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Characterization of closed-doors electrical cabinet fires in compartments
Fire Safety Journal 46 (2011) 243–253
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Characterization of closed-doors electrical cabinet fires in compartments W. Plumecocq n, M. Coutin, S. Melis, L. Rigollet ˆ rete´ Nucle´aire (IRSN), Centre de Cadarache, BP no 3, 13115 St Paul-Lez-Durance Cedex, France Institut de Radioprotection et de Su
a r t i c l e i n f o
abstract
Article history: Received 7 May 2010 Received in revised form 26 January 2011 Accepted 28 February 2011 Available online 29 April 2011
An important cause of fire departure in industrial facilities is due to electrical origin and particularly to electrical cabinets. The investigation of such fires has been scarce up to now and has been investigated ˆ rete´ Nucle´aire (IRSN) exclusively in the nuclear industry. The Institut de Radioprotection et de Su conducted a large number of experiments involving electrical cabinets burning either under a calorimetric hood or inside a mechanically ventilated compartment to investigate this topic. Calorimetric hood experiments demonstrated that the most important parameter is the size of the vents of the cabinet and that the time to flashover depends on many factors and seems somewhat random with regard to the observable parameters. The influence of the compartment on the fire behavior depends on the temperature of the surrounding atmosphere of the cabinet and on the oxygen content in the compartment at the level of the inlet vent of the cabinet. The compartment strongly impacts the pyrolysis of the combustible, affecting the fire duration, but has a weak effect on the Heat Release Rate (HRR). Experiments were usually remarkably reproducible, opening the way to a phenomenological description of this type of fire. A semi-empirical model based on the coupled solution of ventilation limit and excess pyrolysate could then be developed. This model was introduced in a zone code, and an ad-hoc modeling of the fire extinction based on a critical surfacic mass loss rate is proposed. The major features of the compartment fire experiments such as characteristic HRR and fire duration could then be reproduced with acceptable error. The development of such a semi-empirical model is a common practice in fire safety engineering concerned with complex combustibles. & 2011 Elsevier Ltd. All rights reserved.
Keywords: Electrical cabinet Compartment Fires Fire modeling
1. Introduction An important cause of fire departure is reported [1] to be due to electrical origin. In industrial facilities, and, in particular, in nuclear facilities, a large part of this electrical fire hazard is attributed to electrical or electronic cabinets. These devices are usually composed of electrical components (i.e., wires, contactors and circuit breakers) enclosed in a metallic box in which several openings are located at the bottom and top of the door so that cooling of the circuits is ensured by natural convection. Handling of such fuels in fire safety applications is not obvious, however. The first reason for the difficulty encountered lies in the complexity of the fuel, a charring material composed of different polymers, whose composition is uncertain. Characterizing the combustion of electrical cabinets under a calorimetric hood as a function of its geometrical properties can, however, be attempted to cope with this problem [2]. A second difficulty arises when coming to the question of fire safety application in a compartment because the vitiation can in turn influence the combustion: the heat release rate (HRR) measured under a hood cannot be
n
Corresponding author. Tel.: þ33 442 199 657; fax: þ 33 442 199 167. E-mail address: [email protected] (W. Plumecocq).
0379-7112/$ - see front matter & 2011 Elsevier Ltd. All rights reserved. doi:10.1016/j.firesaf.2011.02.006
used directly as an input datum in the calculation methods. This question is the topic of the present study and is supported by a series of real-scale experiments involving electrical cabinets burning either under a calorimetric hood or inside a mechanically ventilated compartment. The facility and the tests are described. Main results are discussed. A semi-empirical model based on the coupled solution of the ventilation limit and excess pyrolysate is developed and applied to the IRSN experiments.
2. Literature review The investigation of electrical or electronic cabinet fires has been scarce up to now and concerned exclusively with the nuclear industry. Most of the data were published in internal reports. However, this topic is likely to be of some interest for many engineering applications. It seemed therefore useful to begin with a brief review of existing experiments and models. The first experiments were carried out at SNL [3,4] using various cabinet geometries, ventilation openings and ignition conditions; the fire load within the cabinet consisted of cables either passing the IEEE-383 fire qualification test or not. The conclusions of this experimental program showed the fire hazard posed by such electrical materials with a peak of heat release rate up to 175 kW
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Greek letters
Nomenclature Acronyms CAA CAB CAO PXF1 PXR1
a b
First series of CARMELA experiments Second series of CARMELA experiments Third series of CARMELA experiments PICSEL_F.1 PICSEL_R.1
g DHc
w e r s
Fire growth factor [W s 2] Empirical constant [W kg 2] Fraction of the fuel submitted to combustion [-] Heat of combustion [J kg 1] Energy fraction [-] Emissivity [-] Density [kg m 3] Stephan–Boltzmann constant [W m 2 K 4]
Symbols Subscripts C cp g h H k L _ m _ 00 m q Q_ t S T Y
Discharge coefficient [-] Specific heat [J kg 1 K 1] Gravity acceleration [m s 2] Convective heat exchange coefficient [W m 2 K 1] Distance between vents of the cabinet [m] Pressure loss coefficient [-] Latent heat [J kg 1] Mass loss rate [kg s 1] Surface mass loss rate [kg m 2 s 1] Mass flow rate [kg s 1] Heat release rate [W] Time [s] Section area [m2] Temperature [K] Mass fraction [-]
for closed-door cabinets and 955 kW for open-door cabinets. Even though the fire propagation to adjacent cabinets was judged unlikely, it was demonstrated that full involvement of the fire load within the cabinet could always be reached, depending on the ignition conditions. The influence of the closed-door electrical cabinet fires on the enclosure environment was also investigated in a 1400 m3 enclosure and revealed a large amount of smoke release along with a limited gas temperature increase up to 100 1C under the ceiling. Calorimetric experiments were conducted by VTT [2,5,6] using real electrical and electronic components as the fire load within closed-door cabinets. The maximum HRR in the range 100–400 kW was measured. The major results are reported in Table 1. The influence of the cabinet size and the surface of the vents was investigated. The size of the openings was shown to be of prime importance. Assuming that the ventilation within the cabinet is driven by buoyancy and that the fire is ventilationcontrolled, Mangs and Keski-Rahkonen [5] developed an expression (hereafter called the ‘‘MKR model’’) for the maximum HRR released by the cabinet sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffiffi 11=t Q_ max ¼ DHair Cout Sout r1 2gH 1=m þ t
ð1Þ
2 Here, m ¼ Sin Cin =Sout Cout represents the vent opening ratio depending solely on the vent geometry through their surface S and their discharge coefficients C. The temperature ratio t ¼Tg/TN is recognized to be a weak parameter in Eq. (1) for t 41.7. According to the authors, the MKR model represents a maximum for the HRR. To give a larger validation to the MKR model, IRSN [7] also conducted a large number of calorimetric experiments with different sizes of vents and using different materials. The main results are summarized in Table 1, and the major conclusions will be recalled in the next section for completeness.
air c cab f g i in max out r v xs w N
Air Combustible Cabinet Flame Gas Incubation Inlet Maximum Outlet Radiation Vaporization Excess pyrolysate Walls of the cabinet Ambient
Avidor et al. [8] also investigated the fire behavior of closeddoor electrical cabinets with the objective of building a fire hazard matrix. They used a propane gas burner and heptane pool as a surrogate of the electrical components within the cabinet. Considering the cabinet as a perfectly stirred reactor surrounded by a thermally thin steel wall, they could calculate the oxygen yield and temperature within the cabinet using a one-zone model.
3. Characterization of electrical cabinet fires under a hood Twenty-five experiments were carried out under a calorimetric hood to provide further validation of the MKR model [2]. The secondary objective was also to gain further knowledge concerning the temperatures reached by the cabinet walls because it is important to assess the possibility that the fire propagates to adjacent cabinets. 3.1. Experiments Cabinets were disposed in a large well-ventilated enclosure to minimize the influence of the environment. Above the cabinet, a 3 m diameter hood collected all the combustion products for filtering and analysis. The cabinets investigated in the IRSN tests were all of the same dimensions (2 m high, 0.8 m wide and 0.6 m deep). Two openings allowed for air admission in the box: the inlet one was located at the bottom of the front wall, and the outlet one was located either at the center of the roof (first series of experiments) or at the top of the back wall (second and third series). During the fire, the door of the cabinet was locked. The first series of tests (CAA series) investigated the geometrical parameters of the cabinets. Five parameters were selected: the ventilation effect (CAA1xx series), the ignition location (CAA201 test), the amount of fuel (CAA301 test), the spatial
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Table 1 Experimental configurations and major results. Vent areas (m2)
Experiment
Cabinet contents (Vol%)
Amount of contents (kg)
Total mass loss (kg)
HRRd (kW)
DHc
ti (s)
(MJ kg 1)
PMMA
PVC
PE
ECc
2.22 2.22 2.22 1.70 1.70 1.70 1.69 1.69 1.04 1.04
– – – – – – – – – –
– – – – – – – – – –
– – – – – – – – – –
100 100 100 100 100 100 100 100 100 100
66.5a 70.7 66.4 61.6 30.4 90.5 5.8 5.8 5.9 5.9
22.0 23.6 14.6 34.8 3.0 38.7 3.5 3.5 3.4 3.4
175 125 100 400 50 200 35 20 40 20
19.6b 11.4 19.7 18.8 8.9 18.3 16.2 11.7 13.9 11.7
– – – – – – – – – –
0.1000 0.1000 0.1000 0.1000 0.0500 0.0500 0.1000 0.1000 0.1000 0.1000 0.1000 0.1000 0.1000 0.1000
1.90 1.90 1.90 1.90 1.90 1.90 1.90 1.90 1.90 1.90 1.90 1.90 1.90 1.90
100 100 100 100 100 100 100 100 100 100 100 100 100 100
– – – – – – – – – – – – – –
– – – – – – – – – – – – – –
– – – – – – – – – – – – – –
10.0 10.0 10.0 10.0 10.0 10.0 10.0 5.0 10.0 10.0 10.0 10.0 10.0 10.0
– – – – – – – – – – – – – –
575 475 275 175 – 125 325 325 325 325 300 325 575 450
– – – – – – – – – – – – – –
425 293 387 234 305 170 458 299 272 268 226 426 285 107
0.0250 0.0250 0.0250 0.0250 0.0250 0.0250 0.0250 0.0250 0.0250
0.1000 0.0500 0.1000 0.1000 0.1000 0.1000 0.1000 0.1000 0.1000
1.70 1.70 1.70 1.70 1.70 1.70 1.70 1.70 1.70
100 100 100 100 100 50 29 29 37
– – – – – 50 71 42 31.5
– – – – – – – 29 31.5
– – – – – – – – –
10.0 10.0 10.0 10.0 10.0 10.0 10.0 10.0 10.0
10.0 10.0 9.7 10.0 9.9 9.3 1.6 0.6 9.7
250 200 225 240 225 200 50 25 225
23.6 23.7 20.5 20.2 23.3 14.6 16.3 15.6 20.6
245 119 183 236 115 294 141 – 294
CAO1 CAO2
0.0250 0.0250
0.1000 0.1000
1.70 1.70
– –
– –
– –
100 100
46.3 46.3
11.5 12.3
225 250
17.5 20.2
600 600
PXF1 PXR1
0.0250 0.0250
0.1000 0.1000
1.70 1.70
100 100
– –
– –
– –
10.0 10.0
5.1 6.8
250 250
10.8 10.8
152 156
Organism
Facility
Test name
Sin
Sout
VTT
Hood
1 2 3 4 5 6 7 8 9 10
0.0097 0.0097 0.0101 0.0403 0.0417 0.0418 0.0055 0.0029 0.0067 0.0034
0.0535 0.0535 0.0535 0.0792 0.0792 0.0792 0.0077 0.0077 0.0119 0.0119
IRSN
SATURNE hood
CAA102 CAA103 CAA104 CAA105 CAA106 CAA107 CAA201 CAA301 CAA401 CAA402 CAA403 CAA501 CAA502 CAA503
0.1000 0.0500 0.0250 0.0175 0.0250 0.0075 0.0250 0.0250 0.0250 0.0250 0.0250 0.0250 0.0250 0.0250
CAB101 CAB102 CAB201 CAB202 CAB203 CAB301 CAB302 CAB303 CAB304
DIVA Multi-rooms
H (m)
a
Estimated from the sum of the total burned mass and mass of contents after the experiment. Total energy release up to 58.7 min divided with corresponding mass loss. c EC: Components of electrical cabinets (see Refs. [2,7] for the nature and proportion). d Maximal HRR for the cabinets with EC contents or HRR corresponding to the plateau during the steady stage of the fire for the cabinets with PMMA, PVC and PE contents. b
arrangement in the cabinet (CAA4xx series) and the cabinet filling (CAA5xx series). A plate of polymethyl methacrylate (PMMA) was used to simulate the fuel: this material burns easily, and its properties are well-known. The plate was 1.5 m 0.7 m 8 mm (half plate for the test CAA301) and was positioned vertically at the center of the cabinet. In the tests CAA401, CAA402 and CAA403, two steel horizontal plates were located inside the cabinet, partially obstructing the flow and allowed a surface occupation rate between 30% and 90%. In the tests CAA501, CAA502 and CAA503, two hollow steel volumes were positioned against the lateral walls inside the cabinet allowing a volume rate of the cabinet between 25% and 75%. Additionally, eight tests were performed for repeatability. The second series of experiments (CAB series) investigated the influence of the fuel. One hundred and five squares of 10 cm 10 cm 8 mm each were mounted on steel shafts and positioned inside the cabinet to form the desired layout: a plate (squares disposed side by side to constitute a vertical plate), shelves, cable-tray (15 squares disposed as a vertical narrow band in front of a 90-squares plate) or dispersed (squares disposed
uniformly all over the volume of the cabinet). Polymethyl methacrylate, polyvinyl chloride (PVC) and polyethylene (PE) were used as combustible materials. Opening sections of the cabinet were 0.025 m2 at inlet and 0.1 m2 at outlet, except for test CAB102, where the latter was 0.05 m2. The third series of experiments (CAO series) consisted of two tests involving two identical relay cabinets. The inlet opening was 0.025 m2 and the outlet was 0.1 m2. The electrical components are circuit breakers, cable trunking, wires, terminal boxes, relays and contactors. The experiments consisted of a piloted ignition of the fuel material inside the cabinet. A linear propane burner generally located at the bottom of the fuel and in the middle of the cabinet provided the pilot flame. Its effective length was 0.7 m except for the test CAB202 (0.1 m) and the CAO series (0.6 m). The burner was shut off after ignition was confirmed (usually after two minutes or so), and the mean propane power output was about 20, 6 and 9 kW, respectively, for the CAA, CAB and CAO series. The fire was then allowed to develop freely until its natural extinction.
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The experiments were instrumented so as to allow a full interpretation of the combustion in terms of heat generation, heat exchanges and release of combustion products. In the exhaust duct downstream of the hood, all the combustion products were sampled and analyzed. The measurements included oxygen, carbon monoxide, carbon dioxide and water vapor concentrations. From these data, the oxygen consumption calorimetry [9] allows the determination of the heat released by the combustion. The HRR uncertainty was evaluated as 10–15%, the highest uncertainty coming from the measure of the flow rate inside the pipe. Thermocouples were placed inside the cabinets and on the outside walls. Radiometers indicated the radiative heat fluxes emitted by the cabinets. In addition, for the second and third series of experiments, an infrared camera gave indications of temperatures, radiant heat fluxes and emissivities on one side of the cabinet. A weighing machine was used to estimate the mass loss rate of the fuel. 3.2. Main results The goal of the analytical experiments of the first and second series was to facilitate the determination of the most influential parameters on the combustion of electrical cabinets. Three target values of interest were chosen: the maximum rate of heat release, the steady-state rate of heat release and the time to reach flashover (see Table 1). These experiments were usually remarkably reproducible. Five stages can be distinguished during the fire (see Fig. 1 for test CAA 104): The incubation stage. This stage starts with the spread of the flame along the fuel and the heating up of gases inside the cabinet. Gas temperatures inside the cabinet increase quickly, initiating a thermal pump, whereas the steel walls of the cabinet remain cold. CARMELA experiments [7] revealed extremely variable time durations of the incubation of the fire, between 107 and 600 s (see Table 1), because it is highly dependent on ignition conditions. The fast spread stage. At some point, the spread of the fire increases drastically, leading to a flashover phenomenon. Flames can be visible at the outlet vent of the cabinet, and increasing wall temperatures indicate that most of the contents of the cabinet are surrounded by hot temperatures. Radiative heat transfer inside the cabinet becomes predominant. This stage is highly oxygenconsuming. 800 Third stage: Combustion outside the cabinet
700
Heat release rate (kW)
600 500 Fourth stage: Steady stage
400 Second stage: Fast spread
300
Fifth stage: Decay D
First stage: Incubation period
200 100 0
0
250
500
1000 Time (s)
750
1250
Fig. 1. Characteristic evolution of the HRR during a test CAA.
1500
Combustion outside the cabinet. As the fuel involved in the combustion increases, the amount of combustible pyrolyzed exceeds the quantity that can react with the oxygen available from venting: the combustion becomes under-ventilated. In some experiments, the products of combustion leaving the outlet vent were hot enough and rich enough in unburned fuel vapors to reignite as they mix with fresh air; a flame then appeared outside the cabinet on the outlet ventilation opening. In the CAB and CAO experiments, no such flame was visible, but the mass loss measurements showed that the combustible was pyrolyzed in excess with regard to the oxygen available; the location of the outlet opening could have been responsible for some additional cooling of the combustion products, thus inhibiting the ignition outside the cabinet. Indeed, the outlet vent was in the ceiling for CAA experiments and at the top of the door for the CAB and CAO experiments. The steady-state stage. The high gas temperatures reached inside the cabinet lead to the relocation of molten materials at the bottom of the cabinet. This phase is closely linked to the combustion of molten materials at the bottom of the cabinet: equilibrium is established between the quantity of oxygen entering the cabinet and the quantity of fuel pyrolyzed. The extinction stage. The fire usually ends due to lack of combustible, with the exception of the test CAA106 where the fall of the plate caused the fire to be extinguished and the tests CAB302 and CAB303 where no ignition could be achieved. A long and smooth decay of the rate of heat released is observed. In the tests CAA501–CAA503, steel volumes were placed in the cabinet, and some influence was observed on the three target parameters in comparison with the reference test CAA104: higher rates of heat were released, and a faster flashover was observed. However this effect is attributed to an insulation effect of the walls and is not judged to be directly linked to the presence of dead volumes in the box. The tests CAB101 and CAB102 are compared, respectively, to the tests CAA104 and CAA106. In the tests CAB101 and CAB102, besides the difference in the outlet vent location, the relocation mode of the combustible was also different because the PMMA squares were allowed to melt and flow down independently from each other, which certainly points to the influence noticed in the peak heat release rate. To complement these experiments, normative ignition tests were performed under a cone calorimeter: the time to ignition of different plastic materials submitted to an imposed radiative flux was measured and proved to be highly variable. The tests CAB302 and CAB303 did not lead to flashover; obviously, there is a crucial effect of the nature of the combustible on the flashover phenomenon. The repeatability tests performed often presented very different times to flashover. One possible reason is the small difference in the ignition, but it can also be attributed to an intrinsic variability. On the contrary, the repeatability of the steady-state HRR, corresponding to the ventilation limit, was rather good (within 10%). The maximum heat release rate repeatability was better for the experiments of the third series involving real cabinets (10%) than for those of the first series (up to 33% for the test CAA503). There seems to be some relationship between the low repeatability and the presence of a large part of the flame burning outside the cabinet. As a conclusion, the analytical experiments demonstrate that the most influential parameter is the size of the openings. The time to flashover depends on many factors and seems somewhat random with regard to the observable parameters. 3.3. Modified MKR model The MKR model provides an estimate of the maximum heat release rate of the cabinets experiencing flashover. However,
W. Plumecocq et al. / Fire Safety Journal 46 (2011) 243–253
given the additional data brought by the experiments, some modifications were done to this model. 3.3.1. Cabinet air flow Firstly, because some of the vent areas studied were rather large compared to the cabinet cross-section, Eq. (1) had to be rewritten. The air flow transiting in the cabinet is due to a balance between the pressure drops caused by the vents and the buoyancy. Assuming rgTg ¼ rNTN, the Bernoulli equation yields q2 q2 q2 Tg 1 1 1 Tg T1 kin 2 in þ kout 2 in þ 1 2 in ¼ 1 r gH 2 Tg 1 Sin r1 2 Sout r1 T1 2 T1 Scab r1 ð2Þ The two first terms of the left-hand side of Eq. (2) represent the pressure drops at the inlet and outlet openings of the box; the corresponding coefficients kin and kout are commonly estimated to be 2.8 for turbulent vent flows [10]. Mangs and Keski-Rahkonen [5] measured kin ¼1.83 and kout ¼2.78; here, the coefficients were derived from the Weisbach equation [11] but only little variations in the range 2.6–2.9 were found. The third term of the left-hand side of Eq. (2) accounts for the acceleration of the fluid due to its heating; this term is neglected in the Mangs and Keski-Rahkonen model but is kept here for generality. The viscous wall friction pressure drops inside the box can be estimated to be less than 5% of the pressure drops due to vents. The right-hand side of Eq. (2) accounts for the buoyancy forces resulting from the density difference between the inside and the outside of the cabinet. As was noticed by Mangs et al. [2], the dependence of Eq. (2) on Tg is weak once Tg 4200 1C. Therefore, a representative value of Tg ¼Tf ¼800 1C is imposed, and one can estimate a steady-state mass flow rate sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffiffi 1ðT1 =Tf Þ q ¼ r1 2gH ðkin =S2in Þ þðTf =T1 Þðkout =S2out Þ þ½ðTf =T1 Þ1ð1=S2cab Þ ð3Þ n
Remarkably, q is representative of the flow inside the box as soon as the gas temperature rises at the outlet. In the experiments, this rise occurs very quickly, well before flashover; the gas burner used for ignition is enough to increase the outlet temperature to about 300 1C. 3.3.2. Energy balance of the cabinet An energy balance equation of the cabinet allows the determination of the gas and cabinet walls temperatures ! dTw Q_ c _ xs Lv MCw ¼ Q_ c q cpg ðTg T1 Þ þm dt DHc 4 4 Sw hw ðTw T1 ÞSw ew sðTw T1 Þ
ð4Þ
In Eq. (4), the left-hand term represents the thermal inertia of the cabinet, with MCw calculated from the total mass of the cabinet and the heat capacity of steel. The first term of the right-hand side represents the power released within the cabinet. The second term represents the heat convected out of the cabinet. The third term is the heat necessary for the pyrolysis of the combustible; it has been decomposed into two parts: the amount of pyrolysis necessary to produce the HRR within the cabinet and an excess pyrolysate that cannot burn inside the cabinet due to the ventilation limit. This last term is modeled in the next section. The fourth term is the heat exchanged by natural convection at the external walls of the cabinet: a classical value of hw ¼12 W m 2 K 1 is used, and the cabinet’s outer exchange surface Sw is calculated from the cabinet geometry. The fifth term is the energy radiated out by the cabinet walls: the value of the emissivity was measured as ew ¼0.7. Some of the terms of Eq. (4)
247
are determined according to further physical modeling: depending on the combustion mode (over- or under-ventilated) the gas temperature Tg is related either to the convected flux as in the open (see Eq. (5a)) or to the temperature of the cabinet Tw by assuming that a thermal equilibrium is achieved (Eq. (5b)). In over-ventilated conditions, convection is the dominant mode of heat transfer, and Tg is calculated as q cpg ðTg T1 Þ ¼ ð1wr ÞQ_ c
ð5aÞ
where wr is the radiated fraction of the combustible burning in the open conditions. In under-ventilated conditions, radiation is the dominant mode of heat transfer, and Tg is calculated as 4 esðTg4 Tw4 Þ ¼ hw ðTw T1 Þ þ ew sðTw4 T1 Þ
ð5bÞ
Eventually, a time-solution of Eq. (4) is provided by an implicit first-order solver. Because the system is represented by a single wall temperature whatever the elevation, the results are to be interpreted as only indicative of the temperature levels achieved. However, the two last terms of Eq. (4) provide an estimate of the thermal stress that a burning cabinet can have on its surroundings. 3.3.3. Pyrolysis rate and heat release The model for pyrolysis and heat release rate is largely based on the theory of a fully developed compartment fire. Two combustion regimes are distinguished: over-ventilated at first and then under-ventilated. The fire spread in the over-ventilated regime is divided in two growth stages. The combustion is modeled by two t-squared laws, which is a simple but practical way to represent complex fires [12]. During the first stage (incubation period, from ignition to time ti) Q_ c ðtÞ ¼ a1 t 2
ð6Þ
During the second stage (fast spread, after ti in the overventilated regime) Q_ c ðtÞ ¼ a2 ðtti Þ2 þ a1 t 2 ð7Þ The parameters a1 and ti were determined for each experiment as the best least-squares unbiased coefficients to fit the experimental data. Values of ti are reported in Table 1. These parameters were extremely variable, and no reproducibility could be achieved by repeating the experiments. However, an order of magnitude for ti of 100–450 s for PMMA tests and 600 s for real electrical components can be a compromise for these data. The growth factor a1 is even more uncertain, with an average of 1.3 W s 2. Some empirical correlation was established from the experiments
a2 ¼ bq2
ð8Þ 2
for the tests involving PMMA, and with b ¼1800 W kg b ¼ 100 W kg 2 for real cabinets is found to give best-fit for the present data. The maximum heat released in the cabinet is determined from the oxygen available, i.e., from qn. The heat released by the combustion of a fuel with a given amount of oxygen is known to be weakly dependent on the nature of the combustible, which leads to the formulation of the ventilation limit in term of power, just as in the original MKR model [5]. Q_ c ¼ q DHair
ð9Þ 1
where DHair ¼3.144 MJ kg is a constant valid over a large range of fuels at standard conditions of pressure and oxygen concentration. As appears in the experiments, the mass loss rate is not strictly correlated with the heat release rate, indicating a strong decrease in the effective heat of combustion during the flashover. To model
W. Plumecocq et al. / Fire Safety Journal 46 (2011) 243–253
this effect, the fire inside the cabinet is considered an underventilated, post-flashover enclosure [13]. Due to the ventilation limit, only a fraction g of the total fuel surface is surrounded by the flame and burns normally, the remaining being exposed to the hot gas temperature leading to its pyrolysis. It is further assumed that the pyrolysis is mainly due to radiation, and an optically thick medium is considered. The pyrolysis of the burning surface is then linked to the oxidant mass flux by stoichiometry 00
_ 1 gSf DHc ¼ gQ_ max q DHair ¼ m
ð10Þ
where Q_ max , the maximum heat release rate of all the combustible burning in free air, is introduced. The excess pyrolysate is obtained by linking the radiative heat fluxes in the flame zone to the amount of pyrolyzed combustible as follows: _ xs ¼ 1g Sf sðTg4 Tv4 Þ=Lv ð11Þ m
800
Characteristic power (kW)
248
Recognizing that Q_ max ¼ Sf sðTf4 Tv4 ÞDHc =Lv , Eqs. (9) and (10) can be recast as =DHC
600
400
200
0 0.00
ð12Þ
3.4. Discussion Similar to the MKR model, the model proposed is first based on the ventilation limit imposed by the geometry of the openings. The experiments carried out with different geometries provided a rather large basis for the validation of this concept. As an example, Fig. 2 compares HRR versus time for three experiments with similar cabinet geometries but different combustibles inside. Obviously, the characteristic HRR is well determined either by Eq. (1) or by the combination of Eqs. (3) and (9). However, the nature and the arrangement of the combustible within the
0.20
200
150
100 PMMA PMMA-PE-PVC Electrical components Agreement
50
0
300
200
100
0 1000 Time (s)
1500
100 150 200 50 Experimental maximum external heat flux (kW)
250
Fig. 4. Calculated versus measured maximum heat flux emitted by the cabinet’s wall to the environment.
CAO2 CAB202 CAB101 Model
400 Heat release rate (kW)
0.15
250
0 500
500
0.10
Fig. 3. Characteristic values of heat release rates for different fuels.
However, as was already noted, the occurrence of a flame outside the cabinet is thought to be specific to the tests of the first series involving PMMA. For the other tests, the excess pyrolysate is quenched as the flow impinges the ceiling of the cabinet and exits through the upper vent, so that little re-ignition was observed outside of the cabinet. However, this excess pyrolysate contributes to decreasing the fire duration without changing the heat release rate.
0
0.05
Intrinsic mass flow rate (kg/s)
Calculated maximum external heat flux (kW)
_ xs ¼ m
Tg4 Tv4 ðQ_ max q DHair Þ 4 4 Tf Tv
PMMA tests PMMA/PVC/PE tests Real Fuel tests VTT tests Theoretical
2000
Fig. 2. HRR of three tests with openings and model limits identical to those of Eq. (9).
cabinet influence the propagation before flashover profoundly and hence the time to flashover. Fig. 3 presents the experimental versus theoretical characteristic HRR at the steady state for all experiments: the deviations can be quite large, but the major part of the experiments collapses clearly against the theoretical line. The under-estimation of the HRR is mostly due to early extinction before flashover is encountered so that the model proposed should be taken as a worst-case calculation. The total (radiative and convective) heat flux from the cabinet walls could be calculated for the second and third series of experiments by integrating the measured heat fluxes over the surface of the cabinet. Because the temperature of the cabinet wall is continuously increasing, these fluxes reach a maximum just before the end of the fire. In addition to the original MKR model, the present calculations include the determination of the cabinet temperature (from the last two terms of Eq. (4)) and therefore provide an estimation of the heat fluxes to the ambient. Fig. 4 compares their maximum values to the calculated ones for the experiments where the measurements were
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249
Fig. 5. View of DIVA facility inside the JUPITER compartment.
available. Only partial agreement to the measured values could be obtained. This discrepancy is certainly due to the isothermal representation of the walls in the model.
4. Compartment fire experiments 4.1. Experimental set-up and conditions The compartment fire experiments were performed in the DIVA facility (see Fig. 5) inside the JUPITER compartment (3600 m3). The experimental device consisted of three rooms of an identical size (120 m3) in succession. Each room opens into a common corridor of 150 m3. Rooms can be connected through doors, calibrated leaks (critical orifices to simulate leaks through doors and transfer grids between rooms) and a ventilation network. The ventilation network consists of two separate circuits (air supply and exhaust circuits) equipped with fans. The inlet and exhaust branches running inside each room have an opening section of 0.3 0.6 m2. Two tests (PXF1 and PXR1) were performed in the DIVA facility. The analytical cabinet is identical to that of the test CAB101. The cabinet was positioned at the center of the fire room (Room 2). The ignition procedure of the cabinet was identical to the one followed for the CARMELA campaigns series. The DIVA facility was used in a single-room configuration for the test PXF1. Doors between the fire room and the other adjacent rooms were all airtight closed, and no calibrated leak was used in this test. Before ignition, the inlet and exhaust renewal rates of the fire room were 2.5 h 1 each, and the initial relative pressure in the room was 120 Pa. The inlet branch was located in the upper part of the room (0.9 m below the ceiling) and the exhaust branch in the lower part (0.5 m above the floor). The mechanical ventilation was maintained throughout the test. The three lower rooms of the DIVA facility and the corridor were used for the test PXR1. Doors between rooms were also airtight closed. The rooms and the corridor were linked together
through the ventilation network and by calibrated leaks (three of 62.05 mm in diameter and one of 68.15 mm in diameter). Before ignition, the inlet renewal rate was 2 h 1 for the rooms and the corridor, and the exhaust renewal rate was 2.5 h 1 for the rooms and 0.8 h 1 for the corridor. The initial relative pressures were 80 Pa in the corridor, 100 Pa in Room 1 and 120 Pa in Rooms 2 and 3. The inlet branches of the rooms and the corridor were located in the upper part of the room. The exhaust branches were positioned in upper part for the corridor and Room 1 and in lower part for the Rooms 2 and 3. The mechanical ventilation of the rooms and the corridor was maintained throughout the test. The compartment fire tests were instrumented to not only access the combustion in the confined environment in terms of heat generation, heat exchanges and chemistry but also pressure variations and smoke propagation in the facility. The gas temperature within the room was obtained from vertical trees of K-type thermocouples. The same type of thermocouples was used to access the temperature of the walls of the cabinet and the DIVA facility as well as the air within the inlet and exhaust branches of the ventilation network. The pressures within the rooms and along the ventilation network were measured with pressure transmitters. On-line gas analyzers were used to determine the concentration of various types of gaseous species (O2, CO, CO2 and H2O) and unburned gazes (CnHm) into the rooms and the corridor and within the inlet and exhaust branches of the ventilation network. The mass concentration of soot in the rooms and within the ducts of the ventilation network was also measured by means of sequential or continuous (TEOM) samplers on filters. The heat transfers to the walls and to targets were obtained from radiometers and total flux meters. The mass loss rate of the combustible was determined with a weighing machine supporting the whole cabinet. The ventilation air flow rate was measured from an average velocity probe considering gas temperature changes. In addition, video films were taken to monitor the behavior of the fire, the smoke propagation and the equipment. From measurements, the HRR was determined by an energy balance in the fire room. This balance takes into account the
250
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40
500 PXF1 PXR1 CAB101
30 Mass loss rate (g/s)
Heat release rate (kW)
400
PXF1 PXR1 CAB101
300
200
20
10 100
0
0 0
200
400
600
800
1000
1200
0
200
400
600 Time (s)
800
1000
1200
Time (s) Fig. 6. Comparison of heat release rates for tests performed in an enclosed atmosphere (PXF1 and PXR1) and in a free atmosphere (CAB101).
thermal inertia of the cabinet, the energy necessary to maintain the pyrolysis of the combustible, the heat transfers between the fire room and the outside environment through leaks and the ventilation network, the variation of energy inside the fire room and the heat transfers with the walls of the fire room. Details of this method and its validation will be published in a separate paper. The HRR uncertainty associated with the estimation of the HRR is evaluated to be around 10%. 4.2. Main results The stages of the fire behavior observed under hood were reproduced under the compartment, except for the combustion outside the cabinet. The test PXR1 did not reveal such a combustion, whereas a slight but not significant combustion outside the cabinet occurred during the test PXF1, caused by a higher production of excess pyrolysate and a lack of leaks in the fire room (see Fig. 6). Assuming that the HRR follows a t-squared law during the fast spread stage of the fire, the fire growth factor was found to be in the order of 2 W s 2 for closed-door electrical cabinet fires in compartments and is characteristic of a slow growth rate. Heat released during the steady-state stage of the fire is weakly dependent on the confinement of the cabinet (see Fig. 6). Comparable characteristic values of the HRR are obtained for electrical cabinets burning under a calorimetric hood (270 kW) and inside a mechanically ventilated compartment (250 kW). The vitiation of the compartment at the inlet vent of the cabinet is weak, with the oxygen molar fraction decreasing by only 5%. Thus, air entering the cabinet is still oxygen rich. High gas temperatures in the upper part of the cabinet are responsible for the relocation of molten materials. The amount of relocated materials and, consequently, the amount of materials participating to the combustion during the steady-state stage are closely dependent on the surrounding atmosphere of the cabinet. Under the hood, the air renewal rate in the cabinet is higher compared to that of an enclosed atmosphere. Gas temperatures inside and outside the cabinet modify the pressure loss at the outlet vents of the cabinet. As a consequence, the air mass flow rate at the inlet vent of the cabinet increases slightly under the hood. The total relocation of the combustible at bottom of the
Fig. 7. Comparison of mass loss rate for tests performed in an enclosed atmosphere (PXF1 and PXR1) and in a free atmosphere (CAB101).
cabinet occurring for electrical cabinets burning under a calorimetric hood leads to an increase of the mass loss rate of the combustible (20 kg s 1 for the test CAB101 against 14 g s 1 for the test PXR1; see Fig. 7) because temperatures are the highest in this part of the cabinet. Fire extinction occurred while some PMMA mass remained in the upper part of the cabinet for electrical cabinets burning inside a mechanically ventilated compartment (see Fig. 8). Temperatures reached inside the cabinet did not allow the relocation of all materials, and extinction occurred when relocated materials were totally consumed, causing an abrupt drop of the mass loss rate, as observed in Fig. 7 at 480 s for the test PXF1. Fresh materials were present at the bottom of the cabinet during the post-test analysis of the cabinet due to the relocation of some molten materials after fire extinction. Some occurrence of combustion instabilities was noticed for the test PXR1, which could be related to the phenomenon already observed by Quintiere near the oxygen-limit extinction [12]. For this test, the pyrolysis of the combustible remained high during the extinction stage of the fire, leading to three sudden restarts of the fire. This phenomenon was not observed in the test PXF1. In total, 51% of the initial combustible inventory was pyrolyzed in the test PXF1 and 69% in the test PXR1. 4.3. Modeling The model presented in the previous section was introduced in the zone code SYLVIA, developed by IRSN for the simulation of phenomena dealing with ventilation networks, fires in a confined atmosphere and transport of contamination within nuclear facilities. The model assumes homogeneous conditions inside the cabinet in terms of gas temperature and gas composition. The two parameters of the model, oxygen content and gas temperature at inlet vents of the cabinet, are linked to the surrounding atmosphere of the cabinet, and they are provided by the code. The walls of the cabinet thermally exchange with the inner and the outer atmospheres of the cabinet by convective and radiative heat exchanges. Fire inside the cabinet is simulated using a pool fire model, characterized by a burning surface area and a reaction of combustion. The model does not take into account the combustion of excess pyrolysate outside the cabinet because this
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251
40
SYLVIA PXF1
Mass loss rate (g/s)
30
20
10
0 0
200
400
600
800
Time (s) Fig. 9. Simulation of the time evolution of the mass loss rate of the fuel during the test PXF1.
400 Fig. 8. Internal view of the cabinet during the post-experimental phase of the test PXF1.
SYLVIA PXF1
Test
ti (s)
a
HRR (W s 2) characteristic (kW)
PXF1 470 2.1 PXR1 470 2.1 CAB101 600 1.2
214 215 222
_ stationary m (kg s 1)
tfire (s)
Total mass loss (%)
15.6 15.7 19.4
686 686 896
34.9 35.1 100
phenomenon has not been fully revealed for electrical cabinets burning inside a mechanically ventilated compartment. Nevertheless, excess pyrolysate (Eq. (12)) is taken into account in the mass balance of species leaving the cabinet because it impacts the pressure loss at the outlet vents of the cabinet significantly. Simulations of the tests PXF1, PXR1 and CAB101 were performed. Two values of the incubation period were set (see Table 2), depending on whether the cabinet was in a confined atmosphere or not. Results in terms of mass loss rate and heat release rate are reported in Figs. 9 and 10 for the test PXF1, and in Figs. 11 and 12 for the test PXR1. Main results of the three simulations are listed in Table 2. The time evolution of the HRR is qualitatively reproduced by the model (see Fig. 10). The t-squared law used in the model can satisfactorily reproduce the flash over phenomenon. The maximum value of the HRR predicted by the model (220 kW) is under-estimated by 30% in comparison with experimental data, partly due to the slight combustion outside the cabinet, which was not modeled. For the test PXR1, for which no combustion outside the cabinet occurred, the maximum value of the HRR predicted is only under-estimated by 14% (see Fig. 12). The slight decrease of the HRR during the steady-state stage is due to an over-estimation of the reduction of the oxygen content of the air at the inlet vent of the cabinet (9%). The abrupt increase of the mass loss rate experimentally observed during the spread stage of the fire (see Fig. 9) and corresponding to a high increase of gas temperatures inside the cabinet is not reproduced by the model (Eqs. (10) and (11)). The model does not
Heat ralease rate (kW)
300 Table 2 Main results of the simulations of tests PXF1, PXR1 and CAB101.
200
100
0 0
200
400 Time (s)
600
800
Fig. 10. Simulation of the time evolution of the HRR during the test PXF1.
simulate the stratification of gases inside the cabinet; it uses a mean gas temperature leading to an under-estimation of the temperature of hot gases in the upper part of the cabinet. High temperatures inside the cabinet lead to a relocation of molten materials at the bottom of the cabinet and therefore to a reduction of the pyrolysis surface of the fuel. These two phenomena are not modeled. However, this limitation has no consequence on the HRR because the quasistationary stage of the fire is controlled by the oxygen content in the compartment at the level of the inlet vents of the cabinet. Thus, the mass loss rate of the combustible during the steady-state stage is reproduced by the model. Fire extinction occurred due to the lack of combustibles at the bottom of the cabinet (combustion gases in the upper part of the cabinet prevented an ignition of the combustion in this part of the cabinet). Although the model does not simulate the relocation of materials, it is, however, possible to define a criterion of extinction of the fire based on a minimum mass loss rate of the fuel. This criterion only takes into account the part of the pyrolysis participating in the combustion, excluding unburned gases. According to tests PXF1, PXR1 and CAB101, the extinction of the fire is
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40
SYLVIA PXR1
Mass loss rate (g/s)
30
20
10
0 0
200
400
600
800
1000
Time (s) Fig. 11. Simulation of the time evolution of the mass loss rate of the fuel during the test PXR1.
400
SYLVIA PXR1
Heat ralease rate (kW)
300
200
100
release rate is tightly linked to the natural ventilation of the cabinet was demonstrated. Simultaneous measurement of the mass loss rate and of the heat release rate also evidenced the importance of nonburning excess pyrolysate, as was observed in fully developed enclosure fires. On the one hand, this excess pyrolysate represents a hazard because it can burn outside of the cabinet. On the other hand, it limits the fire duration because the burning material is vaporized faster. A semi-empirical model could then be developed based on the coupled solution of the ventilation limit and excess pyrolysate. Considering the intrinsic variability of the experiments, it compares favorably to the measurements of the HRR of the steady stage and fire duration but largely fails to reproduce the kinetics of the fire. The development of such semi-empirical models is a common practice in fire safety engineering concerned with complex combustibles. The evolution of the heat release rate can then be introduced in the calculation method to investigate the consequences of such a fire. With regard to the direct use of the experimental HRR value, semi-empirical HRR modeling opens the way to parametric studies. Another advantage is the possibility to account for the feedback that the environment can have on the fire, which was illustrated by conducting two compartment fire experiments involving analytical cabinets. The room temperature was found to influence the buoyancy within the cabinet and therefore the oxygen brought to the fuel by natural ventilation. The combustion regime was therefore changed: it was experienced as a steady under-ventilated regime under hood with full consumption of the combustible while early extinction or unstable combustion was observed in the compartment fires. The semiempirical model was introduced into a zone code, and an ad-hoc modeling of the extinction based on a critical surfacic mass loss rate was proposed. The major features of the compartment fires experiments such as characteristic HRR and fire duration could then be reproduced with acceptable error. However, additional compartment fire experiments are required for a full validation of the model, especially for cabinets with real electrical components. Additionally, all the cabinets used had closed doors, and another model ought to be developed for cabinets with open doors because the ventilation limit is completely changed in this case, yielding larger HRR.
Acknowledgments
0 0
200
400 600 Time (s)
800
1000
Fig. 12. Simulation of the time evolution of the HRR during the test PXR1.
imposed in the code when the mass loss rate per unit surface of the combustible participating to the combustion becomes lower than 8 g m 2 s 1. This value can satisfactorily reproduce the extinction of the fire, as shown in Fig. 10 for the test PXF1. Nevertheless, the combustion instabilities that occurred during the extinction stage of the test PXR1 cannot be modeled by the code. For electrical cabinets burning under a calorimetric hood, and according to the test CAB101, the surface mass loss rate of the combustible remains above this critical value, leading to the total consumption of the combustible.
5. Concluding remarks Additional calorimetric data related to electrical cabinet fires were produced. The use of PMMA as a surrogate for electrical components helped to investigate the different phenomena arising in such fires. In particular, the concept that a characteristic heat
Part of this work was performed within the frame of PICSEL, a common program of research and development between IRSN and AREVA-NC. References [1] V. Babrauskas, Research on electrical fires: the state of the art, in: Proceedings of the 9th International Symposium on Fire Safety Science, International Association for Fire Safety Science, 2008, pp. 3–18. [2] J. Mangs, J. Paananen, O. Keski-Rahkonen, Calorimetric fire experiments on electronic cabinets, Fire Safety Journal 38 (2003) 165–186. [3] J.M., Chavez, An Experimental Investigation of Internally Ignited Fires in Nuclear Power Plant Control Cabinets: Part 1: Cabinet Effects Tests. NUREG/ CR-4527, SAND86-0336 vol1, Sandia National Laboratories, NM, 1987. [4] J.M., Chavez, An Experimental Investigation of Internally Ignited Fires in Nuclear Power Plant Control Cabinets: Part 2: Room Effects Tests. NUREG/ CR-4527, SAND86-0336 vol. 2, Sandia National Laboratories, NM, 1988. [5] J. Mangs, O. Keski-Rahkonen, Full Scale Fire Experiments on Electronic Cabinets. VTT publications 186, Technical Research centre of Finland, 1994. [6] J., Mangs, O., Keski-Rahkonen, Full Scale Fire Experiments on Electronic Cabinets II. VTT publications 269, Technical Research centre of Finland, 1996. [7] L., Rigollet, S., Me´lis, Fires of electrical cabinets, in: Proceedings of the 11th International Topical Meeting on Nuclear Reactor Thermal-Hydraulics (NURETH-11), Avignon, France, October 2–6 2005. [8] E. Avidor, F.J. Joglar-Billoch, F.W. Mowrer, M. Modarres, Hazard assessment of fire in electrical cabinets, Nuclear technology 144 (2003) 337–357.
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[9] M.L. Janssens, Measuring rate of heat release by oxygen consumption, Fire Technology (1991) 234–249. [10] H.W. Emmons, Vent Flows. SFPE Handbook of Fire Protection Engineering, third ed., National Fire Protection Association, 2002, pp. 2-32–2-41. [11] N.P. Cheremisinoff, Practical fluid mechanics for engineers and scientists, Technomic (1990) 210–211.
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[12] B. Karlsson, J.G. Quintiere, Enclosure Fire Dynamics, CRC Press, 2000 pp. 25–46. [13] M.A. Delichatsios, G.W.H. Silcock, X. Liu, G.W.H. Silcock, X. Liu, M. Delichatsios, Y.-P. Lee, Mass pyrolysis rates and excess pyrolysate in fully developed enclosure fires, Fire Safety Journal 39 (2004) 1–21.
Contents lists available at ScienceDirect
Fire Safety Journal journal homepage: www.elsevier.com/locate/firesaf
Characterization of closed-doors electrical cabinet fires in compartments W. Plumecocq n, M. Coutin, S. Melis, L. Rigollet ˆ rete´ Nucle´aire (IRSN), Centre de Cadarache, BP no 3, 13115 St Paul-Lez-Durance Cedex, France Institut de Radioprotection et de Su
a r t i c l e i n f o
abstract
Article history: Received 7 May 2010 Received in revised form 26 January 2011 Accepted 28 February 2011 Available online 29 April 2011
An important cause of fire departure in industrial facilities is due to electrical origin and particularly to electrical cabinets. The investigation of such fires has been scarce up to now and has been investigated ˆ rete´ Nucle´aire (IRSN) exclusively in the nuclear industry. The Institut de Radioprotection et de Su conducted a large number of experiments involving electrical cabinets burning either under a calorimetric hood or inside a mechanically ventilated compartment to investigate this topic. Calorimetric hood experiments demonstrated that the most important parameter is the size of the vents of the cabinet and that the time to flashover depends on many factors and seems somewhat random with regard to the observable parameters. The influence of the compartment on the fire behavior depends on the temperature of the surrounding atmosphere of the cabinet and on the oxygen content in the compartment at the level of the inlet vent of the cabinet. The compartment strongly impacts the pyrolysis of the combustible, affecting the fire duration, but has a weak effect on the Heat Release Rate (HRR). Experiments were usually remarkably reproducible, opening the way to a phenomenological description of this type of fire. A semi-empirical model based on the coupled solution of ventilation limit and excess pyrolysate could then be developed. This model was introduced in a zone code, and an ad-hoc modeling of the fire extinction based on a critical surfacic mass loss rate is proposed. The major features of the compartment fire experiments such as characteristic HRR and fire duration could then be reproduced with acceptable error. The development of such a semi-empirical model is a common practice in fire safety engineering concerned with complex combustibles. & 2011 Elsevier Ltd. All rights reserved.
Keywords: Electrical cabinet Compartment Fires Fire modeling
1. Introduction An important cause of fire departure is reported [1] to be due to electrical origin. In industrial facilities, and, in particular, in nuclear facilities, a large part of this electrical fire hazard is attributed to electrical or electronic cabinets. These devices are usually composed of electrical components (i.e., wires, contactors and circuit breakers) enclosed in a metallic box in which several openings are located at the bottom and top of the door so that cooling of the circuits is ensured by natural convection. Handling of such fuels in fire safety applications is not obvious, however. The first reason for the difficulty encountered lies in the complexity of the fuel, a charring material composed of different polymers, whose composition is uncertain. Characterizing the combustion of electrical cabinets under a calorimetric hood as a function of its geometrical properties can, however, be attempted to cope with this problem [2]. A second difficulty arises when coming to the question of fire safety application in a compartment because the vitiation can in turn influence the combustion: the heat release rate (HRR) measured under a hood cannot be
n
Corresponding author. Tel.: þ33 442 199 657; fax: þ 33 442 199 167. E-mail address: [email protected] (W. Plumecocq).
0379-7112/$ - see front matter & 2011 Elsevier Ltd. All rights reserved. doi:10.1016/j.firesaf.2011.02.006
used directly as an input datum in the calculation methods. This question is the topic of the present study and is supported by a series of real-scale experiments involving electrical cabinets burning either under a calorimetric hood or inside a mechanically ventilated compartment. The facility and the tests are described. Main results are discussed. A semi-empirical model based on the coupled solution of the ventilation limit and excess pyrolysate is developed and applied to the IRSN experiments.
2. Literature review The investigation of electrical or electronic cabinet fires has been scarce up to now and concerned exclusively with the nuclear industry. Most of the data were published in internal reports. However, this topic is likely to be of some interest for many engineering applications. It seemed therefore useful to begin with a brief review of existing experiments and models. The first experiments were carried out at SNL [3,4] using various cabinet geometries, ventilation openings and ignition conditions; the fire load within the cabinet consisted of cables either passing the IEEE-383 fire qualification test or not. The conclusions of this experimental program showed the fire hazard posed by such electrical materials with a peak of heat release rate up to 175 kW
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Greek letters
Nomenclature Acronyms CAA CAB CAO PXF1 PXR1
a b
First series of CARMELA experiments Second series of CARMELA experiments Third series of CARMELA experiments PICSEL_F.1 PICSEL_R.1
g DHc
w e r s
Fire growth factor [W s 2] Empirical constant [W kg 2] Fraction of the fuel submitted to combustion [-] Heat of combustion [J kg 1] Energy fraction [-] Emissivity [-] Density [kg m 3] Stephan–Boltzmann constant [W m 2 K 4]
Symbols Subscripts C cp g h H k L _ m _ 00 m q Q_ t S T Y
Discharge coefficient [-] Specific heat [J kg 1 K 1] Gravity acceleration [m s 2] Convective heat exchange coefficient [W m 2 K 1] Distance between vents of the cabinet [m] Pressure loss coefficient [-] Latent heat [J kg 1] Mass loss rate [kg s 1] Surface mass loss rate [kg m 2 s 1] Mass flow rate [kg s 1] Heat release rate [W] Time [s] Section area [m2] Temperature [K] Mass fraction [-]
for closed-door cabinets and 955 kW for open-door cabinets. Even though the fire propagation to adjacent cabinets was judged unlikely, it was demonstrated that full involvement of the fire load within the cabinet could always be reached, depending on the ignition conditions. The influence of the closed-door electrical cabinet fires on the enclosure environment was also investigated in a 1400 m3 enclosure and revealed a large amount of smoke release along with a limited gas temperature increase up to 100 1C under the ceiling. Calorimetric experiments were conducted by VTT [2,5,6] using real electrical and electronic components as the fire load within closed-door cabinets. The maximum HRR in the range 100–400 kW was measured. The major results are reported in Table 1. The influence of the cabinet size and the surface of the vents was investigated. The size of the openings was shown to be of prime importance. Assuming that the ventilation within the cabinet is driven by buoyancy and that the fire is ventilationcontrolled, Mangs and Keski-Rahkonen [5] developed an expression (hereafter called the ‘‘MKR model’’) for the maximum HRR released by the cabinet sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffiffi 11=t Q_ max ¼ DHair Cout Sout r1 2gH 1=m þ t
ð1Þ
2 Here, m ¼ Sin Cin =Sout Cout represents the vent opening ratio depending solely on the vent geometry through their surface S and their discharge coefficients C. The temperature ratio t ¼Tg/TN is recognized to be a weak parameter in Eq. (1) for t 41.7. According to the authors, the MKR model represents a maximum for the HRR. To give a larger validation to the MKR model, IRSN [7] also conducted a large number of calorimetric experiments with different sizes of vents and using different materials. The main results are summarized in Table 1, and the major conclusions will be recalled in the next section for completeness.
air c cab f g i in max out r v xs w N
Air Combustible Cabinet Flame Gas Incubation Inlet Maximum Outlet Radiation Vaporization Excess pyrolysate Walls of the cabinet Ambient
Avidor et al. [8] also investigated the fire behavior of closeddoor electrical cabinets with the objective of building a fire hazard matrix. They used a propane gas burner and heptane pool as a surrogate of the electrical components within the cabinet. Considering the cabinet as a perfectly stirred reactor surrounded by a thermally thin steel wall, they could calculate the oxygen yield and temperature within the cabinet using a one-zone model.
3. Characterization of electrical cabinet fires under a hood Twenty-five experiments were carried out under a calorimetric hood to provide further validation of the MKR model [2]. The secondary objective was also to gain further knowledge concerning the temperatures reached by the cabinet walls because it is important to assess the possibility that the fire propagates to adjacent cabinets. 3.1. Experiments Cabinets were disposed in a large well-ventilated enclosure to minimize the influence of the environment. Above the cabinet, a 3 m diameter hood collected all the combustion products for filtering and analysis. The cabinets investigated in the IRSN tests were all of the same dimensions (2 m high, 0.8 m wide and 0.6 m deep). Two openings allowed for air admission in the box: the inlet one was located at the bottom of the front wall, and the outlet one was located either at the center of the roof (first series of experiments) or at the top of the back wall (second and third series). During the fire, the door of the cabinet was locked. The first series of tests (CAA series) investigated the geometrical parameters of the cabinets. Five parameters were selected: the ventilation effect (CAA1xx series), the ignition location (CAA201 test), the amount of fuel (CAA301 test), the spatial
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Table 1 Experimental configurations and major results. Vent areas (m2)
Experiment
Cabinet contents (Vol%)
Amount of contents (kg)
Total mass loss (kg)
HRRd (kW)
DHc
ti (s)
(MJ kg 1)
PMMA
PVC
PE
ECc
2.22 2.22 2.22 1.70 1.70 1.70 1.69 1.69 1.04 1.04
– – – – – – – – – –
– – – – – – – – – –
– – – – – – – – – –
100 100 100 100 100 100 100 100 100 100
66.5a 70.7 66.4 61.6 30.4 90.5 5.8 5.8 5.9 5.9
22.0 23.6 14.6 34.8 3.0 38.7 3.5 3.5 3.4 3.4
175 125 100 400 50 200 35 20 40 20
19.6b 11.4 19.7 18.8 8.9 18.3 16.2 11.7 13.9 11.7
– – – – – – – – – –
0.1000 0.1000 0.1000 0.1000 0.0500 0.0500 0.1000 0.1000 0.1000 0.1000 0.1000 0.1000 0.1000 0.1000
1.90 1.90 1.90 1.90 1.90 1.90 1.90 1.90 1.90 1.90 1.90 1.90 1.90 1.90
100 100 100 100 100 100 100 100 100 100 100 100 100 100
– – – – – – – – – – – – – –
– – – – – – – – – – – – – –
– – – – – – – – – – – – – –
10.0 10.0 10.0 10.0 10.0 10.0 10.0 5.0 10.0 10.0 10.0 10.0 10.0 10.0
– – – – – – – – – – – – – –
575 475 275 175 – 125 325 325 325 325 300 325 575 450
– – – – – – – – – – – – – –
425 293 387 234 305 170 458 299 272 268 226 426 285 107
0.0250 0.0250 0.0250 0.0250 0.0250 0.0250 0.0250 0.0250 0.0250
0.1000 0.0500 0.1000 0.1000 0.1000 0.1000 0.1000 0.1000 0.1000
1.70 1.70 1.70 1.70 1.70 1.70 1.70 1.70 1.70
100 100 100 100 100 50 29 29 37
– – – – – 50 71 42 31.5
– – – – – – – 29 31.5
– – – – – – – – –
10.0 10.0 10.0 10.0 10.0 10.0 10.0 10.0 10.0
10.0 10.0 9.7 10.0 9.9 9.3 1.6 0.6 9.7
250 200 225 240 225 200 50 25 225
23.6 23.7 20.5 20.2 23.3 14.6 16.3 15.6 20.6
245 119 183 236 115 294 141 – 294
CAO1 CAO2
0.0250 0.0250
0.1000 0.1000
1.70 1.70
– –
– –
– –
100 100
46.3 46.3
11.5 12.3
225 250
17.5 20.2
600 600
PXF1 PXR1
0.0250 0.0250
0.1000 0.1000
1.70 1.70
100 100
– –
– –
– –
10.0 10.0
5.1 6.8
250 250
10.8 10.8
152 156
Organism
Facility
Test name
Sin
Sout
VTT
Hood
1 2 3 4 5 6 7 8 9 10
0.0097 0.0097 0.0101 0.0403 0.0417 0.0418 0.0055 0.0029 0.0067 0.0034
0.0535 0.0535 0.0535 0.0792 0.0792 0.0792 0.0077 0.0077 0.0119 0.0119
IRSN
SATURNE hood
CAA102 CAA103 CAA104 CAA105 CAA106 CAA107 CAA201 CAA301 CAA401 CAA402 CAA403 CAA501 CAA502 CAA503
0.1000 0.0500 0.0250 0.0175 0.0250 0.0075 0.0250 0.0250 0.0250 0.0250 0.0250 0.0250 0.0250 0.0250
CAB101 CAB102 CAB201 CAB202 CAB203 CAB301 CAB302 CAB303 CAB304
DIVA Multi-rooms
H (m)
a
Estimated from the sum of the total burned mass and mass of contents after the experiment. Total energy release up to 58.7 min divided with corresponding mass loss. c EC: Components of electrical cabinets (see Refs. [2,7] for the nature and proportion). d Maximal HRR for the cabinets with EC contents or HRR corresponding to the plateau during the steady stage of the fire for the cabinets with PMMA, PVC and PE contents. b
arrangement in the cabinet (CAA4xx series) and the cabinet filling (CAA5xx series). A plate of polymethyl methacrylate (PMMA) was used to simulate the fuel: this material burns easily, and its properties are well-known. The plate was 1.5 m 0.7 m 8 mm (half plate for the test CAA301) and was positioned vertically at the center of the cabinet. In the tests CAA401, CAA402 and CAA403, two steel horizontal plates were located inside the cabinet, partially obstructing the flow and allowed a surface occupation rate between 30% and 90%. In the tests CAA501, CAA502 and CAA503, two hollow steel volumes were positioned against the lateral walls inside the cabinet allowing a volume rate of the cabinet between 25% and 75%. Additionally, eight tests were performed for repeatability. The second series of experiments (CAB series) investigated the influence of the fuel. One hundred and five squares of 10 cm 10 cm 8 mm each were mounted on steel shafts and positioned inside the cabinet to form the desired layout: a plate (squares disposed side by side to constitute a vertical plate), shelves, cable-tray (15 squares disposed as a vertical narrow band in front of a 90-squares plate) or dispersed (squares disposed
uniformly all over the volume of the cabinet). Polymethyl methacrylate, polyvinyl chloride (PVC) and polyethylene (PE) were used as combustible materials. Opening sections of the cabinet were 0.025 m2 at inlet and 0.1 m2 at outlet, except for test CAB102, where the latter was 0.05 m2. The third series of experiments (CAO series) consisted of two tests involving two identical relay cabinets. The inlet opening was 0.025 m2 and the outlet was 0.1 m2. The electrical components are circuit breakers, cable trunking, wires, terminal boxes, relays and contactors. The experiments consisted of a piloted ignition of the fuel material inside the cabinet. A linear propane burner generally located at the bottom of the fuel and in the middle of the cabinet provided the pilot flame. Its effective length was 0.7 m except for the test CAB202 (0.1 m) and the CAO series (0.6 m). The burner was shut off after ignition was confirmed (usually after two minutes or so), and the mean propane power output was about 20, 6 and 9 kW, respectively, for the CAA, CAB and CAO series. The fire was then allowed to develop freely until its natural extinction.
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The experiments were instrumented so as to allow a full interpretation of the combustion in terms of heat generation, heat exchanges and release of combustion products. In the exhaust duct downstream of the hood, all the combustion products were sampled and analyzed. The measurements included oxygen, carbon monoxide, carbon dioxide and water vapor concentrations. From these data, the oxygen consumption calorimetry [9] allows the determination of the heat released by the combustion. The HRR uncertainty was evaluated as 10–15%, the highest uncertainty coming from the measure of the flow rate inside the pipe. Thermocouples were placed inside the cabinets and on the outside walls. Radiometers indicated the radiative heat fluxes emitted by the cabinets. In addition, for the second and third series of experiments, an infrared camera gave indications of temperatures, radiant heat fluxes and emissivities on one side of the cabinet. A weighing machine was used to estimate the mass loss rate of the fuel. 3.2. Main results The goal of the analytical experiments of the first and second series was to facilitate the determination of the most influential parameters on the combustion of electrical cabinets. Three target values of interest were chosen: the maximum rate of heat release, the steady-state rate of heat release and the time to reach flashover (see Table 1). These experiments were usually remarkably reproducible. Five stages can be distinguished during the fire (see Fig. 1 for test CAA 104): The incubation stage. This stage starts with the spread of the flame along the fuel and the heating up of gases inside the cabinet. Gas temperatures inside the cabinet increase quickly, initiating a thermal pump, whereas the steel walls of the cabinet remain cold. CARMELA experiments [7] revealed extremely variable time durations of the incubation of the fire, between 107 and 600 s (see Table 1), because it is highly dependent on ignition conditions. The fast spread stage. At some point, the spread of the fire increases drastically, leading to a flashover phenomenon. Flames can be visible at the outlet vent of the cabinet, and increasing wall temperatures indicate that most of the contents of the cabinet are surrounded by hot temperatures. Radiative heat transfer inside the cabinet becomes predominant. This stage is highly oxygenconsuming. 800 Third stage: Combustion outside the cabinet
700
Heat release rate (kW)
600 500 Fourth stage: Steady stage
400 Second stage: Fast spread
300
Fifth stage: Decay D
First stage: Incubation period
200 100 0
0
250
500
1000 Time (s)
750
1250
Fig. 1. Characteristic evolution of the HRR during a test CAA.
1500
Combustion outside the cabinet. As the fuel involved in the combustion increases, the amount of combustible pyrolyzed exceeds the quantity that can react with the oxygen available from venting: the combustion becomes under-ventilated. In some experiments, the products of combustion leaving the outlet vent were hot enough and rich enough in unburned fuel vapors to reignite as they mix with fresh air; a flame then appeared outside the cabinet on the outlet ventilation opening. In the CAB and CAO experiments, no such flame was visible, but the mass loss measurements showed that the combustible was pyrolyzed in excess with regard to the oxygen available; the location of the outlet opening could have been responsible for some additional cooling of the combustion products, thus inhibiting the ignition outside the cabinet. Indeed, the outlet vent was in the ceiling for CAA experiments and at the top of the door for the CAB and CAO experiments. The steady-state stage. The high gas temperatures reached inside the cabinet lead to the relocation of molten materials at the bottom of the cabinet. This phase is closely linked to the combustion of molten materials at the bottom of the cabinet: equilibrium is established between the quantity of oxygen entering the cabinet and the quantity of fuel pyrolyzed. The extinction stage. The fire usually ends due to lack of combustible, with the exception of the test CAA106 where the fall of the plate caused the fire to be extinguished and the tests CAB302 and CAB303 where no ignition could be achieved. A long and smooth decay of the rate of heat released is observed. In the tests CAA501–CAA503, steel volumes were placed in the cabinet, and some influence was observed on the three target parameters in comparison with the reference test CAA104: higher rates of heat were released, and a faster flashover was observed. However this effect is attributed to an insulation effect of the walls and is not judged to be directly linked to the presence of dead volumes in the box. The tests CAB101 and CAB102 are compared, respectively, to the tests CAA104 and CAA106. In the tests CAB101 and CAB102, besides the difference in the outlet vent location, the relocation mode of the combustible was also different because the PMMA squares were allowed to melt and flow down independently from each other, which certainly points to the influence noticed in the peak heat release rate. To complement these experiments, normative ignition tests were performed under a cone calorimeter: the time to ignition of different plastic materials submitted to an imposed radiative flux was measured and proved to be highly variable. The tests CAB302 and CAB303 did not lead to flashover; obviously, there is a crucial effect of the nature of the combustible on the flashover phenomenon. The repeatability tests performed often presented very different times to flashover. One possible reason is the small difference in the ignition, but it can also be attributed to an intrinsic variability. On the contrary, the repeatability of the steady-state HRR, corresponding to the ventilation limit, was rather good (within 10%). The maximum heat release rate repeatability was better for the experiments of the third series involving real cabinets (10%) than for those of the first series (up to 33% for the test CAA503). There seems to be some relationship between the low repeatability and the presence of a large part of the flame burning outside the cabinet. As a conclusion, the analytical experiments demonstrate that the most influential parameter is the size of the openings. The time to flashover depends on many factors and seems somewhat random with regard to the observable parameters. 3.3. Modified MKR model The MKR model provides an estimate of the maximum heat release rate of the cabinets experiencing flashover. However,
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given the additional data brought by the experiments, some modifications were done to this model. 3.3.1. Cabinet air flow Firstly, because some of the vent areas studied were rather large compared to the cabinet cross-section, Eq. (1) had to be rewritten. The air flow transiting in the cabinet is due to a balance between the pressure drops caused by the vents and the buoyancy. Assuming rgTg ¼ rNTN, the Bernoulli equation yields q2 q2 q2 Tg 1 1 1 Tg T1 kin 2 in þ kout 2 in þ 1 2 in ¼ 1 r gH 2 Tg 1 Sin r1 2 Sout r1 T1 2 T1 Scab r1 ð2Þ The two first terms of the left-hand side of Eq. (2) represent the pressure drops at the inlet and outlet openings of the box; the corresponding coefficients kin and kout are commonly estimated to be 2.8 for turbulent vent flows [10]. Mangs and Keski-Rahkonen [5] measured kin ¼1.83 and kout ¼2.78; here, the coefficients were derived from the Weisbach equation [11] but only little variations in the range 2.6–2.9 were found. The third term of the left-hand side of Eq. (2) accounts for the acceleration of the fluid due to its heating; this term is neglected in the Mangs and Keski-Rahkonen model but is kept here for generality. The viscous wall friction pressure drops inside the box can be estimated to be less than 5% of the pressure drops due to vents. The right-hand side of Eq. (2) accounts for the buoyancy forces resulting from the density difference between the inside and the outside of the cabinet. As was noticed by Mangs et al. [2], the dependence of Eq. (2) on Tg is weak once Tg 4200 1C. Therefore, a representative value of Tg ¼Tf ¼800 1C is imposed, and one can estimate a steady-state mass flow rate sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi pffiffiffiffiffiffiffiffiffi 1ðT1 =Tf Þ q ¼ r1 2gH ðkin =S2in Þ þðTf =T1 Þðkout =S2out Þ þ½ðTf =T1 Þ1ð1=S2cab Þ ð3Þ n
Remarkably, q is representative of the flow inside the box as soon as the gas temperature rises at the outlet. In the experiments, this rise occurs very quickly, well before flashover; the gas burner used for ignition is enough to increase the outlet temperature to about 300 1C. 3.3.2. Energy balance of the cabinet An energy balance equation of the cabinet allows the determination of the gas and cabinet walls temperatures ! dTw Q_ c _ xs Lv MCw ¼ Q_ c q cpg ðTg T1 Þ þm dt DHc 4 4 Sw hw ðTw T1 ÞSw ew sðTw T1 Þ
ð4Þ
In Eq. (4), the left-hand term represents the thermal inertia of the cabinet, with MCw calculated from the total mass of the cabinet and the heat capacity of steel. The first term of the right-hand side represents the power released within the cabinet. The second term represents the heat convected out of the cabinet. The third term is the heat necessary for the pyrolysis of the combustible; it has been decomposed into two parts: the amount of pyrolysis necessary to produce the HRR within the cabinet and an excess pyrolysate that cannot burn inside the cabinet due to the ventilation limit. This last term is modeled in the next section. The fourth term is the heat exchanged by natural convection at the external walls of the cabinet: a classical value of hw ¼12 W m 2 K 1 is used, and the cabinet’s outer exchange surface Sw is calculated from the cabinet geometry. The fifth term is the energy radiated out by the cabinet walls: the value of the emissivity was measured as ew ¼0.7. Some of the terms of Eq. (4)
247
are determined according to further physical modeling: depending on the combustion mode (over- or under-ventilated) the gas temperature Tg is related either to the convected flux as in the open (see Eq. (5a)) or to the temperature of the cabinet Tw by assuming that a thermal equilibrium is achieved (Eq. (5b)). In over-ventilated conditions, convection is the dominant mode of heat transfer, and Tg is calculated as q cpg ðTg T1 Þ ¼ ð1wr ÞQ_ c
ð5aÞ
where wr is the radiated fraction of the combustible burning in the open conditions. In under-ventilated conditions, radiation is the dominant mode of heat transfer, and Tg is calculated as 4 esðTg4 Tw4 Þ ¼ hw ðTw T1 Þ þ ew sðTw4 T1 Þ
ð5bÞ
Eventually, a time-solution of Eq. (4) is provided by an implicit first-order solver. Because the system is represented by a single wall temperature whatever the elevation, the results are to be interpreted as only indicative of the temperature levels achieved. However, the two last terms of Eq. (4) provide an estimate of the thermal stress that a burning cabinet can have on its surroundings. 3.3.3. Pyrolysis rate and heat release The model for pyrolysis and heat release rate is largely based on the theory of a fully developed compartment fire. Two combustion regimes are distinguished: over-ventilated at first and then under-ventilated. The fire spread in the over-ventilated regime is divided in two growth stages. The combustion is modeled by two t-squared laws, which is a simple but practical way to represent complex fires [12]. During the first stage (incubation period, from ignition to time ti) Q_ c ðtÞ ¼ a1 t 2
ð6Þ
During the second stage (fast spread, after ti in the overventilated regime) Q_ c ðtÞ ¼ a2 ðtti Þ2 þ a1 t 2 ð7Þ The parameters a1 and ti were determined for each experiment as the best least-squares unbiased coefficients to fit the experimental data. Values of ti are reported in Table 1. These parameters were extremely variable, and no reproducibility could be achieved by repeating the experiments. However, an order of magnitude for ti of 100–450 s for PMMA tests and 600 s for real electrical components can be a compromise for these data. The growth factor a1 is even more uncertain, with an average of 1.3 W s 2. Some empirical correlation was established from the experiments
a2 ¼ bq2
ð8Þ 2
for the tests involving PMMA, and with b ¼1800 W kg b ¼ 100 W kg 2 for real cabinets is found to give best-fit for the present data. The maximum heat released in the cabinet is determined from the oxygen available, i.e., from qn. The heat released by the combustion of a fuel with a given amount of oxygen is known to be weakly dependent on the nature of the combustible, which leads to the formulation of the ventilation limit in term of power, just as in the original MKR model [5]. Q_ c ¼ q DHair
ð9Þ 1
where DHair ¼3.144 MJ kg is a constant valid over a large range of fuels at standard conditions of pressure and oxygen concentration. As appears in the experiments, the mass loss rate is not strictly correlated with the heat release rate, indicating a strong decrease in the effective heat of combustion during the flashover. To model
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this effect, the fire inside the cabinet is considered an underventilated, post-flashover enclosure [13]. Due to the ventilation limit, only a fraction g of the total fuel surface is surrounded by the flame and burns normally, the remaining being exposed to the hot gas temperature leading to its pyrolysis. It is further assumed that the pyrolysis is mainly due to radiation, and an optically thick medium is considered. The pyrolysis of the burning surface is then linked to the oxidant mass flux by stoichiometry 00
_ 1 gSf DHc ¼ gQ_ max q DHair ¼ m
ð10Þ
where Q_ max , the maximum heat release rate of all the combustible burning in free air, is introduced. The excess pyrolysate is obtained by linking the radiative heat fluxes in the flame zone to the amount of pyrolyzed combustible as follows: _ xs ¼ 1g Sf sðTg4 Tv4 Þ=Lv ð11Þ m
800
Characteristic power (kW)
248
Recognizing that Q_ max ¼ Sf sðTf4 Tv4 ÞDHc =Lv , Eqs. (9) and (10) can be recast as =DHC
600
400
200
0 0.00
ð12Þ
3.4. Discussion Similar to the MKR model, the model proposed is first based on the ventilation limit imposed by the geometry of the openings. The experiments carried out with different geometries provided a rather large basis for the validation of this concept. As an example, Fig. 2 compares HRR versus time for three experiments with similar cabinet geometries but different combustibles inside. Obviously, the characteristic HRR is well determined either by Eq. (1) or by the combination of Eqs. (3) and (9). However, the nature and the arrangement of the combustible within the
0.20
200
150
100 PMMA PMMA-PE-PVC Electrical components Agreement
50
0
300
200
100
0 1000 Time (s)
1500
100 150 200 50 Experimental maximum external heat flux (kW)
250
Fig. 4. Calculated versus measured maximum heat flux emitted by the cabinet’s wall to the environment.
CAO2 CAB202 CAB101 Model
400 Heat release rate (kW)
0.15
250
0 500
500
0.10
Fig. 3. Characteristic values of heat release rates for different fuels.
However, as was already noted, the occurrence of a flame outside the cabinet is thought to be specific to the tests of the first series involving PMMA. For the other tests, the excess pyrolysate is quenched as the flow impinges the ceiling of the cabinet and exits through the upper vent, so that little re-ignition was observed outside of the cabinet. However, this excess pyrolysate contributes to decreasing the fire duration without changing the heat release rate.
0
0.05
Intrinsic mass flow rate (kg/s)
Calculated maximum external heat flux (kW)
_ xs ¼ m
Tg4 Tv4 ðQ_ max q DHair Þ 4 4 Tf Tv
PMMA tests PMMA/PVC/PE tests Real Fuel tests VTT tests Theoretical
2000
Fig. 2. HRR of three tests with openings and model limits identical to those of Eq. (9).
cabinet influence the propagation before flashover profoundly and hence the time to flashover. Fig. 3 presents the experimental versus theoretical characteristic HRR at the steady state for all experiments: the deviations can be quite large, but the major part of the experiments collapses clearly against the theoretical line. The under-estimation of the HRR is mostly due to early extinction before flashover is encountered so that the model proposed should be taken as a worst-case calculation. The total (radiative and convective) heat flux from the cabinet walls could be calculated for the second and third series of experiments by integrating the measured heat fluxes over the surface of the cabinet. Because the temperature of the cabinet wall is continuously increasing, these fluxes reach a maximum just before the end of the fire. In addition to the original MKR model, the present calculations include the determination of the cabinet temperature (from the last two terms of Eq. (4)) and therefore provide an estimation of the heat fluxes to the ambient. Fig. 4 compares their maximum values to the calculated ones for the experiments where the measurements were
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249
Fig. 5. View of DIVA facility inside the JUPITER compartment.
available. Only partial agreement to the measured values could be obtained. This discrepancy is certainly due to the isothermal representation of the walls in the model.
4. Compartment fire experiments 4.1. Experimental set-up and conditions The compartment fire experiments were performed in the DIVA facility (see Fig. 5) inside the JUPITER compartment (3600 m3). The experimental device consisted of three rooms of an identical size (120 m3) in succession. Each room opens into a common corridor of 150 m3. Rooms can be connected through doors, calibrated leaks (critical orifices to simulate leaks through doors and transfer grids between rooms) and a ventilation network. The ventilation network consists of two separate circuits (air supply and exhaust circuits) equipped with fans. The inlet and exhaust branches running inside each room have an opening section of 0.3 0.6 m2. Two tests (PXF1 and PXR1) were performed in the DIVA facility. The analytical cabinet is identical to that of the test CAB101. The cabinet was positioned at the center of the fire room (Room 2). The ignition procedure of the cabinet was identical to the one followed for the CARMELA campaigns series. The DIVA facility was used in a single-room configuration for the test PXF1. Doors between the fire room and the other adjacent rooms were all airtight closed, and no calibrated leak was used in this test. Before ignition, the inlet and exhaust renewal rates of the fire room were 2.5 h 1 each, and the initial relative pressure in the room was 120 Pa. The inlet branch was located in the upper part of the room (0.9 m below the ceiling) and the exhaust branch in the lower part (0.5 m above the floor). The mechanical ventilation was maintained throughout the test. The three lower rooms of the DIVA facility and the corridor were used for the test PXR1. Doors between rooms were also airtight closed. The rooms and the corridor were linked together
through the ventilation network and by calibrated leaks (three of 62.05 mm in diameter and one of 68.15 mm in diameter). Before ignition, the inlet renewal rate was 2 h 1 for the rooms and the corridor, and the exhaust renewal rate was 2.5 h 1 for the rooms and 0.8 h 1 for the corridor. The initial relative pressures were 80 Pa in the corridor, 100 Pa in Room 1 and 120 Pa in Rooms 2 and 3. The inlet branches of the rooms and the corridor were located in the upper part of the room. The exhaust branches were positioned in upper part for the corridor and Room 1 and in lower part for the Rooms 2 and 3. The mechanical ventilation of the rooms and the corridor was maintained throughout the test. The compartment fire tests were instrumented to not only access the combustion in the confined environment in terms of heat generation, heat exchanges and chemistry but also pressure variations and smoke propagation in the facility. The gas temperature within the room was obtained from vertical trees of K-type thermocouples. The same type of thermocouples was used to access the temperature of the walls of the cabinet and the DIVA facility as well as the air within the inlet and exhaust branches of the ventilation network. The pressures within the rooms and along the ventilation network were measured with pressure transmitters. On-line gas analyzers were used to determine the concentration of various types of gaseous species (O2, CO, CO2 and H2O) and unburned gazes (CnHm) into the rooms and the corridor and within the inlet and exhaust branches of the ventilation network. The mass concentration of soot in the rooms and within the ducts of the ventilation network was also measured by means of sequential or continuous (TEOM) samplers on filters. The heat transfers to the walls and to targets were obtained from radiometers and total flux meters. The mass loss rate of the combustible was determined with a weighing machine supporting the whole cabinet. The ventilation air flow rate was measured from an average velocity probe considering gas temperature changes. In addition, video films were taken to monitor the behavior of the fire, the smoke propagation and the equipment. From measurements, the HRR was determined by an energy balance in the fire room. This balance takes into account the
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40
500 PXF1 PXR1 CAB101
30 Mass loss rate (g/s)
Heat release rate (kW)
400
PXF1 PXR1 CAB101
300
200
20
10 100
0
0 0
200
400
600
800
1000
1200
0
200
400
600 Time (s)
800
1000
1200
Time (s) Fig. 6. Comparison of heat release rates for tests performed in an enclosed atmosphere (PXF1 and PXR1) and in a free atmosphere (CAB101).
thermal inertia of the cabinet, the energy necessary to maintain the pyrolysis of the combustible, the heat transfers between the fire room and the outside environment through leaks and the ventilation network, the variation of energy inside the fire room and the heat transfers with the walls of the fire room. Details of this method and its validation will be published in a separate paper. The HRR uncertainty associated with the estimation of the HRR is evaluated to be around 10%. 4.2. Main results The stages of the fire behavior observed under hood were reproduced under the compartment, except for the combustion outside the cabinet. The test PXR1 did not reveal such a combustion, whereas a slight but not significant combustion outside the cabinet occurred during the test PXF1, caused by a higher production of excess pyrolysate and a lack of leaks in the fire room (see Fig. 6). Assuming that the HRR follows a t-squared law during the fast spread stage of the fire, the fire growth factor was found to be in the order of 2 W s 2 for closed-door electrical cabinet fires in compartments and is characteristic of a slow growth rate. Heat released during the steady-state stage of the fire is weakly dependent on the confinement of the cabinet (see Fig. 6). Comparable characteristic values of the HRR are obtained for electrical cabinets burning under a calorimetric hood (270 kW) and inside a mechanically ventilated compartment (250 kW). The vitiation of the compartment at the inlet vent of the cabinet is weak, with the oxygen molar fraction decreasing by only 5%. Thus, air entering the cabinet is still oxygen rich. High gas temperatures in the upper part of the cabinet are responsible for the relocation of molten materials. The amount of relocated materials and, consequently, the amount of materials participating to the combustion during the steady-state stage are closely dependent on the surrounding atmosphere of the cabinet. Under the hood, the air renewal rate in the cabinet is higher compared to that of an enclosed atmosphere. Gas temperatures inside and outside the cabinet modify the pressure loss at the outlet vents of the cabinet. As a consequence, the air mass flow rate at the inlet vent of the cabinet increases slightly under the hood. The total relocation of the combustible at bottom of the
Fig. 7. Comparison of mass loss rate for tests performed in an enclosed atmosphere (PXF1 and PXR1) and in a free atmosphere (CAB101).
cabinet occurring for electrical cabinets burning under a calorimetric hood leads to an increase of the mass loss rate of the combustible (20 kg s 1 for the test CAB101 against 14 g s 1 for the test PXR1; see Fig. 7) because temperatures are the highest in this part of the cabinet. Fire extinction occurred while some PMMA mass remained in the upper part of the cabinet for electrical cabinets burning inside a mechanically ventilated compartment (see Fig. 8). Temperatures reached inside the cabinet did not allow the relocation of all materials, and extinction occurred when relocated materials were totally consumed, causing an abrupt drop of the mass loss rate, as observed in Fig. 7 at 480 s for the test PXF1. Fresh materials were present at the bottom of the cabinet during the post-test analysis of the cabinet due to the relocation of some molten materials after fire extinction. Some occurrence of combustion instabilities was noticed for the test PXR1, which could be related to the phenomenon already observed by Quintiere near the oxygen-limit extinction [12]. For this test, the pyrolysis of the combustible remained high during the extinction stage of the fire, leading to three sudden restarts of the fire. This phenomenon was not observed in the test PXF1. In total, 51% of the initial combustible inventory was pyrolyzed in the test PXF1 and 69% in the test PXR1. 4.3. Modeling The model presented in the previous section was introduced in the zone code SYLVIA, developed by IRSN for the simulation of phenomena dealing with ventilation networks, fires in a confined atmosphere and transport of contamination within nuclear facilities. The model assumes homogeneous conditions inside the cabinet in terms of gas temperature and gas composition. The two parameters of the model, oxygen content and gas temperature at inlet vents of the cabinet, are linked to the surrounding atmosphere of the cabinet, and they are provided by the code. The walls of the cabinet thermally exchange with the inner and the outer atmospheres of the cabinet by convective and radiative heat exchanges. Fire inside the cabinet is simulated using a pool fire model, characterized by a burning surface area and a reaction of combustion. The model does not take into account the combustion of excess pyrolysate outside the cabinet because this
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251
40
SYLVIA PXF1
Mass loss rate (g/s)
30
20
10
0 0
200
400
600
800
Time (s) Fig. 9. Simulation of the time evolution of the mass loss rate of the fuel during the test PXF1.
400 Fig. 8. Internal view of the cabinet during the post-experimental phase of the test PXF1.
SYLVIA PXF1
Test
ti (s)
a
HRR (W s 2) characteristic (kW)
PXF1 470 2.1 PXR1 470 2.1 CAB101 600 1.2
214 215 222
_ stationary m (kg s 1)
tfire (s)
Total mass loss (%)
15.6 15.7 19.4
686 686 896
34.9 35.1 100
phenomenon has not been fully revealed for electrical cabinets burning inside a mechanically ventilated compartment. Nevertheless, excess pyrolysate (Eq. (12)) is taken into account in the mass balance of species leaving the cabinet because it impacts the pressure loss at the outlet vents of the cabinet significantly. Simulations of the tests PXF1, PXR1 and CAB101 were performed. Two values of the incubation period were set (see Table 2), depending on whether the cabinet was in a confined atmosphere or not. Results in terms of mass loss rate and heat release rate are reported in Figs. 9 and 10 for the test PXF1, and in Figs. 11 and 12 for the test PXR1. Main results of the three simulations are listed in Table 2. The time evolution of the HRR is qualitatively reproduced by the model (see Fig. 10). The t-squared law used in the model can satisfactorily reproduce the flash over phenomenon. The maximum value of the HRR predicted by the model (220 kW) is under-estimated by 30% in comparison with experimental data, partly due to the slight combustion outside the cabinet, which was not modeled. For the test PXR1, for which no combustion outside the cabinet occurred, the maximum value of the HRR predicted is only under-estimated by 14% (see Fig. 12). The slight decrease of the HRR during the steady-state stage is due to an over-estimation of the reduction of the oxygen content of the air at the inlet vent of the cabinet (9%). The abrupt increase of the mass loss rate experimentally observed during the spread stage of the fire (see Fig. 9) and corresponding to a high increase of gas temperatures inside the cabinet is not reproduced by the model (Eqs. (10) and (11)). The model does not
Heat ralease rate (kW)
300 Table 2 Main results of the simulations of tests PXF1, PXR1 and CAB101.
200
100
0 0
200
400 Time (s)
600
800
Fig. 10. Simulation of the time evolution of the HRR during the test PXF1.
simulate the stratification of gases inside the cabinet; it uses a mean gas temperature leading to an under-estimation of the temperature of hot gases in the upper part of the cabinet. High temperatures inside the cabinet lead to a relocation of molten materials at the bottom of the cabinet and therefore to a reduction of the pyrolysis surface of the fuel. These two phenomena are not modeled. However, this limitation has no consequence on the HRR because the quasistationary stage of the fire is controlled by the oxygen content in the compartment at the level of the inlet vents of the cabinet. Thus, the mass loss rate of the combustible during the steady-state stage is reproduced by the model. Fire extinction occurred due to the lack of combustibles at the bottom of the cabinet (combustion gases in the upper part of the cabinet prevented an ignition of the combustion in this part of the cabinet). Although the model does not simulate the relocation of materials, it is, however, possible to define a criterion of extinction of the fire based on a minimum mass loss rate of the fuel. This criterion only takes into account the part of the pyrolysis participating in the combustion, excluding unburned gases. According to tests PXF1, PXR1 and CAB101, the extinction of the fire is
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40
SYLVIA PXR1
Mass loss rate (g/s)
30
20
10
0 0
200
400
600
800
1000
Time (s) Fig. 11. Simulation of the time evolution of the mass loss rate of the fuel during the test PXR1.
400
SYLVIA PXR1
Heat ralease rate (kW)
300
200
100
release rate is tightly linked to the natural ventilation of the cabinet was demonstrated. Simultaneous measurement of the mass loss rate and of the heat release rate also evidenced the importance of nonburning excess pyrolysate, as was observed in fully developed enclosure fires. On the one hand, this excess pyrolysate represents a hazard because it can burn outside of the cabinet. On the other hand, it limits the fire duration because the burning material is vaporized faster. A semi-empirical model could then be developed based on the coupled solution of the ventilation limit and excess pyrolysate. Considering the intrinsic variability of the experiments, it compares favorably to the measurements of the HRR of the steady stage and fire duration but largely fails to reproduce the kinetics of the fire. The development of such semi-empirical models is a common practice in fire safety engineering concerned with complex combustibles. The evolution of the heat release rate can then be introduced in the calculation method to investigate the consequences of such a fire. With regard to the direct use of the experimental HRR value, semi-empirical HRR modeling opens the way to parametric studies. Another advantage is the possibility to account for the feedback that the environment can have on the fire, which was illustrated by conducting two compartment fire experiments involving analytical cabinets. The room temperature was found to influence the buoyancy within the cabinet and therefore the oxygen brought to the fuel by natural ventilation. The combustion regime was therefore changed: it was experienced as a steady under-ventilated regime under hood with full consumption of the combustible while early extinction or unstable combustion was observed in the compartment fires. The semiempirical model was introduced into a zone code, and an ad-hoc modeling of the extinction based on a critical surfacic mass loss rate was proposed. The major features of the compartment fires experiments such as characteristic HRR and fire duration could then be reproduced with acceptable error. However, additional compartment fire experiments are required for a full validation of the model, especially for cabinets with real electrical components. Additionally, all the cabinets used had closed doors, and another model ought to be developed for cabinets with open doors because the ventilation limit is completely changed in this case, yielding larger HRR.
Acknowledgments
0 0
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Fig. 12. Simulation of the time evolution of the HRR during the test PXR1.
imposed in the code when the mass loss rate per unit surface of the combustible participating to the combustion becomes lower than 8 g m 2 s 1. This value can satisfactorily reproduce the extinction of the fire, as shown in Fig. 10 for the test PXF1. Nevertheless, the combustion instabilities that occurred during the extinction stage of the test PXR1 cannot be modeled by the code. For electrical cabinets burning under a calorimetric hood, and according to the test CAB101, the surface mass loss rate of the combustible remains above this critical value, leading to the total consumption of the combustible.
5. Concluding remarks Additional calorimetric data related to electrical cabinet fires were produced. The use of PMMA as a surrogate for electrical components helped to investigate the different phenomena arising in such fires. In particular, the concept that a characteristic heat
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