Pascal Zavaleta

Sylvain Suard

Laurent Audouin

FIRE SPREAD FROM AN OPEN-DOORS ELECTRICAL CABINET TO NEIGHBOURING TARGETS IN A CONFINED AND MECHANICALLY-VENTILATED FACILITY

Keywords: adjacent cabinet, DIVA facility, electrical cabinet, fire spread, overhead cable trays

Electrical cabinet fire is one of the main fire hazards in nuclear power plants. As part of the OECD PRISME-2 programme, four fire tests were carried out to investigate the fire spread from an opendoors electrical cabinet to overhead cable trays and adjacent cabinets in a confined and mechanicallyventilated facility. These tests, named CFS-5 to CFS-7 and CORE-6, used same both cabinet (fire source) and three overhead cable trays. The trays were filled with a halogenated flame retardant cabletype for CFS-5 and one halogen-free for the three other tests. Moreover, fire dampers were used for CFS-7 test while CORE-6 test implemented two additional cabinets adjacent to the fire source. Measurements such as flame and gas temperature, gas concentration, mass loss rate and heat release rate were performed for investigating the fire spread. Cabinet fire spread to the cable trays for CFS-5 and CFS-6 tests. Three fast and short cable tray fires were showed for CFS-5 while a slow and long cable tray fire was highlighted for CFS-6. In contrast, the fire dampers shutdown for CFS-7 test prevented ignition of the overhead cables. Furthermore, for CORE-6 test, cabinet fire spread to the adjacent cabinets but the upper cables were not ignited.

NOMENCLATURE

Specific area associated with the heat flux sensor ''i'' (m²)

Greek characters

Isentropic constant (equal to 1.4 for diatomic gases) Volume flow rate at the opening i (m 3 /s)

| INTRODUCTION

Nearly 400 fire events in nuclear power plants (NPPs) were recorded since the 1980's to the end of 2010 in the current OECD FIRE Database [START_REF]OECD FIRE Database[END_REF] . From this last one, Werner 2 indicated that one of the main contributions to the fire hazard in NPPs was attributed to electrical cabinets, representing 12 % of all the previous fire events (i.e., 48 cabinet fires). This author also argued that a cabinet fire in a control or switchgear room of NPPs could lead to failure of one safety train or more, if fire spreads beyond the fire cabinet. Consequently, the study of electrical cabinet fires and their consequences in NPPs is a major concern for fire safety analyses. Previous investigations particularly investigated the fire spread within an electrical cabinet and fire ability to spread beyond the cabinet. Chavez 3,[START_REF] Chavez | An experimental investigation of internally ignited fires in nuclear power plant control cabinets, Part II: Room effects tests[END_REF] performed fire tests which involved real-scale vertical and benchboard cabinet types. These tests involved a cabinet used as fire source (i.e., a fire cabinet) and neighboring cabinets such as adjacent cabinets to the fire cabinet. Fire cabinets were most often without door and loaded with IEEE-383 5 qualified or un-qualified electrical cables. Adjacent and fire cabinets were separated by a double wall (a wall for each cabinet and an air gap of 2.54 cm between). All these tests highlighted that the fire can spread fast throughout the fire cabinet without door and whatever the cable-type. Moreover, maximal heat release rate (HRR) was measured higher than 1200 kW for vertical opendoors cabinets and for the two cable-types while it was assessed at 185 kW for closed-doors cabinets. However, despite such HRR for the open-doors cabinets, fire did not spread to adjacent cabinets. The air gap between the cabinets indeed prevented self-ignition of the electrical cables fixed to the side walls of the adjacent cabinets.

Mangs et al. 6 conducted thirteen experiments which involved two real vertical closed-doors electronic cabinets (nine full-scale tests) and two cabinet mock-ups (four reduced-scale tests). The two electronic cabinets used were relay cabinets and circuit board cabinets. Besides relays or circuit boards, these cabinets contained mainly electrical cables made with Poly(vinyl chloride) (PVC) jacket. The cabinet mock-ups contained only cables loosely attached to a holder at their centerline. All experiments were carried out beneath a large hood which collected all combustion products. The electronic cabinets were ignited in all experiments with a propane gas burner. The higher HRR peaks were measured in the 100-350 kW range for all the experiments leading to flashover inside the cabinet. The highest HRR peak was obtained when the door locking mechanism failed during one of the real cabinet fires. Furthermore, for all the nine full-scale tests (real electronic cabinet), an adjacent cabinet, made of 0.5 mm thick steel sheet, was fastened to the fire cabinet. This last one contained samples of PVC cables attached to the wall against the fire cabinet, and also placed at 5 and 30 cm from this last wall. For two full-scale experiments, the adjacent cabinet ignited caused by self-ignition of PVC cables contained inside it.

Coutin [START_REF] Coutin | Phenomenological description of actual electrical cabinet fires in a free atmosphere[END_REF] and Coutin et al. [START_REF] Coutin | Characterization of open-doors electrical cabinet fires in compartments[END_REF] performed eleven real cabinet fire tests. Four cabinet fires were carried out in open atmosphere conditions (under a large-scale calorimeter) while seven fire tests were performed in a confined and mechanically ventilated facility. All the fire tests involved the same real electro-technical cabinet and fire scenarios implied both closed and open doors. A linear propane burner ignited the fuel load at the base of the cabinet, along its entire width. All the open-doors electrical cabinet fires showed significant HRR peaks which were higher in open atmosphere conditions (1200 and 1600 kW) than in confined conditions (from 400 to 950 kW). In contrast, the two closed-doors cabinet fires exhibited HRR lower than 50 kW.

McGrattan and Bareham 9 performed hundred-twelve full-scale experiments under a mediumscale calorimeter. These tests involved eight low voltage control electrical cabinet types which may be found in NPPs. A small propane burner and various size pans of acetone were used as ignition source or preheating source. The main goal of this study was to provide HRR measurements for fire models validation, used in fire safety studies. These numerous experiments confirmed that the cabinet geometry, ventilation, ignition strength (i.e., the amount of energy necessary to start the fire) and combustible loading (i.e., the amount and type of fuel materials) are the key parameters which impact the fire spread and the resulting HRR. Eleven of the cabinet fires had HRR peaks higher than 100 kW and ten of these last ones had a HRR peak lower than 300 kW. Such HRR peaks were obtained for both open-and closed-doors configurations. The highest HRR peak, measured at nearly 600 kW, was obtained for open-doors cabinet loaded with IEEE-383 un-qualified cables, and loosely arranged.

All previous programmes highlighted that cabinet fire may produce HRR higher than 100 kW for a large range of cabinet configuration (open-doors or closed-doors, electronic or electro-technical, vertical or benchboard…). These programmes also showed that for such level of HRR generated, fire may spread from a cabinet fire to an adjoined cabinet (i.e., without air gap). Furthermore, electrical cabinets, in real industrial plants as in NPPs, may be connected each other via electrical cables installed along overhead cable trays or placed inside false floors (or raised floors). Such potential fire spread paths to adjacent cabinets therefore also need to be investigated.

As part of two test campaigns of the OECD PRISME-2 programme 10 , four fire tests were carried out to study the fire spread from an open-doors electrical cabinet to neighbouring targets in a confined and mechanically-ventilated facility. The targets are three overhead cable trays, two adjacent cabinets and a false floor. Three fire tests, named CFS-5 to CFS-7 [START_REF] Zavaleta | Fire spreading from a real open-doors electrical cabinet to overhead multiple cable trays into a confined and mechanically-ventilated[END_REF] , were carried out in the scope of the cable fire spreading (CFS) test campaign. Fire dampers were set up for CFS-7 test and they were shut down early after the starting of the fire. The fourth test, called CORE-6, was conducted as part of the last CORE test campaign 11 which aimed at COmpleting and REpeating some tests of the previous campaigns.

This paper first presents in detail the experimental set-up used for this study. This part thus describes the electrical cabinet fire source, the overhead cable trays, the adjacent cabinets and the false floor, as well as the facility which hosted the tests and finally the instrumentation. The second and main part of this paper first presents how the HRR is assessed for the four tests. Then, the results of CFS-6 test (reference test), dealing with especially the fire spread from the cabinet fire source to overhead cable trays, are detailed. Finally, the effects on fire spread of the fire dampers shutdown and cable-type (installed along the overhead cable trays and in the adjacent cabinets) are especially discussed from the comparison of the results of the reference test with those of CFS-5, CFS-7 and CORE-6 tests.

| DESCRIPTION OF THE FIRE TESTS

Table 1 indicates the main features of CFS-5 to CFS-7 and CORE-6 tests which were carried out in the confined and mechanically-ventilated facility, named DIVA (Fig. 1). These four fire tests used same both open-doors electrical cabinet fire source (Fig. 2), named the fire cabinet, and three overhead cable trays (Fig. 3 and Fig. 4). The cable trays were filled with a halogenated flame retardant (HFR) cabletype for CFS-5 test and a halogen-free flame retardant (HFFR) cable-type for the three other tests. Furthermore, fire dampers were set up for CFS-7 test and they were shut down 2 min 30 s after the starting of the cabinet ignition. Finally, CORE-6 test implemented two additional closed-doors electrical cabinets which were placed against the central electrical cabinet (Fig. 5). These three cabinets were located above a false floor. The two adjacent cabinets only contained PVC trunkings placed against the adjacent side wall and filled with HFFR and HFR cable-types.

| Fire cabinet

The fire cabinets (Fig. 2) used for the four tests were the same twin modules 400 V electro-technical cabinet, as those previously studied by Coutin [START_REF] Coutin | Phenomenological description of actual electrical cabinet fires in a free atmosphere[END_REF] and Coutin et al. [START_REF] Coutin | Characterization of open-doors electrical cabinet fires in compartments[END_REF] . These electrical cabinets, which can be found in French nuclear installations, were fully characterised in these previous experimental works. As part of this study, the electrical cabinets were implemented without door since this configuration is considered in fire safety analyses as the most critical fire scenario. The cabinet width, depth and height are respectively of 1.2, 0.6 and 2.2 m. The main components inside the cabinet (Fig. 2) are transformers (1), terminal blocks (2), motor circuit breakers and contactors (3), relays (4), circuit breakers (5), vertical trunkings (6), horizontal trunkings (7) and electrical cables. Fig. 2 also points out a single pipe (8) with small holes located at the bottom of the fire cabinet. This last one, completed by a spark system (see the enlarged view in Fig. 6), was used as a propane burner which provided a fire power of about 10 kW for a duration between 4 and 5 minutes (depending on the test). For safety reasons, the electrical cabinet was not electrically powered.

| Overhead cable trays

Three horizontal ladder cable trays 5 m long were installed above the fire cabinet and deported beyond its front panel (Fig. 4). The first cable tray was located at 0.3 m from the fire cabinet top. Spacing between cable trays was of 30 cm and the upper cable tray was distanced from the ceiling of nearly 45 cm. The HFFR cable-type which filled the three overhead cable trays for CFS-6, CFS-7 and CORE-6 tests, is named cable A, and the HFR cable-type for CFS-5 test, is called cable B. Cables A and B are control cables found in NPPs and contain four and three main constituent parts, respectively (Fig. 7). Cable A contains Poly(ethylene/vinyl acetate) (EVA) and Polyethylene (PE) as polymeric materials and alumina trihydrate (ATH) as HFFR [START_REF] Meinier | Fire behavior of halogen-free flame retardant electrical cables with the cone calorimeter[END_REF] . Cable B contains plasticized PVC as polymeric material and other loads such as calcium carbonate (CaCO 3 ) [START_REF] Zavaleta | Multiple Horizontal Cable Tray Fire in Open Atmosphere[END_REF] . Chlorine included in the PVC material acts as a flame retardant for cable B. Other characteristics of cables A and B are provided in Table 2. Each cable tray was filled with 32 samples of cable A for CFS-6, CFS-7 and CORE-6 tests and 44 samples of cable B for CFS-5 test. The cable samples 5 m long were all packed loosely along the three trays for all the tests (Fig. 4).

| Adjacent cabinets and false floor (only for CORE-6 test)

Two closed-doors cabinets were placed against the side walls of the fire cabinet only for CORE-6 test (Fig. 5). These adjacent cabinets are cabinet modules identical to the twin modules of the fire cabinet. Each adjacent cabinet was equipped with two trunkings placed against the adjacent side wall (Fig. 8).

The trunkings 1 and 2, included in the left-hand adjacent cabinet (Fig. 8 (a) and (b)), contained samples of a HFR (cable C) and HFFR (cable D) cable-type, respectively. In the right-hand adjacent cabinet (Fig. 8 (c)), the trunking 3 included samples of an HFFR cable-type with 2 cable diameters, 37 and 12 mm for cables E and F, respectively. Finally, the trunking 4 only contained samples of cable F. Characteristics of cables C to F, used as targets in the adjacent cabinets for CORE-6 test, are reported in Table 2. A false floor was located below the fire cabinet and the two adjacent cabinets (Fig. 9 (a)). The central part of the false floor communicated with the above fire cabinet through four cable passages (Fig. 9 (b) and (c)).

| Set-up of the rooms (DIVA facility and ventilation)

The DIVA facility contains five rooms in total (Fig. 1). The ground floor of this facility first comprises three rooms (room 1 to room 3) of similar size arranged in a row separated by fire doors. Moreover, each room also gives access (through fire doors) to a fourth room which is a corridor (room 0 in the background of Fig. 1). Finally, a fifth room is located at the first floor (room 4). CFS-5 to CFS-7 and CORE-6 tests used the rooms 1 and 2 of the DIVA facility (Fig. 10). The fire cabinet and the three overhead cable trays were centred against the west wall of room 1 for the four tests. Room 1 was thus considered as the fire room (FR) and room 2 as the adjacent room (AR). These rooms were 4 m high and had a surface area of 30 m 2 and, they communicated through an open doorway which was 0.79 m wide and 2.17 m high.

The ventilation system of the DIVA facility consists of two separate circuits (one for inlet and one for outlet, respectively shown in blue and red in Fig. 1). For the present study, one inlet duct was implemented in the upper part of the FR and one outlet duct was set up in the upper part of the AR (Fig. 10). For all the tests, before ignition of the cabinet, initial VRR was 15 h -1 for the overall volume of the FR and AR (i.e., initial inlet and outlet flow rates of 3600 m 3 /h). The flow directions are shown in Fig. 10. Furthermore, for CFS-7 test, fire dampers were set up at the inlet and outlet of the FR and AR, respectively. The monitored shutdown of the fire dampers significantly reduced the VRR from 15 to about 1 h -1 .

Cable

| Instrumentation

This section presents the measurements carried out as part of CFS-5 to CFS-7 and CORE-6 tests and discussed in this study.

| Common measurements for the four tests

Eleven thermocouples, separated from each other of 50 cm, were located along each of the three cable trays (Fig. 11). These thermocouples were positioned just above the cables and thus measured either gas temperature or flame temperature when the cables were ignited. Furthermore, twenty-four specific thermocouples were brazed on the walls of the fire cabinet. Six thermocouples were implemented on each of the two side walls and twelve on the rear wall. Two electronic scales were located below the two supports of the three cable trays and thus measured their fuel mass loss (Fig. 3). For same measurement dealing with the cabinets, one electronic scale was positioned below the fire cabinet for CFS-5 to CFS-7 fire tests (Fig. 3) and two were used for the three cabinets involved for CORE-6 test (Fig. 12). Three gas probes, supported by a vertical tree in the south-east (SE) corner and located at 0.7, 2.2 and 3.3 m from the floor (Fig. 13), allowed the oxygen (O 2 ) concentration measurement in the lower, mid and upper parts of the FR, respectively. In addition, unburnt hydrocarbon gases (C n H m ) concentration was also measured in the upper part of the FR (Fig. 13).

| Specific measurements for CORE-6 test

Three thermocouples were located along each of the four trunkings at 50, 100 and 150 cm from the bottom of the adjacent cabinets.

| Measurement uncertainty

The relative uncertainty of a measurement, may be defined as followed:

(1)

Where is the uncertainty (or standard deviation) of the measurement and R its result (expressed in the corresponding unit). is estimated by combining the individual uncertainties, such as the instrument uncertainty and the repeatability, using the usual method called the "law of propagation of uncertainty" 15,[START_REF] Hamins | Verification and Validation of Selected Fire Models for Nuclear Power Plant Applications[END_REF] . Furthermore, the uncertainty is often expressed in terms of an expanded uncertainty, in which the confidence level that the measurement falls within the expanded bounds is high 15 . For an expansion factor (or coverage factor) of 2, considered in this study, the expanded relative uncertainty of the measurement, , is thus related to two relative standard deviations (i.e.,

) and the confidence level corresponds to 95%. Table 3 gives evaluated for the measurements which were carried out as part of CFS-5 to CFS-7 and CORE-6 fire tests and discussed in this paper.

| RESULTS AND DISCUSSION

This section first presents how the HRR is assessed for the four tests. Then, the second part focuses on the results of CFS-6 test (Table 1) which is considered as the reference test. The third part aims at comparing the results of CFS-7 test with those of the reference test for showing the effect of the fire dampers shutdown on fire spread. Next, the effects on fire spread of the cable-type installed along the overhead cable trays are highlighted from simile between the outcomes of CFS-5 test and those of the reference test. Finally, the results of CORE-6 test, regarding the cabinet fire spread to both adjacent cabinets and false floor, are given in detail. This last part also discusses the effect of the adjacent cabinets (CORE-6 vs CFS-6) on fire spread to overhead cable trays.

| Heat release rate assessment

The HRR (or ) assessment for CFS-5 to CFS-7 and CORE-6 tests in the confined and mechanicallyventilated DIVA facility was determined by the thermal method and carbon dioxide generation (CDG) calorimetry method.

The thermal method was developed by Coutin et al. [START_REF] Coutin | Energy balance in a confined fire compartment to assess the heat release rate of an electrical cabinet fire[END_REF] for assessing the HRR of a cabinet fire by establishing an energy balance inside a single room. This method may be extended to test configuration involving several rooms and to other complex fire sources such as the ones implemented for these four tests. In this case, the energy balance is carried out inside the FR and AR of the DIVA facility (Fig. 10). The heat released by the fire was mainly distributed in five main parts:

 heat transfers to steel structures of the electrical cabinet (thermal inertia of the three ladder cable trays is negligible compared with that of the cabinet) ( );  energy per unit of time required for pyrolysis of fuel contained in the electrical cabinet and the three cable trays ( );  heat transfers through the walls of the FR and AR ( );  energy variation per unit of time within the FR and AR ( );  convective heat transfers via the inlet and outlet of the FR and AR, respectively (

).

The fire HRR from the thermal method ( ) is therefore assessed according to:

(

Where:

(

With the cabinet mass (kg), the specific heat value representative of the metallic structures of the cabinet (475 J.kg -1 .K -1 ), the wall temperature of the metallic structures of the cabinet (K). This last one is obtained as the average of the twenty-four wall temperature measurements of the cabinet.

(4)

With

and the latent heat of vaporization of fuel contained in the cabinet and the three cable trays, respectively (both assumed equal to 2.4 MJ/kg), and the mass loss rate of the cabinet and the three upper cable trays, respectively (g/s).

(

) 5
With the measured heat fluxes per unit area through the walls of the FR and AR (kW/m 2 ) and the specific area (m²) associated with the heat flux sensor ''i''. For all the tests, thirty heat flux sensors were stuck over all area of the concrete walls inside the FR and additional thirty ones on walls of the AR.

(6) With the isentropic constant (equal to 1.4 for diatomic gases), V the overall volume for the FR and AR (240 m 3 ) and P the pressure in the FR and AR (Pa). (7) Assuming that the specific heat ( and molar mass of air ( remain constant ( ), the HRR through the opening i = inlet and outlet can be given by: (8) With the volume flow rate at the opening i (m 3 /s), the ideal gas constant (8.314 J.mol -1 .K -1 ) and the pressure at the opening i (Pa).

Alternatively, the fire HRR can also be evaluated from the CDG calorimetry method. For fires with large amount of CO and soot (considered as pure carbon), Tewarson [START_REF] Tewarson | Generation of Heat and Chemical Compound in Fires[END_REF] and Brohez et al. [START_REF] Brohez | Accurate Calculations of the Heat Release in Fires[END_REF] proposed to assess the HRR from the CDG calorimetry method ( ) as follows:

where is the energy produced by mass unit of the generated species j = CO 2 , CO or C (i.e., soot) (kJ/g) and is the mass flow rate of the generated species j = CO 2 , CO or C (g/s). Pretrel et al. [START_REF] Pretrel | Experimental determination of fire heat release rate with OC and CDG calorimetry for ventilated compartments fire scenario[END_REF] proposed to evaluate the fire HRR in multi-rooms facility from Eq. ( 2) and the mass conservation of CO 2 , CO and soot carried out in such facility. This evaluation applied for CFS-5 to CFS-7 and CORE-6 tests in the FR and AR (Fig. 10) may be thus written as follows:

(10)

where is the mass flow rate of the generated species j = CO 2 , CO or C at the outlet of AR (g/s) and and are the mass variation rates of the generated species j = CO 2 , CO or C, in the FR and AR, respectively (g/s). The average values of and are 13.3 and 11.1 MJ/kg, respectively, according to Brohez 20 , and that for is 15.8 MJ/kg according to Pretrel [START_REF] Pretrel | Experimental determination of fire heat release rate with OC and CDG calorimetry for ventilated compartments fire scenario[END_REF] . Fig. 14 shows for CFS-5 to CFS-7 and CORE-6 tests the HRR assessments from both CDG and thermal methods. These two methods give consistent results for all the tests except for CFS-7 test. Uncertainty of the HRR, assessed from CDG method, is acceptable when mass flow rates of the generated species at the outlet of the AR (i.e., terms in Eq. ( 10)) are larger than mass variation rates of the generated species in the FR and AR (i.e., and terms in Eq. ( 10)). Indeed, these last ones give significant fluctuations especially at early stage of the fire when the gas species are filling the rooms and mass flow rates of the species at the outlet are low. For CFS-7 test, the monitored fire dampers shutdown led to nearly zero mass flow rates of the species at the outlet of the AR. Consequently, the HRR evaluation from the CDG method for CFS-7 test gives too large fluctuations (Fig. 14 (c)) and only the thermal method is therefore used for the HRR assessment of this test. In contrast, for each of the three other CFS-5, CFS-6 and CORE-6 tests, final HRR was obtained from the average of the HRR evaluated from the CDG and thermal methods.

The relative uncertainty, , of the final HRR for CFS-5, CFS-6 and CORE-6 tests is estimated from the relative standard deviation of the two HRR assessments obtained from the CDG and thermal methods for each of these three tests (see Fig. 14). In same way as for the measurements, an expanded relative uncertainty, , is considered for these assessments and provided in Table 4. For the HRR of CFS-7 test, which was only evaluated from the thermal method, as previously detailed, is taken equal to the maximal value estimated for the previous tests (i.e., 0.3).

| Fire spread for CFS-6

CFS-6 test (reference test, Table 1) involved an initial VRR of 15 h -1 and used the HFFR cable A (Table 2) installed over the three overhead cable trays. Fig. 15 (a) exhibits, after CFS-6 test, the fire cabinet and three overhead cable trays and Fig. 15 (b) shows an enlarged view of the right-hand side of the cable trays, between the mid and upper cable trays. These figures point out that the fire fully involved the cabinet and also spread to the upper cable trays. Total mass loss (TML) was measured at 35.5 and 39 kg for the cabinet and overhead cable trays, respectively. Furthermore, Fig. 16 indicates that fire only spread on the right-hand sides of the three cable trays. Fig. 17 shows the measured temperatures along the three cable trays for CFS-6 test, just above the cables, the legends indicating the thermocouple positions from the tray centre (in cm). The fire propagation throughout the cabinet led to a first increase of these temperatures up to about 500°C (Fig. 17 showing that the cable trays were ignited. The presence of flame is indeed assumed for a minimal temperature of 600°C according to Drysdale [START_REF] Drysdale | An introduction to fire dynamics[END_REF] . However, given equal to 0.07 for the flame temperature measurement (Table 3), the ignition is certainly obtained when a temperature measurement indicates 645°C for the first time. Indeed, the true temperature, corresponding to this measurement, is ensured to be higher than 600°C since it ranges from 600°C (i.e., 645 -6450.07) to 690°C (i.e., 645 + 6450.07) with a confidence level of 95 % (section 2.5.3). Therefore, considering the 645°C threshold, fire is first ignited at the lower tray at 1237 s (Fig. 17 In the reverse way, it is assumed that cable tray fire is extinguished as soon as all temperatures along the three cable trays are lower than 600°C. However, given equal to 0.07, as previously considered, the extinction is thus certainly achieved as soon as all temperature measurements indicate values lower than 561°C. Indeed, the true temperature, corresponding to this measurement, is surely lower than 600°C since it varies from 522°C (i.e., 561 -5610.07) to 600°C (i.e., 561 + 5610.07) with a confidence level of 95 %. Considering this 561°C threshold, measurements indicate that the cable tray fire ended at 6073 s (Fig. 17 (b2)). This extinction may be corroborated by the other temperature measurements along the cable trays (Fig. 17 (b1), (b2) and (b3)) which are all lower than 400°C when t  6073 s and the MLR of the cable trays which is very nearly zero around 5000 s (Fig. 18). Finally, the fire duration is evaluated as the time difference between the first ignition (i.e., 1237 s) and the extinction time (i.e., 6073 s), which gives 4836 s (about 80 min). a3)). This atypical propagation might have been influenced by the important air volume flow rate (nominal value of 3600 m 3 /h), even if it was directed towards the opposite wall of the fire source (Fig. 10).

Fig. 18 presents the mass loss rate (MLR) for both the fire cabinet and three overhead cable trays as well as the HRR of CFS-6 fire test. The MLR is obtained from time derivation of the mass losses measured separately for the fire cabinet and the three cable trays as early commented (section 2.5.1). The expanded relative uncertainty evaluated for the MLR is 0.3 for MLR  10 g/s, 0.23 for 30 g/s  MLR > 10 g/s and 0.15 for MLR > 30 g/s.

The first and higher HRR peak of 735 kW at 560 s was only caused by the fire cabinet since the MLR for the cable trays was zero at the same time (Fig. 18). Ignition of the overhead cable trays, as above determined, occurred at 1237 s, i.e., 677 s after the cabinet fire peak, when the cable trays MLR reached a value of about 6 g/s (Fig. 18). Furthermore, the last HRR peak of 440 kW at 2400 s was mainly due to the cable tray fire. Indeed, at this time, the cabinet MLR was nearly zero while that of

Left-hand side

Right-hand side

(a) (b)

the cable trays reached its maximal value of 19 g/s (Fig. 18). Thus, it takes about 20 min (1175 s) for the cable tray fire to reach its peak. These outcomes show that the cable tray fire for CFS-6 test mainly grew at the end of the cabinet fire.

| Fire dampers shutdown effects (CFS-7 test vs CFS-6 test)

Both CFS-6 and CFS-7 tests (as for all the tests) used initial VRR of 15 h -1 . For CFS-7 test, the fire dampers were shut down at 2 min 30 s after the cabinet ignition, that led to decrease the VRR to about 1 h -1 . Fig. 19 (a) shows that the fire entirely involved the fire cabinet for CFS-7 test but, unlike CFS-6 test, it did not spread to the overhead cable trays (Fig. 19 (b)). This last figure indeed shows limited impact of the cabinet fire which was focused below the central part of the lower cable tray.

| Fire cabinet

Fig. 20 presents the HRR for CFS-6 and CFS-7 fire tests. The time when fire dampers are shut down is also indicated in this figure. Compared to CFS-6 test, Fig. 20 shows that the HRR of CFS-7 test fast decreased to nearly zero for t ≥ 12 min. Indeed, Fig. 21 first points out that oxygen (O 2 ) concentrations in all the FR were always higher than 14 %Vol for CFS-6 test while those measured in both mid and upper parts were about 5 %Vol around t = 12 min for CFS-7 test. Moreover, O 2 concentration in the lower part of the FR for CFS-7 test was lower than that for CFS-6 test as soon as t ≥ 12 min and then decreased fast down to 10 %Vol (Fig. 21). The fire dampers shutdown thus shortened the cabinet fire

(a) (b)

for CFS-7 test compared with CFS-6 test, due to such critical O 2 concentrations. For that reason, the cabinet TML for the former test (23.9 kg) was lower than for the latter test (35.5 kg). Furthermore, the fast decrease of the cabinet fire for t ≥ 12 min as earlier discussed (Fig. 20) also contributed to the steep MLR decrease of the cable trays located above. The fire dampers shutdown therefore prevented ignition of the overhead cable trays, due to very low O 2 concentration in the upper part of the FR. In contrast, a VRR of 15 h -1 , as used for CFS-6 test, led to suitable O 2 concentration for the ignition of the overhead cable trays.

| Overhead cable trays

(a) (b)

| Cable-type effects on fire spread (CFS-6 vs CFS-5)

The features of CFS-5 test (Table 1) were identical than those of CFS-6 test, except the cable-type installed along the three overhead cable trays. The HFR cable B (Table 2) was indeed used for the former test instead of the HFFR cable A involved in the latter test. Fig. 23 (a) and (b) exhibit the fire cabinet and an enlarged view of the right-hand side of the cable trays (between the mid and upper cable trays) after CFS-5 test, respectively. These figures show that the fire fully involved the cabinet and spread to the upper cable trays. TML was measured at 36.9 and 16.3 kg for the cabinet and the upper cable trays, respectively. Moreover, Fig. 24 indicates that fire spread more over the right-hand sides of the three cable trays. Fig. 25 shows for CFS-5 test the measured temperatures along the three overhead cable trays, just above the cables. Like for CFS-6 test, the fire propagation throughout the cabinet led to a first increase of these temperatures up to about 400 °C (Fig. 25 (b3)). Next, steep peaks of these temperatures up to around 700°C were measured (Fig. 25 (b1) to (b3)). It is assumed, as for CFS-6 test (section 3.2), that when a temperature measurement along a cable tray reaches the 645°C threshold for the first time, the ignition is certainly obtained. After each ignition, it is also considered that as soon as all the temperature measurements of a cable tray indicate values lower than 561°C then the extinction is very likely achieved. b2) and (b3), respectively). The durations of the three cable tray fire stages are deduced from the differences between times to ignition and extinction. This gives fire durations of around 1 min (23, 45 and 89 s for the first, second and third fire stages, respectively), showing three successive short cable tray fires for CFS-5 test. Fig. 26 points out the MLR of both the fire cabinet and overhead cable trays and the HRR of CFS-5 test. Like for CFS-6 test, the first HRR peak at 560 s (700 kW) was only caused by the cabinet fire since the MLR of the cable trays was zero at the same time. This figure next exhibits three successive HRR peaks which match with the three MLR peaks of the overhead cable trays, while the cabinet MLR was in the 15-25 g/s range. These observations corroborate the occurrence of three successive cable tray fires for CFS-5 test and also show that they overlapped with the cabinet fire. These outcomes also exhibit that these fires were fast since their peaks were reached at 860, 990 and 1150 s (Fig. 26) while their ignition occurred at 843, 953 and 1087 s as above determined.

Fig. 27 and Fig. 28 point out for CFS-5 and CFS-6 tests the MLR of the cabinet and the HRR of both the cabinet and overhead cable tray fires, respectively. Given similar MLR of the cabinet for the two tests (Fig. 27), it is thus deduced that the differences showed in the HRR (Fig. 28) confirm two specific cable tray fire behaviors depending on the cable-type as previously discussed; three fast and short CFS-5 cable tray fires (HFR cable B) which overlapped with the cabinet fire and a significantly slower and longer CFS-6 cable tray fire (HFFR cable-type) which mainly grew at the end of the cabinet fire. Clarifications regarding these distinct fire behaviors are discussed hereafter.

Fig. 29 provides for CFS-5 test the O 2 concentration in the upper part of the FR and the MLR for the overhead cable trays. The three successive CFS-5 cable tray fires were short (about 1 min) since the O 2 concentration was fairly low when they occurred (lower than 14 %Vol, Fig. 29), due to the cabinet fire and the fast growth of these fires (HFR cable B). These fast fires indeed led to quick decreases of the O 2 concentration below 13 %Vol (Fig. 29) and such concentration is supposed to extinguish them. In contrast, the CFS-6 cable tray fire was much longer (80 min) since it ignited later (than the CFS-5 cable trays) and grew when the O 2 concentration was higher than 16 %Vol (Fig. 22 (a)), caused by the less powerful cabinet fire and slower growth of this fire (HFFR cable A). Indeed, for this slow fire, and given the air renewal (VRR of 15 h -1 ), the O 2 concentration was still higher than 16 %Vol nearly up to its fire peak (Fig. 22 (a)), which allowed its development. Such differences of cable tray fire behavior, due to the cable-type, finally led to lower TML of the cable trays for CFS-5 test (16 kg) compared with CFS-6 test (39 kg).

| Fire spread for CORE-6

Compared with CFS-6 test, CORE-6 test also implemented two adjacent cabinets placed against the side walls of the fire cabinet (Fig. 5). These adjacent cabinets contained HFFR and HFR cable-types (Table 2) which were installed along PVC trunkings placed against the adjacent side wall (Fig. 8 and Fig. 31 (a)). The results of CORE-6 test highlighted that the fire fully spread within the fire cabinet (Fig. 30), to the two adjacent cabinets and false floor. However, unlike CFS-6 test, fire did not propagate to the three overhead cable trays. The TML was measured at 46.2 and 2.3 kg for the three cabinets and upper cable trays, respectively. Visual comparison between Fig. 8 (c) and Fig. 32 (b) first points out that the trunking 4 and its cable content were removed from their initial location. Indeed, this trunking burnt and its charred part, as well as the samples of cable F contained inside, fell in bottom of this cabinet. Fig. 32 (b) also shows that the trunking 3, which contained the samples of cables E and F, burnt. In contrast, all these samples and also those which fell in bottom of the cabinet (all of same HFFR cable-type) did not burn. Fig. 33 gives the measured temperatures along the trunkings 3 and 4. The results show a steep increase at 26 min 25 s of these measurements at 150 cm high from the bottom of the right-hand adjacent cabinet. At this time, temperatures over the right-hand side wall of the central cabinet (except at the bottom) exceeded 600°C (Fig. 34). Similar wall temperatures are thus assumed on the adjoined side wall of the right-hand adjacent cabinet where the PVC trunkings were fixed. Given that self-ignition of PVC polymeric material may occur when its temperature reaches 400 °C22-23 , it is thus highly probable that the fire that occurred in this cabinet at 26 min 25 s was caused by the self-ignition of PVC trunkings 3 and 4.

Fig. 35 shows the temperature measurements carried out along the trunkings 1 and 2 contained in the left-hand adjacent cabinet. The fast increase of the measurement along the trunking 1 at 100 cm high from the bottom of this cabinet shows that ignition occurred at 38 min 35 s after the CORE-6 test starting. Moreover, Fig. 32 (a) and Fig. 36 show respectively that the PVC trunkings 1 and 2 and the three samples of PVC cable C contained in the trunking 1 (Fig. 8 (a)) partially burnt. In contrast, the smaller samples of cables D (HFFR cable-type) contained in the trunking 2 did not burn.

Works carried out by Chavez 3,[START_REF] Chavez | An experimental investigation of internally ignited fires in nuclear power plant control cabinets, Part II: Room effects tests[END_REF] showed that an air gap of 2.54 cm may prevent fire spread from a cabinet fire to adjacent cabinet. In contrast, the study conducted by Mangs et al. 6 highlighted that fire may spread to an adjoined cabinet (i.e., without air gap) if samples of PVC cables are implemented in this cabinet. The results of the CORE-6 test are thus consistent with the latter work, but they also show that the samples of HFFR cable-types, located inside the PVC trunkings placed against the adjacent side wall of the adjoined cabinets, were not ignited. Fire spread from a fire cabinet to an adjacent cabinet not only depends on the air gap but also on the material type contained inside.

| False floor

Fire spread from the fire cabinet to the false floor (Fig. 9 (a)) via the cables passages (Fig. 9 (b) and (c)) and fully burnt all the cables contained in its central part (Fig. 37).

| Overhead cable trays

For CORE-6 test, temperature measurements carried out along the three overhead cable trays were lower than 600 °C, confirming that fire did not spread over these last ones, unlike CFS-6 test. These two tests were identical except the two adjacent cabinets only implemented for CORE-6 test. Fig. 39 exhibits the HRR of CFS-6 and CORE-6 tests. This figure shows that the cabinet fire growth was slower for CORE-6 test than for CFS-6 test. A plausible explanation of this experimental observation could come from a higher thermal inertia generated by the presence of the three adjoined cabinets in CORE-6 test. Therefore, the heat transferred to the overhead cable trays was lower for CORE-6 test compared with CFS-6 test. As a consequence, the measured temperatures along the cable trays in CORE-6 test were lower compared with those measured in the CFS-6 test (Fig. 40). Pyrolysis of the CFS-6 cable trays therefore started earlier than for CORE-6 test (Fig. 41) when temperatures just above the cables exceeded 400 °C (at 580 s, Fig. 40). This condition was reached later for CORE-6 test but only in the upper cable tray (1080 s, Fig. 40) when unburnt hydrocarbon gases (C n H m ), accumulated under the ceiling, were ignited in same time as the cabinet fire reached its peak (1080 s, Fig. 42). These two concomitant fire events led to decrease the O 2 concentration in the upper part of the FR below 12 %Vol (Fig. 43). Such O 2 concentration was too low for the cable trays ignition, which was highlighted at 16 %Vol for CFS-6 test (section 3.3.2). The higher HRR peak obtained for CORE-6 test (950 kW) compared with that of CFS-6 test (735 kW) was due to the overlapping of the cabinet fire with the C n H m combustion under the ceiling of the FR. The cabinet fire spread to overhead cable trays only for CFS-5 and CFS-6 tests which involved a VRR 15 h -1 . However, two specific cable tray fire behaviors depending on the cable-type were showed; three fast and short cable tray fires for CFS-5 test (HFR cable B) which overlapped with the cabinet fire and a significantly slower and longer cable tray fire for CFS-6 test (HFFR cable-type) which mainly grew at the end of the cabinet fire. These fires for CFS-5 test were short (about 1 min) since the O 2 concentration was fairly low when they occurred, due to the cabinet fire and the fast growth of these fires (HFR cable B). In contrast, the cable tray fire for CFS-6 test was much longer (80 min) since it ignited later (than the CFS-5 cable trays) and grew when the O 2 concentration was higher, caused by the less powerful cabinet fire and slower growth of this cable tray fire (HFFR cable A).

In contrast, cabinet fire did not spread to overhead cable trays for CFS-7 and CORE-6 tests. Indeed, the early fire dampers shutdown for CFS-7 test prevented ignition of overhead cable trays, due to very low oxygen concentration. Moreover, the slower growth of the CORE-6 cabinet fire compared with CFS-6 test, caused by the higher thermal inertia of the three adjoined cabinets, led to lower temperatures of the upper cable trays than the ones required for their ignition (obtained for CFS-6).

Finally, cabinet fire spread to the two adjacent cabinets and false floor for CORE-6 test. Indeed, trunkings, made in poly(vinyl chloride) (PVC), placed against the adjacent side walls and samples of HFR cables (made in PVC) contained inside, had self-ignited. In contrast, all samples of the various HFFR cable-types placed along the trunkings did not burn. These outcomes reveal that the fire spreading to an adjoined cabinet depends on the material-type installed inside. Finally, the long fire duration (about 40 min) in the central part of the false floor led to burn all the cables contained inside.

Fig. 1 :

 1 Fig. 1: Perspective view of the DIVA facility 11 .

Fig. 2 :

 2 Fig. 2: Open-doors electrical cabinet used for CFS-5 to CFS-7 and CORE-6 tests.

Fig. 3 :

 3 Fig. 3: Open-doors electrical cabinet and overhead cable trays for CFS-5 to CFS-7 tests.

Fig. 4 :

 4 Fig. 4: Overhead cable trays used for CFS-5 to CFS-7 and CORE-6 tests.

Fig. 5 :

 5 Fig. 5: Open-doors electrical cabinet, adjacent cabinets and overhead cable trays (CORE-6).

Fig. 6 :

 6 Fig. 6: Zoom on the propane pipe and spark system (metal electrodes) located at the bottom of the fire cabinet.

Fig. 7 :

 7 Fig. 7: Cables placed along the upper cable trays. (a) HFFR cable A for CFS-6, CFS-7 and CORE-6 tests. (b) HFR cable B for CFS-5 test.

Fig. 8 :

 8 Fig. 8: Cable content in PVC trunkings. (a) In left-hand adjacent cabinet: cables C and D. (b) Lefthand adjacent cabinet door-opened. (c) In right-hand adjacent cabinet: cables E and F.

Fig. 9 :

 9 Fig. 9: False floor. (a) Front view of the false floor. (b) and (c) Cable passages between the fire cabinet and the central part of the false floor, as indicated by the extremities of the black arrows.

  Fig. 10: Room 1 (FR) and room 2 (AR) of the DIVA facility used for CFS-5 to CFS-7 and CORE-6 tests. The white arrows indicate the flow directions at the inlet and outlet.

Fig. 11 :

 11 Fig. 11: Thermocouples located just above the cables (circles pointing out their extremity).

Fig. 12 :

 12 Fig. 12: Electronic scales used for CORE-6 test (two below the cabinets and two beneath the supports of the three cable trays).

Fig. 13 :

 13 Fig. 13: Gas probes in the South-East (SE) corner of the FR for measurement of O 2 and C n H m concentrations.

Fig. 14 :

 14 Fig. 14: HRR assessments from the thermal and CDG methods. (a) CFS-5 test. (b) CFS-6 test. (c) CFS-7 test. (d) CORE-6 test.

  (b3)). Next, higher temperatures up to around 800°C are measured (Fig.

17

 17

  (b1) to (b3))

  (b3)), then at the mid tray at 1729 s (Fig.17

  (b2)) and finally at the upper tray at 1878 s (Fig.17

  (b1)).

Fig. 15 :

 15 Fig. 15: After CFS-6 test. (a) Fire cabinet and the three overhead cable trays. (b) Enlarged view of the right-hand side of cable trays, between the mid and upper trays.

Fig. 16 :

 16 Fig. 16: Global view of the three overhead cable trays after CFS-6 test.

Fig. 17

 17 Fig.17also confirms that fire did not spread over the left-hand sides of the cable trays since measured temperatures along these last ones were clearly lower than 600°C (Fig.17(a1) to (a3)). This atypical propagation might have been influenced by the important air volume flow rate (nominal value of 3600 m 3 /h), even if it was directed towards the opposite wall of the fire source (Fig.10).Fig.18presents the mass loss rate (MLR) for both the fire cabinet and three overhead cable trays as well as the HRR of CFS-6 fire test. The MLR is obtained from time derivation of the mass losses measured separately for the fire cabinet and the three cable trays as early commented (section 2.5.1). The expanded relative uncertainty evaluated for the MLR is 0.3 for MLR  10 g/s, 0.23 for 30 g/s  MLR > 10 g/s and 0.15 for MLR > 30 g/s.The first and higher HRR peak of 735 kW at 560 s was only caused by the fire cabinet since the MLR for the cable trays was zero at the same time (Fig.18). Ignition of the overhead cable trays, as above determined, occurred at 1237 s, i.e., 677 s after the cabinet fire peak, when the cable trays MLR reached a value of about 6 g/s (Fig.18). Furthermore, the last HRR peak of 440 kW at 2400 s was mainly due to the cable tray fire. Indeed, at this time, the cabinet MLR was nearly zero while that of

17

 17 Fig.17also confirms that fire did not spread over the left-hand sides of the cable trays since measured temperatures along these last ones were clearly lower than 600°C (Fig.17(a1) to (a3)). This atypical propagation might have been influenced by the important air volume flow rate (nominal value of 3600 m 3 /h), even if it was directed towards the opposite wall of the fire source (Fig.10).Fig.18presents the mass loss rate (MLR) for both the fire cabinet and three overhead cable trays as well as the HRR of CFS-6 fire test. The MLR is obtained from time derivation of the mass losses measured separately for the fire cabinet and the three cable trays as early commented (section 2.5.1). The expanded relative uncertainty evaluated for the MLR is 0.3 for MLR  10 g/s, 0.23 for 30 g/s  MLR > 10 g/s and 0.15 for MLR > 30 g/s.The first and higher HRR peak of 735 kW at 560 s was only caused by the fire cabinet since the MLR for the cable trays was zero at the same time (Fig.18). Ignition of the overhead cable trays, as above determined, occurred at 1237 s, i.e., 677 s after the cabinet fire peak, when the cable trays MLR reached a value of about 6 g/s (Fig.18). Furthermore, the last HRR peak of 440 kW at 2400 s was mainly due to the cable tray fire. Indeed, at this time, the cabinet MLR was nearly zero while that of

  Fig.17also confirms that fire did not spread over the left-hand sides of the cable trays since measured temperatures along these last ones were clearly lower than 600°C (Fig.17(a1) to (a3)). This atypical propagation might have been influenced by the important air volume flow rate (nominal value of 3600 m 3 /h), even if it was directed towards the opposite wall of the fire source (Fig.10).Fig.18presents the mass loss rate (MLR) for both the fire cabinet and three overhead cable trays as well as the HRR of CFS-6 fire test. The MLR is obtained from time derivation of the mass losses measured separately for the fire cabinet and the three cable trays as early commented (section 2.5.1). The expanded relative uncertainty evaluated for the MLR is 0.3 for MLR  10 g/s, 0.23 for 30 g/s  MLR > 10 g/s and 0.15 for MLR > 30 g/s.The first and higher HRR peak of 735 kW at 560 s was only caused by the fire cabinet since the MLR for the cable trays was zero at the same time (Fig.18). Ignition of the overhead cable trays, as above determined, occurred at 1237 s, i.e., 677 s after the cabinet fire peak, when the cable trays MLR reached a value of about 6 g/s (Fig.18). Furthermore, the last HRR peak of 440 kW at 2400 s was mainly due to the cable tray fire. Indeed, at this time, the cabinet MLR was nearly zero while that of

Fig. 17 :Fig. 18 :

 1718 Fig. 17: Measured temperatures along the three overhead cable trays for CFS-6 test. (a1) Upper tray*. (a2) Mid tray*. (a3) Lower tray*. (b1) Upper tray**. (b2) Mid tray**. (b3) Lower tray.** *left-hand side, ** Centre and right-hand side.

Fig. 19 :

 19 Fig. 19: After CFS-7 test. (a) Fire cabinet. (b) Central part of the three cable trays above the fire cabinet.

Fig. 20 :

 20 Fig. 20: HRR for CFS-6 and CFS-7 tests. Fig. 21: O 2 concentration in the FR for CFS-6 and CFS-7 tests.

Fig. 22 (

 22 Fig. 22 (a) and (b) give both the O 2 concentration in the upper part of the FR and MLR of the cable trays for CFS-6 and CFS-7 tests, respectively. Vertical location of this O 2 concentration (3.3 m) is close to the height of the mid cable tray. The lower cable tray ignited at t = 1237 s for CFS-6 test for a MLR at 6 g/s and O 2 concentration in the upper part of the FR at 16 %Vol (Fig. 22 (a)). In contrast, Fig. 22 (b) indicates for CFS-7 test that for the MLR higher than 6 g/s the O 2 concentration in the upper part of the FR ranged from 5 to 6 %Vol. Such O 2 concentration did not allow the ignition of the CFS-7 cable trays and the MLR decreased to zero after its peak (9 g/s, Fig. 22 (b)).

Fig. 22 :

 22 Fig. 22: O 2 concentration in the upper part of the FR and the MLR for the overhead cable trays. (a) CFS-6 test. (b) CFS-7 test.

Fig. 23 :

 23 Fig. 23: After CFS-5 test. (a) Fire cabinet. (b) Enlarged view of the right-hand side of the cable trays, between the mid and upper trays.

Fig. 24 :

 24 Fig. 24: Global view of the three overhead cable trays after CFS-5 test.

  Fig.25shows for CFS-5 test the measured temperatures along the three overhead cable trays, just above the cables. Like for CFS-6 test, the fire propagation throughout the cabinet led to a first increase of these temperatures up to about 400 °C (Fig.25 (b3)). Next, steep peaks of these temperatures up to around 700°C were measured (Fig.25(b1) to (b3)). It is assumed, as for CFS-6 test (section 3.2), that when a temperature measurement along a cable tray reaches the 645°C threshold for the first time, the ignition is certainly obtained. After each ignition, it is also considered that as soon as all the temperature measurements of a cable tray indicate values lower than 561°C then the extinction is very likely achieved. Fig.25(b1) points out that the temperature measurements reach the 645°C threshold at 843, 953 and 1100 s. This figure also indicates that after each ignition, all temperature

Fig. 26 :

 26 Fig. 26: MLR of both the fire cabinet and overhead cable trays and the HRR of CFS-5 test.

Fig. 27 :

 27 Fig. 27: MLR of the fire cabinet for both CFS-6 and CFS-5 tests.

Fig. 28 :

 28 Fig. 28: HRR of CFS-5 and CFS-6 tests.

Fig. 29 :

 29 Fig. 29: O 2 concentration in the upper part of the FR and the MLR for the overhead cable trays for CFS-5 test.

Fig. 30 :

 30 Fig. 30: Fire cabinet after CORE-6 test.

Fig. 31 :

 31 Fig. 31: Inside the right-hand adjacent cabinet. (a) At the starting of the CORE-6 test. (b) Ignition at 26 min 20 s. (c) and (d) 1 and 2 s after ignition, respectively.

  Fire was first observed at 26 min 20 s after the cabinet ignition in the right-hand adjacent cabinet (Fig. 31 (b), (c), and (d)). Fire was next identified at 38 min 35 s in the left-hand adjacent cabinet from a temperature measurement carried out along the PVC trunking 1 as it will be discussed hereafter. Fig. 32 (a) and (b) show the upper part of the left-and right-hand adjacent cabinets, respectively, after the CORE-6 test.

Fig. 32 :Fig. 33 :

 3233 Fig. 32: (a) Upper part of the left-hand adjacent cabinet (b) Upper part of the right-hand adjacent cabinet.Fig. 33: Measured temperatures along the PVC trunkings 3 and 4 of the right-hand adjacent cabinet.

Fig. 34 :

 34 Fig. 34: Wall temperatures over the right-hand side wall of the central cabinet.

Fig. 35 :

 35 Fig. 35: Measured temperatures along the PVC trunkings 1 and 2 of the left-hand adjacent cabinet.

  Fig. 38 provides the temperature measurement located in the central part of the false floor. According to this measurement and the minimal temperature criterion of 600 °C, assumed revealing the presence of flame, it is therefore deduced that fire spread to the false floor at 61 min 50 s and was extinguished at 102 min 22 s. The long fire duration (about 40 min) contributed to consume the total amount of cable fuel mass contained in the central part of the false floor.

Fig. 36 :

 36 Fig. 36: Samples of cable C (HFR cable-type) contained in the trunking 1 (left-hand adjacent cabinet) after CORE-6 test.

Fig. 37 :

 37 Fig. 37: Central part of the false floor. Arrows showing cable conductors coming from the fire cabinet through the cable passages.

Fig. 38 :

 38 Fig. 38: Measured temperature in the central part of the false floor.

Fig. 39 :

 39 Fig. 39: HRR of the CFS-6 and CORE-6 tests.

Fig. 40 :

 40 Fig. 40: Measured temperatures at the centre of the lower tray for CFS-6 test and at the centre of the three trays for CORE-6 test.

Fig. 41 :

 41 Fig. 41: MLR of the CFS-6 and CORE-6 tests.

Fig. 42 :

 42 Fig. 42: HRR and C n H m concentration in the upper part of the FR (CORE-6 test).

Fig. 43 :

 43 Fig. 43: MLR of the cable trays and O 2 concentration in the upper part of the FR (SE corner) for CORE-6 test.

  Heat release rate assessed from the CDG calorimetry method (kW) Heat transfers to steel structures of the electrical cabinet (kW) Energy variation per unit of time within the FR and AR (kW) Heat release rate assessed from the thermal method (kW) Convective heat transfers via the inlet and outlet of the FR and AR, respectively (kW) Energy per unit of time required for pyrolysis of fuel (kW) Heat transfers through the walls of the FR and AR (kW) Heat fluxes measured per unit area through the walls of the FR and AR (kW/m 2 )

	. Mass flow rate of the generated species j = CO 2 , CO or C (g/s) Adjacent room Alumina trihydrate Carbon dioxide generation Ethylene-vinyl acetate Fire room Halogenated flame retardant Halogen-free flame retardant Polyethylene Poly(vinyl chloride) Carbon dioxide generation Soot (considered as pure carbon) Carbon monoxide Carbon dioxide Heat release rate (kW) Latent heat of vaporization of fuel (MJ/kg) Molar mass of air (kg/mol) Mass loss rate (g/s) Cabinet mass (kg) Mass flow rate of the generated species j = CO 2 , CO or C at the outlet of the AR (g/s) Oxygen Pressure in the FR and AR (Pa) Pressure at the opening i (Pa) Heat release rate (kW) Result of the measurement or assessment Ideal gas constant (8.314 J.mol -1 .K -1 ) Room 1 Room 2 South-East Average wall temperature of the metallic structures of the cabinet (K) Total mass loss (kg) Ventilation renewal rate (h -1 ) Uncertainty of the measurement or assessment Relative uncertainty of the measurement or assessment Expanded uncertainty of the measurement or assessment Expanded relative uncertainty of the measurement or assessment C n H Energy produced by mass unit of the generated species j = CO 2 , CO or C (MJ/kg) AR ATH CDG EVA FR HFR HFFR PE PVC CDG C CO CO 2 HRR MLR O 2 P R R g R1 R2 SE TML VRR V Overall volume of the FR and AR (240 m 3 )

m Unburnt hydrocarbon gases Specific heat value of the metallic structures of the cabinet (475 J.kg -1 .K -1 ) Mass variation rate of the generated species j = CO 2 , CO or C in the AR or the FR (g/s)

Table 1 :

 1 Main features of CFS-5 to CFS-7 and CORE-6 tests.

Test ID Open-doors cabinet (fire cabinet) Cable-type in overhead cable trays Adjacent cabinet Fire dampers

	CFS-5	Yes	HFR	No	No
	CFS-6	Yes	HFFR	No	No
	CFS-7	Yes	HFFR	No	Yes*
	CORE-6	Yes	HFFR	Yes	No

*Monitored shutdown of the fire dampers at 2 min 30 s after the cabinet ignition. Abbreviations: CFS, cable fire spreading; CORE, completing and repeating; HFR, halogenated flame retardant; HFFR, halogen-free flame retardant.

Table 2 :

 2 Electrical cables used as targets in CFS-5 to CFS-7 and CORE-6 tests.

	Cable

ID Location Flame retardant Outer diameter (mm) Linear mass density (kg/m) Supplier reference

	Cable A Overhead cable trays	ATH (Halogen-free)	20	540	NU-SHX(ST)HX 1kV 12x1.5 mm 2
	Cable B	(all tests)	Chlorine	14.5	330	SHCVV 8x2 mm 2
	Cable C Cable D	In adjacent cabinets	Chlorine Halogen-free	28 3	2000 21	NYM-J 5x25 mm 2 RM GRAU VARPREN ST 1.5 mm 2
	Cable E	(CORE-6)	ATH	37	3670	CST 74C068 3x95 mm 2
	Cable F		(Halogen-free)	12	250	CST 74C068 3x2.5 mm 2

Abbreviations: ATH, alumina trihydrate.

Table 3 :

 3 Expanded relative uncertainty, , of the measurements discussed in this study.

	Measurement (unit)	Sensor type	(-)
	Flame temperature (°C)	1.5 mm K-type thermocouple	0.07
	Gas temperature in the FR (°C)	1.5 mm K-type thermocouple	0.05
	Gas temperature in the cabinets (°C)	1.5 mm K-type thermocouple	0.02
	Wall temperature of the fire cabinet (°C)	0.5 mm K-type thermocouple	0.08
	Oxygen concentration (%Vol)	Oxygen analyzer	0.02
	Unburnt hydrocarbon gases	Unburnt hydrocarbon gases	0.1
	concentration (%Vol)	analyzer
	Mass (kg)	Electronic scale	< 0.01

Table 4 :

 4 Expanded relative uncertainty, , of the HRR assessed for CFS-5 to CFS-7 and CORE-6 tests.

	Test ID	(-)
	CFS-5	0.2
	CFS-6	0.3 for HRR < 100 kW and 0.2 for HRR  100 kW
	CFS-7	0.3
	CORE-6	0.15

ACKNOWLEDGEMENTS

This work was performed by IRSN in the framework of the OECD PRISME-2 project with the support of the partners: Bel V and Tractebel (Belgium), CNSC (Canada), VTT (Finland), DGA and EDF (France), GRS (Germany), NRA and CRIEPI (Japan), CSN (Spain), SSM (Sweden) and ONR, an agency of HSE (United Kingdom).


Pascal Zavaleta

Olivier Bouygues

Céline Lapuerta

IGNITION OF ELECTRICAL COMPONENTS CONTAINED IN A CLOSED ENCLOSURE ADJACENT TO A CONTROLLED FIRE SOURCE

INTRODUCTION

Nearly four hundred fire events in Nuclear Power Plants (NPPs) were recorded since the 1980's to the end of 2010 in the current OECD FIRE Database [START_REF]OECD FIRE Database[END_REF] . Almost fifty fire events involved electrical cabinets [START_REF] Werner | Enhancements in the OECD fire database -Fire frequencies and severity of events[END_REF] which are commonly arranged in rows of adjacent cabinets. It is also argued [START_REF] Werner | Enhancements in the OECD fire database -Fire frequencies and severity of events[END_REF] that a cabinet fire in a control or switchgear room of NPPs could lead to failure of one safety train or more, if fire spreads beyond the fire cabinet. So the ability of a cabinet fire to spread to adjacent cabinets is a major concern for fire safety in NPPs.

Few studies investigated the fire spread from one cabinet to adjacent ones. Mangs et al. [START_REF] Mangs | Calorimetric fire experiments on electronic cabinets[END_REF] conducted nine full-scale fire experiments which involved two types of vertical closed-doors electronic cabinets. The electronic cabinets were mainly equipped either with relays or circuit boards and they also all contained electric cables with poly(vinyl chloride) (PVC) jackets. The electronic cabinets, called the fire cabinets, were ignited in all experiments with a propane gas burner. A mockup cabinet made of thin steel sheets (0.5 mm thick) was attached to the fire cabinet (ie, there was no gap between the cabinets) to study the response of an adjoining cabinet to the fire. For four fire experiments, the adjacent cabinet contained samples of five PVC cables 400 mm long attached to the wall against the fire cabinet, and also placed at 5 and 30 cm from this wall. For two of these experiments, which led to a flashover in the fire cabinet, the PVC cables contained inside the adjacent cabinet ignited. So the spread of fire to an adjacent and fastened cabinet is possible because of heating of the separating walls.

Furthermore, Zavaleta et al. [START_REF] Zavaleta | Fire spread from an open-doors electrical cabinet to neighboring targets in a confined and mechanically ventilated facility[END_REF] studied the fire spread from an open-door twin modules 400 V electro-technical cabinet to two adjacent closed-door cabinets. The central open-door cabinet, which can be found in French nuclear installations, was fully and previously characterised by Coutin [START_REF] Coutin | Phenomenological description of actual electrical cabinet fires in a free atmosphere[END_REF] and Coutin et al. 6 . The main components inside the central cabinet were transformers, terminal blocks, motor circuit breakers and contactors, relays, circuit breakers, vertical trunkings, horizontal trunkings and electric cables. A propane burner located at the bottom of the central cabinet was used to ignite it. This burner provided a fire power of about 10 kW for a duration of about 5 minutes. For one fire test, two closed-door electrical cabinets were placed against the side walls of the central open-door cabinet without no gap between the cabinets. These adjacent cabinets were cabinet modules identical to the twin modules of the central cabinet and their side walls were 1.5 mm thick. Each adjacent cabinet was equipped with targets composed of PVC trunkings placed against the adjacent side wall and that contained either PVC or halogen free flame retardant (HFFR) cable samples. This fire test showed that the targets composed of PVC trunkings filled with the PVC electric cables burnt while the ignition of the HFFR electric cables was not highlighted. So the spread of fire to a fastened cabinet may depend on the target type contained inside.

Therefore, for completing the previous studies, investigations were carried out for providing new insights on the ignition of electrical components that can potentially occur in an electrical cabinet adjacent to another burning cabinet. This work especially aims at studying the impact on the ignition conditions of the electrical component-type (bundles of electric cables or sets of terminal blocks) contained in the adjacent electrical cabinet and the air gap between the cabinets (in the 1-5 cm range). This study also discusses ignition criteria which are assessed for each electrical component-type. Furthermore, the effects of the ventilation mode of the adjacent cabinet on the ignition conditions are also addressed in this work. Electrical cabinets in nuclear facilities [START_REF] Salley | Refining And Characterizing Heat Release Rates From Electrical Enclosures During Fire (RACHELLE-FIRE)[END_REF][START_REF] Mcgrattan | Heat Release Rates of Electrical Enclosure Fires (HELEN-FIRE)[END_REF] may be indeed either not ventilated or ventilated (naturally or mechanically).

For that purpose, a reduced-scale test device of adjacent electrical cabinets 4 was designed. This test device is composed of two adjacent steel enclosures separated by an air gap. The first enclosure, that contains a gas burner at its bottom, corresponds to a one-quarter scale model of the open-door electrical cabinet studied in previous works [START_REF] Zavaleta | Fire spread from an open-doors electrical cabinet to neighboring targets in a confined and mechanically ventilated facility[END_REF][START_REF] Coutin | Phenomenological description of actual electrical cabinet fires in a free atmosphere[END_REF][6] . The second enclosure, that contains electrical components, considered as targets in this study, reproduces at reduced scale the adjacent cabinets [START_REF] Zavaleta | Fire spread from an open-doors electrical cabinet to neighboring targets in a confined and mechanically ventilated facility[END_REF] .

The first part of this work presents the experimental matrix, the reduced-scale experimental setup, the related instrumentation and the test protocol. The second part of this study first deals with the test repeatability and also aims at assessing the measurement uncertainties. Next, this part, based on a result analysis, discusses the ignition mechanisms and proposes ignition criteria which are assessed for the studied target-types. This part finally presents the effects on the ignition conditions of the air gap and the ventilation mode as implemented in the adjacent enclosure.

FIRE TEST DESCRIPTION

Eleven fire tests were carried out to study the impact on the ignition conditions of the air gap between the two adjacent enclosures (in the 1-5 cm range), the target-type (sets of terminal blocks [TB] or bundles composed of either PVC or HFFR electric cables) and the adjacent enclosure ventilation mode (without ventilation, natural or mechanical), as summarized in Table 1. Furthermore, the setups of T3 and T7 tests are identical to those of the T2 and T6 tests, respectively. The outcomes of these tests will be further used to evaluate the test repeatability.

Experimental setup

The test device (Fig. 1) is composed of two adjacent steel enclosures 0.6 m wide, 0.6 m deep, 1 m high and separated by an air gap (Fig. 2). The design of this test device was based on a scale reduction of adjacent electrical cabinets studied by Zavaleta et al [START_REF] Zavaleta | Fire spread from an open-doors electrical cabinet to neighboring targets in a confined and mechanically ventilated facility[END_REF] . This cabinet set-up (Fig. 3) was composed of a twin modules 400 V electro-technical open-door cabinet, used as the fire cabinet, and two adjacent closed-door cabinets. These last ones, placed against the side walls of the central cabinet, were cabinet modules identical to the twin modules of the central open-door cabinet. The design of the reducedscale test device followed the successive steps:

 provided the symmetry plane of the real-scale cabinet set-up (Fig. 3), only its half part was considered (ie, one module of the open-door central cabinet and one adjacent closed-door cabinet module),  then, a height reduction of a factor two was applied for the two modules.

The left-hand enclosure (Fig. 1), named the fire enclosure (FE), therefore corresponds to onequarter scale model of the central open-door electrical cabinet [START_REF] Zavaleta | Fire spread from an open-doors electrical cabinet to neighboring targets in a confined and mechanically ventilated facility[END_REF] (1.2 m wide, 0.6 m deep and 2 m high). An insulated panel was placed against the internal left-hand side wall of the FE to simulate the symmetry plane. Furthermore, the front panel of the FE was removed (as for the open-door electrical cabinet) and a gravel-packed gas burner 0.5 m x 0.5 m was located at its bottom. This last one provided a constant fire power of 300 kW for all the tests, which corresponds to one fourth of the maximum heat release rate of the studied open-door cabinet [START_REF] Zavaleta | Fire spread from an open-doors electrical cabinet to neighboring targets in a confined and mechanically ventilated facility[END_REF] characterized in open atmosphere 6 . This thus allows to get the same fire power per unit volume for the reduced and real-scales. Furthermore, the work 4 also showed that electrical components contained in the adjacent cabinets ignited when the side wall temperature of the fire cabinet exceeded 600°C, as illustrated in Fig. 4. Numerical simulations were performed with the CALIF3S/ISIS fire field model 9 to confirm that the above design of the FE allowed to reach side wall temperature higher than 600°C. A schematic of the computational domain of dimensions 1.2 m x 1.8 m x 1.5 m is represented in Fig. 5. A symmetry condition was applied for x = 0 and the floor is considered as an adiabatic boundary. The red surface in Fig. 5 represents the gas diffusion burner. A multi-domain approach is used to compute both the flow and the temperature in the steel walls of the fire enclosure. The total mesh number is 31700 for the fluid domain and 100000 for the solid domain which is mainly represented by the FE surfaces. The gas burner was modelled by a fuel injection boundary condition using a mass loss rate determined from the ratio between the fixed fire power and the effective heat of combustion for propane (43.7 x 10 6 J.kg -1 ). In addition, wall-type boundary conditions were considered for the steel walls of the FE. The thickness of the steel walls of the fire enclosure was 0.003 m and the emissivity was fixed to 0.8. Fig. 6 shows the average temperature of the right side wall of the FE calculated for four power values of the gas burner (ie, 100, 200, 300 and 400 kW). The results point out that the wall temperatures calculated for the two higher fire powers are very similar and both exceeding 600°C. Numerical simulations thus confirm that the above criterion is obtained for a fire power of 300 kW.

The right-hand side enclosure is called the adjacent cabinet (AC). The adjacent walls of the FE and AE, named the fire wall (FW) and adjacent wall (AW), respectively, are shown in Fig. 2. The thicknesses of the FW and AW are 3 and 1.5 mm, respectively. The lower thickness was the same as for the side walls of the adjacent electrical cabinets [START_REF] Zavaleta | Fire spread from an open-doors electrical cabinet to neighboring targets in a confined and mechanically ventilated facility[END_REF] . A larger thickness was considered for all the walls of the FE for ensuring its resistance to thermal stresses as long as possible.

Electrical components used as targets

Terminal blocks (TB) and samples of either PVC or HFFR cables were used to compose the three target-types considered in this study:  three sets of forty TB each one (Fig. 7),  a bundle of ten cable samples 1 m long of the HFFR cable-type (Fig. 8),  a bundle of ten cable samples 1 m long of the PVC cable-type (Fig. 9).

Table 2 gives more specifications of the elements of these target-types which are electrical components commonly found close to the side walls of electrical cabinets present in nuclear power plants (NPPs) [START_REF] Zavaleta | Fire spread from an open-doors electrical cabinet to neighboring targets in a confined and mechanically ventilated facility[END_REF][START_REF] Salley | Refining And Characterizing Heat Release Rates From Electrical Enclosures During Fire (RACHELLE-FIRE)[END_REF][START_REF] Mcgrattan | Heat Release Rates of Electrical Enclosure Fires (HELEN-FIRE)[END_REF] . The targets were positioned close to the AW and supported by a grid fixed to the rear panel of the AE (Fig. 7 and Fig. 8). The distance of the targets from the AW was always lower than 1 cm. Therefore, the targets were either very nearly or in contact with the AW.

Left-hand adjacent cabinet

Right-hand adjacent cabinet

Open-door twin modules cabinet Fig. 7: Inside the adjacent enclosure (AW removed).

Ventilation of the adjacent enclosure

The front panel of the AE was equipped with two ventilation grids of unit surface of 0.03 m 2 (Fig. 1).

Gas burner

The studied ventilation mode was either without ventilation (ie, the grids were fully-blocked), natural (ie, the grids were half-blocked) or mechanical (ie, a fan was connected to the bottom grid and the two grids were fully unobstructed), as illustrated in Fig. 9. In the last case, the ventilation flow rate was set at 165 m 3 /h in the AE. This value is indeed about the half of the ventilation flow rate (300 m 3 /h) used in the adjacent electrical cabinets [START_REF] Zavaleta | Fire spread from an open-doors electrical cabinet to neighboring targets in a confined and mechanically ventilated facility[END_REF] (0.72 m 3 in volume) used to design the AE (0.36 m 3 ), as earlier discussed.

Instrumentation

Temperature measurements of the outer FW and outer AW were carried out using six 0.5 mm K-type thermocouples brazed on each of these walls (Fig. 10 and Fig. 11). Temperature measurements in the AE were performed thanks to twelve 1.5 mm K-type thermocouples positioned in the four corners of this enclosure at three heights (100, 500 and 900 mm from the bottom of the AE), as indicated in Fig. 12. Finally, the incident heat fluxes (IHF) in the AE were measured using two Gardon-type heat flux sensors positioned at about 10 mm from the AW (Fig. 7 and Fig. 8). Table 3 summarizes the above measurements and also fully specifies the location of the related sensors in the test device. The relative uncertainties for the instruments, are given in Table 4 * . These values of were deduced from calibration tests conducted for each sensor before its use for the fire tests presented in this work.

Test protocol

The fire tests were conducted under a large-scale calorimeter in open atmosphere conditions. The eleven fire tests, T1 to T11, were carried out with the same test protocol: the controlled fire with a power of 300 kW was maintained until few time after the ignition of the target in the AE. Fig. 13 illustrates the test protocol for example for the T1 test and Fig. 14 shows the FE and AE during the same fire test.

*

The uncertainty (or standard deviation) of the instrument for a given measurement ( , expressed in the same unit as the related measurement) can be deduced from the product of by the value of the measurement.

AW_X_Y

Temperature of the outer AW.

0.5 mm K-type TC X = 100 and 500 mm from the rear panel of the AE. Y = 100, 500 and 900 mm from the bottom of the AE (Fig. 11).

AE_1 to AE_12

Temperature in the AE.

1.5 mm K-type TC AE_1 to 4, AE_5 to 8, AE_9 to 12 located at 100, 500 and 900 mm from the bottom of the AE, respectively (Fig. 12). For each height, the TC were located in the four corners of the AE at about 5 cm from the panels. AE_1, 5 and 9 were placed in the targets.

HF_1 and HF_2

IHF in the AE.

Gardon-type heat flux sensor HF_1 and HF_2 located at 175 mm from the rear and front panel, respectively, at 250 mm from the ceiling and at about 10 mm from the AW (Fig. 7).

Abbreviations: AE, adjacent enclosure; AW, adjacent wall; FE, fire enclosure; FW, fire wall; HF, heat flux; IHF, Incident heat flux.

Fig. 13 especially shows the IHF in the AE measured close to the targets † , and the average temperatures of the outer FW and AW, and , respectively. These average temperatures are calculated as follows: † The heat flux measurements presented in this work were always carried out with the heat flux sensor HF_1 which was located closer to the target than the second one HF_2 (Fig. 7). Where Z = FW or AW, are the six temperature measurements of the outer FW (Z = FW, Fig. 10) or the six ones of the outer AW (Z = AW, Fig. 11), and X = 100 or 500 mm and Y = 100, 500 or 900 mm, are the related coordinates as specified in Table 3. The targets were thus submitted to an increasing IHF emitted by the AW which is heated by the FW itself in contact with the controlled fire. Finally, it can be noticed that is higher than 600°C for the T1 test (Fig. 13) as it was predicted by the 3D simulations for a fire power of 300 kW (Fig. 6). Such values for were also obtained for all the tests as further shown.

RESULTS AND DISCUSSION

The following section first comments test repeatability and also describes how are assessed the uncertainties of the measurements. Next, target ignition criteria are proposed and their variation according to the target-type is discussed. Finally, the impact of the air gap and the ventilation mode of the AE on the wall temperatures, heat flux in the AE and the ignition conditions are especially commented.

Test repeatability and measurement uncertainties

Fire enclosure

Adjacent enclosure

AW FW

The T3 and T7 tests repeated the T2 and T6 tests, respectively, as earlier commented (Table 1). Fig. 15 to Fig. 17 compare time evolutions of , and , respectively, for both the T2 and T3 tests. Furthermore, Fig. 18 to Fig. 20 show also the time evolutions of the three above parameters for both the T6 and T7 tests. The ignition time of the targets ( ) are also indicated in the above figures. All these figures highlight for the T2/T3 and T6/T7 tests a satisfactory test repeatability that can be assumed for all the T1 to T11 tests which followed the same test protocol. The relative uncertainty of a measurement, , may be estimated by combining the individual relative uncertainties, such as the relative instrument uncertainty ( , Table 4) and the relative uncertainty component for repeatability, ‡ , using the usual method called the "law of propagation of uncertainty" 10,[START_REF] Hamins | Verification and Validation of Selected Fire Models for Nuclear Power Plant Applications[END_REF] . is thus evaluated as follows:

(2)

Furthermore, the uncertainty is often expressed in terms of an expanded uncertainty, in which the confidence level that the measurement falls within the expanded bounds is high 10 . For an expansion factor (or coverage factor) of 2, considered in this study, the expanded relative uncertainty of the measurement, is thus related to two relative standard deviations (i.e.,

) and the confidence level corresponds to 95 %. Table 5 provides estimated for , and . First, is assessed as the maximal of the average of the relative deviation between measured for the T2 (

) and T3 ( ) tests and that measured for the T6 ( ) and T7 ( ) tests:

) ( 3
)
Where N and K are the measurement numbers § of / and / , respectively, considered in Eq. (3), is the starting time of the tests, and correspond with the earliest ignition time of the target between the T2 and T3 tests (ie, , Fig. 15) and that between the T6 and T7 tests (ie, , Fig. 18), respectively. and are then estimated in the same way as Finally, is evaluated according to Eq. ( 2) and the estimates of both ** (Table 4) and (Table 5). Table 5

also gives (ie,

) for , and .

Table 5 : Relative uncertainty component for repeatability ( ) and expanded relative uncertainty of the measurements ( ).

Measurement (unit) (-) (-)

Average

Ignition mechanisms and criterion

Fig. 21 shows, for instance for the T2 test, the inside of the AE two hundredth of seconds before the ignition of the target (bundle of ten PVC cables). The following events, successively illustrated in Fig. 22 to Fig. 24, were then observed:

(1) First at , a localized ignition in the central part of the target (Fig. 22) that leads to the sudden increase of the related temperature measurement (AE_5) located nearby (Fig. 25), ‡ While gives the (relative) uncertainty of an instrument in advantageous conditions during calibration tests, aims at quantifying the impact of the fire test conditions on the measurement uncertainty. § One measurement recorded per second. ** Given that and are evaluated from Eq. ( 1), the relative instrument uncertainty of these measurements can be evaluated as follows 10 : where = 0.025 for each temperature measurement of the outer FW and outer AW (0.5 mm K-type TC, table 4).

(2) A couple of hundredth of seconds later, the localized flame on the target ignited the accumulated flammable gases within all the AE volume (Fig. 23), (3) Few tenths of seconds later, fire rapidly spread along the target beyond the ignition zone (Fig. 24). In most tests, fire spread over a significant part or the totality of the target.

The video analysis of all the tests highlights similar ignition mechanisms. Only the localisation of ignition along the target can change according to the test. Furthermore, a question is arisen from the above description: is the ignition either piloted or spontaneous? Piloted ignition is generally achieved through localized heating such as an electric spark, a small flame or a heated wire, and the flame then propagates into the rest of the fuel material [START_REF] Babrauskas | Ignition Handbook[END_REF] . In contrast, spontaneous or self-ignition occurs as a result of raising the bulk temperature of a combustible gas mixture, and does not require any further external heat supply once combustion has started [START_REF] Kanury | Flaming Ignition of Solid Fuels[END_REF] . A spontaneous ignition requires a higher temperature for the same material than a pilot ignition. In the first stage of the ignition process, as above described, there was no piloted source. The AW heated the target that produced flammable pyrolysis gases and ignition occurred when the gas mixture temperature reached the self-ignition temperature (step (1) as above described). Then, the accumulated flammable gases ignited (step ( 2)) and the fire spread over the rest of the target (step (3)). ) accumulated within all the AE (eg, for the T2 test). AE, adjacent enclosure. The most commonly used ignition criterion is that related to the critical temperature [START_REF] Kanury | Flaming Ignition of Solid Fuels[END_REF][START_REF] Vermesi | Pyrolysis and ignition of a polymer by transient irradiation[END_REF][START_REF] Gong | Fire behaviors of flame-retardant cables part I: Decomposition, swelling and spontaneous ignition[END_REF] . These studies indeed mentioned that when the fuel reaches a critical surface temperature or critical ignition temperature, , then it can ignite. Table 6 reports for the T1 to T11 tests , and . The three last parameters are , and measured at , respectively. Given that the ignition can occur anywhere along the target which was either very nearly or in contact with the AW, is assumed to estimate correctly the target temperature leading to its ignition. Furthermore, , measured nearby the target (Fig. 7), is supposed close to the critical incident heat flux, , leading to the ignition of the targets. For each target-type, best estimations of and are thus calculated as the average of respective and measured for the related tests (Table 6):

Where j is the number of the considered tests for a given target-type (j = 4, 2 and 4 for the target-type with PVC cables ‡ ‡ , HFFR cables and TB, respectively), and are and , respectively, for the test k. Moreover,

) and ) can be evaluated from Eq. ( 4) and ( 5), respectively, as follows 10 :

(6) † † The AE_1, 5 and 9 temperature measurements were positioned in the target while the other ones were located outside the target, in the AE.

‡ ‡ For this target-type, only the T1 to T4 tests were taken into account for the evaluation of and . Indeed, the last test with this target-type (ie, the T5 test) involved a mechanically ventilated AE which, as further discussed, affected . In contrast, for the two other target-types, all the tests indicated in Table 6 were considered. (7) Where and (Table 5). Table 7 finally gives and for the three target-types. and are calculated from Eqs. ( 4) and ( 5), respectively, while and are assessed from Eqs. ( 6) and ( 7), respectively. These calculations considered j = 4, 2 and 4 for the target-type with PVC cables, HFFR cables and TB, respectively, as previously explained.

Effect of the target type

The evaluation for the target-type composed with the PVC cables (526 ± 21 °C, Table 7) falls in the 500-600 °C range provided by Gong et al. [START_REF] Gong | Fire behaviors of flame-retardant cables part I: Decomposition, swelling and spontaneous ignition[END_REF] that studied spontaneous ignition of similar cable type (ie, PVC external sheath and XLPE insulation). Furthermore, the piloted ignition temperature was measured at around 300°C [START_REF] Decimus | Study of gases released under incomplete combustion using PCFC-FTIR[END_REF] for electric cables with a PVC external sheath. Accordingly, these outcomes confirm that the target-type composed with the PVC cables spontaneously ignited.

for the two other targets (599 ± 34 °C for the bundle of 10 HFFR cables and 500 ± 20 °C for the sets of TB, Table 7) are also relevant with spontaneous ignition as indicated in the studies [START_REF] Babrauskas | Ignition Handbook[END_REF][START_REF] Kanury | Flaming Ignition of Solid Fuels[END_REF] . These studies indeed report that for a broad variety of natural and synthetic organics solids, falls in the 500-600 °C and 300-450 °C ranges for spontaneous and piloted ignitions, respectively. Moreover, piloted ignition temperature for materials made in polyamide was measured in fire calorimetry apparatus at around 300°C [START_REF] Santamaria | Experimental analysis of the pyrolysis of solids exposed to transient irradiation. Applications to ignition criteria[END_REF] . Furthermore, for similar HFFR electric cable-types as studied in this work (ie, cable-type containing EVA and PE as polymeric materials and ATH as HFFR), the piloted ignition temperature was estimated at about 400 °C18 .

Finally, Table 7 also shows that the values of both and measured for the targets composed with either TB or the PVC cables are similar while these ignition criteria obtained for the target made with the HFFR cables are clearly higher. These results imply that electrical components such as PVC cables or TB contained in a cabinet adjacent to a fire cabinet could ignite earlier than HFFR cables.

Effect of the air gap

Fig. 26 to Fig. 28 show, for the T1, T2, T8, T9 and T11 tests, the air gap effect on time evolutions of and , respectively. is also indicated in those figures. The T1 and T2 tests used the target-type composed with the PVC cables while the T9, T10 and T11 tests involved that made with TB. These figures first exhibit that the air gap increase slows down the raise of both and (Fig. 26 and Fig. 27, respectively). Accordingly, both and are reached later as the air gap increases, which delays . A larger air gap indeed leads to enhance the cooling effect of air flowing between the enclosures and thus delay the rise of both and caused by the controlled fire.

Effect of the ventilation mode

The effect on the ignition conditions of a natural ventilation mode in the AE, was studied using both the T2 and T4 tests. The T2 test involved a not ventilated AE (ie, the grids were fully-blocked) while for the T4 test the AE was naturally-ventilated (ie, the grids were half-blocked § § ), as specified in Table 1 and showed in Fig. 9. Fig. 29 to Fig. 32 show for these two tests , and the average gas temperature in the AE ( ), respectively. was assessed as the average of the twelve temperature measurements located in the AE (Table 3): (8) The above figures exhibit that the natural ventilation as implemented in the AE has no impact on the time evolutions of and (Fig. 29 to Fig. 31, respectively) and very small influence on time evolutions of (Fig. 32). The first outcome also corroborates the good test repeatability previously highlighted with the T2/T3 and T6/T7 tests. Furthermore, for the T2 and T4 tests are very close. This result is relevant with the fact that both and for these two tests are practically identical (Table 6). Furthermore, for the T9 and T10 tests which used identical setups except the natural ventilation mode of the AE implemented for the former test (the grids were halfblocked, Table 1), , and are quite similar (Table 6). This confirms that the natural ventilation mode as used in the AE has no clear impact on the ignition conditions.

Next, the effect of a mechanical ventilation mode in the AE was investigated with the T1 and T5 tests. Indeed, the T1 test involved a not ventilated AE (ie, the grids were half-blocked) while for the T5 test the AE was mechanically-ventilated. For that purpose, a fan was connected to the lower grid (Fig. 9) that imposed a ventilation flow rate set at 165 m 3 /h. Fig. 33 to Fig. 36 show for these two tests time evolutions of , and , respectively. The last figure points out that the mechanical ventilation of the AE slows down the increase of and therefore very likely that of the target temperature. Accordingly, the target ignition for the T5 test is delayed (Table 6). In contrast, the first three figures exhibit that time evolutions of and are very close up to the target ignition time for the T1 test (ie,

). Beyond this time, the three above parameters continue to grow, up to the ignition of the target for the T5 test, as shown in Fig. 33 to Fig. 35. § § The remaining opened grid area was 0.015 m 2 for each ventilation grid (Fig. 9). The impact of the mechanical ventilation mode is due to the significant but nevertheless realistic *** volume flow rate (165 m 3 /h) as used in the AE. In contrast, for the natural ventilation mode as implemented in the AE (the grids were half-blocked and no fan was used) the volume flow rate should be clearly lower.

CONCLUSION

This study was carried out for providing new insights on the ignition of electrical components that can occur in an electrical cabinet adjacent to a burning cabinet. This work aims at investigating the impact on the ignition conditions of the air gap between two electrical cabinets (in the 1-5 cm range), the electrical component-type (bundle of electric cables or sets of terminal blocks) contained in the adjacent cabinet and its ventilation mode. For that purpose, a test device of two adjacent steel enclosures was designed using a fire field model, in order to reproduce at reduced-scale adjacent electrical cabinets used in large-scale fire tests. Eleven fire tests using this test device were carried out. First, the test repeatability was shown satisfactory and the evaluation of the measurement uncertainties was fully detailed. This study next reveals that each electrical component-type (considered as target) spontaneously ignited in the adjacent enclosure when their temperature and the incident heat flux reach critical ignition values. These ignition criteria are assessed similar for the two target-types composed with either the terminal blocks (TB) or the poly(vinyl chloride) [PVC] cables and clearly lower than those estimated for the target-type made with the halogen-free flame retardant (HFFR) cables. These results imply that electrical components such as PVC cables or TB contained in a cabinet adjacent to a fire cabinet could ignite earlier than HFFR cables. The tests also show that the air gap increase slows down the rise of the side wall temperature of the two enclosures, due to the cooling effect of air flowing between the enclosures. Accordingly, the time to ignition of the targets is delayed. This work finally highlights that the mechanical ventilation of the adjacent enclosure cools down the internal gas volume and thus the targets. This extends the required time to reach the ignition conditions of the targets. In contrast, the natural ventilation as implemented in the adjacent enclosure has small impact on these conditions. The impact of the mechanical ventilation mode is due to the significant but nevertheless realistic volume flow rate as used in the AE. In contrast, for the natural ventilation mode as implemented in the AE the volume flow rate should be clearly lower.

Fig. 1 :

 1 Fig. 1: Fire and adjacent enclosures.Fig. 2: Air gap (g) between the two enclosures (eg, for the T11 test, g = 5 cm).

Fig. 2 :

 2 Fig. 1: Fire and adjacent enclosures.Fig. 2: Air gap (g) between the two enclosures (eg, for the T11 test, g = 5 cm).

Fig. 3 :Fig. 4 :Fig. 5 :Fig. 6 :

 3456 Fig. 3: The adjacent electrical cabinet set-up 4 .

Fig. 8 :

 8 Fig. 8: Test device before the side wall of the adjacent enclosure is installed.

Fig. 9 :

 9 Fig. 9: Ventilation mode of the AE. A, No ventilation. B, Natural ventilation. C, Mechanical ventilation. FE, fire enclosure; AE, adjacent enclosure; PVC, poly(vinyl chloride).

Fig. 10 :

 10 Fig. 10: Temperature measurements of the outer fire wall (FW_X_Y).

Fig. 11 :

 11 Fig. 11: Temperature measurements of the outer adjacent wall (AW_X_Y).

Fig. 12 :

 12 Fig. 12: Temperature measurements in the AE (AW removed). A, close to the rear panel. B, close to the front panel. C, inside the targets (eg, in the TB). The cross indicate approximately the TC location (dash-lined cross means that the TC is hidden by an equipment). TC, thermocouple.

Fig. 13 :Fig. 14 :

 1314 Fig. 13: Illustration of the test protocol (eg, for the T1 test). FW, fire wall; AW, adjacent wall; AE, adjacent enclosure. Fig. 14: The fire and adjacent enclosures during a fire test (eg, for the T1 test). AW, adjacent wall; FW, fire wall.

Fig. 15 :

 15 Fig. 15: Average temperature of the FW for the T2 and T3 tests. FW, fire wall.

Fig. 16 :

 16 Fig. 16: Average temperature of the AW for the T2 and T3 tests. AW, adjacent wall.

Fig. 17 :

 17 Fig. 17: IHF † in the AE for the T2 and T3 tests. IHF, incident heat flux; AE, adjacent enclosure.

Fig. 18 :

 18 Fig. 18: Average temperature of the FW for the T6 and T7 tests. FW, fire wall.

Fig. 19 :

 19 Fig. 19: Average temperature of the AW for the T6 and T7 tests. AW, adjacent wall.

Fig. 20 :

 20 Fig. 20: IHF in the AE for the T6 and T7 tests. IHF, incident heat flux; AE, adjacent enclosure.

Fig. 21 :Fig. 22 :Fig. 23 :

 212223 Fig. 21: Before ignition ( ) in the AE (eg, for the T2 test). AE, adjacent enclosure; HF, heat flux; PVC, poly(vinyle chloride); TC, thermocouple.

Fig. 24 :

 24 Fig. 24: Fire spread overall the target ( ) contained in the AE (eg, for the T2 test). AE, adjacent enclosure.

Fig. 26 :Fig. 28 :

 2628 Fig. 26: Average temperature of the FW. A, T1 and T2 tests. B, T8, T9 and T11 tests. FW, fire wall.

Fig. 29 :

 29 Fig. 29: Average temperature of the FW for the T2 and T4 tests. FW, fire wall.

Fig. 30 :

 30 Fig. 30: Average temperature of the AW for the T2 and T4 tests. AW, adjacent wall.

Fig. 31 :

 31 Fig. 31: IHF in the AE for the T2 and T4 tests. IHF, incident heat flux; AE, adjacent enclosure.

Fig. 32 :

 32 Fig. 32: Average gas temperature in the AE for the T2 and T4 tests. AE, adjacent enclosure.

Fig. 33 :

 33 Fig. 33: Average temperature of the FW for the T1 and T5 tests. FW, fire wall.

Fig. 34 :

 34 Fig. 34: Average temperature of the AW for the T1 and T5 tests. AW, adjacent wall.

Fig. 36 :

 36 Fig. 36: Average gas temperature in the AE for the T1 and T5 tests. AE, adjacent enclosure.

Table 1 :

 1 Fire test matrix

	Test ID	Target type	Air gap (cm)	Ventilation mode
	T1	Bundle of PVC cables	1	Without
	T2	Bundle of PVC cables	3	Without
	T3*	Bundle of PVC cables	3	Without
	T4	Bundle of PVC cables	3	Natural
	T5	Bundle of PVC cables	1	Mechanical
	T6	Bundle of HFFR cables	1	Without
	T7*	Bundle of HFFR cables	1	Without
	T8	Sets of TB	1	Without
	T9	Sets of TB	3	Without
	T10	Sets of TB	3	Natural
	T11	Sets of TB	5	Without
	* Repeatability test. Abbreviations: HFFR, halogen-free flame retardant; PVC,
	poly(vinyl chloride); TB, terminal block.

Table 2 :

 2 Specifications of the electrical components composing the targets.

	Electrical PVC electric cable	U1000R02V 3x2.5 mm 2	PVC for external sheath XLPE for insulation
	TB	TB 4/6 1SNA 115 116 R0700	Polyamide

component Supplier specification Chemical composition HFFR electric cable VARPREN ST 1.5 mm 2 Halogen Free EVA, PE and ATH as HFFR Abbreviations: ATH, Alumina trihydrate; EVA, Ethylene-vinyl acetate; HFFR, halogen-free flame retardant; PE, Polyethylene; PVC, poly(vinyl chloride); TB, terminal block; XLPE, Crosslinked polyethylene.

Table 4 :

 4 Relative instrument uncertainty ( .

Table 6 :

 6 Main test results obtained at the ignition time of the targets.

			IHF measured close	Average temperature of the outer walls
	Test ID	time, (s) Ignition	ignition, to the targets at	at ignition (°C)
			(kW/m 2 )	for AW	for FW
	T1	739	36	530	682
	T2	1109	36	520	656
	T3*	1285	34	531	642
	T4	1098	36	523	670
	T5	1121	38	560	662
	T6	1403	51	596	725
	T7*	1321	43	601	699
	T8	496	29	501	658
	T9	811	35	517	658
	T10	884	31	504	630
	T11	979	29	479	636
	* Repeatability test. Abbreviations: AW, Adjacent wall; FW, Fire wall; IHF, Incident heat flux.

Table 7 :

 7 Critical ignition temperature and heat flux according to the target-type.

	Target		Bundle of ten PVC cables	Bundle of ten HFFR cables	Three sets of forty TB
	Critical ignition temperature,	(°C)	526 ± 21	599 ± 34	500 ± 20
	Critical ignition heat flux,	(kW/m 2 )	35.5 ± 3	47 ± 5.5	31 ± 3

Right side wall of the FE