Experimental study of natural roof ventilation in full-scale enclosure fire tests in a small compartment

ARTICLE IN PRESS

Fire Safety Journal 42 (2007) 523–535 www.elsevier.com/locate/firesaf

Experimental study of natural roof ventilation in full-scale enclosure fire tests in a small compartment Bart Mercia,b,, Paul Vandeveldeb,c a

Fund for Scientific Research—Flanders (FWO—Vlaanderen), Belgium Ghent University—UGent, Department of Flow, Heat and Combustion Mechanics, Belgium c WFRGENT NV, Ghent, Belgium

b

Received 29 May 2006; received in revised form 4 December 2006; accepted 12 February 2007 Available online 11 April 2007

Abstract An analysis of full-scale fire test experimental data is presented for a small compartment (3  3.6  2.3 m). A square steady fire source is placed in the center of the compartment. There is an open door and a horizontal opening in the roof, so that natural ventilation is established for the well-ventilated fire. A parameter study is performed, covering a range of total fire heat release rates (330, 440 and 550 kW), fire source areas (0.3  0.3 m and 0.6  0.6 m) and roof ventilation opening areas (1.45  1 m, 0.75  1 m and 0.5  1 m). The impact of the different parameters is examined on the smoke layer depth and the temperature variations in vertical direction in the compartment. Both mean temperatures and temperature fluctuations are reported. The total fire heat release rate value has the strongest influence on the hot smoke layer average temperature rise, while the influence of the fire source area and the roof opening is smaller. The hot smoke layer depth, determined from the measured temperature profiles, is primarily influenced by the fire source area, while the total fire heat release rate and the roof opening only have a small impact. Correlations are given for the hot smoke layer average temperature rise, the buoyancy reference velocity and the total smoke mass flow rate out of the compartment, as a function of the different parameters mentioned. Based on the experimental findings, it is discussed that different manual calculation methods, widely used for natural ventilation design of compartments in the case of fire, under-predict the hot layer thickness and total smoke mass flow rate, while the hot layer average temperature is over-estimated. r 2007 Elsevier Ltd. All rights reserved. Keywords: Full-scale experiments; Natural ventilation; Small compartment fire

1. Introduction



In the fire safety design of buildings, it is common practice to consider a smoke and heat exhaust ventilation system (SHEVS). The choice is made between either natural ventilation, relying on the buoyancy force of the hot smoke in case of fire, or a mechanical ventilation system. We focus on natural ventilation in the present study. Different manual design methods are at hand. In this paper, we consider:





The Belgian design standard method NBN S21-208-1 [1];

Corresponding author. Ghent University—UGent, Department of Flow, Heat and Combustion Mechanics, Belgium. Tel.: +32 9 264 33 14; fax: +32 9 265 35 75. E-mail address: [email protected] (B. Merci).

0379-7112/$ - see front matter r 2007 Elsevier Ltd. All rights reserved. doi:10.1016/j.firesaf.2007.02.003

The European proposed design calculation method CR 12101-5 [2], which is based on [3]; An American calculation method, described in [4].

Where the first two methods use the same assumptions and philosophy as starting point for the design procedure, the third method is somewhat different in the determination of the smoke mass flow rate from the design fire. In the present work, full-scale experiments are discussed of a well-defined, fuel-controlled fire in a small compartment. Openings, present in the roof, serve as natural SHEVS. Specific features of the configuration are:



the fire source size is relatively large, compared to the door opening, roof openings and compartment dimensions;

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Nomenclature Af Ai Aroof Av Ci Cv cp d D Dtc h H DHc HRR k Lfl _f m _s m Nu P

 

fire source area (m2) ventilation inlet area (m2) roof ventilation area (m2) ventilation outlet area (m2) discharge coefficient of inlet ventilation area (dimensionless) discharge coefficient of outlet ventilation area (dimensionless) specific heat capacity (kJ/(kg K)) hot smoke layer thickness (m) fire source hydraulic diameter (m) thermocouple diameter (m) convection coefficient (W/(m2 K)) compartment height (m) total heat of combustion (J/kg) heat release rate (kW) conduction coefficient (W/(m K)) flame height (m) fuel mass flow rate (kg/s) total smoke mass flow rate (kg/s) Nusselt number (dimensionless) fire source perimeter (m)

the heat release rate of the fire source per unit area is relatively high; there is asymmetry in the roof openings’ positions.

The first aim of the study is to investigate the influence of different parameters on the average temperature rise in the hot upper layer, as well as on the hot layer thickness. The investigated parameters are:

  

the total roof opening area; the fire source area; the fire heat release rate.

Correlations are derived as functions of these parameters. The second aim of the study is to investigate to what extent the outcome of the design calculation methods mentioned above, are supported by experimental findings for the configurations considered. Bearing in mind that in the present tests the fire heat release rate per unit area is high (ranging from about 1 MW/m2 to about 6 MW/m2) and the flames are high compared to the compartment height, it is investigated how well their outcome agrees with the experimental data. 2. Experimental set-up Fig. 1 shows the geometry of the experimental set-up. The top picture is a horizontal cross section at floor level. The middle and bottom pictures are vertical cross sections

Pr Q_ f q_ f Q_ s

T T0 Tamb Ttc Tg To Ts,av Ts,max TN Y zo e w ramb rs s

Prandtl number (dimensionless) total fire heat release rate (kW) total fire heat release rate per unit area (kW/m2) convective heat per time unit in the hot smoke layer (kW) temperature (K) temperature fluctuations rms value (K) ambient absolute temperature (K) temperature measured by the thermocouple (K) effective gas temperature (K) absolute reference temperature (K) average hot smoke layer temperature (K) maximum hot smoke layer temperature (K) effective ambient temperature for radiation correction (K) smoke-free height (m) virtual origin height (m) emissivity (dimensionless) factor of incompleteness of combustion (dimensionless) ambient density (kg/m3) smoke density (kg/m3) Stefan–Boltzmann constant ( ¼ 5.67E8 W/ (m2K4))

through the mid-planes of the compartment. They are front and side views, respectively. Objects that are not in the plane are shown as thin dashed lines. The compartment width is 3.6 m (x-direction), the depth is 3.0 m (y-direction) and the height is 2.3 m (z-direction). In the front wall, there is an open door of dimensions 0.9  2.0 m. The thermal properties of the walls and the ceiling, all 15-cm-thick cell concrete, are k ¼ 0.2 W/(m.K) (hardly variable in the wall temperature range of the experiments), cp ¼ 900 J/(kg.K) and r ¼ 2200 kg/m3. There are two openings in the roof (dashed lines in the top picture of Fig. 1) of size 0.75  1 m each. The openings are separated by a 0.25-m-wide beam. They are centrally positioned around x ¼ 1.8 m. The distance between the roof opening centers and the front wall is 1 m (so that the distance from the back wall is 2 m). As mentioned in the introduction, there is thus asymmetry in the openings’ position, being closer to the front wall. It was visually observed that the flames are directed towards the openings near the ceiling (thus: towards the front wall), while they were tilted backwards closer to the burner (due to incoming fresh air through the door opening). In the discussion of the results, this is not seen, since only temperatures and derived quantities are reported. The burner, representing the fire source, is positioned in the middle of the compartment. The burner dimensions are 0.6  0.6  0.4 m (height). The steady fire source is a hexane (C6H14) pool, lying on water, with the pool surface at height 0.3 m (i.e. 10 cm below the upper edge of the burner). The liquid hexane is fed into the burner by a

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1.8m 0.5m

3.0m

0.5m

0.45m

1.2m burner

1.5m 0.6m

1.5m

thermocouple trees

1.5m

0.9m y

0.9m roof openings 0.75m

x

1.8m

0.75m door 3.6m

0.5m

thermocouple trees 2.3m

2.0m z

0.9m

x

burner

3.6m

door

thermocouple trees

2.3m z y

burner

3.0m

Fig. 1. Geometry of the experimental set-up. Top: horizontal cross section; middle: vertical cross section (front view); bottom: horizontal cross section (side view).

volumetric pump. The fuel enters near the bottom of the burner, which is continuously filled with fresh water for cooling purposes. The cooling water (around 0.04 l/s) is supplied at constant temperature (around 15 1C). This water cools the thermocouple trees equipment and the burner. The water leaves the compartment at about 50 1C. Due to its lower density (rhex ¼ 660 kg/m3), the hexane rises to the upper water surface, where it is ignited. The theoretical heat of combustion is DHc ¼ 44 MJ/kg. The total fire heat release rate is then defined as _ f DH c . Q_ f ¼ m

(1)

Three fire heat release rates are imposed, by adjusting the fuel mass flow rate: Q_ f ¼ 330, 440 and 550 kW. In general, the fire heat release rate is defined as Q_ ¼ _ f DH c (e.g. [5,6]), where the factor w accounts for wm incompleteness of combustion. In the experimental set-up here, air is supplied under pressure inside the burner in order to increase completeness of combustion. This air, with mass flow rate in the order of 0.02 kg/s, enters through a series of small holes (diameter equal to 1 mm) at high

525

velocity, so that intense turbulent mixing with the hexane vapor is established. Moreover, the door, serving as inlet ventilation opening area, is large. It is discussed below that hot smoke flows out of the compartment through the upper part of the door opening. Still, under the assumption that only the lower 1.5 m of the door opening serves as fresh air inlet, the heat release rate limit for the fire to become ventilation-controlled is [4]: Q_ f ¼ 1260  ð0:9  1:5Þ 1:50:5 ¼ 2083 kW. The highest total fire heat release rate considered here is only 550 kW. For these two reasons, we assume that combustion is practically complete, so that, in the analysis of the results, we take w ¼ 1. We come back to this point in Section 3.8. In order to examine the influence of the roof opening area, the openings are partly covered for some sets of experiments. For one set, the right opening (looking from the door to the back wall) is completely covered by a plate, so that the area reduces from Aroof ¼ 2  (0.75  1.0 m) ¼ 1.5 m2 to Aroof ¼ 1  (0.75  1.0 m) ¼ 0.75 m2. For the third set, the left opening is also covered over a distance of 0.25 m from the left side, so that the roof opening area becomes: Aroof ¼ 1  (0.750.25 m)  1.0 m ¼ 0.5 m2. The roof covers do not warp as they are heated. Note that, as one opening is covered, there is more asymmetry in the configuration. The flames are sucked towards the uncovered opening, so that the geometrical asymmetry affects the flow field. As mentioned before, this is not directly reflected in the results reported in the present paper. Similarly, in order to study the influence of the fire source area or perimeter, the burner is partly covered. The area remains square and positioned in the center. The two fire source area values are: Af ¼ 0.6  0.6 m and 0.3  0.3 m. The compartment is in the open air, so that influence of wind is in principle inevitable. Care was taken that no experiments were conducted when the wind velocity exceeded 2 m/s. Most of the time, the wind velocity was well below 1 m/s. Moreover, a 3-m-wide wind obstruction was placed, at a distance of 2 m, in front of the door opening, so that the wind could not blow directly into the compartment. As such, the influence of wind is considered to be sufficiently small. Ambient temperature varied between 3 and 7 1C. Temperature measurements are performed in 4 thermocouple trees. One tree is positioned in the symmetry plane x ¼ 1.8 m, at a distance of 0.5 m from the door. The other three trees are positioned behind the burner, at a distance of 0.45 m from the back wall. The middle tree is in the symmetry plane, while the other two trees are at a distance of 0.5 m from the side walls. The coordinates are thus: x ¼ 0.5 m, y ¼ 2.55 m; x ¼ 1.8 m, y ¼ 2.55 m; x ¼ 3.1 m, y ¼ 2.55 m; x ¼ 1.8 m, y ¼ 0.5 m. Each tree contains 10 bare-bead K-type thermocouples, positioned at the following heights: 0.2, 0.7, 1.0, 1.15, 1.3, 1.45, 1.6, 1.75, 1.9 and 2.05 m. The number of thermocouple trees may seem rather limited. However, differences between measured temperatures are, after correction for the radiation effect, small.

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Moreover, numerical simulations confirm that, outside the fire plume, temperature variations in horizontal planes are only moderate (not shown). Therefore, we consider the number of thermocouple trees sufficient. 3. Analysis of the experimental data 3.1. Accuracy The basic quantity examined is temperature. As mentioned, the bare-bead thermocouples are K-type, with an error margin of 72 K. The back thermocouple trees are positioned in such a way that the thermocouples are not exposed directly to the flames. The tree near the door opening, on the other hand, could not be positioned in this manner; these thermocouples are directed towards the flames. The correction for the radiation effect is discussed in the next section. The duration of each experiment is 30 min, so that a thermal steady state is ensured (typically, the steady state was observed after 15 min for the configuration under study). Measurements are taken every 20 s. Only the values during the last 10 min of each experiment are used in the analysis below. As mentioned, the total fire heat release rate is controlled through adjustment of the fuel mass flow rate. To that purpose, the mass loss per minute in a fuel reservoir has been measured. The clock has an accuracy of 1 s while the mass is measured with 0.005 kg accuracy. 3.2. Temperature correction for radiation effect As mentioned, bare-bead thermocouples are used. Consequently, the temperature measurements must be corrected for the radiation effect (e.g. [7,8]). In order to estimate the correction, the steady-state heat balance equation is solved: hðT g  T tc Þ ¼

tc sðT 4tc



T 41 Þ,

(2)

where TN represents the effect ambient temperature for radiation. The convection coefficient is determined from Nu ¼ 2 þ 0:6 Re1=2 Pr1=3 ,

(3)

a correlation for flow around a sphere, reported by Drysdale [5]. The characteristic length scale is the thermocouple diameter (Dtc ¼ 1 mm). Properties for air are used. Numerical simulations reveal that, at the thermocouple positions, the gas velocity varies between 0.6 and 0.75 m/s. In combination with the temperature variations, this leads to convection coefficients in the range of h ¼ 140 W/(m2K), in the cold bottom layer, to h ¼ 180 W/(m2K), in the hot upper layer. The major difficulty in (2) is the estimation of TN. As mentioned in Section 2, the thermocouples in the three back trees are not directly exposed to the flames, which indeed remain between the burner and the back trees (so that the tree construction remains in between the thermo-

couples and the flames). As a consequence, those thermocouples ‘see’ the hot ceiling and the hot upper part of the walls, as well as the cold floor and cold lower part of the walls. The contributions are estimated as 50% each and the temperatures are taken equal to around 250 and 50 1C respectively. The thermocouples in the door tree on the other hand, see the flames directly. The flames are modeled here as a column of 0.6  0.6  1.5 m at 700 1C, positioned above the burner. Numerical simulations support this assumption (not shown here). The thermocouples also see the hot ceiling and the hot upper part of the walls, as well as the cold floor and cold lower part of the walls. For these contributions, the assumptions mentioned above are applied. Note that the door tree thermocouples do not see the door opening, the tree construction being positioned in between. Depending on the configurations, temperature corrections have been performed between +16 1C (in the hot layer) and 11 1C (in the cold bottom layer) for the back tree thermocouples, and between +5 and 34 1C for the door tree thermocouples. Interestingly, after correcting for the radiation effect, the temperatures at the door tree are quite close to the average values of the back tree temperatures. This confirms the already mentioned moderate temperature variations in horizontal planes (outside of the fire plume). 3.3. General observations Results are reported as time averages over the last 10 min for each experiment. Beside mean temperature values, temperature fluctuations are reported. The time-averaged rms fluctuation values are determined as qffiffiffiffiffiffiffi  0:5 rms ðT 0 Þ ¼ T 0 2 ¼ ðT  T amb Þ2  ðT¯  T¯ amb Þ2 , (4) where the over-bar denotes time-averaged values. Note that we do not claim to reproduce actual time-dependent temperature variations. To that purpose, the data sampling frequency would need to be an order of magnitude higher than what is applied, since the puffing frequency in the current configuration is in the order of 2 Hz. Still, the rms value of the temperature fluctuations, as defined in (4), provides insight into the temperature fluctuations with respect to the average value. Thus, while time accuracy is not obtained, the rms values remain relevant. Fig. 2 shows mean temperature profiles for the six configurations investigated. Each figure contains three profiles, one for each of the three total fire heat release rates examined (Q_ f ¼ 330, 440 and 550 kW). All profile shapes are similar, with a plateau of relatively low temperature below the height z ¼ 1.2 m (cold bottom layer). In each figure, there is a clear general shift towards higher temperatures as the fire heat release rate increases. The major reason is the moderate variation of the total mass flow rate of smoke through the openings as Q_ f varies, as discussed below. Comparing the figures in each column,

ARTICLE IN PRESS B. Merci, P. Vandevelde / Fire Safety Journal 42 (2007) 523–535

Af: 0.3m x 0.3m Aroof: 0.5mx 1.0m

350

330kW 440kW 550kW

X

200

X

X

X

X

150 X

100 X

50

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T-Tamb (K)

T-Tamb (K)

300 250

Af: 0.6m x 0.6m Aroof: 0.5mx 1.0m

350 X

527

X

330kW 440kW 550kW

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1.5

0

2

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1

z (m) Af: 0.3m x 0.3m Aroof: 0.75m x 1.0m

T-Tamb (K)

300

X

200

X X X

150 X

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50

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330kW 440kW 550kW

250

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330kW 440kW 550kW

250 X

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1.5

0

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0.5

1

Af: 0.3m x 0.3m Aroof: 1.45m x 1.0m

350 300

1.5

2

z (m)

z (m)

Af: 0.6m x 0.6m Aroof: 1.45m x 1.0m

350 300

330kW 440kW 550kW

250 X

200

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X

T-Tamb (K)

T-Tamb (K)

X

150

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2

Af: 0.6m x 0.6m Aroof: 0.75m x 1.0m

350

T-Tamb (K)

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z (m)

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1.5

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0

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0 0

330kW 440kW 550kW

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X

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z (m)

z (m) Fig. 2. Mean temperature profiles.

a decrease in the mean temperatures is observed as the roof opening increases. This is due to an increase in total smoke mass flow rate out of the compartment. Finally, comparison of the figures in each row reveals a general decrease in the peak temperatures as the fire source area increases, while the temperature profile is flattened near the ceiling. This is most probably due to the fact that, in the current experimental configuration, reaction takes place somewhat closer to the burner as Af increases, because more fresh air is supplied into the reaction zone by convection and diffusion. As a consequence, the fuel is consumed closer to the fire source and the peak temperature is somewhat lower. The effect of the fire source area on the hot smoke layer average temperature is much less clear, see Section 3.5. All the mentioned observations are now quantified in the following sections.

Fig. 3 reports the level of the temperature fluctuations. As mentioned, this is an indication for the local level of large-scale unsteadiness. Although no very clear distinct trends are observed, the highest values are found for most configurations between heights z ¼ 1.3 and 1.7 m. This corresponds to the transition zone between the cold bottom layer and the hot upper layer, which is also the region with the steepest mean temperature gradients (see Fig. 2). As a consequence, variations in time of the instantaneous hot smoke layer thickness result in the largest temperature variations in this region, so that indeed the highest level of temperature fluctuations due to unsteadiness is expected there. Inside each of the layers, where the temperature gradients are less steep, there is in general far less unsteadiness. In this sense, the position of maximum temperature fluctuations can serve as a good indication of the hot layer thickness.

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Af: 0.3m x 0.3m Aroof: 0.5mx 1.0m 330kW 440kW 550kW

20 X

X

X

10 X

X X

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rms(T') (K)

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X

X

10

X

0

X

0

0.5

X

X

X

1

X

1.5

X

2

z (m)

Fig. 3. Profiles of rms values of temperature fluctuations.

The temperature fluctuation intensity, defined as the local rms value of the temperature fluctuations (4), divided by the local mean temperature Ts (in K), does not exceed 5% (with only one exception, for Q_ f ¼ 550 kW, Af ¼ 0.3  0.3 m, Aroof ¼ 1.45 m2, z ¼ 1.75 m). Typical values are well below 3%, with peaks in the region mentioned above (between z ¼ 1.3 and 1.7 m).

ends for all configurations under consideration. Table 1 provides an overview of Ts for all the configurations under study. For each configuration, the influence of Q_ f on Ts is clear. It is seen that the increase of Ts,av with Q_ f is almost linear for Af and all three Aroof values. The correlation reads:

3.4. Influence of fire heat release rate

0:86 ðT s;av  T amb ÞQ_ f .

The influence of the total fire heat release rate Q_ f on the average hot smoke layer temperature is visualized in Fig. 4. This average temperature is determined from the mean temperature values at heights z ¼ 1.15 m and higher. Recall that, in Fig. 2, it was observed that z ¼ 1.15 m is the height where the low-temperature plateau

The reason is that the total mass flow rate of smoke out of the compartment, mainly through the roof openings but also partly through the upper part of the door, is not strongly affected by Q_ f , as discussed in Section 3.8. The exponent is slightly lower than 1, because there is a general tendency towards higher heat losses as Q_ f increases (see

(5)

ARTICLE IN PRESS B. Merci, P. Vandevelde / Fire Safety Journal 42 (2007) 523–535 Af : 0.3m x 0.3m

250

Af : 0.6m x 0.3m

250

200

529

200

150

X

Ts,av (°C)

Ts,av (°C)

X

X 100

50

X

0.5m x 1.0m 0.75m x 1.0m 1.45m x 1.0m

150

X

X X

100 0.5m x 1.0m 0.75m x 1.0m X 1.45m x 1.0m

50

0

0 0

200

400

600

0

200

HRR (kW)

400

600

HRR (kW)

Fig. 4. Influence of total fire heat release rate on hot layer average temperature.

Table 1 Hot smoke layer average temperature (Ts,avTamb) (1C) Af (m2) HRR ¼ 330 kW

HRR ¼ 440 kW

HRR ¼ 550 kW

Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 0.09 0.36

148 139

140 135

125 125

175 175

168 162

Section 3.9, Table 6). In particular, flames were observed through the roof openings from time to time, especially for the higher Q_ f values.

150 145

ðT s;av  T amb ÞA0:24 roof .

3.6. Influence of fire source area In general, a moderate decrease in Ts,av is observed as Af increases. This is due to an increase in total smoke mass flow rate: more air can be entrained in the plume above the fire source, due to the larger fire perimeter. The effect is most pronounced for the highest heat release rate value. Note, however, that the decrease is not spectacular, despite an increase of a factor 4 in burner area (a factor 2 in fire perimeter). For the lowest heat release rate examined, the impact of the fire source area is hardly visible, if not reverse.

182 163

From the mean temperature profiles, the hot smoke layer thickness can be determined as described in [9] RH

(6)

197 184

3.7. Hot layer thickness

3.5. Influence of roof opening area Fig. 5 shows the influence of the roof opening area on the hot layer average temperature rise. Ts,av decreases as Aroof increases, primarily due to the increase in mass flow rate of smoke through the larger roof opening, as discussed in Section 3.8. The relationship can be quantified as

211 203

d¼

0

ðT  T 0 Þ dh . T av  T 0

(7)

This formula transforms the measured temperature curve into a top hat curve, with conservation of the area under the profiles. In (7), T0 is a reference temperature, for which the ambient temperature is used in [9]. In the current experiments there is a global temperature rise due to the high heat release rate values in a relatively small compartment. This explains the temperature rise from z ¼ 0.2 m to the plateau, extending to about z ¼ 1.15 m (see Fig. 2). Therefore we define T0 as the average temperature measured by the thermocouples at z ¼ 1 m, positioned inside the low-temperature plateau region. The result of integration (7) is then given in Table 2. The following observations are made. First of all, there is relatively little influence of the roof opening area on the hot layer thickness. This is somewhat surprising at first sight. Note, however, that the parameter Aroof/Afloor varies only between 4% and 14%, which is a relatively narrow range. Moreover, the presence of the roof opening is felt directly in the compartment only locally under the

ARTICLE IN PRESS B. Merci, P. Vandevelde / Fire Safety Journal 42 (2007) 523–535

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Af : 0.3m x 0.3m X

Ts,av (°C)

200

X

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X

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Ts,av (°C)

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X

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X

0

0 0

1

0.5

1.5

0

0.5

Aroof (m2)

1

1.5

Aroof (m2)

Fig. 5. Influence of roof opening area on hot layer average temperature.

Table 2 Hot smoke layer thickness d (m) Af (m2) HRR ¼ 330 kW

HRR ¼ 440 kW

HRR ¼ 550 kW

Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 0.09 m2 0.87 0.36 m2 0.91

0.83 0.92

0.87 0.96

0.86 0.94

0.86 0.94

0.91 0.96

0.89 0.95

0.92 0.96

0.95 0.97

Table 3 Buoyancy reference velocity (8) (in m/s) Af (m2) HRR ¼ 330 kW

HRR ¼ 440 kW

HRR ¼ 550 kW

Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 0.09 0.36

1.88 1.88

1.80 1.86

1.70 1.85

2.00 2.09

1.97 2.03

openings. This is confirmed by numerical CFD simulations (not shown). Since the thermocouple trees are not positioned directly under the roof openings, this direct effect is not reflected in the measurements, which also explains why the value of d is relatively insensitive with respect to Aroof. Second, the influence of the fire heat release rate on the hot layer thickness is modest, too, although there is a tendency towards thicker hot upper layers for higher HRR values, with fixed Aroof and Af. Finally, there is a general increase in hot layer thickness as Af increases. The reason is the increase in total smoke mass flow rate for larger Af (due to increased entrainment by the larger fire source perimeter, as already mentioned). This is in line with the explanation, given in Section 3.3, for the global decrease of the peak temperature as the fire source area increases. To conclude this section, we remark that the hot smoke layer is clearly much thicker than the distance between the ceiling and the top of the door (equal to 0.3 m). This shows that smoke flows out of the compartment through the upper part of the door. This is confirmed in CFD simulations (not shown).

1.93 1.96

2.18 2.22

2.16 2.15

2.13 2.06

3.8. Estimation of total smoke mass flow rate With the information of Tables 1 and 2, the buoyancy reference velocity can be computed as in [10] sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi gdðT s;av  T amb Þ Ub ¼ . (8) 1 2 ðT s;av þ T amb Þ This is a measure for the upward velocity, generated by the buoyancy force. Table 3 gives the values for all the configurations. The buoyancy reference velocity clearly increases as Q_ f increases. The average correlation, obtained from Table 3, is 0:33 U b Q_ f .

(9)

The velocity slightly decreases as the roof opening becomes larger: U b A0:06 roof .

(10)

Note that the dependence of Ub on Aroof is much weaker than the dependence of Ts,av on Aroof (exponent 0.06 in (10) versus 0.24 in (4)).

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Table 4 Toal smoke mass flow rate (11) out of the compartment (in kg/s) Af (m2) HRR ¼ 330 kW

HRR ¼ 440 kW

HRR ¼ 550 kW

Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 0.09 0.36

0.98 1.01

1.19 1.24

1.85 2.01

0.98 1.03

1.22 1.27

It is remarked that Ub is relatively insensitive to Af. This means that the fire perimeter does not determine the global buoyancy force. Note that the tendencies observed in Table 3 are very similar to what was observed in Table 1, mainly because the variation of the hot smoke layer thickness d (Table 2) is relatively small. From velocity (8), the total smoke mass flow rate through the total opening area, i.e. the roof opening and part of the door opening, can be estimated, as explained in [10] _ s ¼ rs C v A v U b . m

(11)

We use the common value Cv ¼ 0.6. Note, with respect to expression (11), that only outflow was observed through the roof openings during the tests. This has been confirmed in numerical simulations (not shown). We recall that smoke flows out of the compartment through the roof opening(s), but also through the upper part of the door. For the evaluation of Av, we assume here that the hot layer thickness provides an indication for the neutral plane height in the door. For an estimation of the total smoke mass flow rate and the investigation of tendencies, this approximation is considered sufficiently accurate. Table 4 gives the resulting mass flow rates for all the configurations. Table 4 confirms the statements made above:

  

the total smoke mass flow rate slightly increases with Q_ f ; it increases as Aroof increases; it slightly increases as Af increases.

1.97 2.03

_ s cp ðT s;max  T amb Þ, Q_ s ¼ m

(16)

_ s from (11). The maximum temperature difference with m (Ts,maxTamb) is used, because the smoke leaves the roof opening with this temperature. Note that this is not true for the smoke leaving through the door, but on the other hand the temperatures are not measured in the roof opening, but at height z ¼ 2.05 m, so that this compensates the differences to some extent. The remainder of the total fire heat release rate is included in the ‘loss’ term Q_ loss :

 



0:09 _ s Q_ t , m

(12)



_ s A0:64 m roof ,

(13)

_ s A0:06 m f .

(14)

(15)

2.03 2.04

Table 5 provides the maximum temperatures in the hot layer for all configurations. This corresponds to the values at z ¼ 2.05 m. From Tables 4 and 5, an approximate thermal balance can be computed. The theoretical total input of heat per time unit is given by expression (1). Part of this input heat release rate heats up the hot smoke gases. This heat per time unit, convected out of the compartment by the hot gases, is computed as



_ s P0:12 . m

1.26 1.29

3.9. Approximate thermal balance

The roof opening area Aroof is clearly the dominant parameter. The corresponding correlations are:

Correlation (14) can also be expressed in terms of the fire source perimeter:

0.99 1.03

radiative heat losses from the flames and the hot smoke layer due to absorption in the walls and the ceiling; radiative heat losses through the roof opening and the door; note that, from time to time, flames were observed through the roof openings, leading to an increase in this loss term; the heat losses per time unit convected from the hot smoke layer to the walls and the ceiling, approximated as Q_ w;c ¼ hi ðAwalls ðT s  T walls Þ þ Aceiling ðT s  T ceiling ÞÞ; it is not easy to estimate this loss, but it is relatively small, as is easily seen when realistic values are used for the different terms; losses due to incompleteness of combustion: _ f DH c ; ð1  wÞm enthalpy loss due to heating up of the cooling water.

Table 6 provides the different terms for all configurations, as percentage fractions of the total heat release rate (1). It is immediately clear that, within the assumptions made, the heat per time unit convected out of the compartment by the total smoke mass flow rate, increases as Aroof increases, as could be expected from expressions (6) and (13). It is not entirely possible to identify the major term in Q_ loss . It is clear that Aroof determines the difference, but the

ARTICLE IN PRESS B. Merci, P. Vandevelde / Fire Safety Journal 42 (2007) 523–535

532

Table 5 Hot smoke layer maximum temperature (Ts,maxTamb) (1C) Af (m2) HRR ¼ 330 kW

HRR ¼ 440 kW

HRR ¼ 550 kW

Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 0.09 0.36

231 216

208 198

171 172

268 255

241 228

205 193

311 301

286 267

260 238

Table 6 Different heat terms as fractions of Q_ f Q_ f ¼ 440 kW

Q_ f ¼ 330 kW

Q_ f ¼ 550 kW

Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Af ¼ 0.09 m2 Q_ s =Q_ f 0.60 Q_ l =Q_ f 0.40

0.65 0.35

0.85 0.15

0.52 0.48

0.58 0.42

0.81 0.19

0.48 0.52

0.58 0.42

0.82 0.18

Af ¼ 0.36 m2 Q_ s =Q_ f 0.57 Q_ l =Q_ f 0.43

0.64 0.36

0.90 0.10

0.53 0.47

0.56 0.44

0.77 0.23

0.50 0.50

0.53 0.47

0.76 0.24

mechanism behind is not fully clear. In the authors’ opinion, most probably the radiative losses are dominant. These radiative heat losses reduce due to lower gas temperatures by the increase of entrained fresh air into the compartment as Aroof increases. Note that there could also be incompleteness of combustion: as the roof opening increases, more fresh air is entrained through the door, so that w increases. However, recall that the fire is clearly fuel-controlled (see Section 3.3) and that special care has been taken to stimulate reaction through pressurized air supply in the burner, so that this assumption is by far less likely. 4. Relation to manual calculation methods As mentioned in the introduction, one of the reasons for the current study is to test to what extent the philosophy behind some existing manual calculation methods remains valid for the configurations tested. Three calculation methods are considered: the Belgian standard NBN S21208-1 [1]; the European method CR12101-5 [2]; and the American method as described in [4]. It is noted that the fire heat release rate per unit area is high (ranging from about 1 MW/m2 to about 6 MW/m2) and the flames are high compared to the compartment height. Still, the manual calculation methods are sometimes used in practice outside their application range. Therefore, we want to investigate their possible validity and estimate the error. The starting point of the three methods mentioned is the prescription of the smoke-free height Y. For a prescribed design fire, the calculation methods then provide the total area required for natural ventilation. The situation is different here: we have a steady fire source and a fixed

geometrical configuration and determine the corresponding smoke-free height for each method. We compare this to the experimental data. In NBN S21-208-1 and CR12101-5, the fire source perimeter is the major property in the determination of the smoke mass flow rate at a height Y: 1=2

_ s ¼ C e PY 3=2 ; Y p10Af . m

(17)

The value of Ce is Ce ¼ 0.188 in NBN S-21-208-1, while it depends on the configuration in CR12101-5. For the present configuration, it is Ce ¼ 0.19, because the fire source hydraulic diameter (0.3 or 0.6 m) is smaller than one fifth of the largest compartment dimension (3.6 m). We come back to this point below. The hot layer average temperature rise is then computed: T s;av  T amb ¼

0:8Q_ t . _s cp m

(18)

Note that the compartment under study is not sprinklered, which explains the factor 0.8 in the numerator. Also note that the implicit assumption in (18) is that 80% of the total fire heat release rate goes to the hot smoke layer. According to CR12101-5, the total required ventilation area is then determined as Av;tot C v ¼ 

_ s T s;av m  2 1=2 , _ s T s;av T amb =ðAi C i Þ2 2r2amb gdðT s;av  T amb ÞT amb  m

(19) while NBN S21-208-1 stipulates: " #1=2  2 _ s T 2s;av þ Av C v =Ai C i T amb T s;av m   Av C v ¼ , ramb 2gd T s;av  T amb T amb which is solved in an iterative manner.

(20)

ARTICLE IN PRESS B. Merci, P. Vandevelde / Fire Safety Journal 42 (2007) 523–535

Some preliminary remarks may be made:





the fire heat release rate does not enter the expression for the total smoke mass flow rate (expression (17)); the experimental study supports this choice: there is only a _ s on the HRR (see expression very weak dependence of m (12)); the fire heat release rate determines the hot layer average temperature rise (expression (18)); this is again supported by the experimental results, although the relation is not perfectly linear (expression (5)).

In [4], the formula for the smoke mass flow rate is different, based on [11]. This entrainment model accounts for differences in entrainment in the flame region and in the plume region. The flame height is defined as Lfl ¼ 0:235ð0:7Q_ f Þ2=5  1:02D. The total smoke mass flow rate at height z is then determined as 8 _ s ¼ 0:071ð0:7Q_ f Þ1=3 ðz  zo Þ5=3 m > > < z4Lfl ; þ0:00192ð0:7Q_ f Þ; (21) > > :m _ s ¼ 0:0059ð0:7Q_ Þ z ; zoLfl ; f Lfl

where zo is the fire source virtual origin height: 2=5 zo ¼ 0:083Q_ t  1:02D,

(22)

with D the fire source hydraulic diameter. In ISO/FDIS 16734:2004(E) [12], the same formulae are applied as in [4] at and above the flame height. Note that expression (21) implies that only 70% of the total fire heat release rate goes

533

into the smoke layer (compared to 80% in (18)), so that for the same value of total smoke mass flow rate, the temperature rise will be smaller than with NBN S21-2081 or CR12101-5. We also note that expressions (21) and (22) have been derived for entrainment in plumes from a point source assumption, which may be questionable since the fire source area is relatively large compared to the compartment dimensions (particularly the compartment height). The introduction of the virtual origin zo in principle accounts for this effect. Finally, it should be noted that in (21) both the geometry and the heat release rate of the fire determine the total smoke mass flow rate. This is, for the configuration under study, not supported by _ s is relatively unaffected by the experimental results, since m the HRR, as already mentioned. Expression (18) is also applied in [4]. For natural ventilation, the extraction area is determined as " #1=2  2 T s;av þ Av =Ai T amb _s m Av C v ¼ , ramb 2gdðT s;av  T amb ÞT amb =T s;av

(23)

which is, under the assumption that Ci ¼ Cv, identical to expression (20). Now the smoke-free height can be determined for all the configurations under study. With Ci ¼ Cv ¼ 0.6, Ai decreases as Y becomes lower than the door height (i.e. lower than 2.0 m), while Av then increases accordingly. Here, we do not consider tenability conditions with respect to temperature or toxicity [13,14]. Tables 7–9 give an overview of the hot layer thickness, the hot layer average

Table 7 Hot layer thickness (m) with the manual calculation methods Af (m2) 330 kW

440 kW 2

2

Aroof ¼ 1.5 m

Aroof ¼ 0.5 m

NBN S21-208-1/CR12101-5 0.09 0.41 0.30 0.36 0.64 0.55

0.12 0.36

Klote-Milke [4] 0.09 0.45 0.36 0.54

0.17 0.25

Aroof ¼ 0.5 m

Aroof ¼ 0.75 m

2

0.36 0.45

550 kW 2

2

2

Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2

Aroof ¼ 0.75 m

Aroof ¼ 1.5 m

0.42 0.64

0.31 0.54

0.12 0.35

0.43 0.64

0.33 0.55

0.12 0.34

0.49 0.57

0.38 0.47

0.19 0.27

0.51 0.59

0.41 0.49

0.20 0.29

Table 8 Hot layer average temperature rise (Ts,avTamb) (1C) with the manual calculation methods Af (m2) 330 kW

440 kW 2

2

Aroof ¼ 1.5 m

Aroof ¼ 0.5 m

NBN S21-208-1/CR12101-5 0.09 446 409 0.36 271 250

360 214

Klote-Milke [4] 0.09 208 0.36 177

169 144

Aroof ¼ 0.5 m

Aroof ¼ 0.75 m

2

194 165

550 kW 2

2

2

Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2

Aroof ¼ 0.75 m

Aroof ¼ 1.5 m

599 361

550 331

480 283

755 451

693 417

600 352

259 220

239 205

208 179

301 258

281 241

243 211

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Table 9 Total smoke mass flow rate (kg/s) with the manual calculation methods Af (m2) 330 kW

440 kW

550 kW

Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 NBN S21-208-1/CR12101-5 0.09 0.59 0.64 0.36 0.98 1.06

0.73 1.23

0.59 0.98

0.64 1.06

0.73 1.24

0.58 0.98

0.64 1.06

0.73 1.25

Klote-Milke [4] 0.09 1.11 0.36 1.31

1.37 1.60

1.19 1.40

1.29 1.50

1.48 1.72

1.28 1.50

1.37 1.60

1.58 1.82

1.19 1.40

Table 10 Hot layer thickness (m), average temperature rise (Ts,avTamb) (1C) and total smoke mass flow rate with CR12101-5 and Ce ¼ 0.34 Af (m2) 330 kW

440 kW

550 kW

Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 Aroof ¼ 0.5 m2 Aroof ¼ 0.75 m2 Aroof ¼ 1.5 m2 d (m) 0.09 0.36

0.60 0.91

Dt (1C) 0.09 292 0.36 197 _ s (kg/s) m 0.09 0.90 0.36 1.34

0.50 0.84 268 183 0.99 1.44

0.31 0.70 230 160 1.14 1.65

0.60 0.90 389 260

0.50 0.82 357 240

0.90 1.35

temperature and the total smoke mass flow rate for each calculation method. It is immediately clear that the calculated hot layer thickness is much smaller and the average temperature rise is higher than what has experimentally been determined (Tables 1 and 2) for all three calculation methods and for all configurations. The over-prediction of (Ts,avTamb) is primarily due to under-estimation of heat losses: as mentioned, the manual calculation methods assume that 80% or 70% of the total fire heat release rate goes into the hot layer, while Table 5 reveals that in the experiments the percentages are lower, except for the largest roof openings. However, for these configurations, the total smoke mass flow rate is seriously under-estimated in the calculations, as is seen from Tables 4 and 9. The temperature prediction is much worse with NBN S21-208-1 and CR12101-5 than with the method of [4], because of the lower assumed heat loss and the lower mass flow rate predictions. From Tables 2 and 7 the observations for the hot layer thickness d are:

  

d decreases as Aroof increases; this is not in line with the experimental findings; d increases much more strongly with Af than experimentally observed; d is independent of Q_ f in NBN S21-208-1 or CR12101-5 and varies only very little with the method of [4], in line with the experimental results.

0.30 0.66 305 205

0.99 1.47

1.15 1.74

0.61 0.89 491 322 0.90 1.37

0.50 0.81 447 296 0.99 1.48

0.30 0.64 381 252 1.15 1.75

From Tables 1 and 8 the observations with respect to the hot layer average temperature rise are:

  

(Ts,avTamb)strongly decreases as Aroof increases, in line with the experimental findings; (Ts,avTamb)decreases much more strongly as Af increases than what is experimentally observed; (Ts,avTamb)varies practically linearly with Q_ f , in line with the experiments. Finally, Tables 4 and 9 reveal that:



 

_ s is under-predicted with NBN S21-208-1 or CR12101m 5, especially for the smaller Af and the larger Aroof; with _ s is only under-predicted for the the method of [4], m largest Aroof; for each of the calculation methods, the _ s with Aroof is too slow, compared to the variation of m experimental results; _ s increases as Af increases, in line with the experiments; m _ s is practically independent of Q_ f with NBN S21-208-1 m or CR12101-5, in agreement with the experiments; with _ s slightly increases as Q_ f increases. the method of [4], m

To conclude this section, results with CR12101-5 are examined when the value Ce ¼ 0.34 is used. From _ s will increase (for constant expression (17) it is seen thatm smoke-free height), so that better agreement with experimental data can be expected. Note that, particularly for

ARTICLE IN PRESS B. Merci, P. Vandevelde / Fire Safety Journal 42 (2007) 523–535

Af ¼ 0.36 m2, the configuration is relatively close to the situation where Ce ¼ 0.34 is to be used in CR12101-5: there is only ventilation from one side of the fire source (the open door) and the fire source hydraulic diameter is not much smaller than one fifth of the largest compartment dimension. Table 10 shows that agreement with experimental data in general clearly improves, compared to Ce ¼ 0.19: d _ s increases and (Ts,avTamb) decreases. The increases, m general remarks, made above for CR 12101-5, remain true, so that agreement is clearly not perfect. Still, the value Ce ¼ 0.34 seems by far more suitable for the configurations under study. This indicates that further research is required with respect to entrainment modeling for large fires in small compartments. 5. Conclusions Experimental data have been presented for large-scale fire tests in a small compartment with an open door on one side of the compartment and a ventilation opening in the roof. The following aspects have been illustrated:







 

The hot layer average temperature rise (Ts,avTamb): J increases almost linearly with the total fire heat release rate Q_ f , because the total mass flow rate of _ s is not strongly smoke out of the compartment m affected; J decreases as the roof opening area Aroof increases, _ s increases; because m J generally decreases slightly as the burner area Af increases, in particular for the highest Q_ f values, _ s increases; this effect is small and for the because m lower Q_ f values hardly visible. The hot layer thickness d: _f; J is practically unaffected by Q J is practically unaffected by Aroof; J becomes thicker as Af increases, due to increased entrainment. The total smoke mass flow rate out of the compartment _ s: m _ ; J is practically unaffected by Q f J increases as Af increases, due to increased entrainment; Is practically unaffected by Aroof. The temperature fluctuation intensities remain below 5% (and almost always below 3%), with maximum values in the transition region between the cold bottom layer and the hot upper layer, in the zone where the steepest mean temperature gradients are observed.



535

Except for the largest Aroof, more than 35% (and typically more than 40%) of Q_ f is lost and does not go into the hot smoke layer.

Analysis of the manual calculation methods, applied to the configurations studied, reveals that:

  

the hot layer thickness is under-estimated; the hot layer average temperature rise is over-predicted; the total smoke mass flow rate is in general underpredicted.

Finally, it was illustrated that, although the configurations do not fulfill the required conditions, the value Ce ¼ 0.34 is to be preferred in CR12101-5 over Ce ¼ 0.19, because entrainment is seriously under-estimated with the latter value. Acknowledgments This research was partly financed by Ghent University— UGent through BOF project 011/013/04. The authors greatly appreciate the efforts of Martin, Jurgen and Pascal in setting up the experiments. References [1] NBN S21-208-1, BIN-IBN, Brussels, Belgium, 1995. [2] CR12101-5, CEN report, Brussels, Belgium, 2000. [3] Morgan HP, Ghosh BK, Garrad G, Pamlitschka R, De Smedt J-C, Schoonbaert LR. Design methodologies for smoke and heat exhaust ventilation. BRE-FRS; 1999. [4] Klote JH, Milke JA. Principles of smoke management. Society of Fire Protection Engineers; 2002. [5] Drysdale D. An introduction to fire dynamics. West Sussex: Wiley; 2003. [6] Karlsson B, Quintiere JG. Enclosure fire dynamics. London: CRC Press; 2000. [7] Heitor MW, Moreira ALN. Prog Energy Combust Sci 1993;19: 259–78. [8] Brohez S, Delvosalle C, Marlair G. Fire Safety J 2004;39:399–411. [9] Thomas PH, Hinkley PL, Theobald CR, Simms DL. Investigations into the flow of hot gases in roof venting. Fire research technical paper no. 7. London: The Stationary Office, 1963. [10] Etheridge DW. Build Environ 2002;37:1057–72. [11] Heskestad G. SFPE handbook of fire protection engineering. 3rd ed. NFPA; 2002. p. 2-1–2-17. [12] ISO/FDIS 16734:2004(E), Fire safety engineering—requirement governing algebraic formulas—fire plumes. ISO TC 92/SC 4/WG 9, 2004. [13] CIBSE Guide E—fire engineering, 1997. [14] Delichatsios M. Fire Saf J 2004;39:643–62.