# FUEL+ simulation of the HALDEN LOCA test
IFA-650.10

Tony Glantz, Sebastien Belon, Maxime Salvo, Marine Guemas

► To cite this version:

Tony Glantz, Sebastien Belon, Maxime Salvo, Marine Guemas. FUEL+ simulation of the HALDEN LOCA test IFA-650.10. Nuclear Engineering and Design, 2025, 444, pp.114396. 10.1016/j.nucengdes.2025.114396. hal-05446256

HAL Id: hal-05446256

https://asnr.hal.science/hal-05446256v1

Submitted on 7 Jan 2026

HAL is a multi-disciplinary open access
archive for the deposit and dissemination of sci-
entific research documents, whether they are pub-
lished or not. The documents may come from
teaching and research institutions in France or
abroad, or from public or private research centers.

L'archive ouverte pluridisciplinaire HAL, est
destinée au dépôt et à la diffusion de documents
scientifiques de niveau recherche, publiés ou non,
émanant des établissements d'enseignement et de
recherche français ou étrangers, des laboratoires
publics ou privés.

<figure><img src="image_1.png" /></figure>

Distributed under a Creative Commons CC BY 4.0 - Attribution - International License
# FUEL+ simulation of the HALDEN LOCA test IFA-650.10

T. Glantz* ID, S. Belon, M. Salvo ID, M. Guemas

Autorité de Sureté Nucléaire et de Radioprotection (ASNR), France

<figure><img src="image_2.png" /></figure>

## ARTICLE INFO

**Keywords:**

FUEL+
DRACCAR
LOCA
HALDEN IFA 650

## ABSTRACT

FUEL+ is a software platform developed by the French Nuclear Safety and Radioprotection Authority (ASNR) with the objective of modelling the behaviour of nuclear fuel in normal, incidental, and accidental operational conditions, as well as during its storage in deactivation pools and its transportation for reprocessing. To achieve this, the platform brings together several simulation codes and notably eases their coupling. It models the main thermomechanical, thermochemical, and thermohydraulic phenomena occurring within and around nuclear fuel and its cladding at the scale of the fuel rod, assembly, or reactor core.

This paper presents the simulation with the FUEL+ platform of the HALDEN LOCA test IFA-650.10, including:

- A base irradiation calculation of the father rod used to reprocess the rod sample used in the test IFA-650.10 with the FUEL+/FIRST code.

- An irradiation calculation of the rod sample in the HALDEN reactor before the LOCA test with the FUEL+/FIRST code.

- The LOCA transient calculation itself using the FUEL+/DRACCAR code.

- The cladding hydriding evaluation using the FUEL+/SHOWBIZ code coupled to the FUEL+/DRACCAR code.

The calculation results are in good agreement with the experimental one, the cladding temperature evolution, burst time and maximal strain being well predicted by the FUEL+/DRACCAR code.

## 1. Introduction

The in-pile HALDEN LOCA tests (Wiesenack, 2013) are performed on a single fuel rod which can be fresh or irradiated and comes from several types of reactors (PWR, BWR and VVER). The primary aims of these tests are to characterize the possible fuel relocation during a LOCA transient and to evaluate the relocation impact on temperature evolution. Three of the HALDEN LOCA tests have been simulated by the DRACCAR code and presented during the Enlarged HALDEN Programme Group meeting in 2013 (Glantz, 2013).

Since then, new thermomechanical models and improvements in the thermohydraulic models have been implemented and all the HALDEN LOCA IFA-650 tests have been simulated by DRACCAR and are now part of its validation database. Furthermore, DRACCAR integration into the FUEL+ software platform (Salvo, 2024) allows the chaining with the irradiation code FIRST, whose results can be used to initialize DRACCAR calculations, as well as coupling with the SHOWBIZ code (Guémas, et al., 2013), enabling a more detailed resolution of cladding chemical behaviour (oxidation, hydriding) during a LOCA.

This paper presents the simulation with the FUEL+ platform of the HALDEN LOCA test IFA-650.10. An overview of the FUEL+ platform is provided first. Then, a summary of the HALDEN LOCA programme is given, along with a summary of the results from test IFA-650.10. The main modelling assumptions are then presented before detailing the results obtained with the FUEL+ platform and comparing them to the experimental results.

## 2. The ASNR FUEL+ platform

FUEL+ (Salvo, 2024) is the software platform developed by ASNR with the objective of modelling the behaviour of nuclear fuel in normal, incidental, and accidental operational conditions, as well as during its storage in deactivation pools and its transportation for reprocessing. To achieve this, the platform brings together several calculation codes and notably allows the coupling between these different codes. It models all the thermomechanical, thermochemical, and thermohydraulic phenomena occurring within and around nuclear fuel and its cladding at the scale of the fuel rod, assembly, or reactor core.

<footnote>* Corresponding author.</footnote>

<footnote>E-mail address: tony.glantz@asnr.fr (T. Glantz).</footnote>
One of the FUEL+ platform objectives is to simulate the behaviour of fuel rods in the core of a reactor during a Loss-Of-Coolant Accident (LOCA), particularly to evaluate safety criteria (peak cladding temperature, cladding oxidation, hydrogen production). The platform FUEL+ can also be used to determine the failed rod number in a reactor core during a LOCA and fission gas release to the primary circuit at burst from the initial gap inventory, as demonstrated during the R2CA European Union's H2020 project (Taurines, et al., 2021). To fulfil these objectives, the following codes of the FUEL+ platform are used:

* FIRST (Salvo, 2024) is a computer code developed by ASNR used to simulate the behaviour of a fuel rod during its normal operation in a pressurized water reactor. The code gives the temperature, pressure, and deformation of a fuel rod as functions of time-dependent fuel rod power and coolant boundary conditions. All the main phenomena expected during a fuel rod irradiation in a reactor can be modelled, especially the thermal behaviour of the fuel pellet and the cladding, the cladding elastic and plastic deformation, the mechanical equilibrium between the fuel pellet and the cladding, the fission gas production and release, the cladding oxidation and hydriding.

* The DRACCAR code (Glantz et al., 2018) is a multi-rod computational tool developed by ASNR to support the safety analysis of pressurized water reactors. The DRACCAR code is based on a 3D non-structured meshing able to model a simple fuel rod, a partial or a full assembly, as well as a surrounding shroud. The code is designed to simulate any types of loss-of-coolant accidents which could occur in reactor. DRACCAR realistically describes the 3D thermomechanical behaviour of fuel rods composing an assembly during a LOCA, assessing the blockage rate due to rods ballooning and considering both mechanical and thermal interactions between rods. The thermal-hydraulic phenomena involved in the reflooding of such deformed bundle geometry can be simulated by the code accounting for the feedback from rod thermomechanics.

* The SHOWBIZ code (Guémas, et al., 2013) is dedicated to the study of fuel rod cladding throughout its life cycle, i.e. in the reactor, in transport and storage situations. In particular, SHOWBIZ endeavours to model by the finite element method the mechanical and chemical phenomena such as creep, oxidation and hydriding likely to degrade the integrity of the cladding but also its mechanical behaviour. Possible incidents or accidents are considered like Loss-Of-Coolant Accidents (LOCA) or Reactivity Initiated Accidents (RIA). SHOWBIZ gathers several physical models from tools developed in the past at ASNR to study the cladding behaviour in various situations. Recent developments extend the code capability to model the cladding creep behaviour, the transport and diffusion of gases inside a channel interacting with the inner side of the cladding.

In the framework of FUEL+ platform, the fuel performance code FIRST aims to estimate the rod state at a given instant of the irradiation phase. Then this estimation is provided to the DRACCAR code which uses it as initial state for the LOCA transient calculation. DRACCAR allows estimating the peak cladding temperature, the failed rod number, and the fission gas release. The SHOWBIZ code is coupled with the DRACCAR code (in this case, the SHOWBIZ mechanical module is disabled) to assess cladding oxidation and hydriding. This calculation sequence is based on a chaining between FIRST and DRACCAR, as well as a coupling between DRACCAR and SHOWBIZ. The main challenges associated with code chaining or coupling involve, on the one hand, material and fluid properties, which are in most cases specific to each code, and on the other hand, the transfer of information from one meshing to another. The integration of the FIRST, DRACCAR, and SHOWBIZ codes into the FUEL+ platform has made it possible to overcome these issues, as this integration was accompanied by the externalization of all material and fluid properties into a shared database. The platform also provides several tools to facilitate meshing interpolation, whether linear or based on domain overlapping.

All these codes, along with the proposed methodology for their integration, need to be validated and the HALDEN LOCA tests are good candidates to it. Indeed, these tests are dedicated to the study of the behaviour of irradiated fuel during a LOCA. Out of the 16 available HALDEN LOCA tests, it has been chosen to validate the codes of the FUEL+ platform and the proposed methodology on the IFA-650.10 test (Lavoil, 2010), which is presented in the next section.

## 3. HALDEN IFA-650.10 LOCA test

### 3.1. General presentation of HALDEN LOCA facility and tests

HALDEN LOCA tests were performed on a single in-pile fuel rod which can be irradiated or not and comes from several types of reactors (PWR, BWR and VVER). These tests aim to characterize the possible fuel relocation during a LOCA transient and to evaluate the relocation impact on cladding temperature evolution.

For LOCA test performance, the facility is composed of a pressure flask where a single fuel rod is inserted and is connected to a water loop. A low level of nuclear power generation in the fuel rod is used to simulate decay heat. Moreover, the pressure flask includes a cylindrical electrical heater surrounding the rod to simulate the heat coming from surrounding rods.

During an IFA-650 test, the first part of a typical LOCA transient is reproduced by carrying out a rapid depressurization and emptying of the pressure flask and of a section of the water loop. The reflooding phase of the LOCA transient is not performed (not the aim of the HALDEN LOCA test). So, no classical safety injection is done but some liquid water is injected by a spray in some test to limit the temperature increase. Therefore, the sample rod is mainly cooled down by steam causing the rod temperature increase, which leads generally to cladding burst.

### 3.2. HALDEN LOCA test IFA-650.10

The rod sample used in the HALDEN LOCA test IFA-650.10 was a segment from a PWR parent rod, which had been irradiated in the French nuclear power plant of Gravelines 5, a 900 Mwe PWR operated by EDF (Nishi and Lee, 2001). The average burn-up reached after 5 cycles was 56.2 MWd/kgU. The characteristics of the parent fuel rods are summarized in the Table 1. The irradiation history and the power profile at the end of the last cycles are presented respectively in the Fig. 1 and Fig. 2 (extracted from (Nishi and Lee, 2001) – on Fig. 2, the x-axis provides node numbers, which correspond to a discretization of the active fuel length).

A piece with a height of 440 mm, taken from the level of the 3rd intergrid of the parent fuel rod, is used for the refabrication of the

<table><caption>**Table 1**

HALDEN IFA-650.10 parent rod – main characteristic (Nishi and Lee, 2001).</caption><thead><tr><th></th><th>Unit</th><th>Parent rod</th></tr></thead><tbody><tr><td>Fuel material</td><td>—</td><td>UO<sub>2</sub></td></tr><tr><td>Fuel density</td><td>%TD</td><td>95.32</td></tr><tr><td>235U enrichment</td><td>wt%</td><td>4.487</td></tr><tr><td>Active fuel length</td><td>mm</td><td>3846.63</td></tr><tr><td>Pellet external diameter</td><td>mm</td><td>8.192</td></tr><tr><td>Pellet internal diameter</td><td>mm</td><td>0</td></tr><tr><td>Pellet length</td><td>mm</td><td>13.78</td></tr><tr><td>Chamfer width</td><td>mm</td><td>0.56</td></tr><tr><td>Chamfer depth</td><td>mm</td><td>0.16</td></tr><tr><td>Dish depth</td><td>mm</td><td>0.32</td></tr><tr><td>Dish base radius</td><td>mm</td><td>3</td></tr><tr><td>Cladding type</td><td>—</td><td>Zircaloy-4</td></tr><tr><td>Inner diameter</td><td>mm</td><td>8.36</td></tr><tr><td>Outer diameter</td><td>mm</td><td>9.5</td></tr><tr><td>Total free volume</td><td>cm<sup>3</sup></td><td>Unknown</td></tr><tr><td>Filling gas</td><td>—</td><td>100 % Helium</td></tr><tr><td>Filling gas pressure</td><td>MPa</td><td>2.6</td></tr><tr><td>Filling gas temperature</td><td>°C</td><td>25</td></tr></tbody></table>
<figure><img src="image_3.png" /><figcaption>Fig. 1. Power history during base irradiation for IFA-650.10 parent rod.</figcaption></figure>

<figure><img src="image_4.png" /><figcaption>Fig. 2. Axial power profile at the end of life for IFA-650.10 parent rod.</figcaption></figure>

HALDEN rod (Lavoil, 2010). The main characteristics of the latter are presented in the Table 2.

IFA-650.10 test transient can be separated into three phases. The two first phases concern the facility conditioning and the third phase is dedicated to the accidental transient. The first phase allows obtaining a steady state at a pressure close to 70 bar (pressure measurement done at the outlet of the reactor lid). During this phase, the whole facility is filled with liquid water and a forced circulation is maintained in the loop. The rod linear heat generation rate is around 130 W/cm. The flow rate is approximately 0.045 kg/s. The experimental time duration of the first phase is of several days. In the second phase, the forced circulation through the pressure flask is stopped and a natural circulation takes place in this component: liquid water flows up in the space inside the heater wall and flows down in the annular space between the heater wall and the flask wall. The power is decreased to obtain a linear heat generation rate of 14 W/cm. The experimental time length of the second

<table><caption>**Table 2**
HALDEN IFA-650.10 rod sample – main characteristic (Lavoil, 2010).</caption><thead><tr><td></td><td>Unit</td><td>Value</td></tr></thead><tbody><tr><td>Burnup</td><td>MWd/kgU</td><td>61</td></tr><tr><td>Active fuel length</td><td>mm</td><td>440</td></tr><tr><td>External oxide thickness*</td><td>µm</td><td>30</td></tr><tr><td>Hydrogen content*</td><td>ppm</td><td>220</td></tr><tr><td>Lower plenum volume</td><td>cm³</td><td>0.18</td></tr><tr><td>Upper plenum volume</td><td>cm³</td><td>17.42</td></tr><tr><td>Total free volume</td><td>cm³</td><td>17.6</td></tr><tr><td>Filling gas</td><td>–</td><td>95 % Argon/5% Helium</td></tr><tr><td>Filling gas pressure</td><td>MPa</td><td>4</td></tr><tr><td>Filling gas temperature</td><td>°C</td><td>25</td></tr></tbody></table>

phase is around 300 s. The third phase concerns the accidental transient. At the beginning of this phase, the pressure flask is connected to a blowdown tank by opening two valves. The time length of the last phase is nearly 600 s.

### 3.3. Main results of the HALDEN IFA-650.10 test

Upon opening the discharge valve, loop depressurization and rapid vaporization of the coolant fluid occur. The power supplied to the fuel rod and housing leads to a temperature rise towards the target temperature of 850 °C. During this temperature increase, ballooning of the fuel rod occurred, followed by the cladding failure at 249 s after the initial moment of depressurization, at a temperature of approximately 755 °C. The test was concluded at 418 s by reactor shutdown, after the temperature at the lower part of the fuel rod reached a maximum of 841 °C.

A more comprehensive description of the main online measurements as well as post-test measurements of the fuel rod are provided in the references (Lavoil, 2010) and (Oberländer, 2016). The Fig. 3 depicts the appearance of the fuel rod observed in gamma scanning, showing a slight bowing of the fuel rod, a rupture near the median plane, and a small quantity of fuel at the bottom of the test device outside the fuel rod.

The results of post-test examinations reveal a burst opening of 15 mm long and 5 mm width located approximately at mid-height of the active fuel stack (around 24 cm from the bottom of the fuel). The maximum local diameter measured at the opening was 12.5 mm, indicating a maximum local deformation of 31.6 %. The Fig. 4 provides a view of this opening.

About cladding oxidation, the post-examination showed a nominal cladding external oxide thickness of about 30–40 µm, which is 10 µm more than the thickness measured initially. Oxide layer is also observed on the internal face of the cladding, a thickness of about 10–15 µm. At the start of the test, no internal oxide layer is mentioned.

Finally, the hydrogen content in the cladding was measured at several position. Below the burst, the hydrogen concentration is closed to the initial one (around 300 ppm). The maximum value, 767 ppm, were measured just below the burst location.

### 4. Calculation assumptions

The objective of the IFA-650.10 test calculation is to reproduce the thermomechanical behaviour of the fuel rod sample, which is mainly

<figure><img src="image_5.png" /><figcaption>Fig. 3. HALDEN IFA-650.10 – Gamma scans.</figcaption></figure>

<footnote>* measured on a sibling rod (Nishi and Lee, 2001).</footnote>
<figure><img src="image_6.png" /><figcaption>Fig. 4. HALDEN IFA-650.10 – view of the cladding burst.</figcaption></figure>

dependent on the thermohydraulic conditions in the HALDEN device and the loop. Therefore, the DRACCAR code from the FUEL+ platform is the suitable tool for simulating this test as it allows for simulating both the thermomechanical behaviour of the fuel rod, thermohydraulics in the loop, and the heat exchange between the fuel rod and the fluid.

After the fuel rod failure, the test continues with an oxidation phase accompanied by hydrogen pick-up as mentioned in the previous paragraph. Oxidation is managed by DRACCAR in stand-alone mode, but the dedicated model based on oxygen mass and oxide thickness gain correlations is very simple. Moreover, the hydrogen pick-up is not managed by the code, the hydrogen concentration considered in several DRACCAR models is therefore kept constant. To address this phenomenon and to have a more precise evaluation of the cladding oxidation, the FUEL+/SHOWBIZ code, which proposes detailed models to study the influence of O/H species on Zr cladding, is coupled to DRACCAR for the simulation of the HALDEN test.

Finally, to conduct this simulation, the DRACCAR code requires initial data of the rod sample, particularly concerning the fuel (geometry, burnup, gas in free volumes and fuel open porosities and intergranular gas in fuel). To obtain these data, it is necessary first to perform the base irradiation calculation of the parent rod, and then subsequently, the re-irradiation calculation of the sample in the HALDEN reactor. Both simulations are conducted using the FUEL+/FIRST code.

In conclusion, simulating the HALDEN IFA-650.10 test requires the completion of three calculations:

* A base irradiation calculation of the parent fuel rod using the FUEL+/FIRST code.

* A re-irradiation calculation of the rod sample in the HALDEN reactor carried out with the FUEL+/FIRST code, utilizing some of the results from the previous calculation as initial conditions and corresponding to the conditioning phase of the test sample in HALDEN reactor.

* The LOCA calculation itself performed with the DRACCAR code coupled with SHOWBIZ code and using the results from the previous calculation as initial conditions.

The main assumptions of these three calculations are described hereafter.

## 4.1. Base irradiation calculation assumptions

For the base irradiation calculation with FUEL+/FIRST code, the parent rod included into its fluid sub-channel is modelled. The parameters presented in Table 1 are used to define the rod geometry and material. The data needed by the code to define the rod upper plenum free volume are not given in this table because they are not known. These data are relative to the spring located inside the rod upper plenum as well as the plenum lengths. Plenum lengths and volumes of several PWR nuclear fuel rods are provided in reference (Montgomery and Morris, 2019). For the present base irradiation calculation, it has been decided to set the upper plenum length to 180 mm and a plenum volume equal to 8 cm³. Sensitivity calculations were performed using FUEL+/FIRST, considering the height/volume pairs provided in the reference (Montgomery and Morris, 2019), with extrema at 172 cm/7.2 cm³ and

190 cm/8.2 cm³. The impact of the plenum height and volume is negligible on the target results of the present study, that are the burnup, the fuel and cladding geometries, the quantity of gas in the fuel, the cladding hydrogen content, and the oxide thickness. The impact on the rod inside pressure is more important but as the rod sample is refilled with gas during the refabrication process, and before being used in HALDEN facility, it is not considered as an issue.

Concerning the meshing, the father rod active part is separated into 100 axial meshes and only 1 axial mesh is used for the upper plenum. Radially, the fuel pellet is cut into 16 meshes and the cladding into 8 meshes.

The simulated scenario consists of reproducing the 5 cycles that the fuel rod has passed through in the reactor. The power history coming from the reference (Nishi and Lee, 2001) (Fig. 1) is given function of the burn-up. In the FUEL+/FIRST input deck, the power should be set as a function of time. So, average rod power function of average burnup is converted in time evolution of linear heat rate with fitting end of life average burnup (53.8 MWd/kgU (Lavoil, 2010) through a FRAPCON-4.0 calculation (this option is not yet available in the FUEL+/FIRST code). Furthermore, the axial power profile during each cycle would have been necessary for performing the irradiation calculation. Unfortunately, only the axial profile at the fuel rod end-of-life (Fig. 2) is available (Nishi and Lee, 2001). Therefore, it was decided to set this profile throughout the irradiation period.

Finally, the thermohydraulic operating conditions of the reactor in which the fuel rod was located are not provided. Therefore, characteristic operating conditions of a PWR were used, namely a pressure of 155 bar, a core inlet temperature of 290 °C, and a core flow rate of 3750 kg/m²/s.

## 4.2. Re-irradiation calculation assumptions

The re-irradiation simulation has been done with FUEL+/FIRST code. The modelling consists in a single rod included in a fluid-subchannel having a fluid section equal to 242 mm² and a hydraulic diameter equal to 10.4 mm, corresponding exactly to the hydraulic diameter and the fluid section of the fluid channel delimiter by the heater in the HALDEN device that has an internal diameter equal to 20 mm.

The geometry of the active part of the single rod is defined using the results of the base irradiation calculation on the parent rod. All the data of the parent rod (rod geometry, fuel composition, cladding composition, meshing) between the elevations 1280 mm and 1720 mm are used as input data for defining the active part of the rod sample.

Then, the data given in the Table 2 is used to define the lower and upper plena of the rod sample. Similarly, the rod inside pressure and gas composition are set at the values given in the table.

In summary, the HALDEN rod modelled with FIRST consists of a fissile zone 440 mm in length, a lower plenum with a height of 387.5 mm and a volume of 17.42 cm³, and an upper plenum with a height of 28 mm and a volume of around 0.2 cm³. Regarding the meshing, the active zone is divided into 10 axial cells, which exactly matches the number of cells in the corresponding section of the parent rod. One axial cell is used for each of the plena. Radially, the fuel pellet is cut into 16 meshes and the
cladding into 8 meshes as for the father rod.

Finally, the experimental measurements of the rod power, the inlet flowrate, the inlet temperature and the inlet pressure are directly used by the code FUEL+/FIRST as boundary conditions.

The end of the calculation is set at the time 9.433 days, 300 s before the break opening.

### 4.3. LOCA calculation assumptions

LOCA simulation has been performed with FUEL+/DRACCAR using the CESAR module for the thermal-hydraulics calculation. The simulation starts at the ending time set for the previous FIRST simulation.

The modelling used for the calculation is the same as the one presented at the EHPG meeting in 2013 (Glantz, 2013). A portion of the pressure flask which contains the fuel rod is modelled in 3D. The 3D modelling inlet is located at the fluid separation level between the annular space created by the heater external wall and the flask internal wall in one hand and, on the other hand, the annular space created by the internal wall heater and the rod external clad. It ends at the bottom of the flask outlet pipe. It is composed of three main structures:

* The fuel rod, which is constituted by the $UO_2$ pellets, its cladding in Zircaloy-4, an upper plenum and by the inlet and outlet plugs in the same material. The geometrical data used to model the fuel rod are the results of the re-irradiation calculation. This is a difference between the modelling used in 2013 where the parent rod data were directly used to define the fuel rod.

* The zircaloy electrical heater surrounding the fuel rod.

* The pressure flask wall in steel.

Regarding the fuel rod, the DRACCAR representation aims to closely match the geometry of the HALDEN rod, as follows:

* A lower section representing the bottom end plug, with a height of around 28 mm. The outer diameter is taken from the FIRST calculation. The inner diameter was determined to preserve the free volume specified in the HALDEN rod design drawings, i.e., approximately 0.2 $cm^3$, a volume that remained unchanged during the FIRST calculation.

* The fissile height is 440 mm. The inner and outer diameters of the pellets and the cladding are taken from the FIRST calculation and vary along the axial coordinate.

* The upper plenum has a height of 363 mm and is divided into two sections of respective heights 90 mm and 273 mm. The outer diameter of both parts is from the FIRST calculation. The inner diameter was adjusted to preserve the free volumes provided in the HALDEN drawings (the total volume remained unchanged during the FIRST calculation) – approximately 0.70 $cm^3$ for the first part and 15 $cm^3$ for the second.

* The top end plug is 24.5 mm in height. As for the other components, the outer diameter is taken from the FIRST calculation. The inner diameter was recalculated to account for the free volume in the plug (0.24 $cm^3$), as well as the volumes of the pressure measurement device and its connecting line to the HALDEN rod, totalling 1.27 $cm^3$. This approximation should not impact the pressure evolution, as the temperature of these small volumes is, according to the reference (Lavoil, 2010), around 235 °C – very close to the temperature of the top part of the rod during the test (excepting during the blowdown phase).

From a meshing standpoint, the rod is discretized in all three directions and consists of four azimuthal cells along its entire length. A single radial cell is used for the end plugs, the upper plenum, and the cladding. The fuel pellets column is divided into 16 radial cells, as in the FIRST calculation. As for the axial meshing, 24 cells are used for the fissile part, 6 cells for the upper plenum, and 2 cells for each of the end

plugs.

The data transferred from FIRST to DRACCAR include the geometric data for fuel and cladding, the fuel data (axial and radial burnup profiles, ...), the cladding data (oxide thickness) and axial and radial temperature profiles. In the azimuthal direction, the FIRST results are uniformly distributed over the four DRACCAR angular cells. In the radial direction, the same fuel meshing in FIRST and DRACCAR is used to avoid interpolation of the fuel data in this direction. In the cladding, DRACCAR's mechanical model allows only one radial mesh, whereas in FIRST the cladding is divided into 8 radial cells. Interpolation is only needed for the cladding temperature. Since the temperature profile in the cladding is nearly flat at the end of the FIRST calculation (variation below 2 °C), the average value between the inner and outer cladding surfaces is imposed.

In the axial direction, the DRACCAR meshing was refined to better capture the ballooning shape, increasing from 10 axial cells in FIRST to 24 in DRACCAR. Linear interpolation is applied for geometric and temperature data, while other fuel data are interpolated using a volume-overlap weighting method.

The physical models which are activated in the HALDEN IFA-650.10 calculation are:

* The heat exchange by conduction inside the structures (fuel, cladding, heater), by radiation between the structures (fuel and cladding through the gap if it is open, heater and cladding, heater and flask) and by convection between the fluid and the structures.

* The mechanical creep of the cladding.

* The cladding burst which occurs if the temperature criteria for Zry-4 material (Chapman, 1979) is exceeded.

* The fuel relocation. The main input parameter of this model is the filling ratio (or packing fraction), which is defined as the ratio of the volume of fuel to the total available local volume. The correlation given the filling ratio function of circumferential strain for slightly fragmented fuel proposed in reference (Taurines et al., 2021) is used to define this parameter.

* The axial gas transport inside the rod with an effective permeability value equal to $10^{-12} m^2$ in the active fuel part.

* The fission gas release. This model is parametric in the DRACCAR code. It enables the release of a specified quantity of gas based on the fuel temperature. The initial amount of gas in the fuel is the results of the FUEL+/FIRST re-irradiation calculation. Regarding the quantity of fission gas released as a function of temperature, the law proposed in the reference (Taurines et al., 2021) for the test HALDEN IFA-650.14 is used, which involves a linear release from 0 to 15.9 % for a temperature range of 535 to 945 °C.

* The coupling with FUEL+/SHOWBIZ. The oxidation of the cladding external face and cladding hydriding are computed by SHOWBIZ all along the transient and the oxidation of the cladding internal face is activated only after the cladding burst.

The remaining of the HALDEN loop is modelled using 1D elements. The exact geometry of all pipes constituting the HALDEN loop has been replicated. DRACCAR computes only the thermal-hydraulic in this part of the circuit.

As mentioned in reference (Glantz, 2013), one of the main uncertainties in the HALDEN test is the heat losses of the whole facility. The available information is not sufficient to precisely define the heat losses, and yet they have a great importance on HALDEN simulation results, particularly on the blowdown length time. Indeed, the wall temperature depends on the heat losses and this temperature directly impacts the heat transfer with the fluid during the transient and indirectly influences the initial fluid temperature at the opening of the break. Moreover, during the natural circulation phase, the fluid temperature in the pipe where there is no circulation will decrease due to the heat losses – as no fluid measurement in the HALDEN circuit is available, the fluid temperature map in all the pipes just before the opening of the
break cannot be known precisely. So, for the calculation presented hereafter the heat losses have been fitted to obtain a blowdown duration in the calculation close to the experimental value.

The DRACCAR/SHOWBIZ coupling allows for the calculation of oxygen diffusion into the cladding as well as its hydriding. This is a one-way coupling: in its current version, the SHOWBIZ computation has no effect on the progression of the DRACCAR calculation. In particular, it does not modify the cladding properties that could influence its creep behaviour. This coupling is limited to the cladding over the full height of the fissile zone and is applied on a single azimuth, as the DRACCAR results are identical regardless of the azimuth considered. Regarding the meshing, at each axial level, the DRACCAR cladding mesh is refined into 1000 radial cells. At the start of the LOCA calculation, the average hydrogen concentration in the cladding given by FIRST is interpolated using a domain-overlap method to determine the hydrogen concentration to be applied in each radial cell. At each time step, the cladding temperature from DRACCAR is imposed in each radial cell of the corresponding axial level in SHOWBIZ.

# 5. Calculation results

## 5.1. Base irradiation calculation

Using the power history from reference, the total residence time in the reactor is approximately 1500 days. At the end of the FUEL+/FIRST calculation, the rod average burnup obtained is equal to 56.2 MWd/kg, very close to the burnup given in the reference (Lavoil, 2010) (56.2 MWd/kg). The axial burnup profile is shown in the Fig. 5. The segment of the fuel rod used in the HALDEN test is located at the level of the 3rd intergrid between 1.2 m and 1.7 m elevations. The maximum calculated burnup for this segment is 60.9 MWd/kgU, which agrees with the value given in the reference (Lavoil, 2010) (61 MWd/kgU).

The external oxide thickness profile is given in the Fig. 6. The maximum value is equal to around 80 µm at the elevation 3.25 m. At the level of the 3rd grid spacer, the oxide thickness is included between 20 and 30 µm, corresponding of the values measured on the HALDEN rod sample (Nishi and Lee, 2001).

The Fig. 7 presents the hydrogen content in the cladding. The maximum value of approximately 550 ppm is reached at the elevation of 3.25 m. The hydrogen content is comprised between 150 and 220 ppm at

<figure><img src="image_7.png" /><figcaption>Fig. 5. IFA-650.10 parent rod – Burnup axial profile.</figcaption></figure>

<figure><img src="image_8.png" /><figcaption>Fig. 6. IFA-650.10 parent rod – Oxide thickness axial profile.</figcaption></figure>

<figure><img src="image_9.png" /><figcaption>Fig. 7. IFA-650.10 parent rod – Hydrogen content axial profile.</figcaption></figure>

the level of the 3rd grid spacer, values are also in good agreement with the measurements (Nishi and Lee, 2001).

Finally, the axial profile of the gas retention at the end of the calculation is given in the Fig. 8. The values obtained with the code FUEL+/FIRST at the level of 3rd grid spacer are between 1.62 and 1.64 cm³STP/g, within the expected range for UO₂ fuel with a burnup of 60 MWd/kg. Here, STP refers to Standard Temperature and Pressure conditions (0 °C and 1 atm).

To conclude, the main characteristic of the HALDEN rod sample used during the LOCA test IFA-650.10 are indeed reproduced by the FUEL+/FIRST code using the data and assumptions listed in the paragraph 4.1. The results characterizing the rod at the level of the 3rd spacer grid
<figure><img src="image_10.png" /><figcaption>**Fig. 8.** IFA-650.10 parent rod – gas retention in the fuel.</figcaption></figure>

(geometry and composition) will be used as input data for a second FUEL+/FIRST calculation to simulate the re-irradiation phase of the sample in the HALDEN reactor.

## 5.2. Re-irradiation calculation

The re-irradiation in HALDEN reactor last approximately 9 days, the final time being set 300 s before the onset of the LOCA scenario. The Fig. 9 presents the rod inside pressure evolution calculated by FUEL+/FIRST (blue curve, overlapping with the black curve) and the experimental measurement (red curve). The two evolutions are very close, the pressure at the end of calculation being equal to 69.7 bar against 71 bar

<figure><img src="image_11.png" /><figcaption>**Fig. 9.** IFA-650.10 sample rod – internal pressure during the re-irradiation phase.</figcaption></figure>

in the experiment. Concerning the cladding temperature evolution (Fig. 10), a difference of approximately 120 °C is observed between the FUEL+/FIRST calculation result (blue curve) and the experimental measurement (red curve) throughout most of the transient (corresponding to a power plateau of the re-irradiation phase). This gap between the calculation and the experiment is explained by the fact that wall to fluid heat exchanges in the FUEL+/FIRST code can only occur through liquid convection, which is the only type of exchange available in the current version of the code. However, the heat exchanges between the rod and the fluid in the HALDEN device are characteristic of a nucleate boiling regime. Indeed, the main condition to be in this regime is for the wall temperature to be at least equal to the saturation temperature of the fluid (around 245 °C at 70 bar). It is the case in the experiment. In nucleate boiling regime, the wall to fluid heat transfer coefficient is around 15000 W/m²/K, six times higher than the heat transfer coefficient calculated by FUEL+/FIRST under liquid convection regime using the Dittus-Boelter correlation (Dittus and Boelter, 1985). The value determined by this correlation during power plateau of the re-irradiation phase is approximately 2500 W/m²/K. To verify this hypothesis, a sensitivity analysis was conducted by imposing a heat transfer coefficient of 15000 W/m²/K in a FUEL+/FIRST calculation, instead of using the value computed by the Dittus-Boelter correlation implemented in FIRST. The result of this calculation is the black curve on the Fig. 10, which closely matches the experimental curve. A development direction for the FIRST software to broaden its scope of applicability is to incorporate correlations allowing for the calculation of wall to fluid heat transfer in different flow regimes.

Furthermore, it should be noted that the increase in heat transfer between the cladding and the fluid does not alter the evolution of the internal pressure calculated by the code, the black curve, corresponding to the sensitivity calculation with a constant heat transfer coefficient, overlaps with the blue curve from the initial calculation. In fact, in FIRST, instantaneous pressure equilibrium is assumed within the rod. Therefore, to calculate the internal pressure, the ideal gas law is applied, summing the contributions of each axial slice. In the region where the temperature is overestimated, the gap is closed, meaning that the number of moles available in that zone is effectively zero, and thus this region does not contribute to the pressure. As a result, the rod pressure is

<figure><img src="image_12.png" /><figcaption>**Fig. 10.** IFA-650.10 parent rod – cladding temperature during the irradiation phase.</figcaption></figure>
solely determined by the contribution of the cold plena, for which the temperature calculated by FIRST (grey curve in the Fig. 10) matches the experimentally measured temperature (green curve in the same figure).

Thus, the FIRST results presented in the rest of this article are obtained with this imposed heat transfer coefficient.

The Fig. 11 provides calculated axial profiles of external pellet diameters (black points and red line) and internal cladding diameters (green point and blue line) respectively at the beginning and at the end of the calculation. The results obtained are consistent with what is expected, namely an increase in diameters due to a transition from a cold state of the rod (re-fabrication carried out at 25 °C) to a hot state of the rod, the cladding temperature being around 240 °C and the fuel centre temperature being approximately at 270 °C at the end of calculation.

The short duration of the sample's re-irradiation in the HALDEN reactor do not lead to an increase in the oxide layer, as can be seen in the Fig. 12. The same applies to the hydrogen content in the cladding and the amount of gas trapped in the fuel.

To conclude, for this calculation, it has been chosen to impose the experimentally measured power as a boundary condition as well as the thermohydraulic conditions at the inlet (pressure, flow rate, and temperature). The heat transfer coefficient between the rod and the fluid has also been set at a constant value characteristic of a nucleate boiling regime. With these assumptions, the FUEL+/FIRST calculation code successfully calculates the evolution of cladding temperature and rod inside pressure. However, further development is necessary so that the code can calculate by itself the heat transfer coefficient between the rod and the fluid in flow regimes other than liquid convection.

### 5.3. LOCA calculation

The results of the DRACCAR/SHOWBIZ calculation performed with the assumptions presented in the paragraph 4.3 are detailed in the two separate parts, the first one dedicated to the thermal-hydraulic results and the second one to the thermomechanical results.

#### 5.3.1. Thermal-hydraulic results

A comparison between experimental and simulated thermal-hydraulic behaviour in HALDEN loop test is difficult because the loop is

<figure><img src="image_13.png" /><figcaption>Fig. 12. IFA-650.10 sample rod – oxide thickness profiles.</figcaption></figure>

very lightly instrumented. The interesting measurements are the pressure evolution and the evolution of the fluid temperature at the inlet and outlet of the HALDEN device containing the rod sample.

The comparison between the measured and simulated pressures in the HALDEN loop is presented in Fig. 13. In the simulation and in the experiment, the pressure decreases almost instantaneously from nearly 70 bar (conditioning pressure) to about 15 bar when the break is opened (at 300 s). Then, the depressurization predicted by DRACCAR qualitatively matches the experimental one, namely a momentary pause in pressure drops for about 50 s, followed by a new decrease until reaching a final plateau. Quantitatively, DRACCAR underestimates the pressure values measured on both plateau by about 2 bar. In the simulation, the value of the final plateau pressure is equal to the pressure measured at

<figure><img src="image_14.png" /><figcaption>Fig. 11. IFA-650.10 sample rod – fuel external (FDE) and clad internal (CDI) diameters profile.</figcaption></figure>

<figure><img src="image_15.png" /><figcaption>Fig. 13. IFA-650.10 test – loop pressure evolution.</figcaption></figure>
the outlet of the blowdown tank. It is not the case in the experiment, which is surprising considering that the location where the pressure is measured is directly connected to this tank by the discharge line. The available data regarding the HALDEN loop are not sufficient to understand this behaviour and explain the difference between the simulation and the experiment. Moreover, it can be noticed that the blowdown duration is a little shorter in the calculation (88 s) than in the experiment (96.5 s).

The Fig. 14 shows the measured and calculated evolutions of fluid temperatures at the inlet and outlet of the HALDEN device containing the rod. The DRACCAR results generally match the experimental measurements. The main difference is observed after time 500 s, with the increase in inlet temperature of the device occurring earlier in the calculation than in the experiment. However, it should be noted that from this moment, the fluid temperature at the inlet of the device is higher than the outlet temperature in both the calculation and the experiment, indicating that the steam flow cooling down the rod is downward.

### 5.3.2. Thermomechanical and chemical results

Fuel rod and heater behaviour is directly correlated to the circuit depressurization. As the pressure decrease is well estimated by DRACCAR calculation, the start of the cladding and heater temperature increase is well predicted as it can be seen on the Fig. 15 and Fig. 16 respectively.

In the experiment, the clad temperature is measured at two axial positions: close to the bottom of the fuel rod (thermocouple TCC1 at an elevation of 95 mm from the bottom of the fuel active length) and close to the top of the fuel rod active part (thermocouples TCC2 and TCC3 at an elevation of 361 mm from the bottom of the fuel active length). For the upper thermocouples positions (TCC2 and TCC3), the temperature increases until the burst is well estimated by the code (Fig. 15), with the burst occurring at approximately 550 s in both the calculation and the experiment. Just before cladding burst, the temperature increase slows down in the calculation, likely due to the beginning of clad ballooning, which lead to a decrease of the heat exchange between the pellets and the clad. At the burst time (which is correctly estimated by DRACCAR), a slight temperature drop is experimentally observed on these two

<figure><img src="image_16.png" /><figcaption>**Fig. 14.** IFA-650.10 test – fluid temperature evolution at the inlet (TFI) and outlet (TFO) of the HALDEN device.</figcaption></figure>

<figure><img src="image_17.png" /><figcaption>**Fig. 15.** IFA-650.10 test – clad temperature evolution.</figcaption></figure>

<figure><img src="image_18.png" /><figcaption>**Fig. 16.** IFA-650.10 test – heater temperature evolution.</figcaption></figure>

thermocouples, with a more pronounced decrease on thermocouple TCC3. According to the reference (Lavoil, 2010), this temperature drop can be explained by the displacement of “cold” gas, from the upper plenum toward the burst zone. As the gas flows through the region where the thermocouples are positioned, it locally cools the sensors, inducing a transient decrease in the measured temperature. This phenomenon cannot be reproduced by the code, as no energy balance is associated with the gas transport model in DRACCAR. Then, the temperature increases again both in the simulation and in the experiment until reaching a maximum of approximately 850 °C. For the lower thermocouple, the simulation and the experiment are in good agreement, especially before the cladding failure. At the burst time, fuel relocation is slightly more significant in the calculation than in the
experiment, leading to a maximum temperature in the calculation of about 880 °C compared to 850 °C in the experiment.

The temperature measurements on the heater are made at two axial positions, at 95 mm (TCH1) and 195 mm (TCH2) from the bottom of the fuel active length. The calculation results are in good agreement with experimental measurements at both positions (Fig. 16).

Then, the upper plenum temperature evolution is quite well calcu-
lated by DRACCAR as it can be seen on the Fig. 17, both the minimum
temperature reached at the end of the circuit blowdown phase and the
temperature increase occurring just after being underestimated by the
code. Consequently, the inside rod pressure at the end of the blowdown
phase is also underestimated by the code as it can be seen on the Fig. 18.
However, the internal pressure of the rod increases more rapidly in the
calculation than in the experiment from this point onward, although the
opposite trend is observed in the plenum temperature. The evolution of
the rod inside pressure of the rod is strongly linked to the evolution of
the gas temperature in the free volumes, as well as to the release of
fission gas. In the DRACCAR calculation, it has been set that the fission
gas release is possible from a fuel temperature equal to 500 °C, which is
the case few seconds after the end of the blowdown phase. Based on the
experimental internal pressure evolution, it is likely that there is little to
no release of fission gas during this test. Just before cladding burst (at
approximately 550 s), the pressure predicted by DRACCAR is in good
agreement with the experimental measurement, the slight pressure
decrease due to rod ballooning between time 530 s and 550 s is specif-
ically well reproduced by the code.

The burst time can be identified by the sharp pressure drop, which
occurs both in the simulation and in the experiment at 550 s. However, it
is observed that the stabilized pressure value after cladding burst differs
between the experiment and the calculation. Indeed, DRACCAR predicts
a stabilization of the internal pressure at the HALDEN loop pressure,
whereas in the experiment, the rod pressure remains around 12 bar, i.e.,
almost 6 bar higher than the circuit pressure. An overpressure in the rod
relative to the circuit pressure would imply a complete blockage of gas
flow inside the rod between the plenum and the burst location, a
blockage that would persist for the remainder of the test, in particular
during the observable cooling phase which begin at around 650 s.
During this phase, it is highly likely that the pellet-to-cladding gap, if it
had been closed, reopens due to the differential contraction between
UO₂ and the cladding material. Therefore, it is highly likely that a failure

<figure><img src="image_19.png" /><figcaption>Fig. 18. IFA-650.10 test – rod inside pressure.</figcaption></figure>

of the pressure sensor occurred at the moment of cladding burst, a hy-
pothesis that cannot be verified, as no information on this point was
found in the HALDEN documentation related to this test.

The Fig. 19 presents a comparison between calculated and experi-
mental profiles of outer cladding diameters at the end of the test. The
external cladding profile is well reproduced by DRACCAR, except at the
rupture point where the deformation is underestimated by the code.
Only secondary creep is modelled by DRACCAR and so simulation
cannot be used to predict burst opening dimension and the associated
local cladding shape at burst location. As the cladding deformation is
limited in this test, there is little fuel relocation calculated by DRACCAR.
Indeed, as in it can be seen in the Fig. 20, the minimum filling ratio
reached during the calculation is equal to 0.85 corresponding to a strain

<figure><img src="image_20.png" /><figcaption>Fig. 17. IFA-650.10 test – upper plenum temperature.</figcaption></figure>

<figure><img src="image_21.png" /><figcaption>Fig. 19. IFA-650.10 test – rod external diameter.</figcaption></figure>
<figure><img src="image_22.png" /><figcaption>Fig. 20. IFA-650.10 test – calculated filling ratio profile.</figcaption></figure>

equal to 10 % according to the filling ratio law coming from reference (Taurines et al., 2021) and set in the relocation model of DRACCAR. The selected filling ratio law seems to be consistent with experiment, the impact of fuel relocation on the cladding temperature being well reproduced by DRACCAR.

The deformation profile is consistent with the axial cladding tem-
perature profile calculated by DRACCAR just before rod burst, with the
maximum deformation occurring between elevations 1.1 m and 1.2 m,
corresponding to the peak temperature prior to rupture (see Fig. 21).
When compared to experimental measurements taken at three eleva-
tions, it appears that DRACCAR slightly underestimates the cladding
temperature by less than 10 °C. Experimentally, the burst is observed at
an elevation of 1.13 m, which is very close to the value calculated by

DRACCAR, 1.11 m. In both cases, this elevation lies within the calcu-
lated maximum temperature plateau of the profile just before burst.
After the burst (Fig. 22), DRACCAR overestimates the cladding tem-
perature (thermocouple TCC1, TTC2 or TCC3), with a maximum of
25 °C, which is likely due, on the one hand, to an excessive fuel relo-
cation, and on the other hand, to the lack of modelling of the cooling
effect on the thermocouples located at the 1.266 m elevation caused by
the plenum gas flow during the rod depressurization. In contrast, the
plenum temperature (TP) is accurately predicted. Additionally, the
Fig. 22 shows a small experimental difference between the temperatures
measured by thermocouples TCC2 and TCC3 (around 15 °C), thermo-
couples that appear to be unevenly cooled by the gas coming from the
plenum.

The comparisons between experimental and calculated internal and external oxide thicknesses are respectively presented in Fig. 23 and Fig. 24. The coupling between DRACCAR and SHOWBIZ provides a good estimation of the internal oxide thickness, within the range of minimum and maximum values measured at the end of the test. The external oxide layer is underestimated by DRACCAR/SHOWBIZ, especially just below and at the burst position. Moreover, the growth of the external oxide layer in the calculation is only possible if radial cracks in the oxide are defined. One explanation is that the simulation assumes that zirconia is homogeneously distributed on cladding external surface even during deformation. The possible radial cracks appearing during deformation and exposing metal that therefore could be oxidized is not modelled. This assumption seems justified, such cracks being well observed experimentally at the end of the test.

The hydrogen content in the cladding at the end of the test is pre-
sented in the Fig. 25. It turns out that DRACCAR/SHOWBIZ predict an
increase of approximately 50 ppm of hydrogen content in the cladding
during the LOCA transient (green curve compared to red curve). The
hydriding is possible in the calculation due to the hydrogen production
caused by cladding oxidation. Indeed, the hydrogen partial pressure
calculated close to the burst position increase significantly (Fig. 26) at
time 540 s when the cladding temperature is greater than 700 °C. The
maximum hydrogen pressure is approximately 0.05 bar, occurring at the
time of maximum temperature. The DRACCAR/SHOWBIZ coupling re-
produces well the hydrogen peak measured in the experiment. In the
calculation, the hydriding of the cladding is the consequence of the

<figure><img src="image_23.png" /><figcaption>Fig. 21. IFA-650.10 test – cladding temperature profile before burst time.</figcaption></figure>

<figure><img src="image_24.png" /><figcaption>Fig. 22. IFA-650.10 test – cladding temperature profile after burst time.</figcaption></figure>
<figure><img src="image_25.png" /><figcaption>Fig. 23. IFA-650.10 test – Internal oxide thickness at the end of the test.</figcaption></figure>

<figure><img src="image_26.png" /><figcaption>Fig. 24. IFA-650.10 test – External oxide thickness at the end of the test.</figcaption></figure>

hydrogen pickup associated with the oxidation process and the direct hydrogen uptake from the dihydrogen in the gas phase. In this case, the hydrogen uptake from the gas phase at the internal face of the cladding is the predominant phenomenon leading to the formation of the hydrogen peak. This mechanism is possible because the H2 production by the oxidation of the internal face of the cladding is not transported by the coolant, leading to an important concentration of hydrogen inside the rod at the balloon position.

The Table 3 resumes the cladding main characteristics at burst time for the LOCA test HALDEN IFA-650.10. The error between calculation and experiment are lower than 5 % except for the burst elevation (8.40 %) and the maximal external diameter (15.2 %). To conclude, the

<figure><img src="image_27.png" /><figcaption>Fig. 25. IFA-650.10 test – Hydrogen content profile.</figcaption></figure>

<figure><img src="image_28.png" /><figcaption>Fig. 26. IFA-650.10 test – Hydrogen pressure evolution.</figcaption></figure>

<table><caption>Table 3
HALDEN IFA-650.10 tests – cladding characteristics at burst time.</caption><thead><tr><th>Parameters</th><th>Experiment</th><th>Calculation</th><th>Error Calc/Exp</th></tr></thead><tbody><tr><td>Burst time*</td><td>549 s</td><td>551.5 s</td><td>0.46 %</td></tr><tr><td>Burst elevation (from the fuel active length bottom)</td><td>262 mm</td><td>240 mm</td><td>8.40 %</td></tr><tr><td>Burst pressure</td><td>70.5 bar</td><td>68.7 bar</td><td>2.55 %</td></tr><tr><td>TCC1 temperature at burst time</td><td>754.5 °C</td><td>737.6 °C</td><td>2.24 %</td></tr><tr><td>TCC2 temperature at burst time</td><td>750.4 °C</td><td>735.0 °C</td><td>2.05 %</td></tr><tr><td>TCC3 temperature at burst time</td><td>748.8 °C</td><td>735.0 °C</td><td>1.84 %</td></tr><tr><td>Plenum temperature at burst time</td><td>233.5 °C</td><td>223.6 °C</td><td>4.23 %</td></tr><tr><td>Maximum external diameter</td><td>12.5 mm</td><td>10.6 mm</td><td>15.2 %</td></tr></tbody></table>

<footnote>* LOCA transient starts at time 300.s.</footnote>
DRACCAR prediction of the LOCA test HALDEN IFA-650.10 is very accurate.

## 6. Conclusion and perspectives

The test HALDEN IFA-650.10, which reproduces the behaviour of an irradiated fuel during the first phases of a Loss-Of-Coolant Accident, has been calculated with the ASNR codes of the FUEL+ platform.

To simulate this test, three calculations were performed. The irradiation code FUEL+/FIRST has been used for an initial calculation which reproduces the irradiation of the parent rod, from which the sample used for the HALDEN test was taken. The second calculation, performed also with the FIRST code, simulates the conditioning phase that the sample underwent in the HALDEN reactor. The input data for this second calculation are the results obtained from the initial parent rod calculation. The results from this second calculation are then used to initialize the third calculation, specifically dedicated to LOCA transient and performed with the code FUEL+/DRACCAR in charge of simulating the rod sample thermomechanical behaviour and the whole circuit thermal-hydraulic, coupled to the FUEL+/SHOWBIZ, allowing a finer resolution of the thermochemical behaviour (oxidation, hydriding) of the cladding.

The results obtained from these three simulations are very satisfactory. Indeed, the irradiation calculation of the parent rod provides data on the fuel (burnup) and the cladding (oxide thickness, hydrogen content) that are consistent with what was experimentally measured on the sample used for the IFA-650.10 test before the LOCA transient. The second calculation corresponding to the irradiation in the HALDEN reactor of the rod sample coming from the parent rod allows estimating the geometry of the fuel rod sample just before the LOCA transient and showed that the external oxide thickness, hydrogen content, and the amount of gas trapped in the pellet had not changed during this reirradiation phase compared to what had been determined for the parent rod. Finally, the results of the LOCA calculation are in good agreement with the experimental measurements, with discrepancies between calculations and experiments regarding the burst time and the main cladding characteristics at that moment being less than 5 % for most parameters.

The entirety of these results constitutes a positive assessment of the various codes as well as the chaining/coupling methodology used to simulate the HALDEN IFA-650.10 test. This methodology has proven functional and suitable for this type of exercise. The next step will be to apply it to the other tests of the HALDEN LOCA program, all included in the DRACCAR validation database, to improve the code's predictive capabilities.

## CRediT authorship contribution statement

T. Glantz: Writing – original draft, Software, Methodology. S. Belon: Software, Methodology. M. Salvo: Software. M. Guemas: Software.

## Declaration of competing interest

The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.

## Data availability

The authors do not have permission to share data.

## References

Chapman, R., 1979. Multi-rod Burst Test, Program Progress report April-June 1979 USNRC report - ORNL/NUREG/CR-1023, Oak Ridge National Laboratory.

Dittus, P.W., Boelter, L.M.K., 1985. Heat Transfer in Automobile Radiators of the Tubular Type. Univ. Calif. Pub. Eng., 2/13, 443-461, 1930; reprinted in Int. Comm. Heat mass Transfer, 12, 3-22.

Glantz, T., Taurines, T., De Luze, O., Belon, S., Guillard, G., Jacq, F., 2018. DRACCAR: a multi-physics code for computational analysis of multi-rod ballooning, coolability and fuel relocation during LOCA transients Part one: general modelling description. Nucl. Eng. Des. 339, 269-285.

Glantz, T., 2013. Simulation of HALDEN tests IFA-650.4, 650.9 and 650.10 with the DRACCAR code. Enlarged HALDEN programme group meeting. Proceedings of the Fuel and Materials Sessions, Volume 1, Session F1.3, report HPR-278.

Guemas, M. et al., 2013. SHOWBIZ: a multi-physics 3D code to simulate a fuel rod cladding embrittlement in normal reactor operation, transport and storage and during LOCA and RIA transients. In: The 6th Annual Modelling in Nuclear Science and Engineering Seminar, 9-10 November 2023, London, UK.

Lavoil, A., 2010. LOCA testing at HALDEN, the tenth experiment IFA-650.10", OECD HALDEN reactor project, report HWR-974.

Montgomery, R., Morris, R.N., 2019. Measurement and modelling of the gas permeability of high burnup pressurized water reactor fuel rods. J. Nucl. Mater. https://doi.org/10.1016/j.jnucmat.2019.05.041.

Nishi, N., Lee, B.-H., 2001. Summary of pre-irradiation data on fuel segments supplied by EDF/FRAMATOME and tested in IFA-610, 629 and 648. OECD HALDEN reactor project, report HWR-664.

Oberländer, B.C., 2016. PIE on the LOCA test of high burnup PWR fuel in IFA-650.10. OECD HALDEN reactor project, report HWR-1019.

Salvo, M., 2024. FUEL+: an IRSN software platform dedicated to nuclear fuel behaviour. TOPFUEL 2024. 29 September - 3 October 2024 in Grenoble, France.

Taurines, T., Belon, S., Glantz, T., Jacq, F., 2021. New burst criteria, axial fuel relocation and fission gas release options in the DRACCAR code. TOPFUEL2021, TOPFUEL, Oct 2021, SANTANDER, Spain.

Taurines, T. et al. 2021. Rod cladding failure during LOCA- Progress report on experimental database reassessment and model/code improvements, WP 3.2 Evaluation of the failed rod number final report, D3.3 R2CA EU project report, 2021, Reduction of Radiological Consequences of design basis and design extension Accidents European Project.

Wiesenack, W., 2013. Summary of the Halden Reactor Project LOCA Test Series IFA-650, HPR-380.