## Model Scale Tunnel Fire Tests

### Longitudinal ventilation

### Brandforskprojekt 404-011

SP Fire Technology SP REPORT 2005:49### SP Swedish National T

## Model Scale Tunnel Fire Tests

### Longitudinal ventilation

**Abstract **

**Model Scale Tunnel Fire Tests **

**with longitudinal ventilation **

Results from a series of tests in a model tunnel (1:23) are presented. Tests were carried out with longitudinal ventilation under different fire conditions. Wood cribs were used to simulate the fire source, which was designed to correspond to a scaled-down HGV (Heavy Goods Vehicle) fire load. The parameters tested were: the number of wood cribs, type of wood cribs, the longitudinal ventilation rate and the ceiling height. The fire spread between wood cribs, with a free distance corresponding to 15 m in large scale, was also tested. The effects of different ventilation rates on the fire growth rate, fire spread, flame length, gas temperatures and backlayering, were investigated.

Key words: model scale, tunnel fire, fire spread, longitudinal ventilation, exhaust ventilation

**SP Sveriges Provnings- och ** **SP Swedish National Testing and **

**Forskningsinstitut ** **Research Institute **

SP Rapport 2005:49 **SP Report 2005:49 **
ISBN 91-85303-82-8
ISSN 0284-5172
Borås 2005
Postal address:
Box 857,

SE-501 15 BORÅS, Sweden

Telephone: +46 33 16 50 00

Telex: 36252 Testing S

Telefax: +46 33 13 55 02

**Table of Content **

**Abstract ** **2**

**Table of Content ** **3**

**Preface ** **4**

**1** **Introduction ** **7**

**2** **Overview of HRR in HGV trailer tests ** **8**

**3** **Theoretical considerations ** **10**

3.1 Scaling theory 10

3.2 Determination of HRR 10

3.3 Analytical description of the HRR development 12

3.4 Wood crib porosity 13

3.5 Ceiling temperature 18

3.6 Total heat flux 20

3.7 Flame length 20
3.8 Backlayering of smoke 21
**4** **Experimental Setup ** **23**
4.1 Fire load 24
4.2 Instrumentation 25
**5** **Test procedure ** **27**
**6** **Test results ** **29**

6.1 Free burn tests 29

6.2 Heat release rate 30

6.3 Gas temperatures 31

6.4 Total heat flux 31

6.5 Flame length 31

6.6 Backlayering of smoke 33

**7** **Discussion of results ** **35**

7.1 Heat release rate 35

7.2 Influence of ventilation 37

7.3 Gas temperatures 40

7.4 Total heat flux 44

7.5 Flame length 45

7.6 Backlayering of smoke 48

**8** **Conclusions ** **51**

**9** **References ** **52**

**Preface **

This project was financed by the Swedish Fire Research Board (BRANDFORSK). The technicians Michael Magnusson, Joel Blom, Lars Gustafsson and Ari Palo-Oja at SP Fire Technology are acknowledged for their valuable assistance during performance of the tests. They were also responsible for the construction of the test rig. Dr Anders Lönnermark is acknowledged for his valuable discussion of the results and performance of two of the tests.

The advisory group to the project is thanked for their contribution. The advisory group consisted of: Johan Hedenfalk (SL), Lars Aidanpää (LKAB), Magnus Lindström

(Brandkonsulten), Per Walmerdahl (FOI), Staffan Bengtsson (Brandskyddslaget), Anders Walling (Brandforsk), Jan Blomqvist (Cerberus), Anders Berqvist (Stockhol Fire

Brigade), Ovind Engdahl (Norsk Brannvern), Bernt Freiholtz (Vägverket), Odd Lyng, Gunnar Spång (SL), Christer Lindeman (SL), Jenette Stenman (Banverket)

Bo Wahlström (Brandskyddslaget), Katarina Kieksi (Banverket) and Omar Harrami (SRV).

**Summary**

Fire tests were carried out in a 1:23 model scale tunnel. Fire loads corresponding to a HGV trailer were simulated using wood cribs of two different sizes. Longitudinal ventilation was tested under different fire conditions. The parameters tested were: the number of wood cribs, type of wood cribs, longitudinal ventilation rate and the ceiling height. The fire spread between wood cribs, with a free distance corresponding to 15 m in large scale, was also tested. The effects of different ventilation rates on the fire growth rate, fire spread, flame length, gas temperatures and backlayering, were investigated. Longitudinal ventilation was established using an electrical fan attached to the entrance of the model tunnel. The tunnel was 10 m long, 0.4 m wide and with two heights, 0.2 m and 0.3 m, respectively. The corresponding large scale measures are 230 m long, 9.2 m wide, 4.6 m and 6.9 m high, respectively. The lower height obtained using a false ceiling for the exhaust ventilation channel. Longitudinal wind velocities of 0.42 m/s, 0.52 m/s, 0.62 m/s and 1.04 m/s, were used. The corresponding large scale velocities are 2 m/s, 2.5 m/s, 3 m/s and 5 m/s.

The tests show the effect of longitudinal ventilation flow on the speed at which fire spreads between burning items in a model tunnel. The tests show the principal mechanism of fire spread in longitudinal ventilated tunnel fires. Converted to a large scale situation, the tests indicate that a vehicle about 15 m (0.65 m) behind a burning HGV (wood crib) will catch fire within ten minutes (2 minutes in model scale). A third vehicle at a further 15 m (0.65 m) behind the second vehicle will catch fire within only approximately three more minutes (0.63 min), due to heat from the first vehicle. The rate of spread of the fire indicated in these small scale tests is in quite good agreement with the Runehamar large scale tunnel tests [1-3], where 'targets' were placed 15 m from the fire in order to simulate the potential for spread of the fire to nearby vehicles.

The model scale tests show that the initial rate of fire growth increases by about 4-6 times when the longitudinal ventilation velocity increases from somewhat over 0 m/s to 5 m/s (1.04 m/s). On the other hand, for the same increase in ventilation velocity, the maximum fire output does not increase by more than about 1.5 - 1.8 times, which is completely at variance with previously published analyses [4, 5].

**1 **

**Introduction **

Interest in fire safety issues in tunnels has increased dramatically owing to numerous catastrophic tunnel fires and the extensive monitoring of these incidents in the media. A common feature in all of these fires has been the significance of the fire load and of ventilation for the growth of the fire. In several of the fires, the type of load being carried by goods vehicles played an important part in determining the severity of the fire. The main reasons for this are that heavy goods vehicles (HGVs) consist of, or carry, highly flammable materials, and that the fire spreads very rapidly due to the longitudinal ventilation in the tunnel.

The effect of longitudinal ventilation on the fire development in HGV fires is an impor-tant issue. The interaction between the ventilation flow and the HRR was investigated by Carvel et al. [5-7]. Carvel’s results are probabilistic in nature; they say that, for example, that the Heat Release Rate (HRR) of a HGV could increase by much more than a factor of four for a longitudinal flow rate of 3 m/s, but that the most likely HRR enhancement is about a factor of four at 3 m/s and a factor of ten at 10 m/s. They also found that the fire growth rate could increase by a factor of five for 3 m/s and by a factor of ten for 10 m/s. A Bayesian probabilistic approach was used to refine estimates, made by a panel of experts, with data from experimental fire tests in tunnels. Their conclusions were, however, based on rather limited experimental data. The present model scale results in combination with large-scale tests provide new data to investigate the effects of longitudinal ventilation on fire growth rates in tunnels.

Most of our knowledge about smoke and fire spread in tunnels has generally been obtained from large scale testing. Large scale testing is, however, expensive, time consuming and logistically complicated to perform. The information obtained is often incomplete due to the limited number of tests and the limited instrumentation. Large scale testing is, however, necessary in order to obtain acceptable verification of results on a realistic scale. Model scale tests can be used as a complement to large scale testing. They can provide information which is difficult to obtain otherwise as their relatively low cost allows for parameter studies. The tests presented here were performed in November 2002, partly in order to design the large scale experiments carried out in Runehamar tunnel September 2003 [1-3]. The same fire load and concept as in the Runehamar tests were applied. These model scale tests have not been reported until now. In the following the test procedure and set-up of the model scale tunnel is presented.

**2 **

**Overview of HRR in HGV trailer tests **

It is of interest to compare the results obtained in the model scale tests presented here with large scale tests carried out with comparable fire loads. Ingason and Lönnermark [1, 8-12] have collected and summarised data on HRRs from large scale tests with HGV fire loads. A summary of the information obtained from papers [1, 8-12] is given here. In the EUREKA 499 tunnel test programme [13, 14], a test was performed with a HGV loaded with 1994 kg of mixed furniture, where 75 % was cellulose material and 25 % was plastic. The total heat content of the mixed furniture was estimated to be 42.7 GJ and the total heat content of the truck and trailer including mixed furniture, flooring of the trailer, tyres, cab, and fuel was estimated to be 87.4 GJ. The initial longitudinal velocity was approximately 5-6 m/s for 13.5 min from ignition at which point the fan was stopped. The HGV was burning fiercely when the fan was stopped and the estimated HRR was

about120 MW. The fan was restarted after three minutes (16.5 min from ignition) at a lower speed, 2-3 m/s, and the fire soon reached a peak HRR of 128 MW (at 19 min), see Figure 1. The HRR during the three minutes period when the fan was stopped could not be estimated since the HRR method relies on unidirectional flow [15].

0 50 100 150 200 250 0 10 20 30 40 50 60 EUREKA 499 - HGV

EUREKA 499 - simulated truck load Benelux - 36 wood pallets - 0 m/s - (T8) Benelux - 36 wood pallets - 4-6 m/s - (T9) Benelux - 36 wood pallets - 6 m/s - (T10) Benelux - 72 wood pallets - 1-2 m/s - (T14) Runehamar - wood and plastic pallets (T1) Runehamar - Wood pallets - mattrasses (T2) Runehamar - furnitures and fixtures (T3) Runehamar - cartons and PS cups (T4)

H eat R e le as e R a te ( M W ) Time (min)

*Figure 1 The measured HRR for the HGV trailer obtained from large scale fire *

*tests. *

The other trailer test in the EUREKA 499 test programme was performed on a weighing platform using densely packed wood cribs supplemented with rubber tyres and plastic material. The longitudinal ventilation velocity was 0.5 m/s (natural ventilation in the tunnel). The wood cribs consisted of 0.8 m long and 0.04 m thick rectangular rods nailed together in two layers, leaving 50 % air inside the total volume. The wood cribs covered an area of 2.4 m × 2.4 m where one layer consisted of 9 wood cribs (two layers of wood rods). The first wood crib layer was placed on an aerated-concrete blocks. Five layers of wood cribs were followed by one layer of plastic, then three layers of wood cribs, one layer of plastic, and again four layers of wood cribs. The remaining plastic and the rubber tyres were piled on top of this. The total number of layers of wood cribs was twelve and the total weight was 2212 kg (39.6 GJ). In total, 310 kg of plastic and 322 kg of rubber

tyres were used. The fraction of cellulosic material was thus 78 % of the total mass and 22 % were synthetic polymers (rubber and plastic). The total heat content of the trailer load was estimated to be 63.7 GJ. The original plan was to use 5000 kg fire load but that was reduced to 2844 kg due to problems of disposition of the fire source. The HRR is shown in Figure 1. A peak HRR of 16 MW was obtained 13 min after ignition. This peak was reduced to approximately 10 MW, 30 min from ignition after which the fire was relatively constant until it started to decrease after 75 min. The initial peak was probably due to the burning of plastic and rubber tyres in combination with the burning of the wood cribs.

Within in the framework of legal enquiry initiated after the catastrophic fire in the Mont Blanc tunnel 1999, a series of large scale tests were conducted in the same tunnel [16]. A test was carried out in the Mont Blanc tunnel with a real HGV truck and a trailer similar to that which generated the fire 1999 but with a much smaller amount of transported goods. The longitudinal flow at the fire location was about 1.5 m/s. In order to limit the peak HRR tyres had been removed and fuel tank was emptied. Only 400 kg of margarine were stored in the trailer. The total calorific value of the truck and the trailer with its goods was estimated to be 76 GJ. This value can be compared to the real value, which was estimated to be 500 – 600 GJ. The HGV was ignited by successively igniting three small pools filled with a diesel oil and alcohol mixture, respectively, placed in the HGV driver’s cab, behind the cab and between the cab and the trailer. Two HRRs levels were reached in the test. During the first 40 min, the HRR of the HGV fire remained lower than that of the pool fire, about 6 MW. Then the heat release rate reached a level of 23 MW, which can be related to the extensive burning of the HGV trailer.

In the Benelux tunnel test series, three tests consisting of 36 standardized wood pallets (9 in each stack) were performed with different longitudinal velocities (~0.5 m/s (natural ventilation), 4-6 m/s and 6 m/s) and one test with 72 wood pallets and a longitudinal velocity of 1-2 m/s. The outer dimensions of the 36 wood pallet fire load were 4.5 m long, 2.4 m wide, and 2.5 m high. As can be observed in Figure 1, the peak HRR was 13.5 MW without forced ventilation, 19 MW with 4-6 m/s ventilation and 16.5 MW with 6 m/s ventilation. The tests with the 36 wood pallet fire load show that the fire growth rate with ventilation was approximately 4 to 6 times faster than the fire growth rate with-out forced ventilation while the peak HRR was 1.4 and 1.2 times higher, respectively. The peak HRR with 72 wood pallets was 26 MW and the fire growth rate was about 1.9 times faster than the 36 wood pallet fire load with no forced ventilation.

The tests in the Runehamar tunnel were initiated and carried out by SP in co-operation with SPs sister organisations TNO in Holland and SINTEF in Norway. Four different HGVs fire load were burned in the tunnel. Three of the tests used pallet-loads consisting of various mixtures of wood, corrugated cardboard and plastic, while the fourth used furniture and associated items. Two mobile fans were stationed at the tunnel mouth, and together generated an air flow velocity of over 3 m/s through the tunnel. As in the model scale tests presented here, they made it possible to measure the HRR of the fires at the end of the tunnel. The HRR of the first test (wooden and plastic pallets) exceeded

200 MW, i.e. about the same as the theoretical heat release rate of a tanker fire, see Figure 1. This is a 'world record' for measured heat release rate in a tunnel. As can be observed in Figure 1, the HRR of three of four tests in Runehamar exceeded 120 MW [1]. Growth rate was relatively linear from 5 MW up to 100 MW, varying from 17 MW/minute to 29 MW/minute, with the most rapid rate of growth occurring in the second fire, which involved polyurethane mattresses and wooden pallets. The effects of the longitudinal ventilation on the fire growth rate could not be studied as only one velocity was used.

**3 **

**Theoretical considerations **

**3.1 **

**Scaling theory **

The model was built in scale 1:23, which means that the size of the tunnel is scaled geometrically according to this ratio. We neglect the influence of the thermal inertia of the involved material, the turbulence intensity and radiation, but we scale the HRR, the time, flow rates, the energy content and mass, see Table 1. Information about scaling theories can obtained from for example references [17-20].

*Table 1 A list of scaling correlations for the model tunnel. *

Type of unit Scaling model Equation number Heat Release Rate (HRR)

(kW) 2 / 5

### ⎟⎟

### ⎠

### ⎞

### ⎜⎜

### ⎝

### ⎛

### =

⋅ ⋅*M*

*F*

*M*

*F*

*L*

*L*

*Q*

*Q*

(1)
Velocity (m/s) 1/2
### ⎟⎟

### ⎠

### ⎞

### ⎜⎜

### ⎝

### ⎛

### =

*M*

*F*

*M*

*F*

*L*

*L*

*u*

*u*

(2)
Time (s) 1/2
### ⎟⎟

### ⎠

### ⎞

### ⎜⎜

### ⎝

### ⎛

### =

*M*

*F*

*M*

*F*

_{L}

_{L}

*L*

*t*

*t*

(3)
Energy (kJ)
*F*

*c*

*M*

*c*

*M*

*F*

*M*

*F*

*H*

*H*

*L*

*L*

*E*

*E*

,
,
3
### Δ

### Δ

### ⎟⎟

### ⎠

### ⎞

### ⎜⎜

### ⎝

### ⎛

### =

(4) Mass (kg) 3### ⎟⎟

### ⎠

### ⎞

### ⎜⎜

### ⎝

### ⎛

### =

*M*

*F*

*M*

*F*

*L*

*L*

*m*

*m*

(5)
Temperature (K)
*M*

*F*

*T*

*T*= (6)

L is the length scale and index M is related to the model scale and index F to full scale (LM=1 and LF=23 in our case).

**3.2 **

**Determination of HRR **

The heat release rate (HRR),

*Q*

⋅ given in (kW), is assumed to be directly proportional to
the fuel mass loss rate, *m*•

*(kg/s), and can be calculated using the following equation:*

_{f}*T*

*f* *H*

*m*

*Q*⋅ = •

### χ

*(7)*

*where HT* is the net heat of complete combustion (kJ/kg). The fuel mass loss rate,

•

*f*

*m* , is
determined by the weight loss. In fires the combustion of fuel vapours is never complete,
*and thus the effective heat of combustion (Hc*) is always less than the net heat of complete

*combustion (HT*). Further,

### χ

, is the ratio of the effective heat of combustion to net heat ofcomplete combustion, i.e.,

### χ

### =

*H*

_{c}*H*

*[21] (Tewarson calls the ‘effective heat of combustion’ the ‘chemical heat of combustion’).*

_{T}The actual heat release rate,

*Q*

⋅ (kW), at a measuring point downstream the fire can be
obtained by the use of the following equation (without correction due to CO production)
using oxygen consumption calorimetry [22, 23].
### ⎟

### ⎟

### ⎠

### ⎞

### ⎜

### ⎜

### ⎝

### ⎛

### −

### −

### −

### −

### −

### =

• ⋅ 2 2 2 2 2 2### 1

### )

### 1

### (

### )

### 1

### (

### 14330

0, 0,*CO*

*O*

*CO*

*O*

*CO*

*O*

*a*

*X*

*X*

*X*

*X*

*X*

*X*

*m*

*Q*

(8)
where
2
,
*0 O*

*X*

is the volume fraction of oxygen in the incoming air (ambient) or 0.2095
and
2
,
*0 CO*

*X*

is the volume fraction of carbon dioxide measured in the incoming air or
### ≈

2 ,*0 CO*

*X*

0.00033.
2
*O*

*X*

and
2
*CO*

*X*

are the volume fractions of oxygen and carbon
dioxide at the measuring station downstream of the fire measured by a gas analyser (dry). If

2

*CO*

*X*

has not been measured equation (8) can be used by assuming
2
*CO*

*X*

=0. This
will simplify equation (8) and usually the error will not be greater that 10 % for most fuel
controlled fires. In the derivation of equation (8) it is assumed that *m*•

*=*

_{a}### ρ

*and that 13100 kJ/kg is released per kg of oxygen consumed. It is also assumed that the relative humidity (RH) of incoming air is 50%, the ambient temperature is 15o*

_{a}uA_{C, CO}

2 in incoming air is 330 ppm (0.033 %) and the molecular weight of air,

*M*

*, is 0.02895 kg/mol and 0.032 kg/mol for oxygen (*

_{a}2

*O*

*M*

). Further,### ρ

*is the ambient air density,*

_{a}*u*is the average longitudinal velocity upstream the fire in m/s and

*A*

is the cross-sectional area of the
tunnel in m2 at the same location.
The total air mass flow rate,

*m*

⋅ , inside the tunnel (and in the exhaust duct) can be
determined both on the upstream (*m*

*a*

⋅

) and downstream (

*m*

*g*

⋅

) side. The general equation for the air mass flow rate is:

*T*

*u*

*A*

*T*

*m*

_{=}

### ζ

*a*

### ρ

*a*

_{c}### &

*(9) *

The theoretically determined mass flow correction factor (ratio of mean to maximum velocity),

### ζ

, is dependent of the variation of temperature and velocity over the*cross-section of the tunnel, A (or the exhaust duct). In the calculations of the air mass flow rate*a theoretical value of

### ζ

=0.817 was used. The centreline gas velocity (*u*

*) was*

_{c}determined with aid of the measured pressure difference,

### Δ

*p*

, for each bi-directional
probe [24] and the corresponding gas temperature. The diameter of the probes, *D*

, used
was 16 mm and the probe length, *L*

, was 32 mm. The centre line velocity was obtained
from Equation (10):
*a*

*a*

*c*

*T*

*pT*

*k*

*u*

### ρ

### Δ

### =

### 1

### 2

*(10) *

where *k*was a calibration coefficient equal to 1.08. The ambient values used in
equation (10) were

*T*

*= 293 K and*

_{a}### ρ

*=1.2 kg/m3.*

_{a}The mass flow rate was determined according to equation (9) using the centre line velocity both upstream and downstream. The downstream centreline velocity (

*u*

_{g}_{,}

*) was transformed to a “cold” centreline velocity (*

_{c}*u*

_{0}

_{,}

*) with aid of the ideal gas law*

_{c}(
*g*
*a*
*c*
*g*
*c*
*T*
*T*
*u*

*u*_{0}_{,} = _{,} ). The *O*_{2},*CO*and *CO*_{2} were measured at two heights: 0.88

### ×

*H*

and
0.5 *× H*

, where the tunnel height *H*

was 0.3 and 0.2 m, respectively.
The HRR calculated according to equation (8) can be obtained by calculating the air mass
flow rate upstream of the fire and by transform the measured *O*_{2},*CO*and *CO*_{2}to an
average value by using the correlation given by Newman [25], which assumes that the
local gas temperature and the local gas concentration correlate through the average values
over the cross-section:

*h*

*avg*

*h*

*i*

*avg*

*i*

*T*

*T*

*X*

*X*

### Δ

### Δ

### =

_{,},

*(11) *

where

*X*

_{i}_{,}

*is the concentration of species*

_{h}*i*

at height *h*,

*X*

_{i}_{,}

*is the average concentration of species*

_{avg}*i*

, ### Δ

*T*

*is the difference between the temperature at height*

_{h}*h*and the ambient temperature, and

### Δ

*T*

*is the difference between the arithmetic average temperature over the cross-section and the ambient temperature at corresponding location. Equation (11) can be applied directly for*

_{avg}*CO*and

*CO*

_{2}i.e.

*X*

_{CO}_{,}

*and*

_{h}*X*

_{CO}_{2}

_{,}

*and by using oxygen reduction,*

_{h}*X*

_{O}_{,}

_{h}2

### Δ

. Ingason and Lönnermark [1] used this correlation in order to determine the HRR in a series of large scale tunnel fire tests.**3.3 **

**Analytical description of the HRR development **

There are numerous mathematical expressions available in the literature to describe the fire growth and/or the decay period of the HRR. The HRR is either described as a time dependent power law function (

*t*

2,*t*

3) or as an exponential function (*e*

*t*or

*e*

−*t*). None of these approaches are based on derivation from physical processes of combustion or fire spread. If authors want to describe the entire fire development they usually merge two or three mathematical expressions (e.g.

*t*

2, constant and *e*

−*t*).

Numajiri and Furukawa [26] presented an excellent paper, which works out the problem of using many mathematical expressions in order to describe the complete fire

development. Based on their work Ingason [27] presented a mathematical model to describe the HRR development in tunnel fires. The key parameters in the model are the total energy content (

*E*

*) and the maximum heat release rate (*

_{tot}*Q*•

_{max}). Ingason [27] adjusted the basic equation given by Numajiri and Furukawa [26] and rewrote it in the following way:

*t*
*k*
*n*
*t*
*k* _{e}*e*
*r*
*n*
*Q*
*t*
*Q*⋅ = • ⋅ ⋅ ⋅ − −⋅ −1⋅ − ⋅
max (1 )
)
(

*(12) *

where

*Q*

•_{max}

*is the maximum HRR in kW and t is time in seconds. The parameters n*,

*r*

and *k* were defined by Numajiri and Furukawa [26]. They defined

*r*

as the amplitude
coefficient, *k*as the time width coefficient and

*n*as the retard index. Ingason [27] developed relationships for

*r*

and *k*including only

*n*,

*E*

*and*

_{tot}*Q*•

_{max}:

*n*

*n*

*r*

−
### ⎟

### ⎠

### ⎞

### ⎜

### ⎝

### ⎛ −

### =

1### 1

### 1

* (13) *

and
*r*

*E*

*Q*

*k*

*tot*

### ⋅

### =

• max_{ }

_{ }

_{(14) }

_{(14) }

The time to reach maximum HRR is then given as:

*k*

*n*

*t*

_{max}

### =

### ln(

### )

*(15) *

These equations will be used to investigate how well equation (12) fits to experimental data.

**3.4 **

**Wood crib porosity **

The fuel used in the model scale tests presented in this report consisted of wood cribs. The wood cribs were long relative to their height and width. The influence of the longitudinal velocity on the burning behaviour is of interest for the present study. The tests provide new data to investigate the effects of longitudinal ventilation on peak HRR and fire growth rates of fuels in tunnels. The geometrical form and size of the wood cribs is therefore of interest to study.

The exposed fuel surface area and the porosity of the wood cribs used in the tests presented here needs to be determined. Based on work by Gross [28] and Block [29], Heskestad [17] and Croce and Xin [30] have presented a correlation between the reduced free burning rate versus crib porosity for square wood cribs (same length and thickness of all wood sticks but different number of sticks and layers) burning under quiescent

atmospheric conditions. In Figure 2, a plot of the data given on Sugar Pine wood cribs by
Croce and Xin [30] together with a curve fit of the data is shown. It show the correlation
between reduced free-burning rate (*m*⋅ *f*×103/*A _{f}* ⋅

*b*−1/2) versus the crib porosity,

*P*

,
given in [30]. The character for mass burning rate is,*m*

*f*

⋅

, is used here instead of

*R*

*as defined by Croce and Xin. The crib porosity in the abscissa is defined as in reference [17], i.e.:*

_{f}*1/2*

_{s}*1/2*

_{b}*A*

*A*

*P*

*s*

*v*=

*(16) *

where Av is the total cross-sectional area of vertical crib shafts

### (

*l*

### −

*nb*

### )

2,*A*

*s*is the

exposed surface area of crib

### ⎟

### ⎠

### ⎞

### ⎜

### ⎝

### ⎛

_{−}

_{−}

_{−}

*N*

*n*

*n*

*l*

*b*

*blNn*

### (

### 1

### 2

### 1

### 4

. Here*b*is the stick thickness (same width and height),

*l*is the stick length,

*n*is the number of sticks per layer (the

*n*parameter here is different from the one used in Equation (12)) and *N*is the number of
layers [30].
0
0.2
0.4
0.6
0.8
1
1.2
1.4
0 0.5 1 1.5 2

Croce and Xin
Equation 17
m f
x 1
0
3 /A
s
b
-1
/2 (k
g/
s
m
-3
/2 )
crib porosity (P), A_{v}/A_{s} s1/2 b1/2 (mm)

*Figure 2 Reduced free-burning rate versus crib porosity P [30]. *

*As can be seen in Figure 2, when the crib porosity P is less than about 0.7 mm, the *
reduced mass burning rate, * _{m}*⋅

_{×}103/

_{A}_{⋅}

*−1/2*

_{b}*f*

*f* , starts to be influenced by the geometry

of the wood cribs. The fire becomes more and more dependent on the wood stick spacing and the wood crib fire is said to be under-ventilated or ventilation controlled. When the crib porosity is larger than about 0.7 mm, the wood crib fire is well-ventilated. A curve fit of the data given by Croce and Xin [30] gives the following equation (linear regression coefficient R=0.913):

### (

*P*

### )

*f*

*f* *A* *b* *e*

*m*⋅ _{×}_{10}3_{/} _{⋅} −1/2 _{=}_{1}_{,}_{11}_{⋅}_{1}_{−} −6.28⋅

_{ }

_{ }

_{(17) }

This equation can be rewritten in order to calculate the burning rate per square metre
exposed fuel surface area, which is of interest for this study, i.e.:
_{(17) }

### (

*P*

### )

*s*

*f*

*f*

*b*

*e*

*A*

*m*

*m*− − − ⋅ ⋅ ⋅ − × = =

_{1}

_{,}

_{11}

_{10}3 1/2

_{1}6.28 "

*(18) *

If we neglect the term

### (

_{1}

_{−}

_{e}

−6.28⋅_{e}

*P*

### )

_{ in equation (18), which approaches about 0.99 when P }is larger than 0.7, we can compare to the original work by Block [29] for Ponderosa Pine:
2
/
1
3
10
05
,
1
" − −
⋅
×
= *b*
*m* _{f}

*(19) *

This shows that the maximum burning rates in well-ventilated wood cribs is similar in the both cases (1.11 and 1.05, respectively).

*Figure 3 Effects of ventilation on rate of burning of cribs. This is the original *

*graph from Harmathy [31] (here R corresponds to *

*m*

*f*

⋅

*, *

*U *

_{a}*corresponds to *

*m*

*a*

⋅

* and *

*M is the total average mass of wood crib) . *

_{0}

Harmathy [31] presented experiments with square wood and plastic cribs (PMMA, polylite and phenolic) exposed to different ventilation rates. A crib was placed in a ventilation chamber where the air mass flow rate bypassing the wood crib was varied. The crib porosity factor, P, varied from 0.07 mm to 1.58 mm and the mass flow rate of air

*a*

*m*

⋅ between 0.005 kg/s to 0.051 kg/s. In the wood crib test series W-1, P varied between
0.07 mm to 0.98 mm and 0.41 mm to 1.10 mm in wood crib tests series W-2.
Harmathy concluded from his study that the rate of burning of non-charring fuels

(PMMA, polylite), i.e., the majority of synthetic polymers, is ostensibly unaffected by the air flow.

Harmathy also concluded that the rate of burning of charring fuels (wood, phenolic plastic) does exhibit a definite dependence on ventilation. This behaviour can be easily observed from Harmathys study, see Figure 3. The arrows in Figure 3 indicate the air flow rates approximately representing the stoichiometric air requirements for respective materials.

*Figure 4 Effect of ventilation on rate of burning of wood cribs in series W-1 as *

*presented by Harmathy [31] . (Here R corresponds to *

*m*

*f*

⋅

*, *

*U *

_{a}*corresponds to *

*m*

*a*

⋅

* and A*

*0*

* corresponds to A*

*s*

*) . Note that the present *

*author has added the values of the crib porosity, P, as well as n and N *

*in the original graph. *

In the experiments with charring fuels the burning rate first increased with increasing ventilation, reached a maximum burning rate and then decreased slightly as the air flow increased. This can be seen in Figure 4 (tests W-1-1, W-1-2 and W-1-3). Harmathy [31] concluded that the heat released by the oxidation of the char played an important role in the process of pyrolysis. This may explain why the charring fuel is more affected by the ventilation than the non-charring fuels. This conclusion is crucial for the work on effects of ventilation rate on heat release rates in tunnel fires.

As could be observed in the work by Croce and Xin, the wood crib geometry in a free
burn environment (quiescent air around the wood crib) plays an important role whether
the free burn rate is affected or not. When the wood crib porosity is less that about
0.7 mm the mass burning rate is reduced. It is possible, based on the information given in
Figure 4, to compare the reduced burning rate, * _{m}*⋅

_{×}

_{10}3

_{/}

_{A}_{⋅}

*−1/2*

_{b}*f*

*f* , of the work by

Harmathy under well ventilated conditions and the work by Croce and Xin. Values when a peak burning rate, or close to, were obtained, are compared to the results of Croce and Xin. The data points obtained in tests series W-1 by Harmathy, when the ventilation rate is at its highest peak, were plotted together with the data from Croce and Xin. The comparison is shown in Figure 5. This shows that the mass burning rate is similar in both cases, even when the crib porosity is low. This implies that a wood crib with low crib porosity can reach the same burning rate or more when it is exposed to sufficiently high air flow rates. Therefore, the burning rate is highly dependent on the ventilation rate around the wood crib when the crib porosity is low, whereas when it is high it is not so

sensitive. When comparing the influence of ventilation on burning rates of fuels, it is important to be aware of this fact.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 0 0.5 1 1.5 2

Croce and Xin
Harmathy
Eqn 17
m f
x 1
0
3 /A
s
b
-1
/2 (k
g/
s
m
-3
/2 )
crib porosity (P), A_{v}/A_{s} s1/2 b1/2 (mm)

*Figure 5 Comparison of reduced free-burning rate of Croce and Xin and *

*Harmathy. The ventilation flow in Harmatys tests were between *

*0.073 kg/s m*

*2*

* and 0.147 kg/s m*

*2*

*. *

There is a large difference between the wood cribs used in earlier studies [17, 28, 29],
[30] and the ones used here. In Figure 6 a drawing of a wood crib with two different side
lengths is shown. A similar approach has been used as that of Croce and Xin [30] in
defining the exposed fuel surface area (As) and ventilation area (Av) of a square wood
crib. In the following derivation of the equations, *l* refers to the shorter sticks and

*L*

refers to the longer sticks. The free distance between the short sticks is:

⎟⎟
⎠
⎞
⎜⎜
⎝
⎛
−
−
=
1
*L*
*L*
*l*
*n*
*b*
*n*
*l*
*s*

* (20) *

and between the long sticks:

⎟⎟
⎠
⎞
⎜⎜
⎝
⎛
−
−
=
1
*l*
*l*
*L*
*n*
*b*
*n*
*L*
*s*

* (21) *

The exposed fuel surface area of the long sticks becomes:

⎟⎟
⎠
⎞
⎜⎜
⎝
⎛
−
−
−
⋅
⋅
⋅
⋅
⋅
= ( 1 )
2
1
4
,
*L*
*l*
*l*
*L*
*L*
*L*
*s*
*N*
*n*
*n*
*L*
*b*
*n*
*N*
*L*
*b*
*A*

* (22) *

while that of the short sticks is:

### ⎟

### ⎠

### ⎞

### ⎜

### ⎝

### ⎛

_{−}

_{−}

### ⋅

### ⋅

### ⋅

### ⋅

### ⋅

### =

### (

### 1

### )

### 2

### 1

### 4

,*l*

*l*

*l*

*L*

*s*

_{l}

_{l}

*n*

*b*

*n*

*N*

*l*

*b*

*A*

*(23) *

The total exposed surface area is the sum of

*A*

_{s}_{,}

*and*

_{L}*A*

_{s}_{,}

_{l}*l*
*s*
*L*
*s*
*s*

*A*

*A*

*A*

### =

_{,}

### +

_{,}

*(24) *

L
b
b
l
sl
sL
nL
nl
Nl
NL
*Figure 6 Definition of geometrical parameters used to calculate exposed fuel *

*surface area and porosity of a wood crib. *

The ventilation area Av is :

### )

### )(

### (

_{l}

_{L}*v*

*L*

*bn*

*l*

*bn*

*A*

### =

### −

### −

*(25) *

Thus, the crib porosity becomes:
2
/
1
2
/
1 _{b}*s*
*A*
*A*
*P* _{p}*s*
*v*
=

*(26) *

where

*s*

*is chosen as the hydraulic diameter of the Av ;*

_{p}*L*
*l*
*L*
*l*
*p*
*s*
*s*
*s*
*s*
*s*
+
= 2 .

**3.5 **

**Ceiling temperature **

The average fire temperature at the fire place can be obtained from the following energy equation for the bulk flow:

*Q*

_{c}### =

*m*

*a*

*c*

_{p}### (

*T*

_{f}### −

*T*

_{a}### )

⋅⋅

*(27) *

where *Q*⋅ * _{c}*is the convective HRR of the fire (kW),

*m*

*a*

⋅

is the air mass flow rate in the tunnel (kg/s) ,

*c*

*is the heat of capacity for air (kJ/kg K),*

_{p}*T*

*is the average temperature at the fire (K) and*

_{f}*T*

*is the ambient temperature (K). If we assume that 2/3 of the total HRR (*

_{a}*Q*

⋅ ) is comprised of convective energy and that there is a correlation between the
highest obtained average temperature,

*T*

_{max}(K), and the average fire temperature,

*T*

*, the energy equation can be rewritten into a dimensionless form:*

_{f}*a*

*p*

*a*

*a*

*f*

*T*

*c*

*m*

*Q*

*T*

*T*

⋅
⋅
### =

### Δ

_{,}

_{max}

_{max}

### 3

### 2

*(28) *

where

### Δ

*T*

_{f}_{,}

_{max}

### =

### (

*T*

_{f}_{,}

_{max}

### −

*T*

_{a}### )

(K) and the mass flow rate can be expressed as :*ma*=

### ρ

_{a}### ζ

*u*=

_{c}A### ρ

_{a}### ζ

*u*

_{c}HW⋅

* (29) *

Here

*u*

*is the average velocity in the tunnel (m/s),*

_{c}*H*

is the tunnel height (m), *W*is the tunnel width (m) and

### ζ

, is the flow coefficient. We assume that the maximum excess ceiling gas temperature,### Δ

*T*

_{cf}_{,}

_{max}above the fire source can be correlated through equation (28) with aid of a proportionality coefficient which is determined experimentally. Thus, equation (28) can be rewritten in the following way:

*a*
*p*
*a*
*a*
*cf*

*T*

*c*

*m*

*Q*

*T*

*T*

⋅
⋅
### ∝

### Δ

_{,}

_{max}

_{max}

*(30) *

where
*a*

*cf*

*T*

*T*

_{,}

_{max}

### Δ

is the dimensionless excess is gas temperature and

*a*
*p*
*a*

*c*

*T*

*m*

*Q*

⋅
⋅
max _{ is similar to }

the non-dimensional heat flow parameter defined by Newman and Tewarson [32]. Equation (30) will be used to plot the temperature data at the ceiling above the fire source.

The average excess temperature at distance *x*downstream the fire source,

### Δ

*T*

_{avg}### (

*x*

### )

, can be obtained with aid of the following equation.### =

### Δ

### Δ

*mac*

_{p}*hPx*

*f*

*avg*

*e*

*T*

*x*

*T*

### (

### )

*(31) *

Where

### Δ

*T*

_{f}### =

### (

*T*

_{f}### −

*T*

_{a}### )

(K) is the average excess temperature at the fire location,*h*is the lumped heat loss coefficient (convective and radiation) (kW/m2

_{ K) and }

_{P}

_{P}

_{is the perimeter }of the tunnel (m). We assume that the maximum excess ceiling gas temperature,

### )

### (

max,

*x*

*T*

_{c}### Δ

, downstream the fire source at a distance*x*can be correlated through equation (31) with aid of a proportionality coefficient which is determined experimentally. Thus, equation (31) can be rewritten in the following way:

### ∝

### Δ

### Δ

*mac*

_{p}*hPx*

*cf*

*c*

*e*

*T*

*x*

*T*

max
,
max
, ### (

### )

*(32) *

where

### Δ

*T*

_{cf}_{,}

_{max}is obtained from equation (30).

**3.6 **

**Total heat flux **

Based on gas temperatures we can estimate the conditions for a material to ignite at a certain distance from the fire. The possibility of ignition of an object is presently judged by evaluation of whether or not the exposed surface would attain a critical ignition temperature. The critical temperature,

*T*

*, can be estimated as: 600 ºC for radiant exposure and 500 ºC for convective exposure in the case of spontaneous ignition; or, 300 to 410 ºC for radiant exposure and 450 ºC for convective exposure in the case of piloted ignition. Note that these are approximate values, mostly deduced from experiments on small vertical specimens [33].*

_{cr}Newman and Tewarson [32] argue that, at ignition,

*T*

_{cr}### ≈

*T*

*(the average gas*

_{avg}temperature) for duct flow, i.e. when the

*T*

*has obtained a critical ignition temperature, the material at that location will ignite. The critical temperature was estimated from the critical heat flux,*

_{avg}*q*

*"*

_{cr}### ≈

### σ

*T*

*4. This is an over-simplification as it does not take into account the time dependence of the ignition process. According to Newman and Tewarson, the critical radiation for numerous materials is in the range of 12 – 22 kW/m2 and the critical temperature*

_{cr}*T*

_{cr}### ≈

*T*

*is in the range of 680 – 790 K (407 – 517 ºC ). For example, for red oak the critical radiation is 16 kW/m2*

_{avg}_{ (}

*cr*

*T*

=457 o_{C), for PP }(polyethelyene) it is 15 kW/m2

_{ (}

*cr*

*T*

=447 ºC ), for polyurethane foam it is in the range of
16 – 22 kW/m2_{ (}

*cr*

*T*

=447-517 ºC ) and for polystyrene it is in the range of 12-13 kW/m2
(*T*

*=407-417 ºC ). The total heat flux to a wood crib at a given position from the fire can be estimated from the following equation:*

_{cr})
(
)
( 4 4
"
*a*
*avg*
*a*
*avg*
*c*
*flux* *h* *T* *T* *F* *T* *T*
*q* = − +

### εσ

−*(33) *

where

*h*

*is the convective heat loss coefficient in kW/m2*

_{c}_{ K, }

_{F}

_{F}

_{is the view factor, }

_{ε}

_{is the }emissivity and

### σ

is the Stefan-Boltzmann constant of 5.67 x 10-11_{ kW/m}2

_{ K}4

_{. Note that }

*avg*

*T*

and *T*

*must be expressed in degrees Kelvin for this equation to be valid. This equation can be used to roughly estimate the heat flux level at different distances from the fire. The average gas temperature,*

_{a}*T*

*, can be obtained by using equation (31).*

_{avg}**3.7 **

**Flame length **

Rew and Deaves [34] presented a flame length model for tunnels, which included HRR and longitudinal velocity but not the tunnel width or height. Much of their work was based on the investigation of the Channel Tunnel Fire in 1996 and test data from the HGV-EUREKA 499 fire test [13] and the Memorial Tests [35]. They defined the horizontal flame length,

*L*

*, as the distance of the 600 ºC contour from the centre of the HGV or the pool, or from the rear of the HGV. The flame length from the rear of the HGVwas represented by the following equation:*

_{f}### 10

### 120

### 20

− ⋅### ⎟

### ⎠

### ⎞

### ⎜

### ⎝

### ⎛

### ⎟⎟

### ⎟

### ⎠

### ⎞

### ⎜⎜

### ⎜

### ⎝

### ⎛

### =

*Q*

*u*

*L*

_{f}*(34) *

where

*Q*

⋅ is given in MW and *u*in m/s. This equation is a conservative fit to a limited data obtained from the HGV-EUREKA 499 test. Equation (34) tells us that the

longitudinal ventilation shortens the flame length as the velocity increase.

By solving *x* from equation (32), using

### Δ

*T*

_{cf}_{,}

_{max}from equation (30) and assuming that

### )

### (

### )

### (

*x*

*L*

_{f}*T*

_{ft}*T*

_{a}*T*

### =

### =

### −

### Δ

, we obtain the following relationship:### ⎟

### ⎟

### ⎟

### ⎠

### ⎞

### ⎜

### ⎜

### ⎜

### ⎝

### ⎛

### ⎟

### ⎟

### ⎟

### ⎠

### ⎞

### ⎜

### ⎜

### ⎜

### ⎝

### ⎛

_{−}

### −

### ∝

_{⋅}⋅ ⋅

*Q*

*T*

*T*

*c*

*m*

*hP*

*c*

*m*

*L*

*a*

*p*

*a*

*p*

*ft*

*a*

*f*

### )

### (

### ln

*(35) *

where *ma*=

### ρ

_{a}### ζ

*u*⋅

_{c}HWand

*P*

### =

### 2

*HW*

### /(

*H*

### +

*W*

### )

. This equation will be used to plot and compare to visual flame observations. This equation is valid for### (

_{⋅}

### −

### )

### <

### 1

⋅

*Q*

*T*

*T*

*c*

*m*

_{a}

_{p}

_{ft}

_{a}as the flame length,

*L*

*, can not have a negative value.*

_{f}In order to obtain a more simplified correlation an attempt to describe the flame length using the basic parameters

*Q*

⋅ , *u*,

*H*

and *W*will be tested. Therefore, a third alternative would be to plot the experimental data as a function of these parameters but using different experimentally determined exponent coefficients. This means that the flame length is described as:

_{b}

_{c}

_{d}*a*

*f*

*W*

*H*

*u*

*Q*

*L*⋅ ∝

* (36) *

*where the exponent constants a, b, c, d are determined from the experiments. *

**3.8 **

**Backlayering of smoke **

Thomas [36] has established a theoretical correlations between the hot plume flow above the fire and the backflow of hot gases upstream the fire in a longitudinal flow. The equation Thomas established was as follows:

### ⎟⎟

### ⎟

### ⎠

### ⎞

### ⎜⎜

### ⎜

### ⎝

### ⎛

### −

### =

⋅### 5

### 2

### 6

### .

### 0

_{3}

*A*

*u*

*c*

*T*

*Q*

*gH*

*H*

*L*

*p*

*a*

*f*

*b*

### ρ

*(37) *

where

### 2

_{3}

### >

### 5

*A*

*u*

*c*

*T*

*gHQ*

*p*

*a*

*f*

### ρ

,

*L*

*is the length of the backlayering distance (m) and*

_{b}*T*

*is the average fire temperature (K).*

_{f}Vantelon et al. [37] presented a simple model for predicting the length of a backed-up smoke layer,

*L*

*:*

_{b}_{⎟⎟}

### ⎟

### ⎠

### ⎞

### ⎜⎜

### ⎜

### ⎝

### ⎛

### ∝

⋅*H*

*u*

*T*

*c*

*Q*

*g*

*H*

*L*

*a*

*a*

*p*

*b*

### ρ

*(38) *

The proportionality constant for equation (38) has not been determined. Ingason [23] discussed this constant (assuming an exponent of 1/3 instead of 0.3) and found from the data given by numerous model scale tests [19, 37-39] that this constant varied between 0.6 to 2.2.

Equations (37) and (38) will be used to plot the observed smoke backlayering distances using

*u*

### =

### ζ

*u*

*.*

_{c}**4 **

**Experimental Setup **

A total of 12 tests were carried out in a 1:23 scale model tunnel. The parameters tested were: the number of wood cribs, the type of wood crib, the longitudinal ventilation rate and the ceiling height. The fire spread between wood cribs with a free distance of 0,65 m (15 m in large scale) was also tested. Further, the effect of different ventilation rates on the fire growth rate, fire spread, flame length, gas temperatures, radiation and smoke backlayering, was investigated.

*Figure 7 A photo of the 1:23 model scale tunnel. A fan was attached to the *

*tunnel entrance and windows were placed along one side in order to *

*observe the smoke flow. *

Longitudinal ventilation was established using an electrical axial fan attached to the
entrance of the model tunnel, see Figure 7. The fan itself was 0,95 m long with an inner
diameter of 0,35 m and a 0,8 HP motor yielding a maximum capacity of 2000 m3_{/h (at }
1400 rpm and 7,5 mmH2O). The rotational speed, and thereby the capacity, could be
controlled using an electrical device coupled to the motor. In between the fan and the
tunnel entrance, a 0,8 m long rectangular plywood box with the dimensions 0,4 wide and
0,3 m high, was mounted to create a uniform flow at the entrance of the tunnel. The
swirls created by the axial fan, were dampened by filling the plywood box with straw
fibres. Longitudinal wind velocities of 0,42 m/s, 0,52 m/s, 0,62 m/s and 1,04 m/s were
used in the test series. According to equation (3), the corresponding large scale velocities
were 2 m/s, 2,5 m/s, 3 m/s and 5 m/s.

The tunnel itself was 10 m long, 0,4 m wide and with two heights: 0,2 m and 0,3 m, respectively, see Figure 8. The corresponding large scale dimensions were 230 m long, 9,2 m wide, 4,6 m and 6,9 m high, respectively. The lower height (0.2 m) was created by using a false ceiling with the same material.

The model was constructed using non-combustible, 15 mm thick, boards (Promatect H).
The manufacture of the boards provide the following technical data: the density of the
boards is 870 kg/m3_{, the heat capacity is 1130 J/kg K and heat conduction is }

0.175 W/m K. The floor, ceiling and one of the vertical walls, were built in Promatect H boards while the front side of the tunnel was covered with a fire resistant window glaze.

The 5 mm thick window glaze (0,6 m wide and 0,35 m high) was mounted in steel frames which measured 0,67 m by 0,42 m. 300 mm 10000mm 1250mm 800 mm straw model tunnel 20 0 mm 2860mm 950 mm

*Figure 8 A schematic drawing of the model tunnel using longitudinal flow. *

**4.1 **

**Fire load **

The fire load consisted of two types of wood cribs (pine). The first series of tests (test 1 – 9) included a larger wood crib (A). A detailed description of wood crib A is given in Figure 9. In tests 10 – 12, a smaller wood crib was used (B). A detailed description of wood crib B is given in Figure 10. More detailed information about the wood cribs for each test is given in Table 2.

620 mm 140 mm 147 m m sticks 21 x 21 mm 50 m m

*Figure 9 Detailed drawing of the large wood cribs. This is wood crib is defined *

*as A. *

540 mm 120 mm
105 m
m
sticks 15 x 15 mm
50 m
m
*Figure 10 Detailed drawing of the smaller wood cribs. This is wood crib is *

*defined as B. *

The total the weight of wood crib A ranged from 2.21 kg to 2.63 kg (see Table 2). The variation is because each wood crib was manufactured by hand. The free distance

between each horizontal stick was 0.0187 m and the total fuel surface area of wood crib A was estimated to be 0.90 m2 (equations (22-24)) The total weight of wood crib B ranged

from 0.96 kg to 1.24 kg. The free distance between each horizontal stick was 0.02 m and
the total fuel surface area of wood crib B was estimated to be 0.56 m2_{. }

The crib porosity,

*P*

, according to equation (26) is 0.94 mm for wood crib A and
1.24 mm for wood crib B. This means that the wood cribs should not show any type of
under-ventilated tendencies during a test. This is important in order to compare a fuel that
is not under-ventilated during ambient conditions.
**4.2 **

**Instrumentation **

Various measurements were conducted during each test. The first wood crib was placed on a weighing platform (W), consisting of a scale attached by four steel rods to a free floating dried Promatect H board measuring 0.65 m long, 0.35 m wide and 0.12 m thick. In the case when more than one wood cribs was used in the tests, only the first wood crib was weighed. The weighing platform was connected to a data logging system recording the weight loss every second. The centre of the weighing platform was 2.87 m from the tunnel entrance (x=0) and the accuracy of the weighing platform was +/- 0.1 g. The temperature was measured with welded 0.25 mm type K thermocouples (T). The location of the thermocouples is shown in Figure 11 and Figure 12. Most of the

thermocouples were placed along the ceiling at a distance of 0.035 m from the ceiling. A set of thermocouples was placed 6.22 m (stack A in Figure 11) and 8.72 m from the inlet opening (stack B in Figure 11), respectively. The thermocouples in each set were place in the centre of the tunnel and 0.036 m, 0.093 m, 0.15 m, 0.207 m and 0.265 m, respectively, above the floor. These thermocouples are identified as T6-T11 for stack A in Figure 12 and T12 – T16 for stack B. Additional thermocouples were placed at a distance 0.075 m from the Promatect H wall at stack B and at heights 0.036 m, 0.15 m and 0.265 m, respectively, above the floor level, see Figure 11. A thermocouple was attached to the side wall, 0.15 m from the floor and 8.72 m from the tunnel entrance. A plate

thermometer [40] was placed at the ceiling during the final three tests. The location of the plate thermometer was 3.72 m from the tunnel inlet at x=0.

A bi-directional [24] probe (B22) was placed at the centreline of the tunnel 8.72 m from the inlet (at stack B). Another bi-directional probe was placed upstream the fire at the centre of the cross-section and 2.165 m (B21) from the inlet in tests 1 – 8, and then moved further upstream to 1.15 m from the inlet in tests 9 – 24. The pressure difference was measured with a pressure transducer with a measuring range of +/- 20 Pa.

2 65 mm weighting platform Thermcouplepil bi-directional probe Gasanalysis flux meter 1235 mm 1250mm 990mm 380mm 855mm 1250mm 1250mm 1250mm 1250mm Thermocouple K 0.25 mm

Obs: all instrument at the cenre of the tunnel except at pile A see graph to left.

150 m m 36m m 93m m 1 50m m 2 07m m 2 65m m Thermocouple pile B 100 mm 75mm wall thermocouple 700mm X 2865 3355 mm 2500 mm 100 mm 1280mm pile A pile B

*Figure 11 The instrument layout and measures during the tests. *

At three locations and flush to the floor board, water cooled heat flux meters of type Schmidt-Boelter (S) were placed to record the total heat flux. The locations were 3.72 m (Flux 1 or S18 in Figure 12), 6.22 m (Flux 2 or S19 in Figure 12) and 8.72 m (Flux 3 or S20 in Figure 12) from the tunnel entrance (x=0) in tests 1 – 4 and 10-12. During tests 5 – 9 the Flux 2 meter was moved upstream the fire to location 2.165 m. The Flux 1 meter was kept at 3.72 m and Flux 3 at 8.72 m. In tests 10 – 12 the Flux 2 were put back to the first set of locations, 3.72 m (Flux 1), 6.22 m (Flux 2) and 8.72 m (Flux 3).

The gas concentrations (O_{2}, CO_{2} and CO) (G) were measured 8.72 m from the entrance
(at stack B, i.e. G23, G24 and G25) by two measuring probes consisting of open copper
tubes (Ø 6 mm). They were located at three different heights, 0.15 m and 0.265 m above
the floor. The oxygen was measured with an M&C Type PMA 10 (0 – 21 %) and the
CO2 (0 - 10%) and CO (0 – 3 %) was measured with Siemens Ultramat 22. In Figure 12
the number of and identification of the probes used is presented.

W

thermocouple pile velocity

gasanlysis heat flux gage

thermocouple
T1 T2 T3 T4 T5
T17
S18 S19 S20
B21
B22
G23
G24 G25
T=thermocouple
B=bi-directional probe
S=Schmidt-Boelter gage
G=gasanalysis
W=Weightloss
pile B
pile A
pile A pile B
T6
T7 _{T8}
T9 _{T10}
T12
T13
T14
T15
T16

*Figure 12 The channel number and identification of all the instruments is shown *

*in this figure. *

The scale (weighing platform) (W), the thermocouples, the pressure transducers, the gas analysers and flux meters were connected to IMP 5000 KE Solotron loggers. The data was recorded on a laptop computer at a rate of about one scan per second.

**5 **

**Test procedure **

The wood cribs used in each test were dried overnight in a furnace at 60 ºC (<5% moisture). The first wood crib was placed on the weighing platform at a height 50 mm above floor. A cube of fibreboard measuring 0.03 m, 0.03 m and 0.024 m which was soaked in heptane (9 mL) was placed on the weighing platform board at the upstream edge of the wood crib as shown in Figure 13. At 2 minutes from start of the logging system, this cube was ignited. A total of 12 tests were carried out. In the following more detailed information about each test is presented in Table 2.

*Figure 13 A photo from test 5 showing a fully developed fire using a wood crib *

*of type A. The ignition source consisted of a fibreboard cube placed *

*on the weighing platform board at the upstream edge of the wood *

*crib. *

The tests were carried out with the same tunnel width, 0.4 m, but varying tunnel height and type and numbers of wood cribs. In tests with more than one wood crib, the free distance (edge to edge) between the cribs was 0.65 m. This means that the centre to centre distance between the A type cribs were 1.27 m (see Figure 14). An extra Promatect H board was placed under wood crib Nr 2 and wood crib Nr 3 in order to maintain the same distance between the top of the wood crib and the ceiling. At the lower ceiling height, i.e. 0.2 m, the wood crib type B was used as it was not possible to use wood crib type A with that ceiling height.

*Table 2 *

* Summary of tests carried out with longitudinal ventilation *

Test
nr *T*

*a*Nominal longitudinal centreline velocity,

*nom*

*c*

*u*

_{,}Height of tunnel ceiling,

*H*

Height
between
top of
wood crib
and
ceiling
Type
of
wood
cribs
Total
surface
area of
wood
crib, As
Number
of wood
cribs
used
Wood crib
Nr – initial
weight of
each wood
crib
Arrangement
of wood cribs
– free distance
ºC m/s m m m2_{ }

_{kg }

_{m }1 20.0 0.64 0.3 0.1 A 0.90 1 2.292 - 2 17.8 1.04 0.3 0.1 A 0.90 1 2.278 - 3 18.9 0.62 0.3 0.1 A 1.80 2 Nr 1 - 2.296 Nr 2 - 2.210 Serie - 0.65 4 18.6 0.62 0.3 0.1 A 2.70 3 Nr 1 - 2.298 Nr 2 - 2.412 Nr 3 - 2.336 Serie - 0.65 5 20.1 0.62 0.3 0.1 A 0.90 1 2.318 - 6 20.8 0.52 0.3 0.1 A 0.90 1 2.308 - 7 20.7 0.42 0.3 0.1 A 0.90 1 2.394 - 8 20.0 0.47 0.3 0.1 A 0.90 1 2.634 - 9 20.7 0.62 0.3 0.1 A 1.80 2 Nr 1 - 2.592 Nr 2 - 2.626 Parallel - 0 10 20.0 0.62 0.2 0.045 B 0.56 1 1.160 - 11 20.3 1.04 0.2 0.045 B 0.56 1 0.996 - 12 21.9 0.42 0.2 0.045 B 0.56 1 1.144 - 10000mm

straw model tunnel

2860mm 1270 mm 1270 mm

650 mm

Nr 1 Nr 2 Nr 3

650 mm

**6 **

**Test results **

In the following a presentation of the test results is given. A summary of the main results related to maximum HRR, ceiling temperatures, flame lengths and radiation are given. All the detailed test results for each test are given in Appendix A.

**6.1 **

**Free burn tests **

In order to obtain the heat release rate under free burning conditions tests using wood cribs A and B, were carried out under the SBI calorimeter [41]. Each wood crib was placed on four support columns creating a free distance of 50 mm from the floor up to the lower side of the wood crib. No influence of the tunnel walls and ceiling was considered in the free burning tests.

There was only natural ventilation in the free burning tests. The cribs were put into a 60 ºC furnace one night before the test. The ignition source was the same as that used in the tunnel tests but it was put under the centre of the wood crib instead of at the end of it. This was done in order to obtain comparable maximum values of the HRR for the two cribs. When the ignition source was placed in the middle the fire spread symmetrically in both directions of the wood crib length and a peak value was obtained with all the sticks involved. Preliminary tests showed that if the wood crib had been ignited at one end (as in the tunnel tests) the fire would not have peaked with all the sticks involved (part of the crib would have burned out before the fire spread to the other end). In a tunnel this would not be the case since the flames impinge on the ceiling and enhance the fire spread within the wood crib.

The HRR results from these tests are shown in Figure 15. The both wood cribs burn with about the same linear growth rate whereas the peak HRR varies. The linear fire growth rate,

*t*

*Q*

### Δ

### Δ

, can be estimated using values taken from part of the curve showing a

relatively linear behaviour. The linear fire growth rate was found to be about 44 kW/min for both wood crib types.

The peak HRR from wood crib A was 110.4 kW, which corresponds to 280 MW according to equation (1), and for wood crib B, corresponding values were 79.9 kW and 203 MW, respectively. In the free burn tests, the total mass of wood crib A was 2.336 kg and 1.001 kg for wood crib B. The total integrated energy for wood crib A was 38.4 MJ and 17.0 MJ for B. The effective value of heat of combustion,

### Δ

*H*

*, can be calculated based on these tests. For wood crib A, the effective heat of combustion is*

_{c}### =

### Δ

*H*

*3.423 MJ/2.336 kg=16.45 MJ/kg and for wood crib B*

_{c}### =

### Δ

*H*

*17.017 MJ/1.004 kg=16.95 MJ/kg . The average value of both is*

_{c}### =

### Δ

*H*

*16.7 MJ/kg. According to Tewarson [21] the net heat of complete combustion*

_{c}*T*

*H*

Δ for pine is 17.9 MJ/kg and the chemical heat of combustion

### Δ

*H*

_{c}### =

12.4 MJ/kg. In Table 3 a summary of the test results from the free burn tests is given.According to equations (4) and (5), the corresponding values of mass and heat of combustion in the large scale are 28.4 ton and 467 GJ for wood crib A (assuming

### 1

, ,_{=}

### Δ

### Δ

*F*

*c*

*M*

*c*

*H*

*H*

). Corresponding values in the large scale for wood crib B are 12.2 ton and 207 GJ assuming the same ratio of heat of combustion in the large scale.

0
20
40
60
80
100
120
0 5 10 15 20
**time (min)**
**H**
**RR (**
**k**
**W**
**)**

Wood crib A Wood crib B

*Figure 15 Measured HRR from a free burning wood crib of type A and B. *

*Table 3 Summary of test results of the free burn wood cribs A and B. *

Wood
crib
*tot*

*E*

max
,
*f*

*m*

⋅ max
⋅
*Q*

_{ }

*t*

_{max}

### Δ

_{Δ}

*Q*

_{t}

"
max
_{t}

*q*

### Δ

*H*

*c*MJ kg/s

_{kW min }

_{kW/min}

_{kW/m}2

_{MJ/kg }A 38.4 0.0067

_{110.4 3.9 44.2* 122 16.5 }B 17 0.0047

_{79.9 3.4 44.4** 136 17 }* based on values given between 20 kW and 100 kW

** based on values obtained between 20 kW and 60 kW

**6.2 **

**Heat release rate **

In Table 4 the main test results related to the air flow conditions and the heat release rates are given. The test number is given in the first column. The second column shows the average centre line value of the measured longitudinal velocity,

*u*

*, measured upstream of the fire (B21, see Figure 12). The nominal values of*

_{c}*u*can be found in Table 2. The third column show the mass flow rate,

*m*

*a*

⋅

, based on the centreline velocity upstream of the fire. The flow correction coefficient was

### ζ

=0.817. The forth column shows the fuel mass burning rate,*m*

*f*

⋅

, at the peak heat release rate. The fifth column shows the peak heat release rate based on equation (8). In the calculations, a combustion efficiency of

### χ

=0.9 was applied. This value was multiplied with the heat of combustion of 16.7 MJ/kg obtained from the free burn test. The parameter*t*

_{max}is the time in minutes from ignition