B Moudjed
email: brahim.moudjed@cea.fr

J Excoffon

R Riva

L Rossi

EXPERIMENTAL STUDY OF GAS ENTRAINMENT FROM SURFACE SWIRL

Gas entrainment from surface swirls is characterized using water experiments. A free surface shear flow is generated in an open channel flow. A suction nozzle is set at the bottom of the test section to induce a downward flow and provoke gas entrainment. An important originality of these experiments is the possibility to change the inlet condition so as to generate different turbulent shear flows. This is done by adding obstacles of different sizes and shapes at the end of a flat plate separating the inlet flow from a "stagnant" water area. Velocity fields and profiles, measured with the PIV technique, are provided to describe the inlet conditions corresponding to various geometries and flow rates. Gas entrainment mapping are established from direct observations of the different flow configurations. These new results show that the threshold for the suction velocities required to entrain gas are similar for the configurations with small obstacles and the flat plate configuration triggering a standard shear flow. Increasing the size of the obstacles promotes gas entrainment and reduces the threshold values of the suction velocity to trigger gas entrainment. Shadowgraphy with image processing is used to present new results characterizing the geometrical properties of surface swirls and the quantity of gas entrained. Inlet configurations with obstacles generate larger surface swirls which move upstream from the suction nozzle centre whereas they are situated downstream with the flat plate configuration. Moreover, dimensionless power laws are found to be good approximations for the surface swirl width and the quantity of gas entrained. In addition to provide new insights about gas entrainment in analytical configurations relevant to Sodium cooled fast nuclear reactor, these results should provide different test cases for the validation of MCFD codes, e.g. TRIO-U, NEPTUNE-CFD.

Introduction

Gas entrainment has been identified as a possible safety issue in Sodium cooled fast reactors (NaSFR). It can capture gas from the free surface to the bottom of the vessel. This may lead to the accumulation of gas pockets under the core of the reactor [START_REF] Takahashi | Gas Entrainment at Free Surface of Liquid, (I)[END_REF][START_REF] Tenchine | Some thermal hydraulic challenges in sodium cooled fast reactors[END_REF][START_REF] Tenchine | Gas entrainment issues in sodium cooled fast reactors[END_REF]. When these gas pockets get through the core, it can disturb the nuclear reaction and damage the combustible rods. This gas entrainment phenomenon is produced by the combination of existing turbulent structures close to the free surface, such as eddies, and the downward flows created by the suction of hot liquid sodium by the intermediate heat exchangers. Many experimental facilities have been designed to investigate this phenomenon in relation to the nuclear application. Most of them are water experiments and a brief history can be given. [START_REF] Baum | Gas entrainment at the free surface of a liquid: entrainment inception at a vortex with an unstable gas core[END_REF] begins to study the problem through a cylindrical glass apparatus. Here the eddy structure is generated by introducing the flow with a tangential inlet pipe. The liquid exits through an outlet pipe at the centre of the bottom of the cavity. Four liquids were investigated: water, white spirit, Freon 113 and sodium. They determinate the onset of gas entrainment for the four liquids. Flow mechanisms of air entrainment was also studied by [START_REF] Takahashi | Gas Entrainment at Free Surface of Liquid, (I)[END_REF] through an experimental set up close to that used by [START_REF] Baum | Gas entrainment at the free surface of a liquid: entrainment inception at a vortex with an unstable gas core[END_REF] The novelty of their rig and experiments is the possibility to add a forced circulation by turning the cylindrical cavity. In particular, they showed that an air core with break up bubbles changes to a continuous air core with an additional forced circulation. In these experimental investigations, the gas entrainment is maintained continuously whereas it appears intermittently in NaSFR. Moreover, [START_REF] Kimura | Experimental Study on Gas Entrainment at Free Surface in Reactor Vessel of a Compact Sodium-Cooled Fast Reactor[END_REF] have performed experiments in a partial model of the Japan Sodium cooled Fast Reactor1 to show that gas entrainment results from the combination of the wake of obstacles, the presence of shear flows and the suction flow. Then, to study intermittent gas entrainment, new analytical experimental facilities have been designed so as to isolate the key mechanism of turbulent shear flows combined with flow suction, e.g. [START_REF] Ezure | Experimental investigation on bubble characteristics entrained by surface vortex[END_REF], 2008[START_REF] Kimura | Experimental Study on Gas Entrainment Due to Nonstationary Vortex in a Sodium Cooled Fast Reactor -Comparison of Onset Conditions Between Sodium and Water[END_REF][START_REF] Cristofano | Experimental study on unstable free surface vortices and gas entrainment onset conditions[END_REF]. These new design experiments are based on open channel flows. To produce turbulent structures, a shear flow is generated in the middle of the channel. This is realized by separating the main flow from a water dead zone using vertical plates at the inlet and outlet of the channel. The downward flow is generated by flow suction through a nozzle placed at the bottom in the centre of the channel. This combination can produce (or not) gas entrainment through the suction nozzle according to the chosen flow parameters. Such experiments provide the occurrence map of gas entrainment according to flow parameters, i.e. inlet velocity, outlet velocity and suction velocity. Moreover, [START_REF] Kimura | Experimental Study on Gas Entrainment Due to Nonstationary Vortex in a Sodium Cooled Fast Reactor -Comparison of Onset Conditions Between Sodium and Water[END_REF] have investigated both sodium and water fluids using the same test section to demonstrate that water experiments simulated well the onset conditions of gas entrainment in sodium. Gas entrainment experiments have also been performed in liquid metal, e.g. [START_REF] Vogt | Detection of gas entrainment into liquid metals[END_REF] to support the validation of numerical codes. It is worth mentioning that gas entrainment is also a topic relevant to the automobile industry in the context of the water box [START_REF] Recoquillon | Etude expérimentale et numérique des écoulements diphasiques dans la boîte à eau d'un véhicule automobile[END_REF]Recoquillon et al., 2011aRecoquillon et al., , 2011b)).

The present study is an experimental investigation of the gas entrainment occurrence in water. The geometry of the experimental channel is close to that of [START_REF] Ezure | Transient Behavior of Gas Entrainment Caused by Surface Vortex[END_REF][START_REF] Kimura | Experimental Study on Gas Entrainment Due to Nonstationary Vortex in a Sodium Cooled Fast Reactor -Comparison of Onset Conditions Between Sodium and Water[END_REF] The main difference lies in the inlet conditions. To trigger different flow regimes, obstacles of various shapes are added (or not) at the end of the flat plate defining the inlet conditions. The experimental set up and methods are described in section 2. The results are discussed in section 3. New results about gas entrainment introduced for the different inlet conditions are investigated. First, the required condition on the inlet and suction velocities for the occurrence of gas entrainment are studied and compared to previous literature. Second, the geometrical properties of surface swirls are characterized. Third, the projected areas of the air pumped via bubbles entrained through the suction nozzle are measured. This permits to describe and compare the behaviour of surface swirls and the quantity of air entrained according to different inlet configurations.

Experimental approach

Experimental setup

Experiments are performed using an open channel water flow. This channel is integrated in a close loop. A centrifugal pump mounted with an adjustable by-pass is used to generate different regulated flow rates. Flow rates are measured by electromagnetic flowmeters with an accuracy of 0.3%. Moreover, the temperature of the water is regulated at about 20°C using a heat exchanger. The length and width of the test section are respectively 700mm and 108mm. The height of the free surface is 210mm. Figure 1 gives a sketch of the test section of the rig which is named BANGA.

To separate the main free surface flow from a stagnant water area, two vertical plates (300mm apart) are placed in the middle of the channel so that inlet and outlet flows are limited to the half of the channel cross-section. To induce a secondary down-flow, some water is pumped through a suction nozzle. The diameter of the nozzle, noted 𝐷, is of 50mm and it is set at 100mm height, i.e. 110mm below the free surface. The origin of the Cartesian frame is set on the border of the vertical plate at the height of the suction nozzle. The x axis coincides with the horizontal channel flow direction, y and z axis are the horizontal and vertical directions, respectively.

The present experimental set up is relatively close to that of [START_REF] Ezure | Transient Behavior of Gas Entrainment Caused by Surface Vortex[END_REF][START_REF] Kimura | Experimental Study on Gas Entrainment Due to Nonstationary Vortex in a Sodium Cooled Fast Reactor -Comparison of Onset Conditions Between Sodium and Water[END_REF] One of originality of this work is the possibility to introduce obstacles. These obstacles are mounted at the end of the first vertical plate as illustrated by the square in Figure 1. They produce different inlet flow conditions amenable to impact on the structure of the turbulent shear flow so as to provoke different conditions for gas entrainment.

Three geometries of obstacles are selected: square blocks (noted SB), half cylinders (noted HC) and ramps (noted RP). The square block and half cylinder are placed at the end of the vertical inlet Flat Plate (noted FP). The ramps are mounted along the flat plate with the highest part at its downstream end. Figure 2 illustrates these different configurations with the obstacles. The height of the obstacles, noted ℎ 𝑜𝑏𝑠 , are either 12.5mm or 25mm.

Notations

The inlet and suction flow rates (in m 3 h -1 ) are 𝑄 𝑖𝑛 and 𝑄 𝑠 respectively. The inlet and suction velocities are 𝑈 𝑖𝑛 and 𝑉 𝑠 respectively. 𝑈 𝑜𝑏𝑠 is the velocity at the cross section perpendicular to the obstacles.

𝑈 𝑖𝑛 = 𝑄 𝑖𝑛 𝐿 𝑖𝑛 ℎ 𝑤𝑎𝑡 , (1)
𝑉 𝑠 = 𝑄 𝑠 𝜋𝐷 2 /4 , (2)
𝑈 𝑜𝑏𝑠 = 𝑈 𝑖𝑛 𝐿 𝑖𝑛 𝐿 𝑖𝑛 -ℎ 𝑜𝑏𝑠 , (3)
where 𝐿 𝑖𝑛 is the width of the inlet channel (50mm) and ℎ 𝑤𝑎𝑡 is the water depth (210mm). For the flat plate configuration ℎ 𝑜𝑏𝑠 = 0.

Similarly to [START_REF] Chang | Scaling law for air-entraining vortices[END_REF], the diameter 𝐷 of the suction nozzle is chosen as the reference length scale to build dimensionless numbers. In particular, the dimensionless coordinates are noted with a * so that: 𝑥 * = 𝑥 𝐷 ⁄ , 𝑦 * = 𝑦 𝐷 ⁄ and 𝑦 * = 𝑧 𝐷 ⁄ . The inlet and suction Reynolds number, respectively noted 𝑅𝑒 𝑖𝑛 and 𝑅𝑒 𝑠 , are defined as:

𝑅𝑒 𝑖𝑛 = 𝑈 𝑜𝑏𝑠 𝐷 𝜈 , (4)
𝑅𝑒 𝑠 = 𝑉 𝑠 𝐷 𝜈 . ( 5
)
where 𝜈 ≅ 10 -6 𝑚 2 𝑠 -1 is the kinematic viscosity at 20°C. In order to underline the competition between the free surface flow and the downward flow in gas entrainment and surface swirl occurrence, two parameters, 𝜅 and 𝜉 are introduced:

𝜅 = 𝑅𝑒 𝑖𝑛 2 𝑅𝑒 𝑠 . (6) 𝜉 = 𝑅𝑒 𝑠 2 𝑅𝑒 𝑖𝑛 . (7)
Later on 𝜅 is used to study phenomena where 𝑅𝑒 𝑖𝑛 has a stronger influence than 𝑅𝑒 𝑠 . In contrary, 𝜉 is used to describe phenomena where 𝑅𝑒 𝑠 as a stronger influence than 𝑅𝑒 𝑖𝑛 .

Measurements and processing

Flow measurements rely mainly on Particle Image Velocimetry (PIV) measurements to provide details about the inlet conditions, e.g. velocity profiles, and on flow visualization and processing to observe and quantify the gas entrainment phenomenon.

Particle Image Velocimetry (PIV) measurements [START_REF] Raffel | Particle Image Velocimetry[END_REF][START_REF] Rossi | Lamination, stretching, and mixing in cat's eyes flip sequences with varying periods[END_REF] are performed using an Evergreen™ laser with double pulse laser sheets (thickness, wave length, duration and energy: 1mm, 532nm, 5ns, 120mJ). Double frames are recorded using a PowerView™ Plus Charge Coupled Device (CCD) camera of spatial resolution 4872 x 3248 pixels. The time lags of the double frames vary from 1ms to 2.5ms according to the flow rates. PIV computations are performed using Davis™ software from Lavision™. The final correlation windows are square windows of 48 pixels sides. An overlap of 50% is applied to build the velocity grid. Typical maximal displacements are about 10 pixels. The final instantaneous velocity fields are smoothed using a 3x3 mobile average.

The velocity fields are measured in the inlet area within horizontal and vertical planes. The vertical PIV planes are situated at y = 32mm and y = 37.5mm for 12.5mm size obstacles and the flat plate and for the 25mm side obstacles, respectively. These y values correspond to the middle position between the obstacles and the channel wall. The horizontal measurements are performed at z = 50mm (𝑧 * = 1) for all configurations. These latest measurements are performed from the side so as to avoid free surface deformations. To do so, the camera is set with an angle of 30° with the laser sheet. A calibration is applied to PIV frames to correct distortions using a reference grid and the ortho-rectification algorithm of DAVIS™ software.

Visualizations are performed in white backlighting with a Stemmer Imaging™ camera to observe surface swirl occurrence and gas entrainment phenomenon. Frames, with a spatial resolution of 2048 x 1088 pixels frames are acquired at 25Hz during 15min. The physical domain observed is 36cm long and 16 cm wide. It includes the free surface and the suction nozzle.

Image processing toolbox of Matlab™ is used to identify and quantify the occurrence of gas entrainment as follows:

1. The reference background, i.e. without gas entrainment, is subtracted to the original picture (Figure 3a) to improve the contrast. Such results are illustrated by Figure 3b.

2. The edges of the air areas are detected using the Canny method [START_REF] Canny | A Computational Approach to Edge Detection[END_REF] evaluating the gray scale gradient. This step is illustrated in Figure 3c.

3. The edges are dilated to close opened contour. This step illustrated by Figure 3d.

4. The thick lines obtained are filled as illustrated in Figure 3e.

5. The filled shapes are eroded. This processing cleans the dilated border and removes the gray scale gradient of the free surface and the suction nozzle. This step is illustrated in Figure 3f.

6. Iso-countours are extracted (from the black and white image obtained) and are characterized as polygons so as to measure their geometrical characteristics. Resulting contours are illustrated in Figure 3g where they are superposed to the original picture.

Flows explored

The Table 1 gives the flow rates, velocities and Reynolds numbers investigated.

Q in m 3 /h U in m/s Q s m 3 /h V s m/s Re s
Re in obstacles h obs FP 12.5mm 25mm

1.9 0.05 1.9 0.27 13 500 3 333 5 000 2 500

3.8 0.1 2.8 0.4 20 000 6 667 10 000 5 000 3.2 0.45 22 500 5.7 0.15 3.6 0.5 25 000 10 000 15 000 7 500 4.2 0.6 30 000 4.9 0.7 35 000 7.6 0.2 4.2 0.6 30 000 13 333 20 000 10 000 5.7 0.8 40 000 7.1 1 50 000 9.5 0.25 4.2 0.6 30 000 16 667 25 000 12 500 5.7 0.8 40 000 7.1 1 50 000 ⁄ , for the 25mm side square block; ‖𝒖‖ is the velocity amplitude. Figure 4a is measured at y = 37.5mm and Figure 4b 2 is measured at z = 50mm. The two fields are in a good agreement. This highlights the reproducibility of the measurements since they are not made simultaneously. The black vectors indicate the direction of the flow. At the inlet, x < -50mm, the velocities are mainly parallel to the xaxis and their intensities are close to 𝑈 𝑖𝑛 . For x > -25mm, the local speed increases due to the contraction of the test section. For x > 0, a strong transverse velocity gradient occurs between the inlet and the stagnant flows (Figure 4b). This generates a shear flow in the wake of the obstacle.

Figure 5 gives the dimensionless velocity profiles extracted along x = 0. The confidence interval at 95% of these mean velocity profiles is about 1%.

In the vertical plane, the vertical component (Figure 5b) is clearly negligible compared to the axial component (Figure 5a). In the horizontal plane, values of the transverse component (Figure 5d) above y = 25mm for 25mm side obstacles and y = 12.5mm for 12.5mm side obstacles are close to zero. Under these y values, the transverse component is more important due to the wake of obstacle and recirculation flow. Moreover, the axial velocity jumps in Figure 5c correspond to the position of the shear flows between the inlet and the "stagnant" flow. As for the velocity fields, a good agreement appears between the vertical and the horizontal velocity profiles of the axial component (Figure 5a and Figure 5c). The amplitudes of the velocity profiles are ordered similarly for the two sizes of obstacles. The highest velocities are found for the inlet configuration with the 25mm radius half cylinder.

The turbulence intensity levels range from 5% to 11% according to the inlet configurations. The velocities in the region of the suction nozzle are provided to complete the description of the experiments. Figure 6 gives the mean dimensionless velocity field, ‖𝒖‖ 𝑉 𝑠 ⁄ , for the 12.5mm side square block in the vertical plane at y = 5mm. This plane is close to the centre-axis of the suction nozzle. The flow inside the suction nozzle is not symmetric. This indicates that it is sensible to the mean horizontal flow in the channel. In particular, a wake effect appears in left part, for -0.5 < 𝑥 * < 0, when the flow goes into the suction nozzle. Velocity profiles are plotted in Figure 7a for the axial component and in Figure 7b for the vertical component along the white dashed line represented in Figure 6. These profiles highlight the non-symmetrical distribution of the flow inside the suction nozzle. The suction velocity profiles are close to a plateau for -0.2 ≤ 𝑥 * ≤ 0.4 with a ratio of the maximum of velocity to the mean velocity 𝑉 𝑠 varying between 1.2 and 1.4.

Results and discussions

Gas entrainment mapping

Direct observations of the flow are performed to check if some gas is entrained or not. The onsets of gas entrainment are deduced from these observations.

The Figure 8 compares the present results with the ones obtained by [START_REF] Kimura | Experimental Study on Gas Entrainment Due to Nonstationary Vortex in a Sodium Cooled Fast Reactor -Comparison of Onset Conditions Between Sodium and Water[END_REF] in the case of experiments with a flat plate. The results are close.

The Figure 9 provides gas entrainment mapping for the six obstacles according to the inlet velocity, 𝑈 𝑖𝑛 , and the suction velocity, 𝑉 𝑠 . When gas entrainment occurrences are observed the symbols are filled. The boundaries between the domains of gas entrainment and no gas entrainment are represented with red dashed lines. These thresholds change according to the shape and the size of obstacles. It is observed that the large obstacles trigger the occurrence of gas entrainment for lower suction velocities than for the small obstacles for a given 𝑈 𝑖𝑛 . The thresholds of gas entrainment are summarized in Figure 10 for all experimental configurations including the flat plate case. The cases can be ranked according to the 𝑉 𝑠 required to provoke gas entrainment for a given 𝑈 𝑖𝑛 . From the smallest to the highest 𝑉 𝑠 , this ranking is given in Table 3 along with the average ratio, noted ℜ 𝑠 , of the threshold suction velocities to the one of the flat plate configuration. Increasing size of obstacles promotes gas entrainment and reverses the order of obstacles' thresholds in term of suction velocity. In the case of the 12.5mm size obstacles, gas entrainment occurs with a smaller value of 𝑉 𝑠 for the ramp than for the square block and the half cylinder. In the case of 25mm size obstacles, the ramp needs a higher value of 𝑉 𝑠 to induce gas entrainment than for the half cylinder and the square block. Moreover, small obstacles trigger gas entrainment for suction velocity close to the flat plate since values of ℜ 𝑠 are close to 1. However, increasing size of obstacles triggers gas entrainment for smaller 𝑉 𝑠 than the flat plate configuration.

These results show that the occurrence of gas entrainment depends both, on the structure and on the intensity of the turbulent flow produced by the inlet conditions. The differences observed for different inlet configurations may offer different test cases for MCFD codes.

Quantifications of gas entrainment and surface swirls properties

The two modes of gas entrainment observed are:

 air bubbles are pumped through the suction nozzle 3 ,  a continuous surface swirl reaches the suction nozzle. Figure 11 illustrates these two observed modes for gas entrainment.

As the surface swirl is key to trigger gas entrainment, this subsection describes the geometrical properties of the surface swirl generated for the different flow configurations before discussing and quantifying the gas entrainment.

Geometrical characteristics of the surface swirls

The Figure 12 gives three snapshots which illustrate the positions of the surface swirls varying the inlet configurations. The surface swirls positions are upstream or downstream the centre of the suction nozzle according to the inlet configurations.

Using image processing, described in section 2.3. , the spatial positions of the surface swirls are measured at 𝑧 * = 2 (i.e. 10mm bellow the free surface level) and at 3 These bubbles are issued from the breaking of a surface swirls which results in a non-continuous surface swirl as illustrated in Figure 11a.

𝑧 * = 1 as highlighted by the dashed white lines in Figure 12.

The dimensionless position of the center of the surface swirls is introduced as:

𝑥 𝑠𝑠 * = (𝑥 𝑠𝑠 -𝑥 𝑛𝑜𝑧 ) 𝐷 ⁄ , (8) where 𝑥 𝑠𝑠 and 𝑥 𝑛𝑜𝑧 denotes the x-axis positions of the centres of the surface swirls and the suction nozzle, respectively.

The probability distributions of 𝑥 𝑠𝑠 * at 𝑧 * = 2 are given in Figure 13 for different inlet configurations and flow intensities.

These distributions are not far from Gaussian distribution. The mean and the standard deviation of these positions' distribution varies with the obstacles, 𝑅𝑒 𝑖𝑛 and 𝑅𝑒 𝑠 . For the flat plate case (Figure 13a), the surface swirl tends to be downstream of the suction nozzle centre with 𝑥 𝑠𝑠 * ≥ 0. For the cases with obstacles the surface swirl tends to be upstream of the suction nozzle with 𝑥 𝑠𝑠 * ≤ 0. For the 12.5mm size obstacles, the highest probabilities are

(b) (a)

observed around 𝑥 𝑠𝑠 * ≈ -0.5 for 𝑅𝑒 𝑖𝑛 ≤ 13333. However, for higher Reynolds numbers, they are around 𝑥 𝑠𝑠 * = 0. For the 25mm size obstacles, the highest probabilities locations are more sensitive to 𝑅𝑒 𝑖𝑛 values than for 12.5mm size obstacles.

To complement the probability distributions of Figure 13, the temporal average of, 𝑥 𝑠𝑠 * , noted 𝑥 𝑠𝑠 * ̅̅̅̅ , and their standard deviations, /𝐷, are given for 𝑧 * = 2 and 𝑧 * = 1 in Figure 14.

The position 𝑥 𝑠𝑠 * is smaller for large obstacle than for small obstacle. For the inlet configurations with obstacles, 𝑥 𝑠𝑠 * tends to be negative for large obstacle at 𝑧 * = 2. The spreading of 𝑥 𝑠𝑠 * values is larger for 𝑧 * = 2 than for 𝑧 * = 1, in particular for the configuration with obstacles. This is due to the capture of the core of the surface swirls by the suction outlet as illustrated in Figure 12. It can be noted that for the configurations with small obstacles, 𝑥 𝑠𝑠 * tends to be negative for 𝑧 * = 2 and positive for 𝑧 * = 1. For the configuration without obstacles, 𝑥 𝑠𝑠 * > 0 for 𝑧 * = 1 and 𝑧 * = 2. (a)

The thicknesses of the surface swirls are exported at 𝑧 * = 2 and 𝑧 * = 1 to build the Figure 16. It gives the average values of the dimensionless thicknesses, 𝐿 𝑠𝑠 * = 𝐿 𝑠𝑠 𝐷 ⁄ , versus 𝜅 𝜅 𝑐 ⁄ . The parameter 𝜅 𝑐 is a critical value of 𝜅 identified as the transition between two regimes for the growth of 𝐿 𝑠𝑠 * . Its value is about 5500.

The configurations with the large obstacles generate the largest diameters and the one with the flat plate the smallest.

The thickness 𝐿 𝑠𝑠 * increases like power laws of 𝜅 𝜅 𝑐 ⁄ , i.e. 𝐿 𝑠𝑠 * ~(𝜅 𝜅 𝑐 ⁄ ) 𝛽 , when 𝜅 𝜅 𝑐 ⁄ ≤ 1. The mean values of the exponents β are about 1 and 0.7 for 𝑧 * = 2 and 𝑧 * = 1 respectively. Fluctuation of the experimental points around these laws are typically about 8% at 𝑧 * = 2 and 16% at 𝑧 * = 1 which underline the good quality of this power law approximations. The exponents 𝛽 are estimated using least square approximations. They are given in Table 5 for the different flow configurations along with the normalized standard deviations of the experimental points around these laws. For 𝜅 𝜅 𝑐 ⁄ ≥ 1, 𝐿 𝑠𝑠 * does not increase significantly with 𝜅 𝜅 𝑐 ⁄ and it can be described as fluctuations around asymptotic values. Here again, a notable difference appears between the flat plate and the obstacles at 𝑧 * = 2 since power 𝛽 of the flat plate configuration is twice lower than the other configurations. However, at 𝑧 * = 1, the values of 𝛽 are more homogenous between the different configurations. Table 5. Exponents of the power law approximation 𝑳 𝒔𝒔 * ~(𝜿 𝜿 𝒄 ⁄ ) 𝜷 and normalized standard deviations of the dispersion of the experimental points around these laws, noted disp, in percent.

Total area of the entrained gas

The total area, 𝐴 𝑔𝑎𝑠 , of the air entrained through the suction nozzle is measured for the different inlet configurations.

Given the depth of the suction nozzle (100mm) and the range of suction velocities (see Table 1), the 25Hz frequency of the camera is sufficient to capture all bubbles during their crossing of the suction nozzle. However, the air bubbles can collapse and change shape between to frames. This makes bubble tracking methods difficult to implement. Likewise, one difficulty to estimate 𝐴 𝑔𝑎𝑠 is to count all the bubbles and to not count twice the same bubbles in two different frames. Taking into account the area of bubbles included in a subdomain of constant size introduce a bias in the measurement. In particular, considering the entire suction nozzle overestimates 𝐴 𝑔𝑎𝑠 since bubbles are counting twice.

When the bubbles are crossing down the suction nozzle, two opposite forces/phenomena are applied to bubbles: the downward suction flow and the buoyancy force. Moreover, the flow is accelerated at the entrance of the suction nozzle before being quasi-established further downstream. This is particularly important for large bubbles that, for example, can sometimes enter the suction nozzle and leave it from the top without being entrained by the suction flow. Nevertheless, these events are never observed when the bubbles reach the bottom part of the suction nozzle. In consequence, the areas of the air bubbles are counted over a measurement domain starting from the bottom of the suction nozzle with a depth depending on the flow.

Between two frames the displacement of the bubbles is about 𝑉 𝑏𝑢𝑏 𝑓 ⁄ where 𝑉 𝑏𝑢𝑏 is the speed of the bubble and 𝑓 is the frequency of the camera. The domain of the suction nozzle where the area of bubbles is measured is adjusted to the flow conditions by roughly estimating 𝑉 𝑏𝑢𝑏 𝑓 ⁄ . To do so the suction velocity and the terminal velocity of a spherical bubble of equivalent diameter are used. Without the buoyancy effect, bubbles are considered to have a displacement about 𝛾 𝑉 𝑠 𝑓 ⁄ between to frames. The parameter 𝛾 is an empirical coefficient which relates the fact that the velocity is not uniform within the suction nozzle (see Figure 6 and Figure 7). For each suction velocity, 𝛾 is fixed according to the measured velocity profiles to 𝛾 = 1.25. The suction velocity is corrected by the terminal velocity, noted 𝑉 𝑏𝑢𝑜𝑦 . For each bubble, the value of 𝑉 𝑏𝑢𝑜𝑦 is extracted from [START_REF] Haberman | An experimental investigation of the drag and shape of air bubbles rising in various liquids[END_REF] as a function of the bubble radius 𝑟 𝑏𝑢𝑏 . The displacement considered for a bubble is then (𝛾𝑉 𝑠 -𝑉 𝑏𝑢𝑜𝑦 (𝑟 𝑏𝑢𝑏 )) 𝑓 ⁄ .

Hence, the area of a bubble, 𝒜 𝑏𝑢𝑏 , calculated in a frame is taken into account in 𝐴 𝑔𝑎𝑠 calculation if the The total area of entrained air in a recorded frame is defined by: 𝐴 𝑔𝑎𝑠 = ∑ 𝒜 𝑏𝑢𝑏 (9)

Figure 18 provides the dimensionless entrained gas area 𝐴 𝑔𝑎𝑠 𝐷 2 ⁄ versus . The inlet configurations with obstacles entrain more air than the flat plate configuration. Among the cases with obstacles, the quantity of air pumped with large obstacles is greater than with small obstacles. In particular, for the two sizes, square blocks generate more entrained air than the half cylinder. The greater quantity of entrained air is induced by the large ramp whereas the smallest quantity of air is entrained by the small ramp, so that the ranking is inversed between the two sizes. It can be observed that the large obstacles generate both larger surface swirls and higher quantity of entrained gas. Likewise, it is noticed that obstacles with a same size are ranked in the same way in term of threshold and entrained gas area.

Variation of dimensionless entrained gas area 𝐴 𝑔𝑎𝑠 𝐷 2 ⁄ is approximated as power law 𝜉 𝜃 fitted to the experimental measurements with the least square method. Adding obstacles change significantly gas entrainment behaviour since the slope of the flat plate is very larger than cases with obstacles (see Table 6).

The total volume 𝑉 𝑔𝑎𝑠 of entrained air is also estimated adding each measured bubble area to the power of 3/2:

𝑉 𝑔𝑎𝑠 = ∑ 𝒜 𝑏𝑢𝑏 3 2 ⁄ 𝑁 𝑖=1
, with N the number of bubbles. The associated flow rate, 𝑉 𝑔𝑎𝑠 𝑡 𝑣𝑖𝑠𝑢 ⁄ , calculated on the visualisation time 𝑡 𝑣𝑖𝑠𝑢 (15min), is plotted in Figure 19. As for the area 𝐴 𝑔𝑎𝑠 , experimental set of data are fitted with a power law 𝜉 𝛾 which exponents are provided in

Conclusion

The objective of the present study is to investigate gas entrainment through an experimental set up in water. Since wakes of obstacles are expected to be present in Sodium cooled fast nuclear reactor, obstacles, with different shapes and sizes, are added at the channel inlet to produce various inlet conditions. These new results reveal the dependence of the gas entrainment to the triggers generating different structures of the turbulent flows for similar bulk properties, i.e. flow rates. These results complement the existing knowledge issued from the experiments using the flat plate configuration and open new perspectives.

The measurement of the suction velocity threshold required to provoke gas entrainment show that small obstacles behave similarly to the standard shear flow generated using the flat plate configuration. The configurations using the large obstacles tend to provoke gas entrainment for lower values of the suction velocity than the other configurations. Moreover, the ranked of the obstacles in terms of suction velocity required to entrain gas (at a given inlet velocity) is inversed between the two sizes of obstacles.

To study the details of the flow physical mechanisms, the occurrences and behaviour of the surface swirls are described. Moreover, in the context of nuclear safety for RNR-Na, it is important to quantify the volume of argon entrained in sodium through the heat exchanger. Then, an imaging processing tool is developed to detect the contour of surface swirl and entrained bubbles from shadowgraphy.

The different inlet configurations influence the width and the position of surface swirls. To our knowledge, the quantification of geometrical properties of the surface is a new result in the study of gas entrainment in open channel flow experiments. For the configuration with obstacles, they are larger and situated upstream from the suction nozzle. For the flat plate configuration the surface swirls are downstream the suction nozzle. It is found that for all cases, their widths vary with a power law of 𝜅. Surface swirl width increases with 𝜅 for small obstacles and the flat plate until a critical value of 𝜅 at 5500. From this value, it fluctuates around an asymptotic value for large obstacles.

The quantity of entrained gas is also estimated using a method taking into account suction velocity and buoyancy effect on bubbles. Here again, a power law of the entrained gas area was found. It implies the dimensionless parameter 𝜉 where the role of Reynolds numbers is inversed compared to the surface swirl width law. The flat plate differs clearly from the cases with obstacles which are assembled around a same power. Number of bubbles change with values of 𝜉 and several sizes of bubbles are entrained through the suction nozzle.

To conclude, adding obstacles in the channel inlet area modifies significantly the gas entrainment and the surface swirl behaviour. In addition to provide new insights, it is expected that the detailed study of the configurations presented in this manuscript should offer interesting test cases for numerical codes validation.

Figure 1 .

 1 Figure 1. Test section of the rig BANGA: The top picture is the side view and the bottom one is the top view. The origin of the Cartesian frame is indicated with a cross. 𝑄 𝑖𝑛 and 𝑄 𝑜𝑢𝑡 denote the inlet and outlet flow rates, respectively.

Figure 2 .

 2 Figure 2. Illustration of the different configurations with obstacles.

Figure 4

 4 Figure4provides 2D dimensionless mean velocity field, averaged over 250 velocity fields, in the inlet area, ‖𝒖‖ 𝑈 𝑖𝑛 ⁄ , for the 25mm side square block; ‖𝒖‖ is the velocity amplitude. Figure4ais measured at y = 37.5mm and Figure4b2 is measured at z = 50mm. The two fields are in a good agreement. This highlights the reproducibility of the measurements since they are not made simultaneously. The black vectors indicate the direction of the flow. At the inlet, x < -50mm, the velocities are mainly parallel to the xaxis and their intensities are close to 𝑈 𝑖𝑛 . For x > -25mm, the local speed increases due to the contraction of the test section. For x > 0, a strong transverse velocity gradient occurs between the inlet and the stagnant flows (Figure4b). This generates a shear flow in the wake of the obstacle.

Figure 3 .

 3 Figure 3. Illustration of the different image processing steps: (a) original image, (b) background subtraction, (c) edges detection, (d) edges dilatation, (e) filling closed contour, (f) erosion and (g) detected contours.

Figure 4 .

 4 Figure 4. Dimensionless 2D temporal average, ‖𝒖‖ 𝑼 𝒊𝒏 ⁄ , in the inlet area at y = 37.5mm (a) and z = 50mm (b). Experiments with the 25mm side SB, 𝑼 𝒊𝒏 = 𝟎. 𝟐𝟓𝒎 𝒔 -𝟏 and 𝑽 𝒔 = 𝟏𝒎 𝒔 -𝟏 . Black velocity vectors have a unit length to indicate the direction of the flow. White dashed lines represent the location of the profiles plotted in Figure 5.

Figure 5 .

 5 Figure 5. Dimensionless velocity profiles of (a) axial component 𝒖 𝑼 𝒐𝒃𝒔 ⁄ and (b) vertical component 𝒘 𝑼 𝒐𝒃𝒔 ⁄ are plotted at x = 0mm in the vertical planes situated at y = 32mm for the 12.5mm side obstacles and the FP, and y = 37.5mm for the 25mm side obstacles (see Figure 4a). Dimensionless velocity profiles of (c) axial component 𝒖 𝑼 𝒐𝒃𝒔 ⁄ and (d) transverse component 𝒗 𝑼 𝒐𝒃𝒔 ⁄ are plotted few mm after x = 0mm in the horizontal plane situated at z = 50mm (see Figure 4b). The SB, HC, RP and FP are plotted in black, blue, red and green color, respectively (see online version of this article for color coded version of this figure). Thin lines are used for 12.5mm size obstacles and heavy lines are used for 25mm size obstacles. Dashed lines correspond to 𝑼 𝒊𝒏 = 𝟎. 𝟏𝒎 𝒔 -𝟏 , 𝑽 𝒔 = 𝟎. 𝟒𝟓𝒎 𝒔 -𝟏 and solid lines correspond to 𝑼 𝒊𝒏 = 𝟎. 𝟐𝟓𝒎 𝒔 -𝟏 , 𝑽 𝒔 = 𝟏𝒎 𝒔 -𝟏 .

Figure 6 .

 6 Figure 6. Dimensionless 2D temporal average, ‖𝒖‖ 𝑼 𝒊𝒏 ⁄ , in the vertical plane at y = 5mm, namely close to the center of the channel. Experiments with the 12.5mm side SB, 𝑼 𝒊𝒏 = 𝟎. 𝟐𝟓𝒎 𝒔 -𝟏 and 𝑽 𝒔 = 𝟏𝒎 𝒔 -𝟏 . Black velocity vectors have a unit length to indicate the direction of the flow. White dashed lines represent the location of the profiles plotted in Figure 7.

Figure 7 .

 7 Figure 7. Dimensionless velocity profiles of (a) axial component 𝒖 𝑽 𝒔 ⁄ and (b) vertical component 𝒘 𝑽 𝒔 ⁄ are plotted at z = -15mm and y = 5mm. The 12.5mm size SB, HC, RP and FP are plotted in black, blue, red and green color, respectively (see online version of this article for color coded version of this figure). Dashed lines correspond to 𝑼 𝒊𝒏 = 𝟎. 𝟏𝒎 𝒔 -𝟏 , 𝑽 𝒔 = 𝟎. 𝟒𝟓𝒎 𝒔 -𝟏 and solid lines correspond to 𝑼 𝒊𝒏 = 𝟎. 𝟐𝟓𝒎 𝒔 -𝟏 , 𝑽 𝒔 = 𝟏𝒎 𝒔 -𝟏 .

  𝑜𝑏𝑠 = 12.5𝑚𝑚) 1.1

Figure 8 .

 8 Figure 8. Gas entrainment mapping for the flat plate configuration. The present results are plotted with diamonds and the one from Kimura et al., 2009 are plotted with circles for water experiments and with triangles for sodium experiments. Symbols are filled when there is gas entrainment. The dashed line gives the threshold of gas entrainment occurrence for the present study and the solid line is 𝑸 𝒊𝒏 = 𝑸 𝒔 .

Figure 9 .

 9 Figure 9. Gas entrainment mappings. SB of (a) 12.5mm and of (b) 25mm side; HC of (c) 12.5mm and (d) 25mm radius; RP of (e) 12.5mm and (f) 25mm height. Symbols are filled when gas entrainment occurs. The solid line indicates 𝑄 𝑖𝑛 = 𝑄 𝑠 .

Figure 10 .

 10 Figure10. The thresholds of gas entrainment are plotted with solid lines and dashed lines for 𝒉 𝒐𝒃𝒔 of 12.5mm and of 25mm respectively. SB, HC and Ramps cases are highlighted using squares, circles and triangles, respectively. The dashed line without symbol is the FP case.

Figure 11 .

 11 Figure 11. Illustration of gas entrainment through non-continuous (a) and continuous (b) surface swirls. Associated videos can be found on the online version of the paper.

Figure 12 .Figure 13 .

 1213 Figure 12. Illustration of continuous surface swirls for 25mm side SB (a), 12.5mm side SB (b) and PP (c) with 𝑼 𝒊𝒏 = 𝟎. 𝟐 𝒎 𝒔 -𝟏 and 𝑽 𝒔 = 𝟏𝒎 𝒔 -𝟏 . White dashed lines indicates planes 𝒛 * = 𝟐 and 𝒛 * = 𝟏. Corresponding videos are available online.

Figure 15

 15 Figure15gives the dimensionless standard deviation of 𝑥 𝑠𝑠 * ,  * = /𝐷, versus 𝜅 for 𝑧 * = 2 and 𝑧 * = 1. For the configurations with small obstacles and the flat plate,  * is found to increase with 𝜅. For the configurations with the large obstacles,  * seems to fluctuate around a mean value 〈𝜎 * 〉. They are provided in Table4at two positions. Values increase from the flat plate (0.23) to the square blocks (0.4) at 𝑧 * = 2 and are ranked with the same order between the two obstacles' size. Fluctuations are smaller for 𝑧 * = 1 than for 𝑧 * = 2.These results highlight the fact the positions of the centres of the surface swirls strongly depend on the different flow configurations and their underlying turbulent structure.

Figure 14 .Figure 15 .

 1415 Figure 14. The parameter 𝒙 𝒔𝒔 * ̅̅̅̅ is plotted in abscissa and 𝜿 in ordinate at (a) 𝒛 * = 𝟐 and (b) 𝒛 * = 𝟏. The 12.5mm and 25mm size obstacles are respectively plotted with filled and empty symbols. Squares, circles, triangles and stars are used for SB, HC, RP and FP inlet configurations respectively.

  0.76 1.4 0.93 0.88 1.1 0.5 disp % 11.3 6.4 12.1 17 3.2 2.8 5.7𝑧 * = 𝟏 

Figure 16 .

 16 Figure 16. Dimensionless width of the surface swirls, 𝑳 𝒔𝒔 * , versus 𝜿 𝜿 𝒄 ⁄ with 𝜿 𝒄 ≈ 𝟓𝟓𝟎𝟎. Measurements at 𝒛 * = 𝟐 and 𝒛 * = 𝟏 are separated with a red dashed line. The inlet configurations with the 12.5mm and 25mm size obstacles are respectively plotted with filled and empty symbols. Squares, circles, triangles and stars are used for SB, HC, RP and FP cases respectively.

Figure 17 .

 17 Figure 17. Illustration of the calculation of the entrained gas area through the suction nozzle. Bubbles' area is taking into account in 𝑨 𝒈𝒂𝒔 when bubbles are placed under the white dashed line defined by 𝒛 = 𝒛 𝒍𝒊𝒎 with 𝒛 𝒍𝒊𝒎 = (𝜸𝑽 𝒔 -𝑽 𝒃𝒖𝒐𝒚 (𝒓 𝒃𝒖𝒃 )) 𝒇 ⁄ .

Figure 18 .Figure 19 .

 1819 Figure 18. Dimensionless entrained gas area 𝑨 𝒈𝒂𝒔 𝑫 𝟐 ⁄ versus . The inlet configurations with 12.5mm and 25mm size obstacles are respectively plotted with filled and empty symbols. Squares, circles, triangles and stars are used for SB, HC, RP and FP cases respectively.

Table 1 .

 1 Main flow parameters

Table 2 .

 2 Table2gives more details so as to highlight the role of the obstacles in generating turbulence. intensities in percent within the inlet region.

	𝑈 𝑖𝑛 m/s	𝑉 𝑠 m/s	SB 12.5mm size obstacles HC RP	FP
	0.1	0.45	9	8.6	7.7	9.5
	𝑈 𝑖𝑛 = 0.25𝑚/𝑠 𝑉 𝑠 = 1 𝑚/𝑠	7.7	7.3	6.5	7.7
			25mm size obstacles
	𝑈 𝑖𝑛 = 0.1𝑚/𝑠 𝑉 𝑠 = 0.45𝑚/𝑠	11.3	7.6	5.7	-
	𝑈 𝑖𝑛 = 0.25𝑚/𝑠 𝑉 𝑠 = 1 𝑚/𝑠	8.1	6.5	4.8	-

Table 3 .

 3 Threshold of obstacles ranked according to the mean ratio, 𝕽 𝒔 ,

of the threshold suction velocity to the one of the flat plate case.

Table 7 .

 7 As expected, values of 𝛾 in Table7 differs from 𝜃 3 2

⁄ (

Table 6 )

 6 since the number of bubbles changes with 𝜉 and several size of bubbles are entrained.

		12.5mm side	25mm side
			obstacles			obstacles		FP
		SB HC RP	SB HC RP
	𝜽	4.3	4.6	3.8	3.6	2.8	3.7	9.6
	disp	32	34	56	63	33	48	47

Table 6 .

 6 Exponents

, 𝜽, of the approximated power laws giving the area of entrained gas in function of 𝝃, i.e. 𝑨 𝒈𝒂𝒔 ~𝝃𝜽 for the different inlet configurations. disp is the dispersion (in %) of the experimental points around these power laws.

Table 7 .

 7 Exponents, 𝜸, of the approximated power laws giving the volume 𝑽 𝒈𝒂𝒔 of entrained gas in function of 𝝃, i.e. 𝑽 𝒈𝒂𝒔 ~𝝃𝜸 for the different inlet configurations. disp is the dispersion (in %) of the experimental points around these power laws.

		12.5mm side	25mm side
			obstacles			obstacles		FP
		SB HC RP	SB HC RP
	𝜸	5.5	5.8 6.45 6.8	4	4.7 11.2
	disp	37	33	57	85	36	45	49

Their choice of the 1/1.8 scale was based on the works ofEguchi et al., 1994 and Eguchi and Tanaka, 1994.

Note that the boundaries of the velocity fields in Figure4bare not straight because of the correction of the distortions.

Acknowledgements

This work is supported by the R4G program. The authors would like also to acknowledge Murielle Marchand, a CEA colleague, for her involvement in the development of BANGA experimental set up.


F Bertrand

G Mauger

M Bensalah

P Gauthé

Transient behavior of ASTRID with a gas power conversion system

Keywords: SFR, ASTRID, transient analysis, gas PCS, loss of off-site power, depressurization, system regulation

The present article is dedicated to preliminary transient studies carried out for the analysis of the system overall behavior of the ASTRID (Advanced Sodium Technological Reactor for Industrial Demonstration) demonstrator developed in France by CEA and its industrial partners. ASTRID is foreseen to demonstrate the progress made in SFR technology at an industrial scale by qualifying innovative options, some of which still remain open in the areas requiring improvements, especially safety and operability. Among the innovative options, a gas power conversion systems (PCS) is envisaged. In this latter innovative PCS, the working gas is nitrogen whose flow rate delivers power to a turbine driving with the same shaft two compressors (low and high pressure) separated by an intercooler. The other part of the work delivered by the gas is used to drive the alternator that produces electricity. The main objective of such a PCS consists in avoiding physically the possibility of a sodium/water reaction with the secondary circuit but the impact of this PCS on the control of the accidents has also been studied. The main purpose of the studies presented in the paper is to assess the dynamic behavior of ASTRID including a gas PCS with the CATHARE2 code. The first transient presented deals with a loss of off-site power and has been calculated for the gas PCS but also for a classical steam/water PCS for comparison purpose. Then typical transients of gas system have been investigated. Several families of initiating events affecting the PCS are studied: breaches in the ternary circuit, a loss of power supply, a loss of main heat sink and the spurious 1

Introduction

After a brief presentation of the ASTRID reactor main design options, the stress is put on the presentation of the power conversion system (PCS). In this gas innovative PCS designed by ALSTOM, the working gas is nitrogen whose flow rate delivers power to a turbine driving with the same shaft two compressors (low and high pressure) separated by an intercooler. The other part of the work delivered by the gas is used to drive the alternator that produces electricity. Moreover, a recuperator enables to preheat the gas flow entering the gas/sodium heat exchanger with the gas exhausting from the turbine. The main objective of such a PCS consists in physically avoiding the possibility of a sodium/water reaction at the interface between the secondary and the ternary circuit. In the following part of the paper, the main features of the CATHARE2 modelling of nitrogen cycle currently envisaged for ASTRID are detailed as well as the system actions simulated in the CATHARE2 input deck for transient calculations. Finally, the calculation results are presented and the consequences on the reactor (thermal shock, long term behaviour and heat removal process) of the various transients simulated are assessed. First, the seam/water and gas PCS's are compared in case of LOOP. Then the behaviour of ASTRID with a gas PCS is presented for typical transients of gas circuits.

Overview of ASTRID design

ASTRID, standing for Advanced Sodium Technological Reactor for Industrial Demonstration consists in a 1500 MWth SFR pool type reactor of about 600 MWe that is an integrated technological demonstrator designed for industrial-scale demonstration of 4 th generation sodium-cooled fast reactor (SFR) safety and operation (Le [START_REF] Coz | The ASTRID Project: status and future prospects[END_REF]. The main objective of ASTRID is to test advances on an industrial scale in dedicated areas (in particular safety, operability and in service inspection and repair). ASTRID will also be designed to perform waste transmutation in order to demonstrate the industrial feasibility of this technique for reducing volume of end waste and lifetime of ultimate waste. Additionally, ASTRID is being designed to fulfil the GENIV criteria in terms of safety, sustainability, economy and proliferation [START_REF] Chenaud | Status of ASTRID core design studies at the end of predesign phase 1[END_REF]. Various innovative options have been investigated during the first stage of the conceptual design carried out in 2011 and 2012 in order to improve safety on the following points:

-prevention of core degradation and mitigation of its effects (innovative CFV2 core design with a low total voiding effect has been retained); -elimination of the possibility of sodium/water reaction at the interface between secondary loops and ternary circuit (gas PCS whose transient behaviour is presented in this paper); -enhancement of the reliability of the DHR systems.

Core design

The so-called CFV concept (version 1) featured by a low global Na voiding effect has been designed by CEA in order to increase the margin up to sodium boiling in case of unprotected transient [START_REF] Chenaud | Status of ASTRID core design studies at the end of predesign phase 1[END_REF] and also to reduce the severity of a primary power excursion should it occur in case of a severe accident (Lemasson et al., 2013;[START_REF] Bertrand | Comparison of the behaviour of two core designs for ASTRID in case of severe accidents[END_REF]. This core is an axial heterogeneous core (Fig. 1).

The low voiding effect of this core results mainly from the presence of a sodium plenum above the upper fissile zone [START_REF] Sciora | Low void effect core design applied on 2400 MWth SFR reactor[END_REF] combined to the presence of a fertile plate in the inner zone of the core (Fig. 1). The height of the outer fissile zone enables the void reactivity effect to be lowered as well. All these effects lead to a reactivity decrease when the upper part of the core experiences a sodium thermal expansion or sodium boiling. Finally the overall voiding effect for this version of the core is equal to -0,5 $ (one Dollar being the delayed neutron fraction).

Vertical cut

Primary and secondary systems overview

The ASTRID pool type primary circuit includes 3 primary pumps and 4 intermediate heat exchangers (IHX) immersed in the reactor vessel (Fig. 2). Each of the 4 secondary loops delivers a fourth of the core power (375 MWth) to 3 modular steam generators (SG) or to 3 modular sodium/gas heat exchangers (SGHX); the modularity enables the impact of an event in a SG or SGHX to be reduced [START_REF] Coz | The ASTRID Project: status and future prospects[END_REF].

Finally for both options, an argon covered expansion vessel is foreseen in order to protect the loop against an accidental pressure wave. The main features of the reactor operating point are provided in Table 1 for the primary and secondary circuits. The operating point is the same for both PCS options except for the ternary side. The steam/water PCS option classically includes steam modular SGs (three by secondary loop), a steam turbine delivering the reactor power to an alternator, a condenser and finally a feed-water pump.

Considering the operating point and the ternary flow rate (about 680 kg/s), the efficiency of the reactor with the steam PCS would be about 41 % according to first assessments.

Gas PCS

The gas PCS using nitrogen as a working gas is presented on Figure 3 with its associated operating point [START_REF] Laffont | ASTRID power system: assessment on steam and gas options[END_REF][START_REF] Cachon | Innovative power conversion system for the French SFR prototype[END_REF]. The choice of the PCS gas has been performed by taking into account heat transfer, compression work, pressure level and its simplicity of operation. Consideration of all these criteria led to the nitrogen at 180 bars (turbine inlet) as the reference coolant. A Brayton cycle, which has never been implemented in any sodium reactor but has been investigated for High Temperature Reactors (HTR) has been chosen. This kind of cycle provides the best efficiency. The PCS flow rate is about 7160 kg/s and a by-pass line (preliminarily located in red in Fig. 3) enables to stop or to regulate the TM speed by reducing the outlet/inlet pressure ratio, the by-pass opening consisting in an aerodynamic brake. Basically, the PCS includes a turbine driving two compressors mounted on the same shaft 3 . Since the realization of the calculations presented in the present paper, the design of the PCS has evolved and is made now of two shaft-lines, each one including two symmetric turbines arranged in parallel, a low pressure compressor and a high pressure compressor plus an alternator. The efficiency of the whole gas compression is improved by means of a pre-cooler and an intercooler separating the 2 compression stages; these heat exchangers (HX) are cooled by the heat sink of the reactor. Finally, a so-called recuperator HX permits to pre-heat the gas flow before entering the SGHX and to cool it before entering the pre-cooler. The net efficiency of the reactor considering this PCS option would be about 37.5 % according to preliminary assessments (laffont et al., 2013). A multiple pipe design in parallel has been adopted in order to limit the gas velocity in the pipes (minimization of pressure drop) and in the same time to limit the maximum pipe diameter at about 1 m for mechanical and manufacturing reasons. A sensitivity study on the cycle parameters has shown that the turbomachine performance and the pressure losses are the most influent parameters in terms of efficiency. The main operating parameters of the PCS are presented in Primary and secondary circuits

The whole system has been modelled partly by CEA and partly by AREVA and the calculations have been run with CATHARE2 v25_3 mod2.1 [START_REF] Geffraye | CATHARE 2 V2.5_2: a Single Version for Various Applications[END_REF]. The CATHARE2 code is a 6-equations thermalhydraulic system code. Properties of sodium have been implemented in the code but the form of the friction and heat transfer correlations are the same as in the water standard version. The following neutronic effects are modelled: Doppler, sodium density, cladding expansion, fuel expansion, hexcan (HC) expansion, diagrid expansion and finally the reactivity feedback resulting from relative location of control rods within the core. All these reactivity coefficients are included in the point kinetics of neutron physics of CATHARE2 including 8 groups of delayed neutrons and 4 groups of fission products. The core channels, the IHX, the fraction of the primary flow cooling the vessel and the inlet/outlet region of the pumps are modelled thanks to 1-D CATHARE2 axial elements. The flow distribution within the cold and hot collectors is modelled thanks to a dedicated volume decomposition in several volumes whose flow-path has been assessed with CFD calculations performed in nominal regime. The secondary circuit is modelled with a single loop representing the 4 loops, connected to the primary system by means of a single IHX representing the 4 IHXs. A secondary pump is modelled and enables the loss and variation of flow to be simulated.

Nitrogen PCS

The PCS system is fully modelled in the CATHARE2 input deck as represented on Figure 4. This modelling includes a comprehensive TM module whose performance maps are derived from CEA gas fast reactor (GFR2400) performance maps [START_REF] Widlund | Overview of Gas Cooled Reactor applications with CATHARE[END_REF] but adapted to a nitrogen circuit. The parameters of the heat transfer correlation (Dittus-Boelter form) for plate type Na/gas exchangers used in CATHARE2 have been derived from CFD calculations. The next step of modelling will be to include the performance maps elaborated by ALSTOM in order to have a more realistic design and its associated modelling. The SGHX modules are associated to a weight equal to 12 and the recuperator to a weight of 18 in order to cope with the pre-design of the PCS. The pre-cooler and intercooled are respectively divided in 72 and 60 modules represented by a single HX in CATHARE2 associated to their weight. The whole nitrogen inventory of the circuit is around 560 tons. The rotating mass equation is solved on the TM shaft including the turbine, the compressors and the resisting torque of the alternator.

System actions and regulations adopted for transient calculations

This section is devoted to the transient analysis of ASTRID equipped with a gas PCS. In a first general part, the system actions are presented (actions triggered via the input deck after the transient initiation) and further on, the analysis of the reactor behaviour when facing specific transients of the gas PCS are presented. Additionally, the capability to remove decay heat from the system with the gas PCS is also preliminarily investigated. Finally a sensitivity study to the design of the TM by-pass line is also provided, the aim of this design alternative being to improve the capability of the reactor to withstand the transient without scram.

Reactor trip system modelling and other system actions

The transient input deck includes several trip parameters enabling the reactor scram to be actuated and the PCS to be protected (Table 3). Moreover, in case of reactor trip, the rotation speed of the primary pumps is decreased to 25% of its nominal value in 30s and the rotation speed of the secondary pumps is reduced to 15% of its nominal value in 9s.

System regulation

Several regulation actions are implemented in the input deck in order to be able to remove the decay heat and to maintain all the circuit in acceptable conditions. In the frame of those exploratory studies, some regulations have been settled in order to control the power extracted by the ternary circuit (Table 4).

Pressurized PCS transients

The transient investigated in this paper have been distinguished between transients with PCS under pressure and PCS depressurized, those latter occurring in case of breaches. Among the pressurized transients, the black-out transient, the loss of off-site power (LOOP) and the spurious opening of by-pass valves of the TM have been simulated, including a comparison between steam/water and gas PCS. The dedicated DHR loops plugged on the primary circuit of the reactor are not used during the transient but the TM speed is kept as long as possible at 3000 rpm by means of a regulated opening of its by-pass valve (Tab. 4).

LOOP (Loss Of

Unprotected transient

In order to assess and to compare the natural behaviour of the core in case of LOOP depending on the PCS, calculations have been performed without scram in the case of the CFV core (Fig. 5). The core power (Fig. 5), governed by the reactivity feed-back is lower for the water PCS than for the gas PCS. This is due to a core inlet temperature that is higher in the case of the water PCS thus favoring the diagrid expansion. This temperature difference is explained by the difference of power removed by the ternary side depending on the PCS. In the case of the water PCS, the SG is dried out in approximately 30 s after the beginning of the transient, that is, with the same dynamics as the feed water decrease. After its dryout, the SG still removes some power during 900 s but only several Megawatts whereas the gas PCS removes about 100 MW thanks to a substantial residual flow rate of about 250 kg/s driven by natural convection induced by exchanges through the recuperator.

Fig. 5.

Comparison of core power evolution in case of unprotected LOOP for gas and steam/water PCS

Protected transient

The actuation of the scram (Tab. 3) is triggered according to the reach of the threshold of the power to flow rate ratio for the water PCS (at 3.7 s) whereas regarding the gas PCS, the threshold of the TM overspeed is reached (1.3 s) and induces the reactor scram and the opening of the TM by-pass line. The core power is the same for the 2 PCS since the transient occurs in a residual power regime. So it is not influenced by the thermalhydraulic parameters. For the gas PCS, as soon as the by-pass line is fully open (at 4 s), the TM speed decreases up to zero in about 400 s. Despite this absence of rotation of the turbomachinery, the ternary side remains featured by a gas flow rate ranging from about 200 kg/s at the beginning of the transient down to 100 kg/s at 5000 s thanks to exchanges through the recuperator that generates natural circulation. This behavior, already observed in the unprotected case, leads to more moderate heating of the secondary inventory and thus to a moderate primary inventory heating as well (Fig. 6). Even if the difference is about 20 or 30 °C, the impact on the vessel creep could be important for the long term transient.

Gas_PCS Water_PCS

Fig. 6. Comparison of core temperature evolution in case of protected LOOP for gas PCS and for steam/water PCS

Decay heat removal with the gas PCS

During the LOOP transient, emergency electrical supply is available to deliver a back-up flow rate to the primary and secondary pumps as well as to the water-feeding system of the coolers. The aim of this subsection is to present the sensitivity of the system behavior to a back-up flow rate in the coolers 4 .

Conversely to calculations presented in section 5.1.1, in this set of calculation, regulations actions presented in Table 4 are simulated in order to keep the ternary side under operation as long as possible as a temporary heat sink. Figure 7 shows that as soon as a back-up flow rate is available (even only 5 % of the nominal flow rate), the core inlet temperature decreases of at least 150°C in one hour. It is interesting to observe that the minimum flow rate tested (5%) is enough to keep the TM under operation. Therefore, the ternary flow rate and the cold source is enough to cool the core adequately, thus providing a substantial additional delay before switching the decay heat removal to the dedicated in-vessel loops.

Moreover, the rotation of the TM enables to reduce the hot shock on the primary vessel by compensating the heating due to the loss of flow on the primary and secondary circuits (+ 20°C at the IHX outlet instead 4 This study is only performed in order to assess the behavior of the system no matter what the decisions will be taken regarding the systems rescued: in particular, at that time, it is not foreseen to have an emergency electrical supply for the coolers.

Black-out

This transient results from the total loss of electrical supply: there is no back-up flow rate delivered by means of emergency electrical supply and the water feeding of the coolers is unavailable as well. The reactor trip is actuated at the beginning of the transient. The pump run-down is governed by their inertia.

The power balance between the core and the reactor circuits and the temperature evolution across the core and on the interface between the secondary and ternary circuits are presented on Figure 8. A cold shock affects all the circuits because of the heat extraction by the ternary circuit thanks to a substantial flow rate that remains in this circuit thanks to the TM operation and to the gas expansion in it (Fig. 6). This can be observed on Fig. 5 because the nitrogen temperature at the inlet of the SGHX decreases up to 300°C because only the decay heat is transferred to the PCS and yet, its flow rate remains very high as long as the TM is under rotation. After 10 minutes, the balance between the heat transferred to the PCS from the core and the heat released by the coolers leads to the TM rundown (Fig. 9). As a consequence, at this time, the PCS pressure is balanced around 140 (Fig. 9) bars and the secondary and ternary side temperatures begin to increase. At 1 h, a thermal and power balance is reached and the heat removed by the ternary side becomes lower than the decay heat and the primary sodium temperature begins to increase. Actually, even after the primary and secondary pumps shutdown and the TM rundown, natural circulation flows in each circuit drive the whole reactor cooling. However, this result should be to confirm considering the uncertainty on the natural circulation regime calculation (pressure head of the stopped pumps, performance maps of the TM and circuit elevation are not assessed adequately). Nevertheless, this result indicates that the heating rate of the whole system is very low and that thanks to the Brayton cycle and of the inertia of the reactor cooling circuits, the PCS enables the decay heat to be removed more than 6 hours without reaching the temperature level of the nominal state of the reactor.

Spurious opening of by-pass valves

Various flow paths are possible (Fig. 10) in order to by-pass the turbine and practically, a mix of gas taken in the cold leg (opening of by-pass2) and of gas taken in the hot leg (opening of by-pass1) is injected in the exhaust pipe of the turbine (opening of by-pass4). The transients calculated in this paper deal respectively with the spurious opening of the lines by-pass1 + by-pass4 (called BYP1) and of the spurious opening of the lines by-pass2 + by-pass4 (called BYP2). The spurious opening of the valves is realized in 3 s. The line by-pass 4 is open before the transient and the lines by-pass1 and by-pass2 are respectively made of 2 pipes of 60 cm diameter and when the transient occurs, 1 over 2 lines are open. As observed on figure 11, the case "BYP2" leads to a fast rundown of the TM due to the cold temperature of the gas in the cold leg, the flow rate by-passing the turbine being higher than in the case "BYP1". As a result, the flow rate in the PCS is reduced faster in the case "BYP2" than in the case "BYP1" and the heating of the sodium at the core inlet occurs earlier (Fig. 11). The reference design option including 2 by-pass pipes for the parts of the circuit called BYPASS1 and BYPASS2 leads to an over-speed of the TM equal to 114 % of its nominal speed. By keeping the same over-speed, the BYPASS2 lines have been sub-divided into smaller pipes (6 pipes of 25 cm diameter) in order to limit the transient impact of the opening of one single line. In the case "BYP2" the reactor trip threshold are not reached and thus, the reactor is not scrammed, even after a long time.

The electrical power delivered to the alternator remains higher than the minimum value. Therefore, the alternator remains connected and the core power is extracted. However, in the case "BYP1", due a hotter gas intake in the recuperator, the core inlet temperature increases and reaches the trip level (Fig. 12) after 55 minutes. This exploratory calculation has shown that by designing more modular by-pass line, it is possible in the same time to limit adequately the turbine over-speed and to keep the reactor under operation.

PCS depressurization transients

These depressurization transients result from breaches in the PCS and are typical from pressurized gas circuits. Two break locations have been studied: upstream of the (labelled AM_SGHX on fig. 4) and downstream of the turbine (labelled AV_TURB on Fig. 4). Two diameters have been studied as well:

10 inches representing (in order of magnitude) the size of the piping on the main branches of the PCS and 40 inches representing the size of those main branches. The reactor trip signal and the operation of the reactor circuits are the same as indicated in section 4 except the cooler flow rate that is kept at its nominal value because the initiating event does not lead to the loss of the cooling circuits.

10 inch diameter breach

As observed on Figure 13 through the power delivered to the alternator, the TM is kept under rotation a longer time for the break located downstream of the turbine (AV_TURB) than for the case (AM_SGHX). This is due to the increase of the expansion ratio induced by the depressurization of the low pressure part of the PCS. As a result, the transient is compatible with the Brayton thermodynamic cycle for a long time and the low electrical power signal is reached later on (Figure 12). Since the TM keeps running, the compressors decrease the depressurization kinetics. The scram of the reactor occurs at about 8 min instead of 5 min (because of a too high core inlet temperature) because the balance between the power extracted by the PCS and the power released by the core leads later to the core inlet temperature threshold in the case AV_TURB than in the case AM_SGHX (Fig. 14). The primary temperature evolution indicates that in the case of a breach on the high pressure part of the circuit, the heating of the and primary circuits resulting from the TM rundown occurring at 12 min limits the cold shock in the primary circuit. It induces a shift of about 150°C in the primary temperature evolution in the long term transient that is mainly governed by the primary and secondary circuits thermal inertia once the depressurization of the PCS is over (after 1 hour). Thus, depending on the breach location, the grace delay to start the in-vessel DHR systems can be shifted of more than 5 hours (Fig. 14). size leads to a very fast depressurization inducing a flow inversion in pipe portions downstream of the breach including the TM components. So, these results are only preliminary and will be checked with ALSTOM performance map when they will be included in the CATHARE2 input deck. However, the system behavior is consistent with that calculated for more little breaches but with a much faster dynamics. The reactor scram is reached in 6 tenth of seconds in the case AV_TURB because of an out of range electrical power delivered by the turbine (expansion rate and flow increase in the turbine, thus increasing electrical power) and it is reached after 6 s in the case AM_SGHX because of a to low electrical power due to the same factors, but acting inversely. As long as the alternator is connected, the case AV_TURB leads to a cooling of the primary circuit but less pronounced for the 10 inch breach because the ternary flow rate decreases fast due lack of gas inventory in the circuit. So, as far as very large breaches are concerned, the location of the breach has a lower impact on the transient evolution of the reactor (Fig. 15). Once the depressurization is over and once the cold shock is over, the heating rate of the primary circuit is the same for both breach locations.

Conclusions

The comparison of both PCS studied for ASTRID during the pre-conceptual design phase of the project has shown that, in case of loss of site power, the transient behavior of the reactor is not very different between the water/steam PCS and the gas one. However, the cold plenum temperature calculated for gas

Fig 1 .

 1 Fig 1. General CFV core geometry

Fig. 2 .

 2 Fig. 2. Primary system arrangement for ASTRID

Fig. 3 .

 3 Fig. 3. Sketch of the gas PCS (LPC: low pressure compressor; HPC: high pressure compressor; T: turbine; Gen: generator)

Fig. 4 .

 4 Fig. 4. Sketch of the CATHARE2 modelling of the PCS (SGHX: sodium/gas heat exchanger, NSS: Nitrogen Supply System)

  Off-site Power) 5.1.1. Comparison of the reactor behaviour depending on the nature of the PCS As far as the gas PCS is concerned, the initiating event (IE) causing this transient is featured by the alternator disconnection and the instantaneous shutdown of the cooling flow rate on the cold side of the pre-cooler and of the intercooler. Regarding the steam/water PCS, the transient consists in a decrease up to a residual value of the flow rate of feed water of the SG in 30 seconds. The actions triggered by the LOOP on the primary and secondary side are the pump trip according to the speed decrease aforementioned and an emergency supply (by a Diesel engine) of the primary and secondary pumps enabling their rotation at their back-up speed value. The transient effects are investigated first in case of failure of the reactor shutdown (unprotected transient) and then by considering the reactor scram.

Fig 7 .

 7 Fig 7. Core inlet temperature and turbomachine rotation speed for various back-up flow rates of the coolers (from 0 to 25 %)

Fig. 8 .Fig. 9 .

 89 Fig. 8. Power balance and temperature evolutions during the black-out transient

Fig. 10 .Fig. 11 .

 1011 Fig. 10. Sketch of the by-pass line system in the PCS (HP comp and LP comp : high pressure and low pressure compressors; HP and LP recup: high pressure and pressure recuperator sides; SGHX: sodium gas heat exchanger)

Fig. 12 .

 12 Fig. 12. Alternator electrical power and core inlet temperature for by-pass valve opening transients, impact of design improvement

Fig. 13 .Fig. 14 .

 1314 Fig. 13. PCS depressurization transient and electrical power transient for a 10 inch breach

Fig. 15 .

 15 Fig. 15. TM rotation speed and core inlet temperature for a 40 inch breach

  Tab 1. Nominal operating point of the secondary and primary circuits

	Primary Circuit
	Core flow rate (kg/s)	7900
	Fraction of the flow for vessel cooling (kg/s)	600
	Primary flow rate (kg/s)	8500
	Core inlet/outlet temperature (°C)	400/550
	Core inlet/outlet pressure (bar)	4.5/1.9
	Secondary Circuit
	Secondary flow rate (kg/s)	6370
	IHX inlet/outlet temperature (°C)	345/530
	Pump inlet/outlet pressure (bar)	1.9/7.1
	2.3. Power conversion system (PCS)
	2.3.1. Steam/water PCS

Table 2 .

 2

	Tab. 2. Main features of the PCS operating point
	Generator power	580 MWe
	Pre-cooler power	570 MWth
	Intercooler power	340 MWth
	Ternary flow rate	7170 kg/s
	Inlet temperature of SGHX	330°C
	outlet temperature of SGHX	515°C
	Turbine pressure (inlet/outlet)	180 bar/73 bar
	Low pressure compressor (inlet/outlet)	72 bar/112 bar
	High pressure compressor (inlet/outlet)	111 bar/ 181 bar

thrust on it. However, a simple line turbine is modelled in this version of the CATHARE2 input deck.

Table 3 .

 3 Reactor trip parameters and ternary system protection actions

	Parameter	Threshold	Action
	Core power (P)	P/P N > 1.08	Scram
	Core power/Flow rate (P/Q)	(P/Q)/(P N /Q N ) > 1.35	Scram
	Core outlet temperature, T CO (°C)	Tco > 583	Scram
	Core inlet temperature, T CI (°C)	Tci > 420	Scram
	Core Delta T ratio	(Tco-Tci)/(Tco N -Tci N ) > 1.09	Scram
	Alternator electrical power ratio	Pelec/Pelec N > 1.15	Alternator disconnection
	Alternator electrical power ratio	Pelec/Pelec N < 0.1	Alternator disconnection
	Turbomachinery speed (	/ N > 1.1	By-pass opening, scram

Table 4 :

 4 regulation system in order to control the power extracted by the ternary circuit (the regulations implemented depend on the calculated transient)

	Variable to regulate	Targeted value	Action
	Cooler flow rate	20 % of Q N	Boundary condition
	Turbomachine speed	3000 rpm	By-pass valve opening
	Turbomachine pressure inlet	12 bars	Surge valve of NSS opening

CFV stands for « coeur à faible effet de vidange » that means core with low sodium worth effect.

Actually, in order to better balance the forces on the shaft, a double-body turbine is now used in order to reduce the

Acknowledgment

The authors would like to thank their colleagues of AREVA NP for their work on the CATHARE2 input deck, as well as, the sodium fast reactor R&D project and the Gen IV program of the Nuclear Energy Division of CEA that have supported this work.

PCS seems to be lower by about 20 at 30°C than that of the water/steam PCS. Typical transient resulting from typical initiating events of a gas PCS have been investigated. First, the overall reactor behavior when facing PCS transients does not exhibit any specific showstopper in terms of impact of the ternary circuit onto the primary and core temperature evolution compared to a water/steam PCS. Several lessons learnt from the calculations could be used the reactor design and piloting if necessary in order to consolidate the safety design of the reactor or to relax operation constraints dedicated DHR systems.

Regarding the design, the increase of the number of TM by-pass lines allowing the decrease of the flow rate in each line would enable the reactor to withstand a spurious opening of a single by-pass line without exceeding the thresholds of the reactor trip protection system. Another interesting conclusion of the study is the capability of the PCS to remove the decay heat during several hours providing that the TM is kept under operation at the beginning of the transient. This capability is even enhanced by feeding the cold side of the PCS coolers with only 5% of its nominal water flow rate. This could be done by a rescue of the feeding pumps at a low power level requiring only a small power supply from the emergency Diesel engines.


L Desgranges

Th Blay

J Lamontagne

Fission products behaviour during a power transient: Their inventory in an intragranular bubble

The behavior of fission products is a key issue during incidental or accidental sequence for nuclear fuel. Here we characterized how fission products behaved inside a UO 2 pellet during a power ramp. In the pellet center fission products have left the UO 2 ceramic and can be found in bubbles. The composition of the bubble was determined using an original experimental methodology. The existence of separated precipitates made of metallic, for the one, and volatile, for the other, fission products was evidenced. This result is discussed with regards to the behavior of fission products during a power ramp.

Introduction

The behaviour of volatile fission product is key factor when nuclear fuel undergoes an incidental or accidental sequence. The amount of volatile fission products and the associated chemical phases are the primary source term for the chemically induced failure of cladding and the dispersion of radionuclides out of fuel element (see for example [ i ]). In a previous paper we evidenced that the release of volatile fission products, under a chemically aggressive form, could be observed in a fuel element that has been submitted to a power ramp [ ii ]. Coupled thermomechanical and thermodynamic modelling showed that the quantity of aggressive phases formed by released volatile fission products could be a critical parameter for Stress Corrosion cracking of zirconium based alloys in PWR [ iii ]. In order to improve the quality of this modelling we aim at determining how volatile fissions migrate during a power ramp.

On a theoretical point of view, the chemical phases formed when thermodynamic equilibrium is reached can be assessed by thermochemical modelling. But more kinetic data are needed to know when and where these equilibrium phases are actually made. A phase formed inside the nuclear ceramic as precipitates will be immobilised for a longer time than a phase formed outside the nuclear ceramic. There is, up to now, very little knowledge on the behaviour of volatile fission products, which are not gaseous inside the nuclear ceramic. This paper aims at study how volatile fission products behave inside a nuclear fuel submitted to a power transient. More specifically, the chemical phases formed by these volatile fission products in bubbles will be looked for. For that purpose a new experimental procedure was developed to measure the fission products inside a bubble formed during a power transient in a PWR fuel.

Experimental part

Sample history:

A 4 meter long PWR fuel rod, with 4,89% 235 U enriched 8.2 mm diameter pellets and 9.5 mm external diameter 0.57 mm thickness cladding in niobium doped zirconium based alloy, had been previously irradiated in a commercial reactor up to a burn-up of 36 GWd/tU. The released gases were measured by puncturing after irradiation; they correspond to 0.5% of gas inventory. A 29 cm slice was cut from this rod and refabricated as a rodlet. This rodlet was irradiated in OSIRIS reactor. After a conditioning period at 196 W.cm -1 during more than 18 hours, the power was raised up to 470 W.cm -1 with a power increase around 100 W/cm.min and the rodlet was kept at this high power during 12 hours without any detected failure. The released gases during the power transient were measured by puncturing after irradiation; they correspond to 7% of gas inventory.

After power ramp, a slice was cut from the rodlet and prepared for post-test analysis. This slice was embedded in a metallic alloy to ensure the charges evacuation and polished perpendicular to the rodlet axis. The polished section is presented on Figure 1. Characterisation apparatus:

This sample had been examined in the hot cells of the CEA LECA-STAR facility at Cadarache with:

-A shielded electron probe micro-analyser (EPMA) 'CAMECA' SX100R.

[ iv ].

-

SIMS [ v ]
A first examination of the fuel evidenced that volatile fission products formed bigger precipitates in the centre of the pellet, consistently with the peak in Cs radial profile measured by EPMA (Figue1b).

Consequently examination was focused in this area. Moreover EPMA mapping indicated that the precipitates containing volatile fission products would be located inside intragranular bubbles. This hypothesis was difficult to assess because bubbles opened during the sample polishing might have been filled with polishing dusts. Therefore it was necessary to analyse a bubble that had not been previously opened during sample preparation. We developed a dedicated procedure for that purpose that is described below.

Procedure for analysing the Fission products with in one bubble:

First of all, EPMA Xe mapping was used to identify bubbles that were still filled with Xe fission gas underneath the sample surface. One bubble was then selected for further analysis. Its position is shown on Figure 2; though not opened it appears as a darker shadow on the electronic image. The sample was transferred to SIMS. SIMS mapping with low current primary beam was used to find out the selected bubble position. Then high current primary beam was focused at the bubble position while Xe signal was measured. The opening of the gas bubble was detected by a peak in the Xe signal, in a similar manner to [ vi ]. New SIMS mapping with low current primary beam were then performed that confirmed the opening of the bubble and the presence of volatile fission products in it. The sample was transferred back to EPMA and EPMA mapping gave a qualitative composition of the phases that were contained in the bubble. No EPMA quantification was possible because the fission products within the bubble did not lie on a flat surface at a reference position.

Results

The EPMA mappings of Cs, Te, I, Ba, Ru, Tc, Mo, Pd and Cr of the opened bubble are presented on Figure 3. These fission products are characteristic of the volatile (Cs, I, Te) metallic (Mo, Ru, Tc) fission products according to Kleykamp's classification [ vii ]. Cr is a dopant added during fabrication. Ba has a specific status since it can be formed by decay of caesium. Hence it is impossible to state whether the barium observed in the bubble migrate towards the bubble as a barium atom or as a Cs atom that later decayed inside the bubble.

Figure 3: EMPA mapping of several fission products. The yellow circle corresponds to the bubbles identified in Figure 2.

The existence of these fission products in the bubble was also evidenced by EMPA before bubble opening and by SIMS after opening, but with a lower accuracy. In all images, the concentration of fission products outside bubbles is constant, giving a uniform colour on the mappings. EPMA mapping are not quantitative because the experimental background is not subtracted from the measured intensities. The actual concentration of Cs was measured by quantitative EMPA (Figure

Cs I Te Tc

Pd Ru Ba Cr Mo 1b). These concentrations are near to zero within the experimental error bars. So in the central part of the pellet the fission products are either released out of the fuel or trapped in bubbles. Inside other bubbles, visible as darker area, some precipitates are observed or not. They could have not been observed for two main reasons: the precipitates could have been removed during polishing or they could be located at a too large depth.

Discussion:

The experimental results presented are now discussed to derive information on how volatile fission products behave at high temperature in ramped fuels.

Firstly, no new chemical phase is formed inside the fuel ceramic except within bubbles. No grey phase made of barium oxide as observed in FBR fuels [ viii ], no oxidised phase as observed at the periphery of a ramped fuel [ ix ], no caesium iodine or caesium molybdate precipitate (JOG) [ x ] could have been observed in the ceramic outside of the bubbles. This lead to the statement that thermodynamic equilibrium is not achieved within the fuel even though the centre of the pellet was subjected to a high temperature around 2000°C. So the behaviour of the fission products must be considered in a kinetic point of view.

Secondly, the volatile fission products (Cs, I, Te) are in bubbles together with rare gazes, metallic fission products and the doping chromium. It is very unlikely that all these species converge to the same point to create at the same time a common precipitate. In FBR fuel metallic, barium zirconate type oxides and gas bubbles can be observed at different places [ xi ]. So we consider that a defect should exist first, on the surface of which the other fission products could nucleate with a heterogeneous precipitation mechanism. This precipitation mechanism might occur during in commercial operation or during the power transient or during both. The pre-existing defect could be an irradiation induced void as proposed in [ xii ] or a metallic aggregate (made of fission products and/or dopants) that are known to precipitate in reducing conditions in UO 2 (see for example [ xiii ]). Whatever the first phase formed, the precipitation mechanism implies that the precipitated fission products were previously dissolved in the fuel matrix with a concentration higher than a solubility threshold and that they could diffuse in in the fuel ceramic. These requirements imply that, most probably, fission products are in mono-atomic state within UO 2 crystalline structure. The solution energy of different fission products was calculated by first principle calculations [ xiv ]. These results predicted that some fission products precipitated and others did not; the calculated behaviour was in qualitative agreement with the observed behaviour of fission products. From this we conclude that the diffusion of fission products dissolved in monoatomic form are to be considered in order to describe the migration of fission products during a power transient.

Thirdly, the distribution of fission product is not uniform with the bubble. Two groups can be evidenced, metallic fission product and volatile fission products. The metallic fission products are located in two precipitates that are aligned along the vertical direction on Figure . The volatile fission products also form two precipitates by these latter are aligned more or less on a parallel direction of the diagonal of the mapping. This difference in location can be interpreted as a different physical and chemical behaviour of these compounds. First no phase associating volatile and metallic fission product seems to exist, such as caesium molybdate that would have put at the same location caesium and molybdenum. This can be justified by considering that oxygen potential at the pellet centre became more reductive. This point is developed elsewhere [ xv ]. Here we will simply consider that the chromium precipitation indicates oxygen potential, at which all metallic fission products should be in metallic phases. For such oxygen potential, the volatile fission products are expected to form TeI2, CsI, Te, Cs or CsTe2 chemical phases [ii].

These precipitates observed on EPMA mapping were also visible on secondary electron images (Figure 4). They are round-shaped, and one bubble can contain more than one precipitate. The round shape could be associated to a liquid phase that has been rapidly quenched, possibly liquid metallic fission products. Volatile fission products might have later condensed on the round precipitates from the gaseous phase to the liquid and solid phase. The shape of the bubbles that are not all rounded suggests that some bubbles could have coalesced during the power ramp. The existence of two metallic precipitates in the bubble that was analyzed is also consistent with this coalescence mechanism.

Figure 4 secondary electron image of the bubbles and the precipitates they contain (indicated by an arrow for some of them).

Conclusion:

The results presented in this paper provided a new insight in the phenomenology of fission products during a power transient that can be divided in -Insoluble fission products are likely to be in a monoatomic form dissolved in the nuclear ceramic.

-A defect, which could be could be an irradiation induced void or a metallic aggregate, would be more likely to precipitate first. This precipitation may occur during in commercial operation or during transient. -The volatile and gaseous fission products would form bubbles by a heterogeneous mechanism at the surface of this defect. -Inside the bubbles the fission products form chemical phases that correspond to the thermodynamic equilibrium conditions corresponding to the temperature and reductive oxygen potential they are submitted to. -Two bubbles may coalesce during the power transient but their metallic fission products do not necessary coalesce. -Metallic and volatile fission products form solid precipitates after quenching. This phenomenology could be used for further modelling of the behavior of fission products during an incidental or accidental sequence.

  Figure1 : a)optical view of the analysed cross section b) Nd and Cs radial concentration profiles (r/R =1 at periphery, r/R=0 at pellet centre)

Figure 2 :

 2 Figure 2: EMPA a) and Secondary electron images of the sample before b) and after c) sputtering. The bubble position lies inside the yellow circle. The red square in a) correspond to the image frame in b).

Acknowledgements: The authors would like to thank EdF and AREVA for their financial support and fruitful discussions, and Laurence Noirot (CEA) for fruitful discussions.