FIRE DYNAMICS IN FAÇADE FIRE TESTS: MEASUREMENT, MODELING AND REPEATABILITY

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FIRE DYNAMICS IN FAÇADE FIRE TESTS:

MEASUREMENT AND MODELLING

Johan Anderson* & Robert Jansson SP Technical Research Institute of Sweden, Fire Technology Brinellgatan 4, Box 857, S-501 15 Borås

Sweden

ABSTRACT

In a recent paper [1] the fire dynamics was investigated in a test rig for façade constructions according to the test method SP Brand 105 both experimentally and numerically[2,3]. The experimental setup simulates a three-story apartment building (height 6.7 m, width 4 m and depth 1.6 m), with external wall-cladding and a “room fire” at the base. The numerical model was constructed in the CFD program Fire Dynamics Simulator (FDS) [4] with analogous geometry and instrumentation. The general features of the fire test were well reproduced in the numerical model however temperatures close to the fire source could not be properly accounted for in the model. In this paper two additional fire tests including additional instrumentation is modelled. INTRODUCTION

Fire spreading from floor to floor via external walls has been analysed for a long time in a number of different test methods. The main focus is fire spread in or along the façade into rooms and floors above the original room fire. A variety of different test methods have been proposed in order to evaluate different wall claddings, insulations and geometrical aspects [2, 5-13]. The present test method used in Sweden was defined in 1985, SP Fire 105 [3]. The experimental setup described therein is intended for determining the fire behaviour of external wall assemblies and façade claddings, exposed to heat and flames from a fully developed apartment fire. During testing, falling down of parts and the occurrence of burning droplets is recorded. The fire source is a tray (width × length × height: 500mm × 2000mm × 100mm) filled with 60 litres of heptane with a free-board of water to ensure an even level at the base of the tray and facilitate burning across the full width of the tray throughout the test. The burning of 60 litres of heptane with a combustion efficiency of 0.8 corresponds to a fire load of approximately 75MJ/m2 calculated on the total fire room surface area. In order to create a more stable flame, a flame suppressing lattice consisting of a perforated steel sheet with pipes of 25mm diameter is placed in the fire tray. This method has then been presented internationally for different standardization committees. It has been referred to as a valid test method in Sweden, Denmark and Norway. In 2001, the SP Fire 105 test rig was modelled using the CFD code SOFIE [14]. The model was compared with the standard measurements conducted during a fire test. In the standard fire test two thermocouples are placed under an eave, six meters above the fire room and a heat flux meter is placed in a lower fictitious window 2.1 meter above the fire room. The results of the simulation show the response of the two thermocouples and the heat flux meter during 9 minutes of fire exposure. This corresponds to slightly more than half the time of exposure during a real test. The simulation corresponds fairly well with the measurements and a recommendation based on the SP Fire 105 study and another façade test simulation [14] was to use 64 rays in the radiation model when modelling this type of scenario which represented state of the art at that time. In the present study, the default setting of 100 radiation angles in FDS are used. Further, the whole scenario during fire exposure is modelled and additional measurements conducted during an experimental fire test are used for comparison with the simulation. In the previous work, the whole scenario during fire exposure was modelled and additional

measurements conducted during an experimental fire test were used for comparison with the simulation. A good correspondence was found between the shielded thermocouples in the FDS model compared with an experiment in the SP Fire 105 test rig. This confirmed that the design fire based on observations of the fire intensity during the test, previous measurements of heat release rate with the same test setup and the knowledge of the total fire load of 60 litres of heptane was valid. Although a simplified model of the fire test rig was employed the calculated values for thermocouples correspond well with the measurements. The same correspondence was not found for an additional measurement with a plate thermometer in the lower right corner of the entrance of the burning chamber. In this case the model gave substantially higher temperatures. Reasons for this are unclear however it is possibly a consequence of the shape of the flame suppressing lattice in the fire tray that lead to different local effects compared to burning from a liquid surface that is modelled in FDS.

In the present work, we report on two additional fire tests; in the first test, test A, the impact of the fire at the two fictitious windows were carefully monitored using thermocouples, plate thermometers and bi-directional probes and; in the second fire test, test B, accurate measurement of the heat release rate from an inert wall was made and large plates (20x20cm) were placed along the symmetry line of the wall at different elevations. All the physical conditions were taken into account in the improved simulations using the FDS model which are then compared with the experimental results. EXPERIMENTAL SETUP

The experimental setup described in the SP Fire 105 [3] is intended for determining the fire behaviour of external wall assemblies and façade claddings, exposed to heat and flames from an apartment fire. During testing, falling down of parts and the occurrence of burning droplets, is recorded. After fire exposure the construction, typically a combustible inner core protected by plaster, is cut into pieces to assess the internal fire spread in the core. Standard measurements during tests are heat flux to the centre of the lower fictitious window and two temperatures 100 mm under the eave, 100 and 400 mm from the façade measured by 0.25mm thermocouples. The dimensions of the test rig can be found in Figures 1 and 2. The fire source is a tray (width × length × height: 500mm × 2000mm × 100mm) filled with 60 litres of heptane with a free-board of water to ensure an even level at the base of the tray and facilitate burning across the full width of the tray throughout the test. The burning of 60 litres of heptane with a combustion efficiency of 0.8 corresponds to a fire load of approximately 75MJ/m2 total fire room surface. In order to create a more even fire load a flame suppressing lattice consisting of a perforated steel sheet with pipes of 25mm diameter is placed on the fire tray. NUMERICAL MODEL

The numerical work was performed using Fire Dynamics Simulation (FDS) version 5.5.3 [4]. The Navier-Stokes equations in the limit of low-speed, thermally-driven flow with an emphasis on smoke and heat transport from fires are solved by the FDS software. The algorithm used is an explicit predictor-corrector scheme that is second order accurate in space and time where turbulence is treated by means of Large Eddy Simulation (LES) in the Smagorinsky form. This is in contrast to other CFD codes for fire safety engineering where Reynolds averaged Navier-Stokes models are used. The heat transfer by radiation is included in the model via the solution of the radiation transport equation for a gray gas. The equation is solved using a finite volume technique for convective transport, thus the name given to it is the Finite Volume Method (FVM). When using 100 discrete angles, the finite volume solver requires about 20 % of the total CPU time of a calculation, a modest cost given the complexity of radiation heat transfer.

Figure 1. The façade front and side with spatial dimensions are shown from SP FIRE 105 [7].

Figure 2. The spatial dimensions of the fire room from SP FIRE 105 [7].

COMPARISON BETWEEN EXPERIMENTAL AND NUMERICAL RESULTS

In this section we will describe two different experiments (called Test A and B) and

analogous modelling, test matrix shown in Table 1. The first test is with is extra instrumentation in the two fictitious window lower edges and the second is with a Promatect wall cladding with six attached plate thermometers along the symmetry line of façade. In addition, in the second test also the heat release rate (HRR) was measured. In the present article we have focused on the impact on the façade of the fire. Due to the complexity of the geometry of the fire source, in particular the fire tray made of perforated steel sheet; detailed modelling of evaporation of the fuel due to back radiation and subsequent burning is difficult to manage in the models. As the weight loss of the fuel was not recorded during the first test an estimation of the heat release rate was made, in Test A. Visual observation during the fire test showed a clear highest level of fire intensity between 8 to 14 minutes and that the fuel was consumed after approximately 16 minutes, the curve shown in Figure 3 was defined, for Test A. This can be compared with values previously measured by Babrauskas by using the large scale products calorimeter at SP during four SP Fire 105 tests published previously [9]. The average value from these four tests is included in figure 3. As seen in the figure the area under the mean of the measured [by Babrauskas] curves is larger than under the estimated curve. The reason is that the burning of the tested insulation systems on the façades is included during a measurement with the large scale industrial calorimeter. A rough estimation of the energy in the damaged areas reported by Babrauskas shows approximately 5-10% extra energy from the burning of the insulation materials which corresponds fairly well with the difference in area under our estimation and area under the curve from the previous measurements.

Table 1: Tests modelled

HRR in model Instrumentation Thermal properties of

façade surface material

Test A From previous estimations

Plate thermometers, thermocouples and differential pressure probes in the two fictitious windows

From TPS measurements

Test B From measurements

Six attached large plates (20 x 20 cm) with insulation behind and 6 1mm thermocouples along the symmetry line at different heights

From the product developer

Figure 3. Estimated heat release rates from fire tray.

0 5 10 15 20 25

Time [Minutes]

Babrauskas [9]

HRR TEST A

HRR TEST B

As a first step the model described by Jansson and Anderson [1], was re-evaluated using a beta version of FDS 6, re-producing almost consistently higher temperatures for the thermocouples and plate thermometers compared with FDS5.5.3. In the coming version of FDS an improved turbulence model is implemented it is thus of interest to assess the salient features of the previous work using the updated software. However the computation time of this particular model was extended from five days for FDS 5.5.3 to almost nine days for a 5cm grid. In the rest of the paper FDS5.5.3 is used due to the impractically long computation times. In all further simulations a grid size of 5 cm cubes will be used. In [1], a grid sensitivity study was made and it was deemed that a 5cm grid was sufficient to reproduce most of the relevant features. It is of interest to compare the effect on the temperatures at the eave under the influence of different wall claddings such as light weight concrete (Siporex) and a highly insulating wall cladding consisting of Rockwool. Fortunately, experimental findings are available for these cases adapted from Babrauskas [9]. Here, in Table 2 only the maximum temperatures measured at the eave is shown. It should be noted that the temperatures are measured by the inner thermocouple placed 100mm from the façade and 100 mm under the eave, moreover the temperature differences are similar in the experiment and the simulation. In the simulations the materials are characterised by spatially homogeneous physical properties. The properties of interest are the density, heat conductivity and heat capacity as shown in Table 3. Table 2. The effect of wall cladding on the maximum temperature at the eave. Wall cladding\Case Experimental (C˚) (Babrauskas

[9]) FDS (C˚)

Rockwool 292 281 Siporex 260 263

Table 3. Material data used in the numerical analysis. Heat conductivity of Rockwool is taken from SFPE handbook and the density is measured. The Siporex data is found in Ref. [15] and the Promatect

data is found on the suppliers home page (http://www.promat-ap.com/pdf/ph.pdf ). Material\Property Density (kg/m3) Heat conductivity

(W/m K) Heat capacity (kJ/kg K)

Rockwool 85 0.04 0.793 Siporex 500 0.15 1.000 Wall-cladding (Test A) 1115 0.199 0.793 Promatec (Test B) 975 0.242 1.000 During the first test, test A, a rather rapid fire growth was observed in the early stages of the test this dissipated and a shorter colder period commenced followed by the full force of the fire after 8-14minutes. Comparing with the previously defined HRR this does not include this pulsating feature of the fire in this particular test. Nevertheless, useful information could be obtained from studying such a case with this simplified model. In an earlier paper [1] very good correspondence between the test results and the simulations were achieved for the thermocouples whereas some of the plate thermometers where far off, in particular the temperatures measured at a distance from the fire source. The material data of the wall-cladding in Table 2 is found by measurements of the thermal conductivity and the heat capacity utilising the Transient Plane Source (TPS) method. Figure 5 shows a measurement station in one of the fictitious windows.

Figure 4. The measurement station at the lower edge of the fictitious windows in Test A.

In Figure 5, we compare the results from the plate thermometers placed in the lower edges of fictitious windows (along the symmetry line of the façade). We find that in the lower window where the fire impact is significant, the temperatures in the simulation are lower compared to those found experimentally whereas for the upper window there is a good correspondence between the simulation and experiment. When including plate thermometers in an FDS model describing a transient scenario the response time of the measurement device have to be considered, i.e. a thermal model for plate thermometers is needed [16-19]. In the model used in the present study, temperature dependent thermal properties of the components in the plate thermometer were used. The plate thermometer included a metal sheet of Inconel 600 and insulation material in the same manner as in Ref. [1]. A similar result is found for the 3 mm thermocouples placed 50mm off the façade beside the plate thermometers as seen in Figure 6. The temperatures in the lower window are under predicted in the simulation. It is interesting to note that the growth rate of the temperatures is similar except for the peak in the early phase of the test. There are many possible reasons for the discrepancies found between the experimental and numerical results such as radiation exchange, the simplified modelling of the fire source however in general the results are in good agreement at a certain distance from the fire.

Figure 5. Temperatures measured by plate thermometers at the fictitious upper window (UW) and

lower window (LW) during Test A compared to the simulation.

Figure 6. The temperatures at the lower (LW) and upper (UW) windows measured by shielded thermocouples in Test A compared to the simulation.

The velocities of the hot gases along the façade were estimated in test A by the use of bi-directional probes [20, 21], placed at the lower edge of the lower and upper window. The bi-directional probe measures the pressure differences in the gas in the direction of the flow. The measurement has to be compensated for the density change with temperature and the local Reynolds number according to simple relationship,

� = ��(��)�∆ �� . [m/s] [1]

Here k(Re) is a correction coefficient found in Olsson [21], T is the gas temperature, ∆p is the

0 5 10 15 20 25

Time [Minutes]

PT LW Exp

PT UW Exp

PT LW FDS

PT UW FDS

0 5 10 15 20

Time [Minutes]

TC LW 3mm FDS

TC LW 3mm Exp

TC UW 3mm FDS

TC UW 3mm Exp

pressure difference, ρ0 is the ambient density and T0 is the ambient temperature. Using the relation above, the velocities (averaged over 30 s) presented in Figure 7 are found. The upward components of the gas velocities found in the simulation have good correspondence to the experimental values. The correction coefficient is a measured quantity and is approximately constant for a large range of Reynolds numbers (Re). Roughly, in the range Re = 400 – 3800 a correction coefficient of k(Re) ≈ 0.9 is found. Although we did not account for the flange at the edge of the windows in the model the results from the simulations could capture the flow speeds and the gas temperatures rather well. Figure 7. Flow [m/s] along the façade at the two fictitious windows measured by bi-directional probes

compared to the simulation results.

In the second test, Test B, a Promatect wall-cladding was investigated in the rig instrumented with large plates (20x20cm) and thermocouples along the symmetry line at different heights. In this case we have used the measured HRR as an input to the simulation, see Figure 3. As starting point we compare the gas temperatures measured with 1 mm thermocouples placed at three different heights, 538 mm, 1613 mm and 2688 mm above the lower edge of the wall-cladding along the symmetry and 50 mm out from the façade, see Figure 8. The temperatures both in the experiment and the model show significant fluctuating behaviour and thus we have averaged the signals over 30 s. In two of the measurements, TC1 and TC2, the devices failed after a slightly longer time than 15 minutes. As shown in Figure 8 we find a good correspondence between the measured gas temperatures and the simulated cases. Here, it is interesting to note that if the previous estimated HRR input had been used the maximum gas temperatures in the simulations would be considerably lower than the measured ones, in the range 12-14 minutes. To save some computational time we stopped the calculation after approximately 15 minutes.

0 5 10 15

Time [Minutes]

Flow V1 FDS

Flow V2 FDS

Flow V1 Exp

Flow V2 Exp

Figure 8. The temperatures measured along the façade with thermocouples placed 50mm off the façade and at the heights 538mm, 1613mm and 2688mm above the lower edge of the façade wall-

cladding.

Next we investigate the effect on the 20x20cm plates placed along the symmetry line of the test rig at different three heights, 538 mm, 2688 mm and 3763 mm above the lower edge of the wall-cladding. Note that we have removed one plates from the results presented in Figure 9 in order to get good readability. The temperatures measured by PT1 are significantly higher than the simulated values whereas the two other plates are well represented in the simulation. There are two main reasons for this discrepancy namely the radiation exchange and the fact that during mounting of the plates a thin sheet of insulation of different type was placed behind it, the change of insulation could increase the temperature measured in FDS, in particular for the high impact case of PT 1, and could account for some of the discrepancy. Another reason could be the radiation-convection balance for the plates are not fully covered by the FDS model due to limited number of radiation angles and relatively large grid size compared to the plates. Along these lines we have computed the incident radiation on the plates, displayed in Figure 10. In the measurement the radiation is estimated using a constant flow velocity of 6m/s from the relation

��´´ = �� − �� !"# + %��

#%�%&. [2]

The heat transfer coefficient at the plates along the façade is estimated by an expression that is valid for forced convection

ℎ� = 2.4�+,.,-./0� �1 2 � �1 . [3]

Here Ts is the temperature of the plates, d = 0.7 mm is the thickness, x=0.2m is the characteristic length (the plates are 20x20cm and the emissivity ε= 0.9. The estimated values of the incident flux are compared to the computed incident radiation in the numerical model. We find that for the lowest placed plate the measured radiation is much higher than the numerical model accounts for, nevertheless, higher up on the façade for PT2 and PT3 a better correspondence between the two is found. Possible sources of error in this simplified calculation are deviations from constant temperature over the whole plate and the use of the measured temperature, by the 1mm thermocouples, 50mm from the surface.

0 5 10 15 20

Time [Minutes]

TC1 (30s)

TC2 (30s)

TC3 (30s)

TC1 Exp (30s)

TC2 Exp (30s)

TC3 Exp (30s)

Figure 9. The temperatures measured by the plate thermometers along the façade at the heights 538mm, 2688mm and 3763mm above the lower edge of the façade wall-cladding.

Figure 10. The incident heat flux [kW/m2] estimated from plate measurements.

DISCUSSION AND CONCLUSIONS

A good correspondence could be found between the shielded thermocouples in the FDS model compared with two experiments in the SP Fire 105 test rig. In the first test a design fire based on observations of the fire intensity during the test is used in the simulation. The used HRR is validated by previous measurements of HRR with the same test setup and the knowledge of the total fire load of 60 litres of heptane. In the second test, good agreement between the experimental data and the numerical model was found where the measured HRR was used as an input in the simulations. In

0 5 10 15 20

Time [Minutes]

PT 1 Exp

PT 3 Exp

PT 4 Exp

PT 1 FDS

PT 3 FDS

PT 4 FDS

0 5 10 15 20

Time [Minutes]

HF1 Exp (30s)

HF2 Exp (30s)

HF3 Exp (30s)

HF1 (30s)

HF2 (30s)

HF3 (30s)

both cases, discrepancies close to the fire source was found. Although, during the first test a rather rapid fire growth was observed in the early stages of the test this dissipated and a shorter colder period commenced, followed by the full force of the fire after 8-14minutes. Since the difficulties of the geometry, i.e. the flame suppressing lattice consisting of a perforated steel sheet with pipes of 25mm diameter placed in the fire tray, the numerical model could not resolve the flows close to the fire tray. Note that good information on the actual HRR in the test is invaluable information in reproducing right dynamics of the test rig, in particular where a pulsation in the fire intensity is obtained during test. Although a simplified model of the fire test rig was employed the calculated values for thermocouples correspond well with the measurements. The same correspondence was not found for the large plates (20x20cm) closer to the fire source at the lower window where substantially higher temperatures were sometimes found in the measurements. A possible source of this deviation is that the flanges at the lower edges of the fictitious windows were not included in the numerical model. The flow velocity along the façade was measured by bi-directional probes and the compensated velocities could be well reproduced by the simulation.

ACKNOWLEDGMENTS

The authors are grateful to the SP technicians for their help during instrumentation of the

façade experiment and Michael Rahm for sharing the experimental data from the second test. REFERENCES

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FIRE DYNAMICS IN FAÇADE FIRE TESTS: MEASUREMENT, MODELING AND REPEATABILITY

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