at the containment wall. For this purpose the calculation of a laminar type of combustion is sufficient. The GASFLOW code is used since it takes account of the heat transfer to the walls (in contrast to COM3D described below).

3 PAGE : 3 / 12 e) If fast deflagration cannot be excluded by the application of the criteria, an explicit calculation of combustion and the resulting pressure development is necessary. COM3D, a computer code suitable for the analysis of fast deflagrations, is used for this purpose. To perform COM3D calculations, one or more time instants must be chosen for ignition, typically by finding a compromise between the decreasing likelihood of flame acceleration (increasing homogenisation of the atmosphere) and the increasing quantity of hydrogen in the containment (potential energy). For the instant of ignition, the results of the GASFLOW calculations (gas and temperature distribution) are used as the input data for the COM3D calculation. Additionally a location for ignition must be chosen considering the quality of the gas mixture in the vicinity and the pathways available for flame acceleration. The combustion process itself must be analysed for a period of approximately one second. Dynamic effects with high pressure peaks are usually observed in the first tenth of a second. There is no need to calculate the total maximum pressure of the whole combustion process using COM3D as it is covered by the AICC pressure which, similar to the rate of expansion (sigma criterion), is calculated on the basis of the gas and temperature distribution as a function of time. This section gives the results of the calculations of the gas distribution, pressure and temperature loads for the representative and bounding scenarios selected (see 2.1 within this Sub-chapter). In the first instance, the arrangement of the recombiner units is assessed. The calculations presented in this section are based on a one-room containment. However, it can be shown that, with the means briefly described in S , the two room containment can be transformed into a one room containment sufficiently rapidly to make the one-room calculations valid, in principle. Different configurations based on the opening of convection foils (initiated by temperature) and rupture foils (initiated by pressure difference) in the ceiling of the SG compartments, and mixing dampers (initiated by absolute pressure and pressure difference) between the airspace of the IRWST and the lower annular rooms, were evaluated to model the transformation from a tworoom containment to a one-room containment. In the case of a large break LOCA both, rupture foils and convection foils above the SG open together with the mixing dampers, providing the largest possible flow cross section early in the accident. In case of a SBLOCA only convection foils and dampers will open. Opening of the convection foils occurs first above the affected loop, but later at the beginning of core degradation, also above the SGs of the other loops. In order to show that the calculations performed for the one-room containment remain valid, the small break LOCA scenario with late depressurisation was re-calculated with GASFLOW and COM3D to assess the loads due to combustion. In conclusion, it was shown that technical solutions exist, based on convection foils, rupture foils and mixing dampers, which allow transformation to a configuration in which gas mixing, hydrogen recombination and combustion loads are quasi-identical to what would occur in a single zone containment. Apart from the rupture foils, convection foils and the mixing dampers, the catalytic recombiner units are the only technical devices fitted in the containment to control the hydrogen. They remove hydrogen sufficiently quickly, so that at the time of reactor vessel failure, the quantity of hydrogen in the containment is limited to a few hundred kg. In this case, the ex-vessel ignition of hydrogen will not lead to fast deflagration. In addition, the recombiner units help to homogenise the atmosphere. The general arrangement of the recombiner units has been studied using GASFLOW to analyse the behaviour of each individual recombiner unit, examining for different arrangements the effectiveness of the unit and its box temperature, for example. The calculations shown in this section are based on the mass and energy data produced by the MAAP calculations of the scenarios which are shown in the previous sections.

4 PAGE : 4 / Arrangement of the recombiner units The locations of the 47 recombiner units are shown in table F.2.4 TAB 1. This arrangement results from the study of a range of configurations, analysed for the same scenario. The recombination performances of the different configurations were analysed on the basis of the following results: - overall results: mass of hydrogen in the containment, mass of recombined hydrogen, pressure in the containment, history of the recombination rate, etc; - local results: hydrogen distribution in the containment and study of the deflagration risk, temperatures of the recombiner units, concentration of hydrogen at the recombiner unit intake, etc. In the adopted configuration the recombiner units are preferentially installed in the primary component compartments Pressure loads Outline of the scenarios analysed. In order to assess the pressure changes, various representative and bounding scenarios were analysed by following the procedure described in 2.1 within this Sub-chapter. The AICC pressure, which is a purely theoretical upper limit for pressure, is available for all the scenarios while the combustion process itself has only been calculated for the most important scenarios considering the risk of the flame acceleration. The table below lists the scenarios analysed for the pressure and temperature loads. Row one to five show the representative scenarios (framed with a blue line), the three remaining rows show the bounding scenarios (framed with a red line). For some scenarios the thermal loads of the liner was also determined with the laminar combustion model of GASFLOW. Although the main aim of the calculations is to quantity the heat loads, they also highlight the margins to AICC pressure.

5 PAGE : 5 / 12 Scenario AICC Pressure for laminar combustion Dynamic pressure coming from fast deflagration Hydrogen combustion ex-vessel Heat loads due to recombination Heat loads due to combustion PTAEE [LOOP] x 2 cold leg r.s.r* x x x x x 2 hot leg r.s.r* x 3 pressuriser r.s.r x x 2 cold leg r.s.p* x x x x PTAEE [LOOP] with reflooding 2 cold leg r.s.r* Reflooding 2 cold leg r.s.p Delayed depressurisation x x x x x x x x x * rapid secondary cooldown ** partial secondary cooldown Hydrogen distribution The temperature and gas distributions, which are a prerequisite for evaluating combustion loads, were calculated for all the scenarios with the CFD code GASFLOW. The modelling of the containment consists of roughly 100,000 non-equal cells in cylindrical coordinates. GASFLOW has been developed and validated by FZK. The table below gives an overview of the main characteristics of the scenarios studied. The maximum quantity of hydrogen in the containment varies between 450 and 860 kg. Due to the recombiner units, this value is always lower than the quantity of hydrogen released by the primary system. The extent of the recombination depends on the hydrogen release flow rate: if it is released slowly, for example a 2 break LOCA in the cold leg with fast secondary cooldown, the maximum quantity of hydrogen in the containment is considerably lower than the amount released. In the case of rapid release, e.g. small break LOCA with late primary system depressurisation the difference is small. The maximum hydrogen concentration also depends on the steam concentration: it is relatively high for scenarios with a low steam concentration, such as scenarios with fast secondary cooldown; it is relatively low for scenarios with a high steam concentration, as in the case of partial secondary cooldown. The maximum average hydrogen concentration value does not exceed 10.%vol in any of the scenarios. The size of the high concentration cloud (13.%vol) gives an initial indication of the risk of fast deflagration. A maximum cloud volume of m 3 is obtained for two scenarios: 3 break LOCA in the pressuriser and 2 break LOCA with reflooding. In both scenarios, the combustion process is analysed using the COM3D code (see within this Sub-chapter). Generally, the maximum volume of a 13 %vol cloud is reached shortly after the start of hydrogen release when the hydrogen distribution is still relatively heterogeneous, well before the end of release (when the maximum volume of the cloud at 4 %vol is reached). There is no risk of combustion in the case of PTAEE [LOOP] or PTAEE with reflooding because of the small quantity of hydrogen in the now inert containment. These scenarios are therefore not shown in the table below.

6 PAGE : 6 / 12 Maximum mass of H 2 present in the containment (kg) Average H 2 concentration (%vol) Maximum volume of the H 2 13 %vol cloud (m 3 ) 2 hot leg r.s.r* 450 5, cold leg r.s.p* PZR 3 r.s.r cold leg r.s.p r.s.p 2 /D r.s.p 2 /R 860 9, * r.s.r: rapid secondary cooldown ** r.s.p: partial secondary cooldown *** Time expressed on the basis of the start of ex-vessel hydrogen release AICC pressure and predicted combustion pressure The combustion of hydrogen can follow different combustion regimes. For a hydrogen concentration of up to approximately 10 %vol, the expected flame speed is subsonic and the combustion is laminar. The maximum possible pressure peak of a laminar combustion process is supplied by the AICC pressure. However, it is difficult to actually reach this pressure in a containment building, because: - since the speed of the flame is low, a significant fraction of the heat will be transferred to the structures, to the inert gases like steam and to the droplets (departure from an adiabatic condition); - if the hydrogen concentration is lower than 8 %vol, combustion is not complete. The complex geometry of the containment causes incomplete combustion in the case of deflagration. Therefore the AICC pressure and the laminar combustion pressure were calculated. Laminar combustion leads to a quasi-static pressure increase, which means that this increase is isotropic and steady. For a hydrogen concentration above 10 %vol flame acceleration up to the speed of sound has been observed in many experiments. Flame acceleration is supported by structures which generate turbulence (i.e. obstacles in the shape of a ring in tube experiments). The fast combustion of hydrogen gives rise to dynamic pressure loads, or in other words, anisotropic loads which are very variable in time and location. In the worst case the rapid flame may become a detonation, the phenomenon known as deflagration detonation transition DDT. Directly triggering a detonation is not possible within the containment because the high amount of required energy. The pressure peaks are comparable to those of fast deflagration, but the frequency of the pressure waves is higher. The phenomenon of hydrogen combustion in reactor containment is well understood nowadays due to the many experiments which have been performed over the past few years for different conditions. Consequently, criteria are now applicable to directly predict the possible combustion mode based solely on the gas and temperature distributions:

7 PAGE : 7 / 12 - Flame acceleration can only occur if the expansion rate, which represents the ratio of the gas mixture density before combustion to that after combustion, exceeds a threshold value (the "sigma" criterion). The threshold value is worked out directly from the experiments and depends only on the gas concentration and the temperature. - To reach the speed of sound, flame acceleration must last a sufficiently long period of time. In other words, path of the flame must be long enough. This so-called runup distance is a new idea and is consequently not yet well established; it has not been applied in these analyses (this is a conservative assumption). The minimum distance to reach the speed of sound also depends on the gas concentrations. - Finally, the DDT (deflagration to detonation transition) can only take place if the characteristic length of the compartment (or compartments) taken into consideration is greater than seven times the size of the detonation cell which in turn also depends on the mixture quality (lambda criterion). The factor 7 has also been derived from experiments. These criteria define the conditions necessary to produce the phenomena of flame acceleration or DDT. Since they were derived from experiments using closed tubes with a more or less onedimensional geometry, they fail to take account of gas relief perpendicular to the main axis of the flame propagation. When this effect was assessed experimentally using vented tubes, a correction factor for the sigma criterion was developed on the basis of the relative surface area of the vertical openings (not considered in these analyses). This is why these criteria are considered to be very conservative when applied to the geometry of a reactor building. Due to the high level of conservatism described above, the criteria are often violated in the severe accident case, at least locally and temporarily. As a consequence, the combustion process has been calculated for four scenarios using the COM3D computer code, developed by FZK to describe flame accelerations and fast deflagration: Two representative scenarios: - 2 break LOCA in a cold leg with fast secondary cooldown, - 3 break LOCA in the top of the pressuriser with fast secondary cooldown. and two bounding scenarios: - 2 break LOCA with partial cooldown and late depressurisation, - 2 break LOCA with fast secondary cooldown and reflooding. These calculations are described in within this Sub-chapter. In addition to the two bounding scenarios, four representative scenarios were analysed for the AICC pressure: - 2 break LOCA in a cold leg with fast secondary cooldown, - 2 break LOCA in a hot leg with fast secondary cooldown, - 3 break LOCA in the top of the pressuriser with fast secondary cooldown, - 2 break LOCA in a cold leg with partial cooldown only (for the ex-vessel stage).

8 PAGE : 8 / 12 A PTAEE [LOOP] scenario with the loss of all the diesel generators was analysed using the GASFLOW code. However, in the case of scenarios with depressurisation at 650 C, the steam concentration was too high for a global combustion. In addition, two bounding scenarios were analysed: 2 break LOCA with partial cooldown and late primary system depressurisation and small break 2 LOCA with fast secondary cooling and reflooding in the reactor vessel. For all the scenarios, the pressure calculated in the containment in the case of combustion does not exceed the design basis pressure of 5.5 bars. In a single very pessimistic case of small break LOCA (bounding scenario), the AICC pressure may briefly exceed the design basis pressure, but remains in all cases lower than the containment leaktightness pressure limit of 6.5 bars. NB: the AICC pressure curves, unlike the curves relating to the calculated pressure, do not describe the development of combustion, but simply indicate the AICC pressure which would result from combustion at a given moment Evaluation of the combustion mode The possible combustion mode was analysed for all the scenarios selected. The risk of fast combustion is analysed using the sigma criterion, which specifies that there is no risk of flame acceleration as long as the "sigma index" is less than 1. In the case of containment zones with a sigma index greater than 1, flame acceleration cannot be excluded by this approach alone, but does not necessarily occur. The geometric conditions must also favour flame acceleration (obstacles, absence of venting). All the calculations made using COM3D for situations with a sigma index greater than 1 show that large margins remain before getting to rapid combustions with significant loads on the containment wall or even to a DDT. The most severe scenarios are the following: - 3 break LOCA in the pressuriser compartment with fast secondary cooldown (representative scenario), - 2 break LOCA in the cold leg with fast secondary cooldown and in-vessel reflooding at the worst moment (bounding scenario). On the basis of the results of the combustion mode analyses, the cases in need of detailed analysis using the combustion code to determine whether flame acceleration and fast deflagrations actually occur, have been selected. These calculations are shown in within this Sub-chapter Effect of the use of the spray system (EVU [CHRS]) The effect on the hydrogen risk of containment spraying was studied for a bounding scenario (2 break LOCA with late depressurisation of the primary system) using the CFD code GASFLOW. Spray was activated at an unfavourable time when the quantity of hydrogen and the risk of flame acceleration were close to the maximum. This analysis indicates that the use of the spray system does not negatively affect the hydrogen risk.

9 PAGE : 9 / 12 Moreover, the beneficial aspects appear to dominate despite the increase of the hydrogen volume concentration: The homogenisation of the atmosphere (and the reduction of the mass of hydrogen by the recombiners) produces favourable sigma and lambda values in the case of spraying. This means that the risk of flame acceleration appears to be lower in the case of spraying. In addition, the presence of droplets, which is not explicitly taken into account in the sigma-lambda criteria formulae or in the COM3D calculations, helps to further reduce the risk, because: - the flame velocity is lower, - droplets can absorb energy in the case of combustion, lowering the pressure. The main results of this study of the small break LOCA with late primary system depressurisation are as follows: - the use of the spray system reduces locally the risk of hydrogen, because of the homogenisation of the H 2 -air-steam mixture ; - Complete combustion is impossible in both cases, with or without spraying; - The maximum average hydrogen concentration value is lower than 10 %vol with or without spraying. The risk of DDT is only slightly reduced in the case of spraying. With or without spraying, DDT is only possible in a part of the loop compartments for approximately 3 minutes during hydrogen release. During this period of time, the 16 %vol cloud is smaller than 1000 m 3. The cloud size with violated sigma criterion decreases rapidly with the use of the spray system and the period during which the sigma index is greater than unity is shortened by approximately 100 s. While the maximum size of the cloud approaches 12,000 m 3 without spraying, it reaches only 9,000 m 3 with spraying. The consequences of spraying have been studied for only one scenario. Further scenarios will be assessed in detail in future studies. In conclusion, the negative aspect of an increase in hydrogen concentration, which is a slow process limited by the recombiners, is offset by the beneficial effects of better atmosphere mixing Dynamic pressure loads The consequences of fast combustion for the containment shell are analysed using the COM3D computer code. Combustion is modelled on the basis of an eddy break-up model. The initial conditions of the code are drawn from the appropriate GASFLOW calculation; the time and location of hydrogen ignition are chosen on the basis of the flame acceleration criterion. Since simulation of turbulent combustion requires a detailed spatial resolution of the reacting flows, the containment is modelled with a more precise resolution than that needed for GASFLOW (distribution and laminar combustion calculations). COM3D uses the Cartesian cubic calculation mesh with a constant cell size (41*41*41 cm 3 ) to achieve this high level of calculation resolution. The total number of cells in the whole containment is greater than 1,000,000. Four scenarios have been analysed in order to envelope to some extent all the other scenarios involving less risk of flame acceleration according to the criteria:

10 PAGE : 10 / 12-2 break LOCA in the cold leg. Justification: little steam, large quantity of hydrogen. Generally speaking, the combustion is a mild fast deflagration which begins is a relatively violent manner within the affected loop compartment, i.e. the area where the hydrogen concentration is highest, and is then damped as it develops towards the walls of the containment because of the lower concentration of hydrogen in the dome. - 3 break LOCA at the top of the pressuriser. Justification: little steam, stratification. With the exception of short duration pressure peaks of 4 bars, close to the ignition zone, no dynamic effects have been observed on the containment wall, and hence the combustion pressure is enveloped by the AICC pressure for this scenario (4.3 bars). - 2 break LOCA in the cold leg and late depressurisation. Justification: high output rate of hydrogen release and hence temporarily a high local concentration of a large quantity of hydrogen and several release locations. No local high level pressure peaks appear in the containment because the acceleration of the flame inside the steam generator compartment is limited by the lateral venting and deceleration in the dome. The dome seems to have a significant effect, since the risk that the hydrogen combustion could result in unacceptable loads for the containment shell is limited to the period following the start of hydrogen release, when hydrogen distribution is not yet homogeneous and areas of high concentration exist in the primary equipment compartments whilst there is little hydrogen in the dome. Afterwards, when the concentration in the dome increases, the general gas distribution is quite homogeneous with a hydrogen concentration considerably lower than 10 %vol. This effect is attributable to natural and forced convection, the recombiner units being a contributor. - 2 break LOCA in the cold leg with reflooding. Justification: most conservative scenario which combines a low concentration of steam in the containment (because of the fast secondary cooldown) with a high hydrogen concentration peak (because of the reflooding). The dynamic effects are limited and are comfortably below the AICC pressure. Despite a relatively slow pressure development on the containment wall, the pressure differences on the internal structures are quite large. The maximum pressure difference across a wall is 0.4 bar Thermal loads The elimination of hydrogen by recombination or combustion can be the cause of thermal loads to the structures, in particular to the containment liner. The following paragraphs show typical results drawn from calculations carried out using GASFLOW Thermal loads due to recombination In the absence of combustion, the temperatures of the containment walls are determined by the recombination of the hydrogen and the thermal hydraulic processes in the containment atmosphere.

11 PAGE : 11 / 12 For all the scenarios studied (PTAEE [LOOP] with reflooding and small break LOCA with reflooding), the conclusion of the preliminary analysis of thermal loads due to recombination is as follows: although high local gas temperatures may have been observed at the recombiner unit outlets, around 800 C, and the extended areas of the dome may be subjected to gas temperatures of around 200 C, the temperatures on the containment steel liner are in a range which is comfortably lower than the steam saturation temperature at the design pressure Thermal loads due to combustion In the case of hydrogen combustion, higher local temperatures at the containment walls can be expected because the combustion heat may be released in a shorter time period than that of hydrogen recombination. Ignition can occur accidentally. The recombiner units, or sparks from electrical equipment, may be seen as possible combustion triggers. Analysis using the GASFLOW code is based on three scenarios. The first two scenarios deal with the combustion of hydrogen generated in-vessel, while the third concerns the combustion of hydrogen produced ex-vessel in the form of a standing flame above the non-flooded corium: - 2 break LOCA in a cold leg and fast secondary cooldown In all studied cases, combustion is incomplete (less than half the released hydrogen is burned). Consequences of incomplete combustion: containment gas temperature and pressure are well below the values for complete adiabatic combustion. The maximum temperature peak for the containment wall surface is reached in the dome with 145 C, corresponding to a rise of 85 C. - 2 break LOCA in a cold leg with partial secondary cooldown and late depressurisation. The maximum temperature of the containment wall surface is 220 C. - 2 break LOCA in a cold leg with secondary partial cooldown The highest temperatures of the containment wall surface are found at the top of the containment above the compartments of the broken loop. Temperature peaks are lower than 325 C Summary and conclusion The concept of EPR hydrogen control comprises - recombiner units, essentially distributed in the compartments of the primary system complemented by recombiner units fitted in the dome and near the service floor area, - convection foils, rupture foils and mixing dampers allowing the transfer of the two zone containment configuration to a one zone containment configuration. It has also to be pointed out the positive effect of the primary system discharge into the containment atmosphere which results in higher steam concentration and better mixing of gases in the reactor building.

12 PAGE : 12 / 12 The hydrogen control concept was assessed in a consistent manner using up-to-date CFD codes with an adequate nodalisation and by incorporating current knowledge about the combustion phenomena using criteria relating to flame acceleration or DDT occurrence. A number of representative and bounding scenarios were specified to justify the concept of hydrogen control. The main results are: - the average hydrogen concentration never exceeds 10 %vol for all scenarios; - the AICC pressure never exceeds 5.5 bar for representative scenarios; - for a bounding scenario, the AICC pressure may briefly exceed the design basis pressure, but remains in all cases lower than the containment leaktightness pressure limit of 6.5 bar. The actual pressure is much lower since combustion is neither complete (lower hydrogen concentration) nor adiabatic (relatively slow combustion); - the recombination rate is largely independent of the arrangement of the recombiner units; - thermal loads on the containment wall resulting from the recombination process are lower than 120 C; - thermal loads on the containment wall due to combustion of in-vessel generated hydrogen are below the design-basis temperature (170 C) for representative scenarios. The combustion of ex-vessel generated hydrogen may cause the containment wall temperature to rise to 220 C because of the assumed continuous combustion; - dynamic loads on the wall of the containment are low, the AICC pressure covers the actual pressure. Although the combustion flame accelerates in the steam generator compartment, it decelerates in the dome giving rise to small dynamic loads on the containment wall; - according to the initial analyses, the favourable effect of the spray system (enhanced atmosphere mixing) appears to overcompensate the negative effect (increased hydrogen concentration due to steam condensation).