Lucas Terrei

Zoubir Acem

Véronique Marchetti

Paul Lardet

Pascal Boulet

Gilles Parent

In-depth wood temperature measurement using embedded thin wire thermocouples in cone calorimeter tests

Keywords: Temperature measurement, cone calorimeter, mass loss rate, wood degradation, wood charring

In-depth temperature measurements of wood samples using thermocouples are commonly performed in re research at all scales. However, the size or the type of thermocouples or how they are placed in the sample can lead to the measured temperature being signicantly dierent. In this work, an original method combining thin wire thermocouples with a meticulous implantation, was used to provide an accurate measurement of the in-depth temperature of wood samples during their degradation. Wood samples were thus equipped with both embedded thin wire thermocouples (0.1 mm diameter) and sheathed thermocouples (1 mm diameter) installed perpendicularly or in parallel to the conductive heat ux to compare the results between the two kinds of measurements. Afterwards, samples were exposed to the cone calorimeter at dierent heat uxes (from 16.5 to 93.5 kW.m -2 ) in order to measure both in-depth temperature and mass loss during the degradation. The smallness and high exibility of thin embedded wire thermocouples enabled in-depth temperature measurement without aecting the mass loss measurement. This is usually not possible when using sheathed thermocouples, because of their high stiness. Results showed that it is crucial to rigorously control the thermocouple implantation, as an underestimation up to 400 • C was observed for thermocouples put near the surface exposed to the heat ux. There are several reasons for this dierence: the heat sink due to conductive heat ux through the thermocouple body, a poor contact between the sensor and the sample and uncertainty about the actual position of the thermocouple. The char front evolution deduced from temperature measurement by wire thermocouples was found to comply with direct measurements of the char layer performed on samples. Finally, the charring rate obtained from in-depth temperature as a function of the heat ux was found to follow an ane law.

INTRODUCTION

Wood materials such as cross-laminated timber (CLT), glued laminated timber (glulam) or laminated bamboo, to cite just a few, are increasingly popular for use as structural materials in buildings because of their low carbon footprint. These materials also have good mechanical properties and are convenient to machine, shape and implement in constructions. For all these reasons, wood-based materials are now increasingly used in building facades and structures. However, this combustible material degrades in a re situation. Consequently, a good understanding of wood Preprint submitted to Int. j. therm. sci. re resistance is important to correctly size wood structures in buildings, for reasons of re safety and prevention. This sizing mainly requires knowledge of the extent to which wood materials can withstand re and how they degrade in re situations. Previous studies on the wood degradation process showed a complex coupling between chemical, thermal and mechanical reactions [START_REF] Bartlett | A review of factors aecting the burning behaviour of wood for application to tall timber construction[END_REF][START_REF] Roberts | Problems associated with the theoretical analysis of the burning of wood[END_REF]. Thus, reliable in-depth temperature data are important for the development of models whose aim is to predict wood behavior. Among other consequences, the pyrolysis of wood results in the formation of a char layer whose mechanical properties are much poorer than those of wood. On the other hand, the char layer acts as a protective layer which slows down the in-depth degradation of wood.

At a small scale, several experimental studies aiming at determining the charring rate of wood samples have been carried out in the past. Xu et al. [START_REF] Xu | Combustion and charring properties of ve common constructional wood species from cone calorimeter tests[END_REF] exposed ve wood species horizontally to the cone calorimeter at dierent heat uxes and exposure times. After cutting each sample, Xu et al. measured the char layer and deduced the charring rate for dierent wood species. This work showed that the increase of the charring rate according to the heat ux is related to the wood species. Other studies have shown empirical equations describing the charring rate as a function of the heat ux [START_REF] Maciulaitis | Singularity and importance of determination of wood charring rate in re investigation[END_REF]. The recent state of the art report on the factors which aect the burning behavior of wood by Bartlett et al. [START_REF] Bartlett | A review of factors aecting the burning behaviour of wood for application to tall timber construction[END_REF] showed a charring rate varying from 0.4 mm. min -1 for hardwood to 0.8 mm. min -1 for softwood. The value generally used in re science engineering is 0.65 mm. min -1 [5, 6]. The wood density, species or heating method (a heat ux increasing over time, constant heat ux, furnace) are parameters which aect the charring rate [START_REF] Lizhong | The inuence of dierent external heating ways on pyrolysis and spontaneous ignition of some woods[END_REF][START_REF] Frangi | Charring rates and temperature proles of wood sections[END_REF]. Several empirical models are used to determine the char rate like the nonlinear char rate model used in engineering [START_REF] White | Fire safety, Wood handbook: wood as an engineering material[END_REF]. The charring front has also been regularly deduced from the 300

• C isotherm measured by thermocouples in dierent types of tests [START_REF] Bartlett | A review of factors aecting the burning behaviour of wood for application to tall timber construction[END_REF][START_REF] Lizhong | Predicting charring rate of woods exposed to time-increasing and constant heat uxes[END_REF][START_REF] Njankouo | Experimental study of the charring rate of tropical hardwoods[END_REF]. Measurement of in-depth temperature has commonly been performed in cone calorimeter tests by using thermocouples [START_REF] Tsai | Orientation eect on cone calorimeter test results to assess re hazard of materials[END_REF][START_REF] Wasan | Study of pyrolysis and upward ame spread on charring materialspart i: Experimental study[END_REF]. Some papers dating back to the 1960's recommended putting thermocouples perpendicular to the exposed heat ux (or parallel to isotherms) [START_REF] Beck | Thermocouple temperature disturbances in low conductivity materials[END_REF][START_REF] Brewer | Eect of thermocouple wire size and conguration on internal temperature measurements in a charring ablator[END_REF]. The present paper is devoted to wood, but most of the observations made would apply when dealing with other low thermal conductivity materials like concrete or polymers. For practical reasons, sheathed thermocouples are generally placed into wood samples by drilling boreholes parallel to the heat ux, i.e. perpendicular to the exposed wood surface. However, it is dicult to make long holes with a small diameter in the wood. Therefore, drilling is generally carried out using a drill bit with a diameter in the millimeter range and the thermocouples put into the material are also millimeter-sized. Also, when using sheathed thermocouples, mass loss cannot be measured accurately because the thermocouple stiness aects the mass measurement. Thus, tests must be doubled in order to get the in-depth temperature and the mass loss rate, but not simultaneously [START_REF] Reszka | In-depth temperature measurements in wood exposed to intense radiant energy[END_REF]. Moreover, this method can lead to mismeasurements due to the thermal sink along the thermocouple, the inertia of the sheath, the poor contact between the sensor and the sample or even because of the inaccurate location of sensors. Fahrni et al. conducted two interesting studies related to in-depth temperature measurement in wood samples [START_REF] Fahrni | Correct temperature measurements in re exposed wood[END_REF][START_REF] Fahrni | Investigation of dierent temperature measurement designs and installations in timber members as low conductive material[END_REF]. In these works, up to nine dierent combinations of types of thermocouple/ways of setting-up were studied -sheathed or wire thermocouple, hot junction nature (welded or twisted), thermocouple inserted in boreholes or inlaid in the sample, diameter of the thermocouple wires, maximum diameter of the thermocouple (sheath diameter or diameter of the wires including electrical insulation), diameter of the borehole (if applicable) and orientation of the thermocouple. A comprehensive designation to describe all these parameters was proposed.

In reference. [START_REF] Fahrni | Correct temperature measurements in re exposed wood[END_REF], tests were performed both inside a furnace following the ISO 834 re exposition and by submitting the sample to a constant heat ux equal to 55 kW.m -2 . In reference. [START_REF] Fahrni | Investigation of dierent temperature measurement designs and installations in timber members as low conductive material[END_REF], only the furnace was used. The main nding of these two works is that measurements can be very dierent according to the design used and that discrepancies reaching several hundred degrees were observed. The designs in which the thermocouple body (wires or sheath) is oriented parallel to the isotherms were always found to give the highest temperature values which are undoubtedly the most reliable. The dierence was mainly attributed to heat transfer along the thermocouple body. The secondary cause leading to the under-measurement of temperature was attributed to the thermal contact between the thermocouple and the sample. Convection inside the borehole, when this was signicantly larger than the thermocouple, was found to be negligible. Cooling of the thermocouple surrounding was also observed.In the present work, spruce wood samples were exposed to dierent heat uxes (16.5, 38.5, 60 and 93.5 kW.m -2 ) with a cone calorimeter oriented vertically, in order to measure the in-depth temperature evolution. This measurement was performed with two types of thermocouple implantations. In the rst setup non-insulated very thin wire thermocouples (0.1 mm diameter) were embedded in the sample after specic machining aimed at controlling the position of the measurement. Such small thermocouples have the advantage of being minimally intrusive compared to the thermocouples that are usually used which are commonly millimeter-sized. Following the designation proposed by Fahrni et al., this design would be named K-w-e-0.1/0.1/in-pa. Compared to the parallel design studied by Fahrni et al., the present one involved thinner wire diameter (0.1 mm vs 0.2 mm) and thinner maximum thermocouple diameter (0.1 mm vs 1.2 mm, because of the lack of electrical insulation). The accuracy provided by our design is thus expected to be even better, and less intrusive. In the second method, 1.2 mm diameter holes were drilled from the sample back side and 1 mm diameter sheathed thermocouples were inserted inside with the depth position adjusted directly from the hole depth. The designation of this second implantation is thus K-s-i-0.2/1/1.2-pe. The temperature evolutions obtained with the two types of thermocouple implantation are compared each other. Afterwards, the measured in-depth temperatures from both techniques were used in order to track the char front position from the 300

• C isotherm. Results were compared with direct measurements from samples cut in their middle after being exposed to the same heat uxes for dierent durations.

Finally, a study of the charring rate, calculated from the char front positions, was also performed.

MATERIALS AND METHODS

Experimental setup

Tests were performed with a cone calorimeter oriented vertically. Figure 1 The experiment consists of exposing a sample to a radiative heat ux coming from an electrically heated conical coil. The radiative heat ux emitted by the cone was measured before each test using a Schmidt-Boelter uxmeter (Medtherm). The heat ux was considered correct when its value was ± 0.5 kW.m -2 from the desired ux. The distance between the sample and the heater was 25 mm. The spark igniter was not placed between the cone and the sample so that the wood ignition was not explicitly piloted. The selected samples were of spruce wood which were 100×100 mm 2 in size with a 50 mm thickness and the heat ux was perpendicular to the wood grain.

The sample average density was about 490 kg.m -3 for an average moisture content around 9.5 %.

Wood samples were exposed to dierent constant incident heat uxes equal to 16.5, 38.5, 60 and 93.5 kW.m -2 . This large range of heat uxes made it possible to study the in-depth temperature time evolution either with sample aming auto-ignition (for high enough heat uxes) or without ignition. The experiment lasted for between 15 and 60 minutes according to the imposed heat ux.

Setting up thermocouples in the sample

In-depth temperature measurements by thermocouples are commonly performed in cone calorimeter tests. As aforementioned, the type of thermocouples used and the way they are implemented can greatly aect the measured in-depth temperature. In this work, a method was developed aiming at embedding very thin (0.1 mm diameter) wire thermocouples inside the sample. Eleven K-type thermocouples were placed inside the wood sample. Figure 2 shows the meticulous ma-chining steps used for embedding wire thermocouples in the sample. There are four main steps to the proposed protocol:

• Step 1: The sample was cut in the middle, perpendicularly to the exposed surface.

• Step 2: 0.2 mm depth and width square-grooves were precisely machined (with a 0.2 mm thickness circular milling cutter) on the surface of the wood at the desired locations (every 2 mm from 2 to 20 mm depth from the exposed surface, one at a 30 mm depth and a last one at a 40 mm depth).

• Step 3: After machining, wire thermocouples were put in the square-grooves. In this picture red dots represent the position of the thermocouple junction which was homemade welded where the temperature is measured. The alumel and chromel cables were welded by autogenous electric welding (without adding metal) end to end.

• Step 4: Samples were nally glued with Melamine-urea-formaldehyde (MUF) and pressed for 6 hours.

Setting up wire thermocouples in this way takes three hours per sample from grooving to bonding and including thermocouple welding. This process can be carried out on a larger scale, but it becomes more dicult the more the sample size increases. For example, it was also used by the authors for Single Burning Item (SBI) tests with samples as large as 150 × 50 cm. It appears that MLR evolutions are very similar. Tests with embedded wire thermocouples can be thus considered to be carried out in similar conditions as without. Compared to existing ways of proceeding, the thermocouple implantation proposed in the present paper has numerous advantages for achieving more accurate and reliable in-depth temperature measurements namely:

-The thermocouple wires are very thin which minimizes the conductive heat ux along them and the intrusiveness of the wires in the sample (very tiny matter removal, very low quantity of glue).

-The wires are very supple which means that mass loss measurements can be performed at the same time as temperature measurements.

-The chromel and alumel wires extend from each other in one straight unit and therefore no electrical insulation is required. These two points mean that thinner grooves can be used to put them inside the wood -The autogenous electric welding is very small (no additional matter).

-The position of the thermocouple junction is very accurate.

-The thermal resistance between the thermocouple and the sample is negligible.

-The wires are parallel to the isotherms thus minimizing the conductive ux as will be shown later and this is undoubtedly the main positive point of this method.

RESULTS AND DISCUSSION

Results and discussion are dissociated in three parts. The rst part presents in-depth temperatures measured by embedded wire thermocouples at dierent heat uxes. The second section is devoted to a comparison with the other kind of thermocouple implantation. A numerical simulation was performed showing that the conductive heat ux through the thermocouple can actually lead to very large mismeasurements. Finally, the 300

• C isotherm deducted from temperature measurements are compared with the directly measured charring rate.

3.1. In-depth temperature measured by thin embedded wire thermocouples

Thermograms

The wood samples equipped with embedded wire thermocouples were exposed to the cone calorimeter at several heat uxes from 16.5 to 93.5 kW.m -2 . Results obtained for tests performed at 16.5, 38.5, 60 and 93.5 kW.m -2 are presented in gures 4 to 7. In gures 4 and 6, surface temperature is added to in-depth themogrammes. This temperature was measured with an infrared camera equipped with a lter at 3.9 µm wavelength (outside emission bands of combustion gases).

At the cone calorimeter scale and especially with a vertical orientation, the optical thickness of the ame is small enough to consider that there is a low level of emission by soot and consequently to assume that the ame is almost transparent outside gas emission bands [START_REF] Acem | Surface temperature of carbon composite samples during thermal degradation[END_REF]. The surface temperature measurements are described with more details in [START_REF] Terrei | Experimental tools applied to ignition study of spruce wood under cone calorimeter[END_REF]. As a general observation, the consecutive in-depth temperature curves are ordered according to the thermocouple depth and their spacing is rather regular. The surface temperature curve is higher than in-depth temperature curves and its variation is consistent with in-depth curves. In particular there is no odd gap between surface temperature and 2 mm in-depth temperature. The prole temperature curves presented in gure 7 were also found to have very likely behavior with no spurious steps or slope changes which made us very condent about the quality of the performed measurements. As the exposure time increases, some measured in-depth temperatures become noisy and somewhat erratic, sometimes exceeding the surface temperature. This happens when the thermocouple is no longer embedded inside the sample, due to the char combustion and thus the thermocouple is directly submitted to the ux from the heater and to the ame turbulent hot gases. For these two low heat uxes auto-ignition did not occur. For the test performed at 16.5 kW.m -2 , surface temperature rst increases fast and asymptotically tends toward a rst plateau at around 400 • C or little below. However, at a time equal to 20 minutes, the temperature rises fast again before reaching a new equilibrium temperature near 580

• C. This change in behavior is due to the smouldering combustion of char beginning which brings an additional heat source. This transition was already noticed in papers devoted to the study of the surface temperature of wood samples in cone calorimeter tests [START_REF] Terrei | Experimental tools applied to ignition study of spruce wood under cone calorimeter[END_REF][START_REF] Boonmee | Glowing ignition of wood: the onset of surface combustion[END_REF]. Figure 5 shows infrared pictures taken at dierent times during a test performed at a 16.5 kW.m -2 heat ux. These pictures show the smouldering combustion onset at 20 minutes (picture 2). The beginning of the smouldering combustion results in glowing hot zones (pink color areas in the snapshot) which then spread over the entire surface.

The inuence of char combustion can also be seen in depth particularly for depths less than or equal to 6 mm. The beginning of the second temperature rise is delayed compared to the surface temperature, and, logically, the deeper the position, the larger this delay. The new equilibrium temperature is also lower when the depth increases. For depths greater than 6 mm, the eect of char combustion becomes less and less remarkable, since thermocouples are either far from the regressive smouldering surface or even from the charring front. For the test performed at 38.5 kW.m -2 , there is no ame ignition either. Smouldering combustion occurred too but at a shorter time for which temperature increases still fast, so its inuence is not so obvious. Surface temperature reaches a plateau fast and, the higher the heat ux, the higher this value is logically. As a consequence of the low thermal diusion of the sample, the larger the depth, the lower the plateau value and the later the plateau occurs. The higher the heat ux, the faster the temperature rises in the wood. For instance, the temperature at 18 mm depth reaches 300 • C in 18 min 30 sec for a 60 kW.m -2 incident heat ux, whereas it takes only 14 minutes for a 93.5

kW.m -2 incident heat ux. In the case of thermocouples close to the surface exposed to the heat ux, the temperature increases rapidly followed by a plateau. The temperature of thermocouples close to the exposed surface (from 2 to 6 mm depth) increases fast at the beginning of the test.

At 2 mm depth, the temperature reaches 600 • C in 107 and 88 seconds for 60 and 93.5 kW.m -2 respectively. For large depths (10 mm and above), there is a delay before the temperature begins to rise. Then, the temperature increase slows down. When the temperature reaches around 150

• C the slope increases again. samples degrade and crack, the rst wire thermocouples were no longer embedded in the wood and were directly exposed to cone calorimeter. When this occurred, the measured temperature becomes unstable and the values can no longer be considered reliable.

Inuence of the thermocouple setting-up on in-depth temperature measurements

Some tests were also performed with 1 mm-diameter thermocouples placed perpendicularly to isotherms after drilling the back of the sample. This kind of implantation is typically employed in most re science experiments since it is quite convenient to setup. One millimeter diameter thermocouples are the standard thermocouple size. Smaller diameter sheathed thermocouple are available but they are more fragile and larger diameter holes would need to be drilled in any case because smaller drill bits are not strong enough to be used. The aim of this part of the study was to assess discrepancies which could arise in measurements carried out with the usual thermocouple implementation. In the following wire thermocouples will denote the thin wire thermocouples embedded inside the sample parallel to the surface exposed to the heat ux, as presented in this paper. Sheathed thermocouple will denote the 1 mm diameter sheathed thermocouples put inside the sample perpendicular to the surface exposed to the heat ux. The dierences between the two types of measurements can be attributed at rst sight to both the dierence in diameter and the dierence in orientation. A COMSOL simulation will allow the respective importance of these two parameters to be assessed.

Experimental study

Four sheathed K-type thermocouples (1 mm diameter) were placed respectively at 2, 4, 6 and 8 mm depth of the exposed surface. The position of the sheathed thermocouples is shown in gure 1. These thermocouples were placed perpendicularly to the surface exposed to the heat ux after drilling the back of the sample. As suggested by Reszka et al. [START_REF] Reszka | In-depth temperature measurements in wood exposed to intense radiant energy[END_REF], sheathed thermocouples were spaced by 20 mm in order to limit mutual disturbance Figure 8 It should be noted that the temperature evolutions for the two types of thermocouple are very dierent. For tests performed at 16.5 kW.m -2 , discrepancies reaching 100 • C or more can be observed between the temperature measured by the two kinds of thermocouples. The beginning of smouldering combustion is observed for both techniques but it is not so clear with the sheathed thermocouples. For tests carried out at 38.5 kW.m -2 , the temperature measured by the sheathed thermocouples are still largely underestimated and the dierence can reach more than 200

• C. The dierence between the two kinds of measurements can mainly be explained by the large heat sink through the thermocouple sheath. Indeed, the sheathed thermocouple has a high conductivity and large diameter. This means there is a large conductive heat ux along the thermocouple whereas the heat brought to the thermocouple apex is weak due to the low conductivity of wood (or of char) and the poor thermal contact between the thermocouple and the sample. Figure 9 presents results obtained at higher uxes equal to 60 and 93.5 kW.m -2 . Three tests were carried out at 60 kW.m -2 and gure 9a shows the mean temperature obtained from these three tests with error bars corresponding to the standard deviation between the tests. Once again, the temperature evolutions for the tests performed at 60 and 93.5 kW.m -2 are underestimated with sheathed thermocouples for the same reasons as for lower heat uxes. Temperatures obtained from wire thermocouples increase fast before reaching a plateau whereas temperature evolutions provided by sheathed thermocouples are more linear and gradual. Temperature dierences between the measurements made by the two thermocouple setups can reach several hundreds of degrees which is considerable. For the two heat uxes, the temperature discrepancy reaches up to 400

• C at the beginning of the test, during the fast increase of temperature. After 9 minutes of exposure, the deviation is about 200

• C, which is still very high. Moreover, error bars for curves obtained by sheathed thermocouples are larger than those corresponding to wire thermocouple measurements. This is evidence of the inferior level of repeatability of measurements performed by sheathed thermocouples. This can be explained by larger sources of inaccuracy, including: exact position of the thermocouple (drilling process), thermal resistance between the thermocouple and the wood, the hole having to be larger than the thermocouple, etc. As a consequence, a large underestimation of the temperature is observed (which will be illustrated in next section with a COMSOL multiphysics R simulation). Furthermore, the position of sheathed thermocouples can be inaccurate particularly if these are perpendicular to the heat ux for the purposes of the study. Indeed, the long and thin drill bits (typically 1.2 mm diameter and several dozens of millimeters of drilling depth) get bent easily when encountering timber heterogeneities (e.g. knots or growth rings) during the drilling process. Embedded wire thermocouples thus appear preferable to achieve accurate in-depth temperature measurements. Moreover, thanks to their high exibility, such ne wire thermocouples allow MLR measurement during tests, which is not possible with the usual large diameter sheathed thermocouples.

Simulation of the heat sink created by the thermocouple implantation

Simple simulations with COMSOL multiphysics R software were performed in order to illustrate the eect of a thermocouple implantation in a wood sample. The aim was to demonstrate that the perturbation created by inserting a material having a high conductivity in a low conductivity material can be very signicant rather than to actually present a predictive model providing an accurate temperature evolution. For that purpose, the sample was modelled to the cone test dimensions (i.e. 100 × 100 × 50 mm) and a thermocouple was placed at 2 mm away from the exposed surface.

The sample mesh for the two thermocouple orientations is shown in Fig. A.15 in appendix. The mesh has 24,000 nodes for the conguration where the thermocouple is perpendicular to the surface exposed to the ow and 20,000 nodes when the thermocouple is oriented parallel to the surface.

The sample was considered inert, which means that the heat transfer inside it is purely diusive.

Temperature at a large time interval (10 minutes) was considered so that the char depth is higher than 2 mm where the thermocouple junction is placed. The sample density and thermal properties were then chosen to be equal to the char properties -135 kg.m -3 for the density, 1650 J.kg -1 .K -1 for the specic heat and 0.073 W.m -1 .K -1 for the conductivity. The thermocouple was modelled as a homogeneous cylinder of inconel having a density, a specic heat and a thermal conductivity equal to 8250 kg.m -3 , 444 J.kg -1 .K -1 and 14.8 W.m -1 .K -1 respectively. Two orientations were considered namely parallel or perpendicular to the heat ux. For the parallel conguration, the thermocouple was put at the center of the back of the sample and its apex set at a distance equal to 2 mm from the front surface. Two diameters (1 mm and 0.1 mm ) were considered. For the perpendicular conguration, the thermocouple was set perpendicular to one of the side faces, at the middle of its width and 2 mm from the exposed surface. A perfect thermal contact between the thermocouple and the wood was assumed. A convective exchange coecient was considered on all non-exposed surfaces equal to 10 W.m -2 .K -1 and 15 W.m -2 .K -1 on the exposed surface to the heat ux [START_REF] Staggs | Convection heat transfer in the cone calorimeter[END_REF][START_REF] Quintiere | A theoretical basis for ammability properties[END_REF]. An imposed radiative heat ux of 93.5 kW.m -2 was applied to the exposed surface. The temperature prole within the sample width 2 mm depth is presented in gure 10 for each conguration at time t= 600 s. At a same depth, a large drop can be observed between the temperature at the thermocouple location and the temperature far from the thermocouple. This drop reaches 280

• C which is in the same order of magnitude to what is observed experimentally. This large dierence is due to the conductive heat ux along the thermocouple sheath. Moreover, the temperature drop extends for several millimeter inside the sample itself, conrming the cooling of the thermocouple surrounding noticed by Fahrni et al. [START_REF] Fahrni | Correct temperature measurements in re exposed wood[END_REF] for instance. If we consider a 0.1 mm diameter thermocouple parallel to the conductive heat ux, the temperature underestimation is now only 60

• C because the conductive heat ux along the thermocouple is considerably less due to its smaller section. For the thermocouple implemented perpendicular to the heat ux, the underestimation is equal to 8

• C for a 1 mm diameter thermocouple and only 0.3 • C for a 0.1 mm diameter thermocouple (this case corresponding to the thin embedded wire thermocouple used in the present work). Although this simulation was not intended to be quantitative because a lot of phenomena were not taken into account (pyrolysis, change of thermal properties of materials, thermal contact resistance and so on), it does demonstrates that very large temperature dierences can be observed depending on the thermocouple size and the way it is implanted inside the sample. The thermocouple orientation is found to be the most inuencial parameter because the conductive heat ux along the thermocouple wire (or sheath) is proportional to the temperature gradient. Even with a large diameter thermocouple, the temperature underestimation is only 8

• C when it is oriented perpendicular to the heat ux. For a given orientation, the larger the thermocouple diameter, the larger the conductive heat ux and thus the temperature underestimation. Almost negligible temperature dierence is obtained using a very thin thermocouple oriented perpendicular to the heat ux.

Adding a thermal resistance between the sample and the thermocouple would result in an even higher underestimation of the temperature.

Comparison between 300 • C isotherm and char front

It is often reported that the progress of the char front within the sample follows the 300

• C isotherm [START_REF] Tran | Burning rate of solid wood measured in a heat release rate calorimeter[END_REF]. Thus, the new thermocouple implantation was used to study the char layer progression in the sample. To validate this approach, the position of the char front was also determined directly -wood samples were submitted to a given heat ux (60 and 93.5 kW.m -2 ) for dierent durations and then cooled fast with liquid nitrogen to stop pyrolysis reactions. These samples were cut in their middle and the position of the char front was measured with a rule. In Figure 11: Char front positions measured with a rule and deduced from the 300 • C isotherm provided by sheathed and wire thermocouples. For measurement marks having error bars, three measurements were performed and error bars correspond to standard variations. For single marks, only one measurement was performed.

Figure 11 shows a good level of compliance between direct measurements of the char front position and the position obtained from the 300

• C isotherm measured with the embedded wire thermocouples. This means we can have even more condence about the accuracy of the measurements that were made in this way. By contrast, when considering the 300

• C isotherm measured
with sheathed thermocouple, a large discrepancy can be noticed. For instance, for a test performed at 60 kW.m -2 , the char front is located at a 6 mm depth after 180 seconds according to embedded wire thermocouple measurements while 620 seconds are needed when considering sheathed thermocouple data. A power t was performed on wire thermocouple measurements. It can be seen that the power law ts pretty well with the data and this kind of t makes it possible to calculate an analytical charring rate. Figure 12 presents the charring rate deduced from the char front position.

The charring rate for wire thermocouples was deduced with centered nite dierences for values calculated from measured data (green squares) and with analytical derivative of the power law (dot-dashed curve). Figure 12: Charring rates deduced from the char front measured with a rule and deduced from with the 300 • C isotherm of wire thermocouples.

The charring rate evolutions deduced from rule measurements and wire thermocouples isotherm comply well with the two heat uxes. Using the position of the 300

• C isotherm, measured by the embedded thermocouples, makes it possible to obtain the char front position and the charring rate from a single test, which is not possible using the direct method which provides only discrete values with each value coming from a dierent test. The charring rate was found to reach a peak value around 2.5 mm. min -1 in a transient part and to decrease to a quasi-steady state around 0.8 mm. min -1 . In the following, the variation of the charring rate with the submitted heat ux was also investigated.

Time (min)

Charring front (mm) The char front location as a function of time for dierent uxes is presented in Fig. 13 which also shows the power law ts performed on measurements data. To take into account the delay before the char process begins when the heat ux is small, a power law including a time shift t 0 was chosen according to X(t) = a(t -t 0 ) b , X(t) corresponding to the char front position at time t. For the test performed at 16.5 kW.m -2 , two power laws were used -the rst for the beginning of the test, before the onset of smouldering combustion and a second after this.

The charring rate was calculated from the analytical derivative of the ts. Given the time shift, comparing the charring rate at a same time for the dierent heat ux would have no meaning. The charring rate was thus calculated for a xed position reached by the char front. It is illustrated in Fig. 13 for a position equal to 10 mm (horizontal dashed line). The charring rate at this depth corresponds to the slope of the dierent curves as represented in the gure. Results are presented in Fig. 14a for depths equal to 5, 10, 15 and 20 mm. As expected, the higher the ux, the higher the charring rate, and the deeper the position, the lower the charring rate. For a given heat ux, as the depth increases, the variation of the charring rate tends to be increasingly lower which is in accordance with the decrease with time of the slope of the rating rate in Fig. 12 (the charring rate tends toward a quasi steady state). The charring rate varies in a linear manner with the incident radiative heat ux, which conrms the ndings of Tran and White works [START_REF] Tran | Burning rate of solid wood measured in a heat release rate calorimeter[END_REF][START_REF] White | Charring rate of wood exposed to a constant heat ux[END_REF]. As proposed by Tran and White [START_REF] Tran | Burning rate of solid wood measured in a heat release rate calorimeter[END_REF], an average charring rate can be calculated as the depth reached by the char divided by the time needed to reach this depth. Such a calculation was performed with a 15 mm depth. For low heat ux, the time taken into account for this calculation is the time minus the time delay t 0 needed for the charring process to begin. Results are presented in Fig. 14b. A quasi-linear dependence is observed for the average charring calculated in this way.

CONCLUSION

In this paper a new technique involving the meticulous implantation of thin wire thermocouples embedded inside wood samples was developed in order to measure the in-depth temperature of wood during its degradation. This was carried out using a specic process, involving cutting the sample, grooving it, setting up the wire thermocouples inside the square-grooves, and reassembling the sample using MUF glue. Numerous thermocouples can be put inside the sample without signicantly modifying either the sample thermic or the degradation process. This technique meant the mass loss and the temperature evolution during tests could be measured simultaneously. After setting up the experiment, wood samples were exposed to dierent heat uxes from 16.5 to 93.5 kW.m -2 . For the lowest heat ux, the thermocouple set up permitted us to observe the beginning of smouldering combustion which is particularly visible for measurements made close to the exposed surface. Tests were also performed with sheathed thermocouples placed at 2, 4, 6 and 8 mm indepth perpendicularly to the isotherm. The results show that signicant temperature discrepancies are obtained according to the kind of thermocouple implementation studied. At the same depth and time, the dierence between the temperatures measured either by thermocouples perpendicular to the exposed surface or by embedded wire thermocouples parallel to the surface can reach up to 400

• C. Numerical results show that the main reason for this deviation can be attributed to the large heat sink (conductive heat ux) along the thermocouple wire or sheath. This heat sink is found to be greater when the thermocouple or the sheath diameter is large and the thermocouple is oriented parallel to the conductive heat ux in the sample (i.e. perpendicular to isotherms).

The deviation may also be explained by the thermal contact resistance between the sample and the thermocouple and by a possible error in the actual position of the thermocouple. The comparison between the char front and the 300

• C isotherm validates the temperature measurement with wire thermocouples. The charring rate derived from the char front increases in a linear manner with the imposed heat ux from 0.5 mm. min -1 for 16.5 kW.m -2 to 1 mm. min -1 for 93.5 kW.m -2 .

In-depth temperatures obtained with the method proposed here are more accurate and reliable than the measurements usually made. This machining made it possible to put thermocouples in the wood sample with the least possible intrusion for temperature measurements. Such reliable data are of primary importance as input for comprehensive models aiming at describing/predicting the thermal degradation of wood.

When possible, putting thermocouples perpendicular to the temperature gradient is by far the preferable solution, and the smaller the thermocouple diameter, the lower the temperature error. Depending on the nature of the samples, this can be done during the manufacture of the samples or otherwise by cutting and gluing. Alternatively, thermocouples can be placed after drilling but perpendicularly to the heat ux. However, this method can lead errors on the position of the thermocouples and cannot be performed for very large panels (long thin drill bits are not available). If putting the thermocouples perpendicular to the heat gradient cannot be done, the use of a thermocouple with a minimum diameter is recommended and the hole should be lled if its diameter is larger than that of the thermocouple.

  presents the experimental setup.

  (a) Sample equipped with sheathed thermocouples. (b) Sample equipped with wire thermocouples.

Figure 1 :

 1 Figure 1: Experimental setup.

Figure 2 :Figure 3 :

 23 Figure 2: Wire thermocouple implantation in wood sample: main steps of the setting-up.

  16.5 kW.m -2 . (b) 38.5 kW.m -2 .

Figure 4 :

 4 Figure 4: In-depth temperature evolutions measured by embedded wire thermocouples located from 2 to 40 mm depth to the exposed surface.

  The maximum temperature reached at 40 mm depth is around 100 • C for 16.5 kW.m -2 and 250 • C for 38.5 kW.m -2 after one hour of exposure.

Figure 5 :

 5 Figure 5: Surface temperature eld in Kelvin and studied area (white circle) at dierent exposure time for 16.5 kW.m -2 .

Figure 6

 6 Figure 6 presents typical in-depth temperature evolutions for tests at 60 and 93.5 kW.m -2 , for which wood auto-ignition occurred. For the heat ux equal to 60 kW.m -2 three tests were performed and errors bars correspond to standard variations.

Figure 6 :

 6 Figure 6: In-depth temperature evolutions measured by embedded wire thermocouples from 2 to 40 mm depth to the exposed surface.

3. 1 . 2 .

 12 Temperature prolesIn-depth temperature measurement provided by embedded thermocouples were used to draw temperature proles within the sample. Results for dierent heat uxes and at dierent times are presented in gure 7.[START_REF] Reszka | In-depth temperature measurements in wood exposed to intense radiant energy[END_REF].5 and 38.5 kW.m -2 at 600, 1200 and 1800 sec. 60 and 93.5 kW.m -2 at 60, 300 and 600 sec.

Figure 7 :

 7 Figure 7: Temperature proles at dierent times and dierent heat uxes.

  presents results obtained for heat uxes equal to 16.5 kW.m -2 and 38.5 kW.m -2 . 38.5 kW.m -2 .

Figure 8 :

 8 Figure 8: Mean in-depth temperature evolution for tests performed with sheathed (S) and wire (W) thermocouples.

  93.5 kW.m -2 .

Figure 9 :

 9 Figure 9: Mean in-depth temperature evolution for tests performed with sheathed (S) and wire (W) thermocouples.

  (a) Parallel to the heat ux. (b) Perpendicular to the heat ux.

Figure 10 :

 10 Figure 10: Temperature proles in the sample along a line passing through the thermocouple apex (2 mm depth) -Two diameters (0.1 and 1 mm) and two orientations relative to the heat ux.

gure 11 ,

 11 this char front position is compared to the one obtained indirectly from the position of the 300 • C isotherm. The error bars for the char front deduced from thermocouples (wire and sheathed) correspond to the standard variations obtained from three tests for each heat ux. 93.5 kW.m -2 .

  93.5 kW.m -2 .

Figure 13 :

 13 Figure 13: Charring front as a function of time for dierent heat uxes. Marks: experimental data Dot-dashed curves: power law ts Plain lines: curve tangent at position X = 10mm.

  Charring rate as a function of heat ux at dierent depths. Average (calculated from 0 to 15 mm depth) charring rate as a function of heat ux.

Figure 14 :

 14 Figure 14: Charring rates for dierent heat uxes.

ACKNOWLEDGEMENTS

The authors thank the LEMTA's technical sta, Jean-Yves Morel, Sébastien Lejeune, Jérémy Bianchin, Mathieu Weber and Jamal Ouhajjou, for their contribution during samples machining and instrumentation.


Benjamin Richard
email: benjamin.richard@irsn.fr

Marie-Cécile Robin-Boudaoud

Emmanuel Viallet

Sefano Cherubini

Julien Berger

Ibrahim Bitar

Julien Clement

Alexis Courtois

Olivier Fabre

Reine Fares

Loic Le Gratiet

Fabien Grange

Thomas Heitz

Nicolas Ile

Maria Lancieri

Thomas Langlade

Natacha Suin

François Voldoire

OVERVIEW OF THE RESEARCH ACTIVITIES IN EARTHQUAKE ENGINEERING AND SEISMIC RISK ASSESSMENT WITHIN THE JOINT FRAMEWORK CEA-EDF-FRAMATOME-IRSN

The French nuclear community is highly active in the field of earthquake engineering and more precisely, regarding seismic risk assessment. Among the existing safety-oriented frameworks, there is one which involves not only nuclear operators such as Electricité de France (EDF), the French Sustainable Energies and Atomic Energy Commission (CEA) and FRAMATOME, but also the French technical support organization which is the Institute for Radiological Protection and Nuclear Safety (IRSN). Within this framework, research activities are conducted, aiming to enhance knowledge in order to improve nuclear safety of either existing or new facilities. These activities deal with topics and challenging issues that appear to be significant in the seismic risk assessment process. More precisely, they aim to better assess the seismic margins when available, to better understand the beyond-design behavior of nuclear buildings or related equipment and to improve the necessary knowledge to produce robust seismic hazard assessment studies. The main objective of this paper is to give an overview of the past, ongoing, and future joint activities carried out by CEA-EDF-FRAMATOME-IRSN and will especially highlight main findings and future challenges and opportunities in the field of seismic safety of nuclear facilities. To reach the objective, technical focus will be made during the presentation.

INTRODUCTION

The French nuclear industry is highly active in the field of earthquake engineering and more precisely, regarding seismic risk assessment [START_REF] Berge-Thierry | Toward an integrated seismic risk assessment for nuclear safety improving current French methodologies through the SINAPS@ research project[END_REF]. The major nuclear energy operators such as EDF, CEA and FRAMATOME have gathered their research efforts and scientific skills with the French technical support organization (TSO), namely the IRSN, in order to make the knowledge move forward on several topics Berlin/Potsdam, Germany, [START_REF]th International Conference on Structural Mechanics in Reactor Technology[END_REF] Division V related to seismic risk assessment and earthquake engineering. The framework which has been setup to drive and guide the research activities on the aforementioned topics is linked with a project titled Mechanics, Structure, Earthquake project (MSE project) which was setup in 2015. The main objectives of the MSE project are (i) to tackle technical and scientific issues related to methodological practices which may appear in technical safety demonstrations, (ii) to provide to the engineering community recommendations and guidelines approved jointly by the four partners. The MSE project deals with several subjects in the field of seismic risk assessment and earthquake engineering, namely hazard assessment, soil dynamic behavior and soil-structure interaction (SSI), response of structures, systems, and components (SSCs) under seismic loading. In addition, transverse topics are addressed such as the verification and the validation of advanced modeling technics (mostly nonlinear) of SSCs or the updating of the conventional assessment approaches to take into account the experience feedback from either in situ measurements or experimental evidence.

Regarding the tasks related to seismic hazard assessment, we can mention, among others, (i) the maintenance of a French seismic database used to assess the local seismicity, (ii) research works on "kappa" attenuation parameter and the definition of reference rock for site amplification studies or (iii) exchanges to homogenize the tools for probabilistic seismic hazard assessment. The aim is to share basic knowledge between organizations. Regarding the tasks related to structural assessment, several issues are investigated.

Ongoing activities related to the soil-structure interaction (SSI) or to the soil behavior itself can be mentioned. They aim to improve the understanding of this complex phenomenon, especially when soil nonlinearities appear in the vicinity of building foundations. In addition, civil engineering or equipment behavior is also investigated through dedicated tasks. They aim to better assess the structural behavior of complex buildings or equipment when subjected to a beyond-design seismic loading in order to quantify the margins provided by design basis approaches. Some of these tasks are performed within an international context (OECD/NEA), especially when best practices are to be identified by means of benchmarks. Finally, transverse tasks, such as verification/validation methodologies for best estimate approaches or robust updating methodologies to consider the seismic experience feedback and the experimental evidence when dealing with conventional assessment approaches are also ongoing.

The activities include experimental tasks thanks to access to the well-known TAMARIS experimental facility which is operated by CEA in Saclay, France [2]. The objective of this paper is to give an overview of the activities carried out within the framework of the MSE project. To reach this objective, the MSE project is first presented. Then, the recent and planned activities are introduced.

THE MECHANICS, STRUCTURE, EARTHQUAKE (MSE) PROJECT

Topical overview

The seismic probabilistic risk assessment (PRA) is based upon the knowledge of two contributions: the seismic hazard which is described by a hazard curve and acceptability thresholds controlled by dynamic responses of SSCs which is represented by a fragility curve. More precisely, the seismic failure probability is defined as follows:

𝑅 = -∫ 𝑑𝐻(𝑎) 𝑑𝑎 . 𝑓(𝑎)𝑑𝑎 ∞ 0
where 𝑅 is the seismic failure probability (the seismic risk being the convolution of failure probabilities and their consequences), 𝑎 an intensity measure of the seismic input, 𝐻(𝑎) the hazard curve (i.e. the annual frequency of shaking exceeding intensity 𝑎 and 𝑓(𝑎) the fragility curve (i.e. the probability that failure occurs given an intensity measure 𝑎). These ingredients are expressed according to appropriate confidence level values. The hazard curve, is computed using the classic Cornell -McGuire approach [START_REF] Mcguire | FORTRAN computer program for seismic risk analysis[END_REF], [START_REF] Cornell | Engineering seismic risk analysis[END_REF], [START_REF] Mcguire | Computations of seismic hazard[END_REF]; Berlin/Potsdam, Germany, July 10-15, 2022 Division V based on the following four steps: identification and parameterization of the seismic sources; model of earthquake magnitude recurrence for each source; application of ground motion prediction equations and their uncertainty; integration of uncertainties in earthquake location, earthquake magnitude and ground motion prediction, within the combination through a logic tree of the hazard equations. The input used in hazard assessment (fault structure, seismicity catalogues, maximum magnitude assessment, soil properties at the site of interest, etc.) need continuous efforts to be improved and updated, especially in low-moderate seismicity seismotectonic context like Metropolitan France. Regarding the fragility curve, four ingredients are needed in order to estimate this quantity in a satisfactory manner: 1) a random model which takes into account both aleatory and epistemic uncertainties; 2) a methodology to propagate uncertainties that permits to optimize the computational demand and does not introduce any statistical bias in the fragility curve; 3) mechanical models to assess the SSCs response from the needed seismic motion description (time histories, etc.), especially in their "beyond design" range; and 4) a failure criteria and related thresholds necessary to define acceptability criterion and safety domain. In addition to the aforementioned needs and topics, the partners have also decided to gather their efforts to deal with two transverse actions. The first one aims to establish a verification and validation procedure to ensure the reliable character of the advanced methodologies to describe the beyond-design behavior of SSCs. The second one aims to setup a framework permitting to take into account the experience feedback from both the experimental evidence and the postearthquake survey. The recent and ongoing actions are summarized in Figure 1.

Programmatic overview

Two project reviews are organized each year to follow-up the research activities conducted within the framework of the MSE project, to identify difficulties and, if necessary, to re-schedule some of them. The first one usually takes place in April and is centered on exchanges on each activity. The second one occurs in November and is aimed to draw the main conclusions of each activity, and to confirm the work-plan and the tasks to be accomplished in the following years. Besides these two formal meetings, three to four technical meetings are organized per year; they gather the partner together and promote exchanges on the ongoing work, the obtained results, and on the new directions to be taken, if necessary. Despite the fact that the activities are planned each year, partners define work programs for four to five years before starting any new activity.

INSIGHTS ON CURRENT ACTIVITES

ISSNL: Nonlinear soil-structure interaction

Under seismic loading, shallow foundations apply significant stress on the footing-soil contact surface inducing strong nonlinearity in the soil. Hence, it leads to the occurrence of nonlinear phenomena as rocking, uplift, slide, and settlement of the foundation while reducing the bearing capacity of the soil. Continuous efforts are ongoing to improve modeling techniques approximating rocking and uplifting in inelastic soil behavior ([6], [START_REF] Gajan | Effect of critical contact area ratio on moment capacity of rocking shallow footings[END_REF]). However, tedious calibration of numerical models is required in order to capture these complex phenomena. The ISSNL activity aims to improve the nonlinear SSI tools (Figure 2) of "best-estimate" type thanks to a multi-annual research and development program. The state-of-the-art has shown the existence of numerous numerical and analytical studies and a lack in the experimental analyses. In order to propose improvement to foundation-soil interface numerical modeling, controlled experimental tests in-situ and in laboratory are essential. However, difficulties arise on the limitations in the dimensions of the specimen and therefore its representativeness. Compliance with Laws of similarity requirements is indeed not easy to ensure. Therefore, in the frame of this work, the project aims to develop a methodology to measure "more directly" parameters as Impedance Function (IF) for rigid slab type foundation that would be an input in Finite Element analysis (Figure 3). The principle of this experimental study lays on the application of uniaxial mono-frequential sinusoidal forces. Hence, setting the equilibrium of forces and momentum applied to the foundation for one particular direction of solicitation and one frequency permits to assess the value of the impedance.

NLTA: Damping modeling in time history nonlinear analysis of nuclear buildings: effects on the behaviour of structure and equipment

Energy dissipation in structures subjected to seismic loading is typically modeled using Rayleigh, Caughey or modal (Wilson-Penzien) damping. It is well known that these techniques are not based on the appropriate energy dissipation mechanisms in structures. The use of these approaches and, in particular the classical Rayleigh damping in nonlinear analyses, is questionable. Indeed, as many studies have shown [START_REF] Zareian | A practical method for proper modeling of structural damping in inelastic plane structural systems[END_REF], [START_REF] Hall | Problems encountered from the use (or misuse) of Rayleigh damping[END_REF], [START_REF] Charney | Unintended consequences of modelling damping in structures[END_REF], [START_REF] Chopra | Modeling viscous damping in nonlinear response history analysis of buildings for earthquake excitation[END_REF], making use of this type of damping for a relatively high level of seismic loading often leads to inaccurate estimates of displacements and internal forces. These inaccurate estimations are related to the fact that unrealistic viscous forces are generated by the proportional Rayleigh damping, in particular by the stiffness proportional term, when cracks open. The proportion of energy dissipated by the viscous damping model should therefore be better controlled. Despite the fact that the calibration of the proportional Rayleigh damping associated with the implementation of dissipative processes constitutive modelling within FE analyses is not the result of scientific consensus, it is considered that this way of modelling damping will continue to be used in the near term, in parallel with the development of new model formulations of constitutive laws. The research activities, heavily relying on the feedback from experimental campaigns on regular scale reinforced concrete structures, aim at:

➢ providing guidance on the best way to use Rayleigh and modal damping. This part of the research should make it possible to identify the limits of commonly used models and to develop a set of recommendations for the best use of current methods, depending on the type of structure, seismic level, constitutive modeling, etc.; ➢ calculating energy balance in order to better understand how the energy is transmitted to the system during the transient response: how much dissipation is provided by the viscous damping, compared to the one associated with the hysteretic cyclic constitutive law used in the analysis; ➢ developing new approaches for damping modeling in nonlinear analysis. This is obviously a more difficult task than the previous one, since it involves developing consistent nonlinear cyclic constitutive laws accounting for damping, which aims to improve the predictive capabilities of nonlinear concrete models.

On the other hand, another action is added to this activity and consists in identifying the contribu tion of energy dissipated at the steel-concrete interface of a structure subjected to dynamic loading among the other sources of energy dissipation: viscous dissipation, numerical dissipation due to the time integration scheme, and mainly material dissipation. More precisely, the main objective is to propose a new simple modeling strategy to take into account the behaviour of the steel-concrete bond. To this end, we define the relationship between steel and concrete displacements at the interface in terms of kinematic relationships that depend on the loading level increment. This work fits into the framework of Maryam Trad's PhD thesis, which started in 2020.

CP2R: Seismic behavior of cranes bridges

Within the context of Level 1 Probabilistic Safety Assessment (PSA) studies (which deals with occurrence of core damage) performed on nuclear power plants, crane bridges failure have been identified as a significant contributor in the probability of core meltdown. Depending on the reactor type and on the age of the reactor design, one of the significant failure mode can be related to anchorage failure during earthquake. Therefore, the issue of dynamic behavior of crane bridges needs to be considered within the global framework of the safety demonstration of a plant. In this context, it is necessary to improve Berlin/Potsdam, Germany, July 10-15, 2022 Division V knowledge about the dynamic behaviour of this equipment in order to fully understand how failure would occur in the case of beyond design loadings induced by most severe earthquakes (that are considered in risk assessment). Therefore, it is important to estimate accurately the forces transmitted to the anchorages. In addition, the incorporation of different sources of uncertainties through a fragility curve still raises several questions such as: what are the main variables to be considered as random? What are the failure criteria to be used? Does the hypothesis of a lognormal distribution remain justified for seismic inputs for which the intensity is in the beyond-design range?

In order to provide answers to these aforementioned questions, a benchmark named SOCRAT endorsed by the Working Group on Integrity and Ageing of Components and Structures (WIAGE) of OECD/NEA/CSNI started in 2020. The main objectives of this action are (i) to identify best practices to model seismic behaviour of crane bridges; (ii) to identify relevant failure criteria.

An experimental campaign on a scaled model of an overhead crane bridge was carried out in 2015 on the AZALEE shaking table of CEA in Saclay, France and the results have been gathered in a large database. The crane bridge mock up is shown in Figure 2. On one hand, some of these data will be used by participants to characterize and calibrate their models and, on the other hand, some other data will be used to assess the predictive capacity of the mechanical models. The benchmark is concluded by a restitution workshop in which the different participants have met to exchange and discuss about their models and results they have obtained. As a result, best practices for modeling overhead cranes under seismic loading were identified and will be published in a forthcoming paper with a synthesis of lessons learnt and recommendations based on findings of the benchmark analysis. Figure 3 gives an overview of the benchmark schedule. For more details, see [START_REF] Bitar | Dynamic behaviour of crane bridges -International benchmark[END_REF]. For those interested in SOCRAT benchmark, there is a special session during the SMiRT26 confere nce on this topic entitled "Overview of the work done in the OECD SOCRAT benchmark dedicated to the beyond design seismic behavior assessment of crane bridges".

V&V: Verification and validation of nonlinear techniques

Due to the increasing complexity of numerical models and methodologies in computational structural mechanics, it is more and more difficult to ensure the validity of simulation results in the expected use range. For instance, a model developed by academic research might be discarded because it does not offer sufficient experience feedback for large-scale industrial applications. In the same time, the safety requirements for nuclear structures and equipment become stricter making it essential to have tools to verify and validate theoretical models and their numerical implementation. In order to allow for the use of cutting edge mechanical models in nuclear energy structure assessment, the MSE project wishes to develop a methodological framework for the verification and validation of nonlinear models. Such frameworks have already been proposed in a more general way following different strategies ( [START_REF] Roy | A comprehensive framework for verification, validation, and uncertainty quantification in scientific computing[END_REF], [START_REF] Roy | Review of code and solution verification procedures for computational simulation[END_REF], [START_REF] Ladevèze | Validation of structural dynamics models containing uncertainties[END_REF], [START_REF] Sargent | Verification and validation of simulation models[END_REF], [START_REF] Rebba | Validation and error estimation of computational models[END_REF]) or more specifically in other domains, such as soil mechanics [START_REF] Mestat | Référentiel de tests pour la vérification de la programmation des lois de comportement dans les logiciels d'éléments finis[END_REF]. However, while general recommendations are found in the IAEA specific safety guide [START_REF]Deterministic safety analysis for nuclear power plants[END_REF], no guidelines are provided to meet these requirements. The goal of V&V activity is to transpose the frameworks developed in the literature to the specific challenges of the nuclear industry [START_REF]Qualification des outils de calcul scientifique utilisés dans la démonstration de sûreté nucléaire[END_REF]. Validation and verification are both strongly related not only to uncertainty quantification but also the quality assurance process of the whole simulation solution. One could define the validation process as the demonstration of the acceptability of uncertainties related to the theoretical model (model form uncertainty), while the verification is the demonstration of the acceptability related to the numerical implementation of this model for representative physical cases (discretization, convergence and round-off uncertainties). Both processes require reference data which can be either theoretical, experimental or numerical (provided that it comes from a validated and verified model). Hence, the building of an exhaustive V&V data library associated with a methodological flowchart for practitioners is of primary importance for this activity. The collaboration between nuclear operators and IRSN is an opportunity to mutualize the knowledge and expertise of the different partners. The V&V actio n will be valorized by the drafting of a technical report in 2023.

REX: Integration of the experience feedback

Conventional approaches can lead to results sometimes very different from post-seismic observations and/or require disproportionate engineering efforts to assess SSCs outputs in a consistent way. In that context, the exploitation of post-seismic or experimental feedback data is a key point in the process of improving their robustness and their accuracy. Then, the objectives are to have a better understanding of the differences between the results of the current engineering methods and the observations, to q uantify and compare those gaps regarding the uncertainties. In order to achieve this goal, IRSN and EDF have proposed to use feedbacks from the Le Teil earthquake which occurred on the 11 th of November 2019 close to the CRUAS nuclear power plant (NPP). More precisely, organizing an international benchmark under the umbrella of OECD/NEA is planned. The above-mentioned Benchmark will start in early 2023 and continue until 2024 in order to assess the adequacy of engineering practices for calculating the seismic response of the structure, especially for the CRUAS NPP isolated seismic base, and to share the experience with the wider community. For more information about the benchmark, please refer to the SMiRT 26 special session entitled "2019-11-11 Le Teil Earthquake: A special earthquake occurred close to a special NPP" in which a presentation will be made.

CONCLUSIONS AND OUTLOOK

In this paper, an overview of the research and development activities carried out by CEA, IRSN, EDF and FRAMATOME within the framework of the MSE project has been presented. Their main objective is to reduce the uncertainties when assessing the seismic safety. To reach this objective, the topics of ongoing and planned activities have been briefly presented. Among the ongoing tasks, we mentioned specific effort dedicated to knowledge and modeling improvement related to the seismic behavior of soil and SSCs. Berlin/Potsdam, Germany, July 10-15, 2022 Division V Especially, one of the activities aims to better understand and assess the seismic behaviour of crane bridges. This action is carried out under the umbrella of OECD/NEA in order to share with the international community on this topic. This research and development initiative has been supported by seven member States and will lead to organize an international benchmark over the period 2020 to 2021. Two transversal activities have been initiated in 2019. The aim of the first one is to define and to apply a probabilistic framework to update the input parameters of nonlinear models in order to take into account the experience feedback. In addition, the second transversal activity lies in setting up a framework based upon the available regulatory corpus to verify and validate the best-estimate techniques commonly used to assess the safety margins when exist. In 2020 four new joint actions have been initiated within the framework of the MSE project, each of them related to different steps of seismic hazard assessment: source, propagation, site response, PSHA.

Fig. 1 -

 1 Fig. 1 -Summary of the research activities setup within the MSE project.

Fig. 2 -

 2 Fig. 2 -Problem definition of nonlinear SSI taking into account rocking and uplifting of the footing.

Fig. 3 -

 3 Fig.3-Principle of the mechanical system of the experimental system for measuring IF.

  Conference on Structural Mechanics in Reactor Technology Berlin/Potsdam, Germany, July 10-15, 2022 Division V

Fig. 2 -

 2 Fig. 2 -Overhead crane bridge mock up put on the AZALEE shaking table.

Fig. 3 -

 3 Fig. 3 -SOCRAT Benchmark agenda.

Table 1 -

 1 Overview of the ongoing activities

	Activity label	Title	Starting date	Ending date	Objectives

ACKNOWLEDGEMENTS

The Authors would like to express their most grateful thanks to EDF, CEA, FRAMATOME and IRSN for the financial supports granted.


Orianne Dumas

Régis Matran

Farid Zerimech

Brigitte Decoster

Helene Huyvaert

Ismaïl Ahmed

Nicole Le Moual

Rachel Nadif

Occupational exposures and fluorescent oxidation products in 723 adults of the EGEA study

ou non, émanant des établissements d'enseignement et de recherche français ou étrangers, des laboratoires publics ou privés.

Occupational asthma can be induced by a variety of agents, including high and low molecular weight sensitisers, and respiratory irritants [START_REF] Tarlo | Occupational asthma[END_REF]. The role of exposure to cleaning products and disinfectants in work-related asthma is increasingly recognised, although the specific substances that increase asthma risk are not well identified [START_REF] Siracusa | Asthma and exposure to cleaning productsa European Academy of Allergy and Clinical Immunology task force consensus statement[END_REF]. Some of the numerous agents contained in these products are chemical sensitisers, but most are hypothesised to act as respiratory irritants [START_REF] Siracusa | Asthma and exposure to cleaning productsa European Academy of Allergy and Clinical Immunology task force consensus statement[END_REF]. While high molecular weight sensitisers are known to cause occupational asthma through a typical allergic response, the pathophysiological mechanisms involved in occupational asthma induced by low molecular weight (LMW) chemicals, and in irritant-induced asthma, remain poorly understood [START_REF] Tarlo | Occupational asthma[END_REF][START_REF] Maestrelli | Mechanisms of occupational asthma[END_REF][START_REF] Tarlo | Irritant-induced asthma in the workplace[END_REF].

Oxidative stress is one of the potential mechanisms causing epithelial injury, which may be especially relevant in irritant-induced asthma [START_REF] Tarlo | Occupational asthma[END_REF][START_REF] Tarlo | Irritant-induced asthma in the workplace[END_REF]. In this context, we sought to investigate the associations between occupational exposure to potentially asthmogenic chemicals and irritants, and the level of fluorescent oxidation products, a global marker of damage due to oxidative stress [START_REF] Wu | Plasma fluorescent oxidation products as potential markers of oxidative stress for epidemiologic studies[END_REF], in adults from the Epidemiological study on the Genetics and Environment of Asthma (EGEA).

The French EGEA combines a case-control and family-based study [START_REF] Siroux | Phenotypic determinants of uncontrolled asthma[END_REF]. The baseline study (EGEA1; 1991-1995, n=2047) included cases with asthma, their first degree relatives, and population-based controls. The study protocol was approved by the relevant institutional review board committees (Cochin Port-Royal Hospital and Necker-Enfants Malades Hospital, Paris, France) and all participants gave written informed consent. The current analysis used data from the follow-up study (EGEA2; 2003-2007) and included only adults who had never had asthma (n=888), in order to distinguish occupational exposures from the oxidative stress generated by the disease. 723 participants (328 men and 395 women) had fluorescent oxidation products measurements available and complete data for smoking habits, an environmental exposure related to oxidative stress. Levels of fluorescent oxidation products were measured in plasma as described previously [START_REF] Wu | Plasma fluorescent oxidation products as potential markers of oxidative stress for epidemiologic studies[END_REF]. Briefly, plasma was extracted with 3:1 by volume ethanol/ether and was measured using a spectrofluorometer (360nm excitation wavelength, 430nm emission wavelength). The fluorescence was determined as relative fluorescent intensity units per millilitre of plasma (RFU•mL -1 ).

Occupational history was recorded by questionnaire, and an asthma-specific job-exposure matrix ( JEM) was used to evaluate exposure to agents potentially at risk for asthma or respiratory health [START_REF] Kennedy | Development of an asthma specific job exposure matrix and its application in the epidemiological study of genetics and environment in asthma (EGEA)[END_REF]. Hospital workers (including healthcare workers or cleaners in hospitals) were further asked to complete a job-specific questionnaire regarding exposure to cleaning/disinfecting tasks and products. An expert assessment was also conducted to evaluate exposure to 18 specific products in hospital workers [START_REF] Donnay | Under-estimation of self-reported occupational exposure by questionnaire in hospital workers[END_REF]. We chose to study the following occupational exposures, in the current or last job held, in relation to levels of fluorescent oxidation products: 1) exposure to cleaning/disinfecting products evaluated both by the JEM (a highly specific assessment in which only jobs with high probability of exposure to industrial cleaning products were classified as exposed), and by the expert assessment with a focus on cleaning/disinfecting products which are more likely to be respiratory irritants; and 2) exposure to LMW agents evaluated by the JEM.

Associations between occupational exposures and levels of fluorescent oxidation products were evaluated separately in men and women. We used linear regression of log-transformed levels of fluorescent oxidation products, adjusted for age and smoking habits. The coefficients obtained from the log-transformed fluorescent oxidation products models were back-transformed to obtain geometric mean (GM) ratios.

The GM (first quartile-third quartile) of fluorescent oxidation products level in all adults was 94.4 (81.9-108.1) RFU•mL -1 . As expected [START_REF] Wu | Plasma fluorescent oxidation products as potential markers of oxidative stress for epidemiologic studies[END_REF], levels of fluorescent oxidation products increased with age ( p<0.0001), were higher in women (age-adjusted GM (95% CI) 95.7 (93.7-97.7) RFU•mL -1 ) than in men (92.9 (90.8-95.0) RFU•mL -1 ; p=0.06), and increased with current smoking. Among women, levels of fluorescent oxidation products were increased in current smokers (age-adjusted GM (95%CI): 100.2 (95.5-105.2) RFU•mL -1 ; p=0.004) and ex-smokers (99.6 (95.6-103.7) RFU•mL -1 ; p=0.002) compared with nonsmokers (92.2 (89.7-94.8) RFU•mL -1 ). Among men, levels of fluorescent oxidation products were increased in current smokers (98.8 (94.0-103.7) RFU•mL -1 ; p=0.005), but were not significantly higher in ex-smokers (92.4 (88.7-96.3) RFU•mL -1 ; p=0.51) compared with nonsmokers (90.7 (87.6-93.4) RFU•mL -1 ). Occupational exposure to irritant cleaning products and to LMW agents, especially highly reactive chemicals, were significantly associated with higher levels of fluorescent oxidation products in men (table 1). In women, an association was suggested ( p=0.08) between occupational exposure to irritant cleaning products and levels of fluorescent oxidation products.

To our knowledge, this study is the first to show an association between occupational exposures to potentially asthmogenic chemicals and irritants, and levels of fluorescent oxidation products. Several studies have linked irritant exposures, such as cigarette smoke, to levels of fluorescent oxidation products [START_REF] Wu | Plasma fluorescent oxidation products as potential markers of oxidative stress for epidemiologic studies[END_REF][START_REF] Yang | Association between global biomarkers of oxidative stress and hip fracture in postmenopausal women: a prospective study[END_REF][START_REF] Jensen | Fluorescent oxidation products and risk of coronary heart disease: a prospective study in women[END_REF]. By contrast, a study of 236 male workers in the truck industry did not find evidence for any association between occupational exposure to particulate matter and levels of fluorescent oxidation products, but the authors acknowledged limitations in the study power and occupational exposure assessment method [START_REF] Hart | Plasma fluorescent oxidation products and short-term occupational particulate exposures[END_REF]. In a study investigating asthma among 92 Spanish cleaning workers, no differences in levels of exhaled breath condensate 8-isoprostanes, a specific marker of lipid peroxidation, were observed according to case/control status or to self-reported occupational exposures to cleaning products [START_REF] Vizcaya | Functional and biological characteristics of asthma in cleaning workers[END_REF]. However, levels of fluorescent oxidation products, as a nonspecific but global systemic measurement, has been suggested to be a better marker of oxidation as it measures products resulting from oxidation of lipids, proteins, amino acids, and DNA [START_REF] Wu | Plasma fluorescent oxidation products as potential markers of oxidative stress for epidemiologic studies[END_REF][START_REF] Jensen | Fluorescent oxidation products and risk of coronary heart disease: a prospective study in women[END_REF].

This study included a relatively large sample of participants compared with the existing literature, although numbers were low for some exposure categories. One of the strengths of this study is the good characterisation of occupational exposures using expert assessment and an asthma-specific JEM, two complementary methods which do not rely on the workers' own report [START_REF] Dumas | Occupational exposure to cleaning products and asthma in hospital workers[END_REF]. However, detailed assessment of exposure to the multiple chemicals contained in cleaning and disinfecting products is challenging, and the sample size did not allow us to study associations between levels of fluorescent oxidation products and specific agents, or a potential dose-response effect according to duration or timing of exposure.

For both exposure to irritant cleaning products and LMW chemicals, the associations were stronger and reached statistical significance in men ( p-value for interaction ranging from 0.03 to 0.28). Results were similar after adjustment for body mass index (data not shown). Among exposed subjects, the main gender difference in occupations was that women were more often employed as personal care workers than men. We previously found, in EGEA2, that cleaning exposures among female healthcare workers were associated with current asthma, with a potential role for both irritants and sensitisers [START_REF] Dumas | Occupational exposure to cleaning products and asthma in hospital workers[END_REF]. A potential association between occupational exposures and oxidative stress may be more difficult to detect in these women experiencing exposure to multiple cleaning and disinfecting products, especially when further considering domestic exposures [START_REF] Moual | Domestic use of cleaning sprays and asthma activity in females[END_REF][START_REF] Moual | Use of household cleaning products, exhaled nitric oxide and lung function in females[END_REF]. Gender differences may also be explained by the different ability to maintain redox status between men and women [START_REF] Malorni | Redox features of the cell: a gender perspective[END_REF]. Larger studies are needed to further investigate the role of exposure to specific chemicals in women.

Irritant-induced asthma was first described in relation to a single exposure to high concentrations of irritants. There is increasing recognition that repeated workplace exposures to moderate levels of irritants also has an adverse effect in asthma, although whether such chronic exposure can cause new-onset asthma remains unclear [START_REF] Tarlo | Irritant-induced asthma in the workplace[END_REF]. Although the biological processes by which irritant cleaning products induce asthma onset or exacerbations are unknown, it has been proposed that irritants may induce injury of the bronchial epithelium, triggering release of inflammatory mediators and/or increasing epithelium permeability [START_REF] Tarlo | Occupational asthma[END_REF][START_REF] Tarlo | Irritant-induced asthma in the workplace[END_REF]. Our results are consistent with the hypothesis that oxidative stress is one of the mechanisms through which respiratory irritants may cause epithelial damage.

In conclusion, the results suggest that occupational exposures to asthmogenic chemicals and irritants, in particular cleaning and disinfecting products, generate oxidative stress. Future research should investigate oxidative stress as one of the potential mechanisms of chemical-or irritant-induced occupational asthma.

Associations between occupational exposures to asthmogenic chemicals and irritants and oxidative stress were found http://ow.ly/K6RSt

Orianne Dumas 1,2,3,4 , Régis Matran 5 , Farid Zerimech 6 , Brigitte Decoster 5 , Helene Huyvaert 5,6 , Ismail Ahmed

Eur

  Respir J 2015; 46: 255-258 | DOI: 10.1183/09031936.00047115 | Copyright ©ERS 2015 Occupational exposures and fluorescent oxidation products in 723 adults of the EGEA study To the Editor:

  7,8 , Nicole Le Moual1,2 and Rachel Nadif 1,2 1 INSERM U1168, VIMA: Aging and chronic diseases, Epidemiological and public health approaches, Villejuif, France.2 Univ Versailles St-Quentin-en-Yvelines, UMR-S 1168, Montigny le Bretonneux, France.3 Channing Division of Network Medicine, Dept of Medicine, Brigham and Women's Hospital, Boston, MA, USA.4 Dept of Emergency Medicine, Massachusetts General Hospital, Boston, MA, USA.5 Univ Lille Nord de France, Lille, France.6 Laboratoire de Biochimie, Centre de Biologie Pathologie, CHRU de Lille, Lille, France.7 INSERM UMR 1181 «Biostatistics, Biomathematics, Pharmacoepidemiology and Infectious Diseases» (B2PHI), Villejuif, France.8 Univ Versailles St-Quentin-en-Yvelines, UMR 1181, B2PHI, Montigny le Bretonneux, France.

Correspondence: Rachel Nadif, INSERM U1168, VIMA: Aging and chronic diseases, Epidemiological and public health approaches, F-94807, Villejuif, France. E-mail: rachel.nadif@inserm.fr Received: Sept 26 2014 | Accepted after revision: Feb 04 2015 | First published online: April 02 2015

TABLE 1

 1 Age-and smoking-adjusted associations between current occupational exposure and levels of fluorescent oxidation products in men and women Jobs classified as exposed to irritant cleaning products (assessed among hospital workers only) included various healthcare-related occupations (e.g. nursing professionals, personal care workers (mainly women), medical doctors, physiotherapists and life science technicians). Jobs classified as exposed to highly reactive chemicals included chemists, biologists or pharmacologists, and jobs such as varnishers/painters, embalmers or hairdressers. GM: geometric mean; JEM: job-exposure matrix; LMW: low molecular weight. # : nonexposed to any agent potentially a risk for asthma or respiratory health (asthma-specific JEM); ¶ : including bleach, ammonia, decalcifiers, solvents, formaldehyde, glutaraldehyde, alcohol, quaternary ammonium compounds, chloramine T, ethylene oxide, and disinfection or bio-cleaning tasks; + : including industrial cleaning products, highly reactive chemicals, metals and diisocyanates.

	Occupational exposure		Men			Women		p-value for
								interaction
		Subjects	GM ratio	p-value Subjects	GM ratio	p-value
		n	(95% CI)		n	(95% CI)
	Non exposed # (reference)	180	1		289	1
	Cleaning products
	1) Industrial cleaning products (JEM)	1			11	1.06 (0.93-1.21)	0.36
	2) Irritant cleaning products	10	1.24 (1.07-1.43)	0.004	51	1.06 (0.99-1.13)	0.08	0.05
	(expert assessment) ¶
	1) or 2)	11	1.24 (1.08-1.42)	0.002	52	1.05 (0.99-1.12)	0.12	0.03
	LMW agents ( JEM)
	Any LWM agent +	26	1.10 (1.02-1.18)	0.01	27	1.01 (0.92-1.09)	0.89	0.14
	Highly reactive chemicals	13	1.12 (1.01-1.23)	0.03	24	1.02 (0.93-1.11)	0.66	0.28
	Metals	11	1.05 (0.95-1.17)	0.34	1

Acknowledgements: The authors thank all those who participated in the study and in the various aspects of the examinations and all those who supervised the study centres. The authors are grateful to the three CIC-Inserm units at Necker, Grenoble and Marseille (France), which supported the study and where subjects were examined. They are also grateful to the three biobanks in Lille (CIC Inserm), Evry (Centre National de Genotypage) and Annemasse (Etablissement Français du Sang; France) where biological samples are stored. The authors thank Sylwester Karpiel (INSERM U1018, Centre for research in Epidemiology and Population Health (CESP), Respiratory and Environmental Epidemiology Team, Villejuif, France) for his contribution to this work. They are indebted to all the individuals who participated, without whom the study would not have been possible. The EGEA cooperative group are as follows. Coordination: V. Siroux (epidemiology, PI since 2013); F. Demenais (genetics); I. Pin (clinical aspects); R. Nadif (biology); F. Kauffmann (PI 1992-2012). Respiratory epidemiology: Inserm U 700, Paris: M. Korobaeff (Egea1) and F. Neukirch (Egea1); Inserm U 707, Paris: I. Annesi-Maesano (Egea1-2); Inserm CESP/U 1018, Villejuif: F. Kauffmann, N. Le Moual, R. Nadif, MP. Oryszczyn (Egea1-2) and R. Varraso; Inserm U 823, Grenoble: V. Siroux. Genetics: Inserm U 393, Paris: J. Feingold; Inserm U 946, Paris: E. Bouzigon, F. Demenais and M.H. Dizier; CNG, Evry: I. Gut (now CNAG, Barcelona, Spain) and M. Lathrop (now McGill University, Montreal, Canada). Clinical centres: Grenoble: I. Pin and C. Pison; Lyon: D. Ecochard (Egea1), F. Gormand and Y. Pacheco; Marseille: D. Charpin (Egea1) and D. Vervloet (Egea1-2); Montpellier: J. Bousquet; Paris Cochin: A. Lockhart (Egea1) and R. Matran (now in Lille); Paris Necker: E. Paty (Egea1-2) and P. Scheinmann (Egea1-2); Paris-Trousseau: A. Grimfeld (Egea1-2) and J. Just. Data and quality management: Inserm ex-U155 (Egea1): J. Hochez; Inserm CESP/U 1018, Villejuif: N. Le Moual; Inserm ex-U780: C. Ravault (Egea1-2); Inserm ex-U794: N. Chateigner (Egea1-2); Grenoble: J. Quentin-Ferran (Egea1-2).

This analysis was funded by AstraZeneca. funded by AstraZeneca.

Conflict of interest: None declared.