fire in deep enclosures

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FIRE AND MATERIALSFire Mater. 2009; 33:157–185Published online 20 November 2008 in Wiley InterScience (www.interscience.wiley.com). DOI: 10.1002/fam.986

An experimental study of fire development in deep enclosures

and a new HRR–time–position model for a deep enclosurebased on ventilation factor

Khalid A. M. Moinuddin∗,† and Ian R. Thomas

Centre for Environmental Safety and Risk Engineering, Victoria University, P.O. Box 14428,

Melbourne, Vic. 8001, Australia

SUMMARY

The behaviour of fire within a deep (high-depth-to-height ratio) enclosure with various openings at oneend has been studied experimentally. Sixteen fuel trays were placed within an 8 .0m long×2.0m wide×0.6m high steel enclosure. The experiments confirmed previous smaller-scale experiments, which showedthat the fires in deep enclosures are strongly influenced by the ventilation and are not at all uniformthroughout the depth of the enclosure. The severity of exposure of structural members is much moresevere close to the ceiling near the front of the enclosure compared with the back of the enclosure.Based on the test data that can be found in this report, as well as other well-known tests, a heat releaserate–time–position model is proposed for fire development in deep enclosures. The same sets of test dataare also used to evaluate other model predictions. Finally, a recommendation is made to choose a modelbased on the shape of the enclosure (i.e. deep, wide, square, etc.). Copyright q 2008 John Wiley & Sons,Ltd.

Received 15 November 2007; Revised 12 June 2008; Accepted 9 September 2008

KEY WORDS: burning rate; mass loss rate; deep enclosure; ventilation factor; fire severity

1. INTRODUCTION

Enclosures in buildings occur in a wide range of sizes and shapes. Many rooms in residential

accommodation and some commercial accommodation (offices, etc.) are roughly cube shaped ( D,

W , H all similar magnitude). However, in many buildings the spaces can be very wide, deep

or both in comparison with their height. For example, many open plan office storeys can have a

∗Correspondence to: Khalid A. M. Moinuddin, Centre for Environmental Safety and Risk Engineering, VictoriaUniversity, P.O. Box 14428, Melbourne, Vic. 8001, Australia.

†E-mail: [email protected]

Copyright q 2008 John Wiley & Sons, Ltd.



158 K. A. M. MOINUDDIN AND I. R. THOMAS

Figure 1. Enclosure details.

floor to ceiling height of about 2.4 m, and a floor-to-floor height of about 3.0 m, but may be up to40 m or even 60 m square in plan. If, in a fire situation, they are ventilated only from one side,

the depth-to-height ratio can be of the order of 10–25. Ceiling spaces can be even more extreme.

This means that the behaviour of fires in wide and deep enclosures needs to be understood and

appropriate fire severity correlations of heat release rate (HRR) vs time vs position need to be

developed to make an input to the zone fire models such as the CFAST [1].The following terminology and nomenclature are used for the clear internal dimensions of the

enclosure (Figure 1):

• Width (W )—horizontal dimension parallel to the plane of the ventilation opening.

• Depth ( D)—horizontal dimension perpendicular to the plane of the ventilation opening.

•Height ( H )—vertical dimension from the bottom surface to the top surface.

The following terminology and nomenclature are used for the dimensions of the ventilation opening:

• Opening width (w)—the clear horizontal dimension.

• Opening height (h)—the clear vertical dimension.

The current experimental program was designed to investigate the influence of variation in the

opening size and shape on fire in a deep enclosure ( D/ H ∼13), and to assess the fire severity (in

relation to damage to structural members, etc.) at various locations in the enclosure. An effort has

also been made to develop correlation based on the shape, size and ventilation of the enclosures

using a number of experimental studies reported in [2–4]. The deep enclosures can be classified

into two categories:

•Deep enclosure D/ H 2 and D/W 2.

• Deep but square/wide D/ H 2 and D/W <2.

The main focus of this paper is on the first category; however, the latter category will also be

partially discussed.

In CIB tests [2], the ranges of D/W were 0.5, 1 and 2 and that of D/ H were 1, 2 and 4. On the

contrary Cardington tests [3] had D/W of 4 and D/ H of ∼8. The research programme reported

Copyright q 2008 John Wiley & Sons, Ltd. Fire Mater. 2009; 33:157–185

DOI: 10.1002/fam



AN EXPERIMENTAL STUDY OF FIRE DEVELOPMENT 159

in this paper extended experimental research reported previously [4], which were conducted in

an enclosure 1.5 m×0.6m×0.3 m high ( D/W of 2.5 and D/ H of 5) and each test had a single

ventilation opening with the full height of the enclosure and either 33, 50 or 100% of the width of

the front of the enclosure. It was found that the behaviour of a fire in an enclosure with a full width

(100%) opening is substantially different from that in an enclosure with a partial width opening.

None of the studies reported in [2–4] had the HRR measurements using oxygen calorimetry [5].Furthermore, during Cardington tests [3], the mass loss rate was also not recorded.

It was considered important to conduct deep enclosure tests at a large scale with higher D/ H

and more variations in opening width and measure the HRR using oxygen calorimetry.

Although the current tests have been simulated numerically using Fire Dynamics Simulator

version 4 (FDS4) [6], due to the space constraint the results are not discussed in this paper. A

huge uncertainty is observed in FDS4 modelling of these ventilation-controlled fire experiments,

especially while the HRR is not prescribed. In the follow up of this paper, the experimental data

will be compared with the results of the numerical simulation.

2. EXPERIMENTAL PROGRAMME

The enclosure was 8.0 m long, 2.0 m wide and 0.6 m high ( D/W of 4 and D/ H of ∼13) and the

enclosure was constructed of 1.3 mm thick sheet steel except as follows. One 2.0 m side of the enclo-

sure (Figure 2) was used as the opening (for ventilation) by varying the construction of this wall.

Thus, scaling parameters Afuel-surface/ Afloor-area [7], D/W , D/ H ,W / H and At/Openingfully open

of the current large-scale test are 1.7, 1.6, 2.7, 1.7 and 2.2 times greater than those of previous

small-scale experiments (where, At= total inside surface area of the enclosure). All other wall and

Figure 2. Schematic diagram of experimental test configuration.


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Table I. Schematic of deep enclosure test openings.

OpeningTest name configuration

Fully open end ( A= H ×W )

Half open top ( A= H /2×W )

Half open bottom ( A= H /2×W )

Half open side ( A= H ×W /2)

Half open centre ( A= H ×W /2)

Half closed centre ( A= H ×W /2)

Quarter open centre ( A= H ×W /4)

Quarter open side ( A= H ×W /4)

fuel properties (and related dimensionless parameters) were identical. The sides other than the

opening side were completely closed; however, there were three windows (of fire-resistant glass)

in the walls that allowed details of the shape and behaviour of the fire to be viewed.

The opening size was varied with eight different test configurations as shown in Table I. Thevariation is designed for both height and width of the opening. While the opening height was varied

(50 or 100% of the height of the front of the enclosure), the opening width was kept constant i.e.

full width of the front of the enclosure. On the other hand, while the opening width was varied (25,

50 or 100% of the width of the front of the enclosure), the opening height was the full height of

the enclosure. This was done to isolate the effect of variation in one dimension (i.e. either height

and width). In addition, the location of the opening was also varied to examine the effect of such

variation. The specified opening existed throughout a test allowing free passage of entering air and

the outgoing products of combustion, which were then collected in the hood of an oxygen (O 2)

calorimeter [5] for HRR estimation. In the O2 consumption system, the oxygen, carbon dioxide

(CO2) and carbon monoxide (CO) metres were first zeroed with pure nitrogen gas. The metres

were calibrated at approximately 80% with a gas mixture (known as span or calibration gas) to

the scaled range of 0–25, 0–10 and 0–1% for O 2, CO2 and CO, respectively.Four-type K thermocouples were placed above the centre of each tray as shown in Figure 2.

The thermocouples on the enclosure roof were spot-welded to the steel and were intended to

measure the steel temperature while the other thermocouples were used to record gas tempera-

tures. No correction of the gas temperature, related to radiation, was incorporated. Three water-

cooled Gordon-type infrared radiometers were placed on the centreline of the floor as shown


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in Figure 2. Each radiometer was equipped with a sapphire window for elimination of convective

heat flux. The view angle of the radiometers, with the sapphire window installed, was 150◦. The

tests were also video recorded.

The fuel used was commercial grade methylated spirits, a liquid consisting of 97% ethanol and

3% water. During the test, 16 fuel trays were placed within the enclosure in eight rows and two

columns as shown in Figure 2. Trays 15 and 16 were located adjacent to the opening (front of the

enclosure) and these formed Row 1. The fuel trays were 0.90 m×0.90m×0.05 m deep and were

constructed of 3 mm thick steel. Each tray was rested on a weigh scale so that the mass of the fuel

could be measured throughout each test. Ignition was primarily at the front of the opening (Trays

15 and 16). However, some of the tests were repeated with ignition at the rear of the enclosure (on

Tray 2) using a gas flame inserted via a small hole that was quickly sealed following ignition. Fuel

was ignited at the rear of the enclosure to confirm the behaviour of fire progression towards the

front of the enclosure without burning all of the fuel that was observed in [3, 4]. Before ignition,

the height and mass of fuel in each tray were approximately 6 mm and 4 kg, respectively. The total

mass of fuel during each test was approximately 64 kg and the fuel was at ambient temperature

before ignition.

3. EXPERIMENTAL RESULTS

3.1. Burning behaviour

The burning behaviour in these deep enclosures was found to be identical to that reported in [4].In general, when the fuel in a rear tray (Tray 2) was ignited the flame-front moved quickly to the

front of the enclosure. This behaviour varied slightly depending on the ambient temperature at the

commencement of the test. In tests conducted with the ambient temperature below about 15 ◦C

it was possible to ignite the fuel in an individual tray (Tray 2), and the flame could be observed

moving across that tray, then to Tray 1, and once trays (Trays 1 and 2) across the width of the

enclosure were ignited the flames moved rapidly towards the opening (the front of the enclosure).In tests with the ambient temperature above about 20◦C the introduction of a flame into the rear

of the enclosure caused very rapid propagation of flames through the enclosure and these very

rapidly subsided except at the opening where burning continued. In either case, the process took

no more than 2 min.

In both cases the flame-front rapidly established itself at the front edge of the front two trays

(Trays 15 and 16) with little of the fuel having been burned in the process. When the fuel in the

front trays was exhausted, the flame-front usually jumped to the front of the second row of trays

(Trays 13 and 14) where it burned until the fuel in these trays was exhausted. It then jumped to

the front of the next rearward row (Trays 11 and 12) and so on until finally the fuel in the rear

most row of trays was exhausted. In some cases the flames stayed at the opening even after the

front trays were empty, and in fewer cases even after several rows were empty.

During the previous tests [4] three phases of burning were identified. They were also observedin these experiments as follows:

Phase I: Rapid transition of burning up to the front edge of the front two rows (usually Trays

13–16; for opening on the side cases Trays 11–16). When the opening was not across the full

width of the enclosure, burning in the front two rows took place across the width of the opening(s),

not across the width of the trays or the enclosure. In these cases, as the fuel in the front trays


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(Trays 15 and 16) burned out and the flame-front moved to the second row (Trays 13 and 14)

and the flame-front extended across the full width of the enclosure. Full width burning was then

maintained as the fire moved back over the other rows of trays. As noted above in some cases (but

this was the exception rather than the usual situation) the flame-front remained at the opening for

much of the test even though the fuel in the front rows was exhausted.

Phase II: Relatively stable burning from the end of Phase I up to the burning of fuel trays in the

sixth row (Trays 5 and 6). During this phase the flames and smoke flows in the enclosure were

essentially two dimensional (i.e. when viewed from the side, the shape of the flames was the same

across the width of the enclosure). There were two sub-phases as reflected by the HRR curves:

initially the HRR was fairly constant, but often then decreased fairly linearly while burning in the

rearward trays.

Phase III: A period of more vigorous burning then took place, generally when the flame-front

reached the back two trays (Trays 1–4). In some cases the fuel was expended in the back row

before the flame-front reached these trays (see mass loss discussion below), and the final phase of

burning was actually in the second back row.

These phases can be readily identified in the HRR vs time profiles shown in Figure 3. A summary

of test results is presented in Table II, where t flameout and t 500 represent the time from ignition tofinal flame out and the time for which the maximum gas temperature in enclosure remained above

500◦C, respectively. It is observed that, t flameout≈1.03× t 500.

3.2. Heat release rate

The HRR vs time profiles from all of the experiments mentioned are plotted in Figure 3. In this

figure the HRR–time profile for the fully open end test is plotted as well as the HRR–time profile

from other tests as follows:

Group 1: Half open top and half open bottom.

Group 2: Half open side, half open centre and half closed centre.

Group 3: Quarter open side and quarter open centre.

In the fully open end case when the flame-front reached the front of the enclosure the burning

quickly became stable. In this case, as in all of these experiments, the measured HRR varied

considerably as the test proceeded. In this case after an initial very sharp peak, the burning (and

HRR) was reasonably stable with the HRR gradually falling as the flame-front moved from row

to row. Then as the flame-front reached the last row the HRR rose steeply and then fell sharply

when the fuel was exhausted (Figure 3).

Figure 3 (bottom) presents data from the tests included in Group 1. In Group 1 tests, it was

notable that when the flame-front reached the front edge of the enclosure the burning was rather

unstable but eventually became stable as the flame-front moved back through the enclosure. This

was particularly so for the half open top case—the first fire self-extinguished shortly after reaching

the front of the enclosure. The second attempt continued to burn as shown in Figure 3, but fora period was very unstable as reflected in the HRR. It is clear from this diagram that when the

opening size was reduced to half height keeping the opening width constant, the duration of

burning is approximately doubled (and was about 10% shorter for the bottom opening compared

with the top opening), both rather less than was calculated using the assumption of proportionality

of burning rate to the ventilation factor A√

h, in which case the burning duration is predicted to


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Figure 3. HRR from deep enclosure tests.


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Table II. Summary of experimental results (HRR and burning duration).

Duration of Max. HRR (kW) Average HRR (kW) burning (min)

Test name Phase I Phase II Phase III Overall Phase I-II Phase II-III t flameout t 500

Fully open end 1170 420 470 372 543∗ 311 76 70

Half open top 605 350 350 183 212† 153 138 134

Half open bottom 600 285 300 164 215† 116 151 147

Half open side 710 408 430 276 373† 181 96 93

Half open centre 560 235 381 253 306‡ 207 108 106

Half closed centre 1070 310 395 254 317‡ 200 102 100

Quarter open centre 510 275 260 175 222§ 126 154 150

Quarter open side 430 250 250 177 215¶ 124 148 144

∗first 24 min.†first 50 min.‡first 75 min.§

first 80 min.¶ first 88 min.

increase by 2.8 times. Based on the HRRs shown in Figure 3, reducing the opening height by half

(with width unchanged) reduced the average burning rate to 49% (top open) and 44% (bottom

open) of that for the fully open case rather more than expected using the A√

h ventilation factor,

which predicts 35% of the fully open case for both.

Referring to the centre diagram in Figure 3 (opening height constant, opening width reduced

to half) the duration of burning was increased in the range of 33–51%. According to the A√

h

assumption it should have simply doubled. The lateral variation of location of the opening in this

case also produced up to 14% difference in burning duration. Based on the HRRs, reducing theopening width by half the average burning rate was reduced only to 74% of the fully open case

for the half open side case, and 68% for half open centre and half closed centre cases—which

are much greater than the 50% figure expected based on A√

h. Burning at the opening was more

stable in Group 2 cases than in Group 1 cases but some fluctuation occurred while the flame-front

was in the opening.

In the two tests with the opening size reduced to one quarter (Figure 3 top), although the A√

h

assumption suggests a fourfold increase in burning duration, in reality it approximately doubled.

Variation in the location of the same shaped opening only caused approximately 5% difference in

burn out time. Furthermore, Group 3 cases (quarter width, full height) A√

h comparison leads to an

expected HRR of 25% of the fully open case, much lower than 47–48% obtained experimentally.

It is clear that the opening factor A√

h does not provide a particularly good prediction of the

average burning rate based on these experiments.In view of the complexity of the HRR–time curves in Figure 3, the average burning rate may

not be a good measure of HRR. The HRR generally fell significantly as burning progressed except

for the final peak that occurred in all cases. In the fully open case the average burning rate during

the first 24 min was 543 kW compared with 311 kW during the remaining 51 min. Similarly, there

was a decrease in all cases as shown in Table II.


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3.3. Mass loss

The mass loss from each tray is presented in Figure 4(a) for the fully open end case. The trends

in the mass loss during the other tests were similar. Note that scales 1 and 2 are trays at the rear

of the enclosure, through to scales 15 and 16 for front row trays (close to the opening).

In Figure 4 it can be seen that fuel is being lost from all trays (with fuel remaining) throughoutthe test. However, the rate of mass loss is greatest for the trays adjacent to the flame-front and

for them it is fairly constant once burning in the tray is established (see Figure 4(b)). It can also

be seen that the mass loss rate for the back row of trays (Trays 1 and 2) is greater through much

of the test than for trays between the back two trays and the burning trays. In this case (but not

in every case) this results in the back row of trays being depleted of fuel before the second back

row (Trays 3 and 4). For the remaining (intermediate) rows the mass loss rate (the slope of the

mass loss–time curves shown) decreases with distance from the opening and the flame-front. Close

correspondence between the measured HRR and the total mass loss rate was obtained in all tests

except for the first 1.5 min of the tests (see Figure 4(c)). Heat of combustion is considered to be

25930 kJ/kg taking into consideration of the water component of the fuel.

3.4. Gas temperature

Figures 5–7 show the time–temperature curves measured 25 mm below the top surface for all

tests as one moves from the opening towards the end of the enclosure. Data from the tests

included in Group 1 of Section 3.2 are presented in Figure 5, whereas Groups 2 and 3 data are

illustrated in Figures 6 and 7, respectively. These data are the average of measurements taken by

two thermocouples located above two adjacent trays in the same row (such as Trays 15 and 16).

It is observed that during first few minutes (say 5 min), almost identical temperatures are recorded

throughout the test rig. The variation starts at the locations close to the opening and gradually

occurs at other locations also. It is also observed that when the flame-front propagates to the deeper

into the rig, the location of the high temperature also shifts deeper.

Figure 5 reveals that temperature data at various locations are almost identical for the tests

included in Group 1 during Phase II. Temperatures recorded by the thermocouples are also foundto be identical for the tests included in Group 3 during Phases I and II (Figure 7). For these two

groups, during Phase I data measured over last five rows are not only identical, but also very close

to the data recorded during the fully open end test.

Overall temperature data measured during fully open end test are closer to the data recorded

during the tests included in Group 2 (Figure 6). This was also the case for HRR data presented in

Figure 3. Up to 10 min identical temperature data were recorded over the first three rows of the

tray. Moving further deep into the rig, this period of identical temperature gradually increases up

to the end of Phase II. Unlike Groups 1 and 3, during Phase II data were not identical over Trays

7–12. However, over the first two rows of trays, data presented for all tests in Figure 6 including

fully open end tests are found to be quite similar up to 75 min.

Table III shows that the spot peak gas temperature attained during the tests varied from 793

to 958◦C, the highest and lowest being for fully open end and half closed centre, respectively.Considering averaged temperatures (presented in Figures 5–7), the highest average peak was

recorded during fully open end test (862◦C) and the lowest was during quarter open side test

(710◦C). The spot peak gas temperatures are generally recorded above the second row of the trays

from the opening. However, for quarter open centre test, it was recorded above a tray at the first

row. Detailed measurements during all tests are presented in Appendix B of [8].


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0.2

0.4

0.6

0.8

1.0

1.2

M a s s L o s s R a t e ( k g / m i n )

dm/dt-16

dm/dt-15

dm/dt-14

dm/dt-13

dm/dt-12

dm/dt-11

dm/dt-10

dm/dt-9

dm/dt-8

dm/dt-7

dm/dt-6

dm/dt-5

dm/dt-4

dm/dt-3

dm/dt-2

dm/dt-1

0.0

200.0

400.0

600.0

800.0

1000.0

1200.0

1400.0

1600.0

Time (min)

H R R ( k W )

HRR-Mass Loss Rate

HRR-Calorimetry

-4.5

-4.0

-3.5

-3.0

-2.5

-2.0

-1.5

-1.0

-0.5

0.00.0 10.0 20.0 30.0 40.0 50.0 60.0 70.0 80.0

Time (min)

M a s s L o s s ( K g

)

Scale 3

Scale 4

Scale 1

Scale

2

Scale 8

Scale 7

Scale 10

Scale 9

Scale 11

Scale 12

Scale 15

Scale

16

Scale

14

Scale

13 Scale 5

Scale 6

0.0 10.0 20.0 30.0 40.0 50.0 60.0 70.0

0.0 10.0 20.0 30.0 40.0 50.0 60.0 70.0 80.0

Time (min)

80.00.0

(a)

(b)

(c)

Figure 4. Fuel mass/HRR–time relationship for fully open end test: (a) fuel mass vs time for each trayof fuel; (b) mass loss rate vs time for each tray of fuel; and (c) HRR comparison.


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Figure 5. Gas temperature–time relationships during fully open end, half opentop and half open bottom tests.

Figure 8 shows the gas temperatures at various levels as well as steel temperatures at the ceilingabove all trays as two-dimensional contour plots during the fully open end test. Time is taken as

horizontal axis and row position as vertical axis, and temperature is taken as amplitude. It is clear

from these figures that the temperature near the ceiling above the front row of trays was much

higher for most of the test duration than the temperature above the rear trays. For further clarity,

the gas temperatures above the front two rows and back two rows (25 mm below the top surface)


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Figure 6. Gas temperature–time relationships during fully open end, half open centre,half open side and half closed centre tests.

during the same test are plotted together in Figure 9. Similar trends were seen during the other tests

(please refer to Appendix B of [8]). This indicates that the severity of the exposure of structural

members is much more severe near the ceiling close to the opening of the enclosure than near

the back of the enclosure. If the area under the temperature–time curve is taken as indicative of


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Figure 7. Gas temperature–time relationships during fully open end, quarteropen centre and quarter open side tests.

severity of exposure then the severity at the second front row is over twice as severe as the back

row. Alternatively, if the duration of time above a temperature high enough to cause structural

weakening, etc. is used as the criterion the difference is far greater. For example, if the duration

above 600◦C is taken as the criterion, then the second front row location is more than five times

as severe as the back row.


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Table III. Summary of experimental temperature results.

Max. temperature (◦C)

Test name Gas Steel

Fully open end 958 (2) 471 (2)Half open top 797 (2) 414 (2)Half open bottom 846 (2) 452 (2)Half open side 905 (2) 466 (2)Half open centre 907 (2) 490 (2)Half closed centre 793 (2) 412 (2)Quarter open centre 820 (1) 431 (2)Quarter open side 812 (2) 415 (2)

Row numbers are given in parenthesis.

3.5. Steel temperature

Table III shows that the peak steel temperatures in all tests varied between 412 and 490◦C. Even

with above 900◦C gas temperature the corresponding steel temperature did not exceed 490◦C at

which the structural steel would retain significant strength. This lower temperature is attributed to

the excessive heat loss to the surrounding as well as very high expose area to mass ratio.

As shown in Figure 8 (fully open end test data) the peak steel temperatures are recorded above

the second row of the trays from the opening. This trend is the same for all the tests (for details

please refer to Appendix C of [8]), which is not clearly observed for the gas temperatures. The

trend of recording the peak steel temperatures above the second row of the trays from the opening

clearly shows that at this location structural weakening is far more severe than any other part of

the enclosure. This indicates that less protective construction can be used in areas away from the

potential enclosure openings while increasing in areas close to the potential openings.

3.6. Radiation heat flux

The radiation heat flux data from two representative tests (fully open end and half open centre) are

plotted in Figure 10 (for other test data please refer to Appendix D of [8]). Once the fuel close to

the opening of the enclosure is alight, due to short-lived simultaneous burning of fuels of adjacent

tray, it resulted in very high radiation heat flux readings by the front radiometer in initial period

of burning (up to few minutes). The values went up to ≈19kW/m2 and then the readings settled

down to the much lower values.

The radiation heat flux data show that once the flame-front jumps back to the second row of

trays and gets settled there, the front radiometer reads up to ≈17kW/m2 heat flux. As the flame-

front moves away from these two trays, this radiometer records lower heat flux values. However,

while the flame-front moves to the other trays located close to the middle and rear radiometers,high heat flux recording like the front radiometer did not take place. Occasionally the middle and

rear radiometers recorded up to 10 and 6kW/m2 heat flux, respectively. Low value of these two

radiometers (especially during fully open end test) may be attributed either to the lower flame

temperatures and shorter burning duration of the fuel over corresponding trays (see Figure 7 for

Trays 7–8 and 1–2) or to the deposition of soot particles on the sapphire window.


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Time (min)

R o w

0 10 20 30 40 50 60 70 80

1

2

3

4

5

6

7

8

9 900

8 800

7 700

6 600

5 500

4 4003 300

2 200

1 100

Level Temp

Time (min)

R o w

0 10 20 30 40 50 60 70 80

1

2

3

4

5

6

7

8

9 900

8 800

7 700

6 600

5 500

4 400

3 300

2 200

1 100

Level Temp

Time (min)

R o w

0 10 20 30 40 50 60 70 80

1

2

3

4

5

6

7

8

9 900

8 800

7 700

6 600

5 500

4 400

3 300

2 200

1 100

Level Temp

Time (min)

R o w

0 10 20 30 40 50 60 70 80

1

2

3

4

5

6

7

8

9 900

8 800

7 700

6 600

5 500

4 400

3 300

2 200

1 100

Level Temp

(a)

(b)

(c)

(d)

Figure 8. Temperature contour plots with respect to row location and time in fully open end test. (a), (b),(c) represent gas temperatures at levels 200, 115 and 25 mm below the ceiling and (d) represents steel

temperature. White areas have no data.


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Temperature 25mm below top surface

0

100

200

300

400

500

600

700

800

900

0

Time (min)

T e m p e r a t u r e ( C )

Row-1

Row-7

Row-2 Row-8

10 20 30 40 50 60 70 80 90

Figure 9. Gas temperature–time relationship above front and back rows (fully open end).

4. FIRE SEVERITY AND VENTILATION FACTOR

The test results reported in this paper show that fire severity should be characterized by the size

of the fire (in terms of HRR, Q̇) and the duration of this fire, t b rather than temperature–time

combination, which is also echoed in [9]. This is because if the area under the temperature–time

curve is taken as indicative of severity of fire then the severity close to the opening is much intense

than that at the back of the enclosure. Therefore, it is impossible to characterize fire severity by a

single temperature–time curve, which is definitely possible by a single HRR–time relationship.

It is now known that both total enclosed area and its servicing vent or opening are important to

determine the mass loss rate ( R) or HRR [9–14]. The combination of enclosure shape, size and

opening is traditionally determined by using a parameter called ventilation factor, w×h×√

h/ At.

Several models for determining averaged HRR (HRRavg) have been proposed by various researchers

based on this ventilation factor [13]. However, in this study effort is made to develop a model forobtaining HRR based on regression of all geometrical parameters providing equal emphasis on

each parameter.

The simplest way to determine t b is the time from ignition to final flame out (t flameout). However,

as the growth and decay period of fire are highly variable, Thomas et al. [10] considered to limit the

t b to the time for which the maximum gas temperature in enclosure remained above 500◦C (t 500).

It was thought that structural materials would not be affected significantly below this temperature,

either in the growth or the decay phase. For a fire severity model once the HRR avg is determined,

the burnout time (t b) can be found using the following relationship:

t b=F l(MJ)

HRRavg(s) (1)

Here F l is the total fire load in the enclosure, Q× D×W / H c (kg) or Q× D×W (MJ)

Q = Fire load per unit area,mass× H c

D×W (MJ/m2)

H c = Heat of combustion of the fuel (MJ/kg);(for wood 17 and liquid fuel 27)


DOI: 10.1002/fam




0

2

4

6

8

10

12

14

16

18

20

0

Time (min)

R a d i a t i o n F l u x ( k W

/ m 2 ) Rad_front

Rad_middle

Rad_rear

0

2

4

6

8

10

12

14

16

18

20

0Time (min)

R a d i a t i o n F l u x ( k W / m 2 )

Rad_front

Rad_middle

Rad_rear

20 40 60 80 100 120

20 40 60 80 100 120

Figure 10. Radiation heat flux data from fully open end and half open centre tests.

4.1. Model based on current deep enclosure tests

The only input variables throughout tests reported in this paper have been the geometry of the enclo-

sure and vent. For ventilation control fire, the effect of wall materials is insignificant. Considering

a relationship of the form

HRRavg

=c

×W cW

× Dc D

× H c H

×wcw

×hch (2a)

a least-squares regression may be used to evaluate the terms c,cW ,c D,c H ,cw and ch .

If this is done on the data as a whole a good correlation between the actual and predicted HRR

(in MW) results:

HRRavg=1.161×W 0.726× D−0.516× H 0.766×w0.521×h1.071 (2b)


DOI: 10.1002/fam




VU Model (Eq. 2b) Correlation

0

50

100

150

200

250

300

350

400

0

Experimental data (kW)

P r e d i c t e d d a t a ( k W )

50 100 150 200 250 300 350 400

Figure 11. Correlation between experimental (deep enclosure) and predicted (VU model) data.

Figure 11 shows a comparison of the actual values of the HRR and that predicted by this expression

(hereafter denoted as Victoria University or VU model). VU model is ideally applicable for the

deep enclosures (where D/ H 2 and D/W 2). Now this correlation is applied to the CIB test

data—data from an international test programme conducted under the auspices of CIB by eight

laboratories in several countries [2] and the correlation is shown in Figure 12. It is observed that

the VU model also works quite well for an enclosure where D/ H 2 and D/W =1, i.e. deep

enclosures with square configuration. It is to be noted that in the CIB tests with the same enclosure

and ventilation configuration, a large variation of mass loss rate data was recorded. Data outside

the 90% confidence interval (±1.65) are discarded.

In CIB tests, wood cribs with a variety of wood species, stick thicknesses and spacings were

burned in enclosures of various shapes and sizes. The enclosures were 0.5, 1, 2, 3 and 6 m wide,

0.5, 1, 1.5, 2 and 6 m deep and 0.5, 1.0 and 1.5 m high. The ranges of D/W were 0.5, 1 and 2and that of D/ H were 1, 2 and 4. Each test had a single ventilation opening with the full height

of the enclosure and either 25, 50, 75 or 100% of the width of the front of the enclosure. The fire

load density was equivalent to 10, 20, 30 or 40kg/m2 of wood over the floor area.

The VU model has now then applied (first to determine the average HRR and then calculate t busing Equations (2b) and (1), respectively) to the Cardington test data [3] and shown in Table IV

where the bold values represent prediction within ≈30%). Similar to the current and CIB tests,

Cardington tests had single opening with various sizes. In these tests, wood cribs were used as

fuel and the fire load density was equivalent to 20, 26 or 40kg/m2 of wood over the floor area.

Application of two other models to the Cardington test data is also presented in Table IV, where

the italic values represent deep but square enclosure test.

For tests 1, 3, 4, 5 and 6, the VU model has predicted the burnout time quite well (vary from

4 to 23%). In these five tests, except for test 1, w/W is less than 1. Prediction is also goodfor tests 2 and 9. On the contrary, for tests 7 and 8 the prediction is not as good. It is worth

mentioning that during test 7, the enclosure shape was D/ H 2 and D/W =1 and during test 8

fire growth was delayed due to release of water vapour from the plasterboard walls at an early

stage. It is recommended that the VU model be only used for the deep enclosures (where D/ H 2

and D/W 2).


DOI: 10.1002/fam




CIB Test Data (D/H=1) Against VU Model

0

1

2

3

4

5

6

7

8

9

10

Predicted data (MW)

E x p e r i m e n t a l d a t a ( M W )

CIB Test Data (D/H=2 & 4 and D/W=1) Against VU Model

0

1

2

3

4

5

6

7

8

9

10

0

Predicted data (MW)


CIB Test Data (D/H=2 & 4 and D/W=2) Against VU Model

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0

Predicted data (MW)

E x p e r i m e n t a l d a t a

( M W )

0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

1 2 3 4 5 6 7 8 9 10

0 1 2 3 4 5 6 7 8 9 10

(c)

(b)

(a)

Figure 12. Correlation between CIB test and predicted (VU model) data: (a) where D/ H 2 and D/W 2;(b) where D/ H 2 and D/W =1; and (c) where D/ H =1 & D/W =0.5.


DOI: 10.1002/fam




T a b l e I V .

T e s t i n g c a r d i n

g t o n t e s t d a t a w i t h V U ,

P r o j e c t 3 C I B a n d S F P E m o d e l s .

V U m o d e l

P r o j e c t 3 C I B m o d e l

S F P E m o d e l

F u e l l o a d

E x p t t ∗ b

H R R

t b

H R R

t b

H R R

t b

W

D

H

w

h

( k g / m

2 )

( h )

( M W )

( h )

% d i f f .

( M W )

( h )

% d i f f .

( M W )

( h )

% d i f f .

T e s t 1

5 . 5 9

5 2 2

. 8 5 5

2 . 7

5 0

5 . 5 9

5

2 . 7 5

0

4 0 . 0 0

1 . 8

5

1 2 . 6 8 1

1 . 9

0 5

2 .

9 6

1 0 . 7 7 5

2 . 2 4 2

2 1 .

1 9

1 6 . 5

6 6 †

2 . 0 4

2 ‡

1 0 .

3 6

T e s t 2

5 . 5 9

5 2 2

. 8 5 5

2 . 7

5 0

5 . 5 9

5

2 . 7 5

0

2 0 . 0 0

1 . 2

8

1 2 . 6 8 1

0 . 9

5 2

− 3 4 . 4

0

1 0 . 7 7 5

1 . 1 2 1

− 1 4 .

0 2

1 6 . 5

6 3 †

1 . 0 2

1 ‡

− 2 5 .

3 7

T e s t 3

5 . 5 9

5 2 2

. 8 5 5

2 . 7

5 0

5 . 1 9

5

1 . 4 7

0

2 0 . 0 0

1 . 9

1

6 . 2 3 8

1 . 9

3 6

1 .

0 4

3 . 2 4 0

3 . 7 2 7

9 5 . 4 5

5 . 1 0 §

2 . 0

8 9

9 .

3 6

T e s t 4

5 . 5 9

5 2 2

. 8 5 5

2 . 7

5 0

5 . 1 9

5

1 . 4 7

0

4 0 . 0 0

3 . 1

7

6 . 2 3 8

3 . 8

7 2

2 2 .

1 5

3 . 2 4 0

7 . 4 5 4

1 3 5 . 8 9

5 . 1 0 §

4 . 1

7 8

3 1 . 8 9

T e s t 5

5 . 5 9

5 2 2

. 8 5 5

2 . 7

5 0

2 . 1 3

9

1 . 7 3

0

2 0 . 0 0

2 . 5

8

4 . 6 7 8

2 . 5

8 2

0 .

0 7

1 . 7 8 9

6 . 7 5 2

1 6 1 . 6 1

2 . 6

6 §

4 . 0

1 3

5 5 . 3 2

T e s t 6

5 . 5 9

5 2 2

. 8 5 5

2 . 7

5 0

5 . 1 9

5

0 . 3 7

5

2 0 . 0 0

7 . 5

0

1 . 4 4 4

8 . 3

6 2

1 1 .

5 0

0 . 2 7 7

4 3 . 5 8 0

4 8 0 . 3 0

1 . 8 6

6 ¶

6 . 4

7 6

− 1 5 . 5 3

T e s t 7

5 . 5

9 5 5

. 5 9 5

2 . 7

5 0

1 . 3

7 0

2 . 7

5 0

2 0 . 0

0

0 . 5

8

1 2 . 5 9 4

0 . 2

3 5 − 1 4 7 . 0

6

2 . 6 3 8

1 . 1 2 1

9 3 . 2 8

9 . 5

6 4 †

0 . 4 3

3 ‡

− 3 3 . 9 7

T e s t 8

5 . 4 6

5 2 2

. 8 5 5

2 . 6

8 0

5 . 0 6

5

2 . 6 8

0

2 5 . 9 6

2 . 2

5

1 1 . 2 8 9

1 . 3

5 6

− 6 5 . 9

3

9 . 3 1 2

1 . 6 4 4

− 3 6 . 8 6

1 5 . 1

5 †

1 . 3 8

1 ‡

− 6 2 . 9 2

T e s t 9

5 . 5 9

5 2 2

. 8 5 5

2 . 7

5 0

5 . 1 9

5

2 . 7 5

0

2 0 . 0 0

1 . 3

3

1 2 . 2 0 0

0 . 9

9 0

− 3 4 . 3

6

1 0 . 0 0 5

1 . 2 0 7

− 1 0 .

1 9

1 6 . 2 7 6 †

1 . 0 3

9 ‡

− 2 8 .

0 1

∗ E x p e r i m e n t a l t i m

e i s t a k e n a s t h e t i m e u p t o w h i c h i n

t e r n a l g a s t e m p e r a t u r e w a s a b o v e 5 0 0 ◦ C .

†

L a w m e t h o d .

‡

t b i s c a l c u l a t e d b y m u l t i p l y i n g E q u a t i o n ( 5 ) v a l u e w i t h 1 . 4 .

§

M & T m e t h o d .

¶

L i e m e t h o d .


DOI: 10.1002/fam




Having established the validity of Equation (2b), HRR vs time correlations for liquid and cellu-

losic fuels have also been derived from the current test data considering a sixth-order relationship

of the form:

HRR

HRRavg =a×t

t b

6

+b×t

t b

5

+c×t

t b

4

+d ×t

t b

3

+e×t

t b

2

+ f ×t

t b+g

4.1.1. HRR vs time correlation for liquid fuel fire. A regression analysis of the average of all eight

test data gives the following relationship for liquid fuel fire:

HRR

HRRavg= 75×103×

t

t b

2

for 0<t /t b0.0054 (represents 807kW fire in 24.56 s

for fully open end test; a rate much faster than ultra-fast fire [9])

= 5636159.13×

t

t b

6

−3649832.94×

t

t b

5

+904299.33×

t

t b

4

−109064.64×

t

t b

3

+6647.65×

t

t b

2

+187.95×

t

t b

+3.16 for 0.0054<t /t b0.15

=−190.85×

t

t b

6

+753.84×

t

t b

5

−646.22×

t

t b

4

+336.27×

t

t b

3

−79.96×

t

t b

2

+6.43×

t

t b

+1.276 for 0.15<t /t b0.625

= 2925.1× t

t b

6

−11057× t

t b

5

+16688× t

t b

4

−12574× t

t b

3

+4736.6×

t

t b

2

−720.53×

t

t b

+4.511 for 0.625<t /t b0.95

=−26199618118.21×

t

t b

6

+15276365206.43×

t

t b

5

−37109745892.67×

t

t b

4

+48073844555.28×

t

t b

3

−35027290664.7×

t

t b

2

+13609976052.95×

t

t b

−2203187439 for 0.95<t /t b1.0 (3a)

Based on the ventilation and fuel load of a deep enclosure, a single HRR–time correlation repre-

sented by Equation (3a) is sufficient to characterize the fire severity. Note that HRRavg and t b are to

be calculated using Equations (2b) and (1), respectively. Equation (3a) can be used as prescriptive

fire severity for FDS [6] and CFAST [1] analysis. It is proposed that fire is to be distributed in eight

strips (for CFAST, interconnected open sub-enclosures) in succession of time. Thereby, we get an


DOI: 10.1002/fam




0.000

1.000

2.000

3.000

4.000

0.000 0.125 0.250 0.375 0.500 0.625 0.750 0.875 1.000

Time/ Burnout Time

H R R / H R R_ a v g

Fully open(n-d) half open top (n-d)half open bottom (n-d) half open side (n-d)half open centre (n-d) half closed centre (n-d)quarter open centre (n-d) quarter open side (n-d)Equation (3a)

0.116

0.324

0.599 0.7390.865

0.929

0.973

Plan view of enclosure

opening

Figure 13. Proposed HRR vs time vs position model for zone and field model.

HRR vs time vs position correlation for enclosure with single opening as shown in Figure 13. This

figure also shows the times at which the fire propagates from the opening to the rear. These times

are determined by averaging data from all eight tests.

The first propagation phase from the rear to the front (in the cases of ignition is done at the

rear) is not considered. As the duration of this phase is very short (less than 2 min), it is considered

that this phase will not have any significant effect either on the structure or tenability within the

enclosure.

4.1.2. Cellulosic fire. Similar enclosure fires conducted with the wood crib were described in

[3, 15]. When the fire was ignited at the rear of the enclosure in all cases it established itself asclose to the ventilation opening as possible, and then burned back through the enclosure consuming

the fuel closest to the vent first. Readers are requested to refer to the temperature data in Figures

34–38 and 41 of [3] to visualize this phenomenon.

To apply the above fire curve (Equation (3a)) for wood and other cellulosic material fire, we

need to modify the fire growth and decay to a slower rate. It is recommended that the fire engineer


DOI: 10.1002/fam




should make a choice from ultrafast, fast, medium and slow [9] rate and calculate appropriate

growth time (t g) using the following expression:

t g

=HRRmax

1/2

where is the fire growth parameter kW/s2 and HRRmax can be taken as 2.17×HRRavg as found

in Equation (3a). This will lead to the following expression for the growth period:

HRR=(kW/s2)× t 2b (s2)×

t

t b

2

for 0<t t g

Similarly the decay rate is also needs to be altered. As per [9] the decay period starts when 80%

of the fuel has been consumed (approximately at time, t decay= t g+0.8× t b) and decays at a linear

rate ( Drate) determined experimentally or which can be justified in other way.

HRR=HRRat t =t decay− Drate× t for t decay<t t HRR=0

For wood and other cellulosic material fire, Equation (3a) can be modified to the following

relationship:

HRR

HRRavg= (kW/s2)× t 2b (s2)

HRRavg(kW)×

t

t b

2

for 0<t t g

= 5636159.13×

t −t g

t b+0.0054

6

−3649832.94×

t − t g

t b+0.0054

5

+904299.33×

t − t g

t b

+0.0054

4

−109064.64×

t − t g

t b

+0.0054

3

+6647.65×

t −t g

t b+0.0054

2

−187.95×

t − t g

t b+0.0054

+3.16

for t g<t 0.1446t b+t g

=−190.85×

t −t g

t b+0.0054

6

+753.84×

t − t g

t b+0.0054

5

−646.22

×t −t g

t b +0.0054

4

+336.27

×t − t g

t b +0.0054

3

−79.96×

t − t g

t b+0.0054

2

+6.43×

t −t g

t b+0.0054

+1.276 for 0.1446t b+t g<t 0.6196t b+t g


DOI: 10.1002/fam




= 2925.1×

t − t g

t b+0.0054

6

−11057×

t − t g

t b+0.0054

5

+16688×t

−t g

t b +0.0054

4

−12574×t

−t g

t b +0.0054

3

+4736.6×

t −t g

t b+0.0054

2

−720.53×

t − t g

t b+0.0054

+4.511 for 0.6196t b+t g<t 0.8t b+t g

= HRRavg at t =0.8t b+t g − Drate×(t − t g−0.8t b)/HRRavg for 0.8t b+t g<t t HRR=0 (3b)

Fire growth stage can be placed in the sub-enclosure where the fire is considered to have started

and the decay stage should be placed at the sub-enclosure next to the back wall. Equation (3b) for

the HHR(t ) for cellulosic fire is primarily a suggestion as the HRR(t ) or mass loss rate (t ) data

are not available from [3, 15] and no wood crib test was conducted with the test rig described inthe current study. This suggestion will be validated in the near future with further tests with the

wood crib.

The HRR vs time vs position correlation for wood and other cellulosic material (based on

fully open end enclosure with medium fire growth rate and a decay rate of 1.3 kW/s) is shown in

Figure 14.

It is to be noted that the above procedure for both liquid and solid fuels should be used for

fire development in the deep enclosure situation only. For the analysis of fire in cubic enclosure,

please refer to [16].

4.2. Application of other models on current test result

4.2.1. Model based on CIB tests (FCRC Project 3). In the late 1990s the Fire Code Reform Centre(FCRC) of Australia undertook a study to evaluate fire resistance level of building materials, which

0

0.5

1

1.5

2

2.5

H R R / H R R_ a v g

t g 0.8t b t decay

Figure 14. Proposed HRR vs time vs position model for cellulosic materials.


DOI: 10.1002/fam




Table V. Testing VU test data with Project 3 CIB and SFPE models.

SFPE model (M&T) Project 3 CIB model

Expt HRR HRR % HRR %W D H w h F i (kW) (kW) overprediction (kW) overprediction

Fully open 2 8 0.6 2 0.6 46.05 372 605.66 62.81 248.62 −33.17Half open top 2 8 0.6 2 0.3 132.06 183 211.17 15.40 71.40 −60.99Half open bottom 2 8 0.6 2 0.3 132.06 164 211.17 28.76 71.40 −56.47Half open centre 2 8 0.6 1 0.6 93.38 254 298.64 17.58 124.31 −51.06Half closed centre 2 8 0.6 1 0.6 93.38 253 298.64 18.04 124.31 −50.87Half open side 2 8 0.6 1 0.6 93.38 276 298.64 8.20 124.31 −54.96Quarter open centre 2 8 0.6 0.5 0.6 188.06 175 148.30 −15.26 62.15 −64.48Quarter open side 2 8 0.6 0.5 0.6 188.06 177 148.30 −16.22 62.15 −64.88

is known as FCRC Project 3. Researchers of this project developed two sets of correlation: the

first based on the data of [4] and the second based on the CIB tests data [2]. In both cases, foran enclosure with single opening, the steady state/average HRR (in MW) was derived based on a

least-square regression of the experimental data.

While the first set of correlations (based on [4] data) have been applied to the test data reported

in this paper overpredictions of HRR by 55–90% are observed. From the CIB tests the following

relationship was suggested:

w/W =1 HRRmax=0.53×w×h1.8(r 2=0.97) (4)

These are the maximum values of HRR. To get an average value the maximum value needs to be

divided by 1.7. Equation (4) is now denoted as Project 3 CIB model—the name based on CIB

tests and has been applied to the Cardington test data and shown in Table IV. In Column 12 of

this table HRR is obtained by Equation (4)/1.7 and in Column 13, t b is obtained by Equation (1).Project 3 CIB model reasonably predicted the test 1, 2 and 9 results. Its prediction is within 40%

of burn out time of test 8. This model grossly failed in predicting test 6 data. When Project 3 CIB

model is applied to the current test data, it underpredicts HRR from 33 to 65%, i.e. overpredicts

the burnout time (see Table V).

4.2.2. Model based on SFPE engineering guide. SFPE Engineering Guide [13] reviewed a number

of correlation models developed by various researchers. However, none of these models made

any attempt to correlate HRR/burning rate and hot layer temperature. Correlations proposed for

predicting the burning rate and hot layer temperature are independent of each other. These two

parameters are separately calculated based on enclosure size, shape and ventilation factor.

The above publication recommended to choose the correlation from three methods based

on inverse ventilation factor, F i= At/w×h×√ h. The suggested method that corresponds to

At/w×h×√

h is as follows:

• Law [11] : ∼18

• Magnussion & Thelandersson (M&T) [12]: 45–85.

• Lie [14] : ∼345


DOI: 10.1002/fam




Law method :HRR=3.06×wh×

hW

D×(1−e−0.036( At/wh1.5)) (5)

Magussion & Thelandersson (M&T) :HRR=1.743× D×W /F i (6)

Lie method :HRR=1.564×w×h1.5 (7)

It is also recommended that while using Law method the burning duration (Equation (1)) to be

multiplied by a factor of 1.4. Furthermore, an M&T (also known as the Swedish method) method

is based on the value of H c, 15MJ/kg for wood and other cellulosic material.

An SFPE model (precisely M&T method) has been applied to the current test data and the result

is presented in Table V, which shows that it well predicts the data (8–30%) except the fully open

test where its prediction is within 65%. Therefore, the applicability of an M&T method can be

extended to F i =190.

Then an SFPE model is applied to CIB test data and shown in Figure 15. It is observed that for

the enclosures of 6m( D)×6 m(W )×1.5 m ( H ) size, the predictions are not quite good.

An SFPE model is mainly based on the Cardington test data and the model is now applied to

the Cardington test data itself and shown in Table IV. Unlike two other models discussed above,an SFPE model has predicted the burnout time reasonably well for the deep but square test (test 7).

However, its prediction of deep enclosures is not better than that by the VU model. Its prediction

is a little better for tests 2 and 9, but significantly worse for test 5. Like the VU model, the SFPE

model also failed to reasonably predict test 8.

4.3. Model adoption

Table VI shows how well three discussed models fit the data from the current tests, CIB tests and

Cardington tests.

It is observed that although VU model and SFPE model both predict deep enclosure fire severity

quite good, all models have difficulty in predicting Cardington test 8 (during which fire growth

was delayed due to release of water vapour from the plasterboard walls at an early stage). Another

SFPE Model

0

1

2

3

4

5

6

7

8

9

10

0

Predicted data (MW)


1 2 3 4 5 6 7 8 9 10

Figure 15. Correlation between CIB test and predicted (SFPE model) data.


DOI: 10.1002/fam




Table VI. Comparison among all models.

Deep enclosure Deep but square/wide( D/ H 2 and D/W 2) ( D/ H 2 and D/W <2)

VU model GOOD MODERATE

Current : 0.5–13% (Figure 11) and fails some CIB tests (Figure 12(b))CIB: Generally good (Figure 12(a)) and Cardington test 7 (Table IV)Cardington: 3 tests under 3%, 4 test 12–35%,test 8: 66% (Table IV)

Project 3 CIB POOR MODERATEmodel Current : 33–65% (Table V) fails Cardington test 7 (Table IV)

Cardington: fails tests 3, 4, 5, 6, 7test 8 36%, 3 other tests 10–21% (Table IV)

SFPE model GOOD GOODCardington: 6 tests 10–32%, Cardington test 7; 33% (Table IV)test 5: 55%, test 8: 63% (Table IV) and some CIB tests (Figure 15) failCurrent with M & T method up to F i =190; 8–30%except fully open end test; 62% (Table V)

least-squares regression of both current test data (liquid fuel) and Cardington test data except test 7

(solid fuel) for Equation (2a) results:

HRRavg=1.161×W 0.645× D−0.562× H 0.985×w0.604×h0.818 . . . in MW (8)

Equation (8) has been applied to the current test data and Cardington test data (except test 7)

and are presented in Figure 16. It is found that Equation (8) well predicts (within 25%) both sets

of data. Therefore, we propose that for deep enclosure ( D/ H 2 and D/W 2), Equation (8) be

used. This can now be regarded as the VU model rather than Equation (2b).

For deep enclosure but square/wide shaped we propose to use an SFPE method. However, tospread VU model’s applicability to deep but square/wide enclosure, some tests with square/wide

configuration ( D/ H 2 and D/W <2) need to be conducted.

5. CONCLUSIONS

The experimental programme was designed to investigate the influence variation of opening size and

shape on a fire in a deep enclosure ( L/ H ∼13). In relation to potential damage to structural members

the variation in fire severity at different locations in the deep enclosure was also investigated.

The experiments confirmed previous smaller-scale experiments, showed that the fires in deep

enclosures are strongly influenced by the ventilation and are not at all uniform through the depth of

the enclosure. The severity of exposure of structural members is much more severe in the vicinityof the ceiling near the front of the enclosure compared with the back of the enclosure. If the area

under the temperature–time curve is taken as indicative of severity of exposure then the severity

at the front is over twice that at the back, but if the duration of time above a temperature high

enough to cause structural weakening is used (for example 600◦C) the difference is far greater,

about five times as severe in some cases.


DOI: 10.1002/fam




VU Model (Eq. 8 ) on VU data

0

50

100

150

200

250

300

350

400

Experimental data (kW)

0

P r e d i c t e d d a t a ( k W )

VU Model (Eq. 8) on Cardington Data (Deep Enclosure)

0

1

2

3

4

5

6

7

0

8

Experimental data-Burnout time (hr)

P r e d i c t e d d a t a - B u r n o u t t i m

e ( h r )

1

50 100 150 200 250 300 350 400

2 3 4 5 6 7 8

(a)

(b)

Figure 16. Correlation between experimental data and predicted (Equation (8)) data: (a) current test data

and (b) cardington test data except test 7.

The correlation between the ventilation factor A√

h and the averaged HRR data are examined.

It is observed that the experimental data deviate from the classical relationship. Therefore, a new

HRR–time model for field model and for the first time an HRR–time–position model for zone

model are proposed for fire development in deep enclosure based on the test data reported in this

paper. This is denoted as the VU model. Its averaged HRR/burning rate version (Equation (8)) has

been well validated against current and Cardington tests data. However, to validate the proposed

empirical relationship for the deep enclosure fire, further tests with other fuel packages including

wood cribs are required. Two other model (Project 3 CIB model and SFPE model) predictions have

also been evaluated against the same set of test data as well as CIB test data. It is recommended

to choose model based on the followings:

• VU model (Equation (8)) for deep enclosure ( D/ H 2 and D/W 2).

• For deep but square/wide-shaped enclosure, use Project 3 CIB model or SFPE method.

To spread VU model’s applicability to deep but square/wide enclosure, some tests with

square/wide configuration ( D/ H 2 and D/W <2) are also proposed.


DOI: 10.1002/fam




NOMENCLATURE

A opening area (w×h)

At total inside surface area of the enclosure

D enclosure depth

F l total fire load in the enclosure (MJ) H enclosure height

HRR heat release rate (kW)

H c heat of combustion of the fuel (MJ/kg)

h opening height

Q Fire load per unit area (MJ/m2)

W enclosure width

w opening width

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DOI: 10.1002/fam

fire in deep enclosures

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