Tubos Sob Fogo - Davies

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17th International Conference on

Offshore Mechanics and Arctic Engineering

Copyright 0 1998 by ASME

OMAE98-3095

HEAT TRANSFER ANALYSIS OF GRP PIPES

IN FIRE-WATER SYSTEMS EXPOSED TO FIRE

J Michael Davies

The Manchester School of Engineering

The University o f Manchester, UK

Hong-B0 Wang

The Manchester School of Engineering

The University of Manchester, UK

ABSTRACT

This paper presentsa one-dimensionalnumerical heat ransfer

analysis used to develop predictive capabilities for the thermal

responsesof Glass-Reinforced Plastic (GRP) pipes exposed o

cellulosic or hydrocarbon fires. The behaviour patterns which

are usually associatedwith organic material exposed to fire,

such as thermal decomposition and expansion, are included in

the numerical model. The accuracy of the model is evaluated

by comparing the analytical and experimental temperature

history inside GRP pipes subject to standard ire tests. Empty

pipes and pipes filled with both stagnantand flowing water are

considered.

INTRODUCTION

In the offshore industry, it is necessary to design key

componentsof an installation to withstand the effects of fire.

One such component s the fire-water system which supplies a

deluge of water in the event of a fire-related emergency.

Traditionally, fire-water systemshave used steel pipework but

there is considerab le interest in replacing s teel with GRP

becauseof its corrosion resistanceand lighter weight. In order

to prove the viability of GRP dry riser pipes in this context, it

is necessary o demonstrate hat they can withstand the effect of

the fire until such time as the pipe is full of flowing water and

the deluge system s in operation.

This problem has been approached by a combination of

numerical analysis and tire testing.

A one-dimensional

analytical model for GRP pipes using polar coordinatessubject

to a prescribed time-temperature history applied during a

furnace fire test has been developed. The explicit finite

difference method was employed to solve the transient heat

conduction equation. Particular attention was given to the

effect of decomposition (charring and ablation) of the resin at

high temperature.

In order to verify the computational model, experimental

investigations of the performance of Ameron glass-reinforced

epoxy pipes and BP glass-reinforced phenolic pipes were

carried out. For tests on empty pipes, the pipe sections were

first cut to about 5OOmmength. Before placing each piece of

pipe inside thefurnace, the ends were blocked by packing them

with ceramic fibre. The furnace was then program med to

follow either the standard cellulosic or hydrocarbon fire

conditions [l] . In the simulated hydrocarbon fire test, a

notional dry time of live minutes was considered so that the

furnace was run for this length of time before the gas supply to

the furnace was switched off.

Water-filled pipes were tested in a specially-constructed

outdoor furnace which was also capable of following a

hydrocarbon time-temperaturecurve. A comprehensiveseries

of tests were carried out incorporating varying periods of dry

pipe or stagnant water conditions before initiating water flow

[2]. The results of two representative tests are chosen for

comparison with the theoretical results later in this paper.

A hydrocarbon fire features a rapid initial temperature

increase accompaniedby fierce burning and a high release of

heat flux. Under such circumstances, in which the average

furnace temperaturesoars o 944 C in 5 minutes, the GRP pipe

will undergo radical physical changes. In the initial stage,

when the outside surface of the pipe is exposed o an incident

heat flux, the temperature rise is a function of the boundary

conditions and the rate of heat conduction into the material.

As

the furnace temperature increases, the surface temperature

reachesa certain level (usually around 250C) beyond which

decomposition begins to occur and the resin components

degrade o form gaseousproducts and a carbonaceous esidue.

Extra energy will then be required in order to break the

constituent chemical bonds. The gaseousproducts, which are

initially trapped within the composite matrix owing to its low


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permeability, attain high internal pressuresand induce the solid

matrix to expand. In the meantime, the heat releasedby the

combustion of these escaping volatiles will feed back to the

pipes.

An advantageof the thermosetting resins which are normally

used in GRP pipes is that they do not melt when heatedowing

to their highly cross-lined chemical structures. As thermal

decomposition proceeds, a residual char layer builds up as the

pyrolysis front moves urther into the virgin solid. Initially, the

char layer provides an increasing thermal resistancebetween he

exposedsurface and the pyrolysis front as a consequence f its

low thermal conductivity and because t can only be ignited

with difficulty at normal oxygen concentrations. This is one

reason why GRP can provide a useful fire barrier performance,

despite the fact that it is an organic material. However, after

this initial phase, at very high temperature, the char is then

gradually oxidized and eroded away. Finally the heat resistance

of char layer will be totally lost and the wound glass-fibre

remains alone.

Over the past 50 years, a number of analytical models have

been proposedwhich consider the heat transfer in decomposing

materials [3 to 81.

They differ in their assumptions,

approximations, the phenomena ncluded and the property data

used. Although most of these studies were concentrated on

timber, they provide a theoretical basis for further investigation

of the fire performance of other combustible materials.

The

present study is along the lines of these previous works.

The heat transfer across the cross-sectionof an empty GRP

pipe in a furnace tire environment is first investigated. The

numerical results given by this analysis are successfully

correlated with the temperaturemeasurements ecorded during

the fire tests. The study is then extended o include the effect

of stagnantand flowing water in the pipe.

THEORETICAL FORMULATION

Since the physical and chemical processes re concurrent, as

described above, the problem of interpreting and predicting

observed thermal responses is a task of considerable

complexity. It is considered that the main task in the current

state of the art is not to form a complete model due to the

almost total absenceof information concerning the majority of

the material parameters involved.

The most promising

approach is to form a relatively simple model with the

minimum number of input parameterswhich can capture the

main features of the decomposition processand the consequent

heat transfer behaviour. Thus, several idealizations are made:

1)

The GRP material is assumed o be homogeneous nd

the transport of heat and mass is along the radial

direction of the cross-sectionof the pipe only, ie, the

problem is assumed to be one dimensional. It is

presumed hat the heat loading of the outside the pipe

is uniform in both the axial and circumferential

directions.

2)

There is

thermal equilibrium between the

decomposition gasesand the solid material and there

is no accumulation of these volatile gases n the solid

material.

3)

The rate of decomposition is assumed o conform to

a mean reaction which is described by a single tirst-

order Arrhenius function.

The governing principles on which the analytical model has

been developed are a combination of the principles of

conservation of mass and conservation of energy.

The one-

dimensional energy equation in a pipe undergoing thermal

decompositionexpresses balancebetween he transient energy

accumulation rate, with the sum of the rates of conduction,

pyrolysed convection, and the energy sink due to pyrolysis and

the heat feedbackby volatile combustion

$(ph) =

in which R, 5 r 5 R, , for t > 0, and where

;:

is the density (kg/m3)

is the solid enthalpy (J/kg)

t

is time (s)

r

is the polar coordinate (m)

T is the temperature C)

kr

is the thermal conductivity in the radial

direction (WlmC)

h*

is the enthalpy of gas (J/kg)

mg

is the mass lux of gas (kg/m*-s )

Q

is the heat of reaction (J/kg)

R, and R, are the inside and outside radii respectively.

Probably as a consequence f the pipe samplesbeing enclosed

within the furnace, it was found that the influence of heat

releaseby the combustionof volatiles was significant here.

The

heat of reaction Q was therefore defmed as the net sum of the

endothermal heat of decomposition and the heat feedback by

volatile combustion.

The specific enthalpies of the solid and volatile are

h = j-$dT ;

where

T, is the ambient reference temperature.

Equation (1) must be solved simultaneously with equations or

the rate of decompositionand the mass lux of the volatile. The

rate of decomposition, according to assumption 3, is

dp

L=

dt

(3)

where

pr is the instantaneousdensity of partially pyrolysed

resin

EA is the activation energy (J /mol)

R is the gas constant (8.314JKmol)

and

T is the temperature K).

The constantA is known as he pre-exponential factor and has


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units of s-l. The relationship between he p and pr is:

P

= P# -v) + P,V

(4)

where pp is the density of the glass-fibre and Y is the volume

fraction of glass-fibre. The glass-fibre is assumed o be intact

in the time zone of interest under tire.

In equation (3), it is assumed that the resin pyrolysis is

continuous until it is totally consumed. Some nvestigatorshave

included the final density of the char in expression (3).

However, this might cause two problems in application. One

is that the precise definition of the final char status s difficult

to defme. Another point is that yet another expression or char

pyrolysis will then be required if it commences its final

breakdown within the time zone of interest. Although much

researchhas been carried out into the thermal decompositionof

polymers, in view of the chemical complexity involved,

combined with problems of interpreting data from a variety of

sourcesand experimental assemblies, he available data is still

very limited and not in a form suitable for adopting in a more

complex treatment.

If the accumulation of gases s ignored, the conservation of

massmay be written approximately as:

3

ap

ar = -at

(5)

and the mass lux, %, at any spatial location and time can be

calculated by integration of equation (5).

Equation (1) is modified to its final form by expanding the

first three terms, substituting in the specitic heat and the

continuity equations, and rearranging. This results in

6)

:i(kmr$) - maCPgz - $JQ + h - hg)

where

C, is the specific heat of the material (J/kgC)

C, is the specific heat of the gas (J/kgC) .

Equations (3),(5) and (6) form a set of non-linear partial

differential equations which may be solved simultaneously for

p, mg and T respectively.

The boundary conditions on the exposedand unexposedsides

of pipe may be either a prescribed temperatureor a combined

radiative and convective boundary condition. To exclude the

uncertainties of the heat transmission rate from the fire to the

samplesunder test, the measured empera ture was used as the

boundary condition on the exposed outside surface. On the

unexposed nside surface, for empty pipes, an adiabatic surface

was assumed n the basis that the heat capacity of the inside air

is negligible and on the assumption of a uniform temperature

distribution on the inside surface.

Equations (3) to (6) were solved by a simple and effective

numerical scheme,namely the explicit finite difference method.

By using the energy balance technique, the finite difference

equation (7) for a typical interior node within the pipe wall can

be obtained.

Since the thermal expansion of the composite is included in

the mode l, the spatial intervals between the nodes change with

temperature,although they are uniform at first.

Six mesh nodes

across he thickness of pipe were found to be sufficient.

It was

assumed hat the inside surface did not move during expansion

at high temperature.

It was found that the second erm on the

right side of equation (6) can be ignored under the powerful

heat impingement condition of a standard ire test.

C

= T;+

2At

pCp[(rm - >)*rl + (rm + :)*rI

&[$ - O.,)(T;-i - T;)] - 8pQ;ct - hg)

(7)

For node 1, on the inside surface of an empty pipe, the finite

difference formulation is

Tt

= T;+

2k,, (ri + 0.5 &) At

p C ,

(ri + 0.25 &r)(&)*

(T; - T;)

(8)

~P(Q + h - hg)

G

THERMAL AND KINETIC PROPERTIES OF THE

MATERIAL

Usually, the thermal response of a composite material is

significantly sensitive to the thermal properties and also to the

stageof decomposition, e, the rate of density change.

This is

to be expected since the rate of energy consumption and the

thermal and transport properties of the material are functions of

the rate of pyrolysis.

The values of the pyrolysis reaction

kinetic parametersA and EA will determine the intensity and

duration of decompositioncorresponding o a given intensity of

heat flux. Since the accurate m-situ measurement of the

parameters A and EA (also the other thermal properties) is

almost impossible, the current practice is to initially assume

reasonable values based those available in the literature.

Subsequently, hesevalues are further refined by comparison of

the test and analytical results.


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Firstly, two lengths of Ameron 2000M glass-reinforcedepoxy

pipes were tested with 75 mm and 100 mm inside diameters

respectively.

The average hickness of the 75 mm pipe was 5.7

mm and that of 100 mm pipe was 5.4 mm.

For these pipes,

the kinetic parameterswhich were used n the calculation were:

A = 800 s-

EA = 56,000 J/mol for a cellulosic tire

and

A = 1200 s-

E, = 52,000 J/mol for a hydrocarbon fire.

The thermal conductivity K was set as 0.24 W/mC until the

char was totally lost.

Then the conductivity was assumed o be

given by the equation

k = 0.24 + 1.2 x 10e3 W/mC

(9)

The increase of conductivity at this stage s due to the char

being lost and cracking of the wound glass-fibre reinforcement.

The glass-libre content in the pipe by volume was about 4 1%

The variation of specific heat with temperaturewas

5

= 1270 + 0.230 x T (J/kg=C)

(10)

The heat of reaction Q was chosen as -3 x lo4 J/kg and the

coefficient of thermal expansion was 5 X lOa (per C).

Subsequently, hree 50 mm diameter BP phenolic pipes were

tested with thicknesses of 5.6 mm, 7.7 mm and 9.5 mm

respectively. The values of the kinetic parameters or these

pipes which were used in calculation under hydrocarbon fire

conditions were A = 1500 S and EA = 50000 J/mol. The

thermal conductivity was set as 0.3 W/mC, until the char was

oxidized and eroded away. Then, from that point, the thermal

conductivity was assumed o be

k, = 0.3 + 1.6 x 10m3T W/mC (11)

As with the epoxy pipes, the conductivity will increase as

temperature ncreasesdue to the loss of char and the cracking

of glass-fibre reinforcement. The glass-fibre content in the

pipes was 57% by volume and the density of the resulting

composite was 1900 kg/m3. The specific heat of the GRP

material was taken to be a bulk value which was weighted by

the fractional proportions of resin and fibre and was also

assumed o be a linear function of temperature:

C, = 1115 + 0.484 x T (J/kgC) (12)

The heat of reaction Q was -5 X lo4 J/kg and the coefficient

of thermal expansion was 4 X 10e4per C).

PIPES WITH STAGNANT AND FLOWING WATER

For pipes with stagnant or flowing water, the major change

required to the analysis describedabove s the internal boundary

condition. When a pipe is filled with flowing water, forced

convective heat dissipation is encountered. For flow inside a

circular tube, the rate of heat exchange between the water and

the pipe dependson whether the flow is laminar or turbulent.

In most fire-water systems, the nature of the flow inside the

pipe is turbulent.

An empirical formula which was proposedby

Nusselt is employed to calculate the internal heat transfer

coefficient h(T) from the pipe to the water [9]:

1

h(T) = 0.036R,0.8

0.0% k

6

(13)

for 10 < k < 400

where R,

is the Reynolds number

L

is the length of the pipe

D

is the inside diameter of the pipe

k

is the conductivity of the fluid

P,

is the Prandtl number of the fluid.

If the pipe is filled with s tagnantwater, the problem becomes

more complicated because t is difficult to describe the internal

circulation of the water and the heat transfer process is a

mixture of both conduction and convection. Since the heat

transfer coefficient under such conditions is unknown, our

treatment is to incorporate the effect of convection into the

conduction process. Using this equivalent conduction method,

it is shown later that the numerical calculation can provide a

good agreementwith the measured esults.

COMPARISON OF ANALYTICAL AND EXPERIMENTAL

RESULTS FOR DRY PIPES

As describedabove, both the hydrocarbon or cellulosic time-

temperature curves, as defined in BS476 [l] were used in the

fire tests.

In the hydrocarbon fire tests on dry (empty) pipes,

the gas supply to the furnace was switched off after following

the time-temperature urve for 5 m inutes. The outer and inner

surface temperatures of the pipes were measured during the

testsby several hermocouples. Although the variations among

them were small, it is the averageof them that was used in the

numerical comparison.

For the two Ameron epoxy pipes, the times for the inner

temperature o reach 200C were 3.2 minutes for the 75 mm

pipe in a cellulosic tire test and 1.7 minutes for the 100 mm

pipe in a hydrocarbon fire test respectively. The comparisons

of calculated and measured emperatureson the inner surface

for two pipes are shown in Figs. 1 and 2. The dashed ines

represent the transient temperature given by the numerical

analysis. A good correlation between both sets of results is

obtained.

For the BP phenolic pipes with three different thicknesses, he

corresponding imes for the inside temperature o reach 200C

were 1.86, 2.75 and 3.5 minutes respectively under simulated

hydrocarbon ire conditions.

The comparisonsof calculated and

measured emperatureson the inner surface or two of the three


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different pipes are shown in Figs. 3 and 4. The third pipe gave

similar results and correlation with the theory. Good agreement

between calculated and measured esults is again obtained.

Although, in order to keep the solution stable and to give

good accuracy, the time step At is chosen as 0.1 set, the

running time for the complete analysis is less than 1 minute on

a PC486. It is clear that the cost for numerical simulation is

insignificant.

There are, nevertheless, some discrepanc ies which exist

between he test and analytical results, especially at the cooling

stage. These might have been caused by following factors.

First, the numerical model does not precisely and explicitly

account for all of the physical and chemical processeswhich

occur within the materia l under fire attack, such as the heat

exchange between the volatile and solid and the effect of

blisters on the inside surface of the pipe at the final stagesof a

test. Secondly, most of the input data of the thermal properties

of the material was deducedon the basis of a limited number of

test cases and did not account thoroughly for the difference

between the heating up and cooling down stages. It is well

known that the values of thermal properties play a crucial role

in the numerical simulation and, unfortunately, there are

considerabledifficulties in the explicit determination of material

parametersunder fire conditions. Furthermore, at least part of

the part of limited published data can not be used directly

becauseof d ifferences in the preparation of the test sample, he

test assemb ly and the heat loading level. Nevertheless, the

results given in this paper have demonstrated hat consistent

results for GRP pipes with different thicknessesand diameters

under furnace fire conditions can be obtained using a single set

of material parameters.

The experimental and numerical study revealed that the

temperature ncrease in a plain dry GRP pipes under the fire

impingement is rapid and that failure o f unprotected dry pipes

invariably occurred within the first five minutes of testing. To

overcome this problem, it has been shown that additional

protection, such as a coating of fire-retardant intumescent, a

wrapping of mineral or ceramic wool, or filling the pipe with

stagnant or flowing water, can considerably reduce the fire

damage o GRP pipes and retain their function for a satisfactory

period.

COMPARlSON OF ANALYTICAL AND EXPERIMENTAL

RESULTS FOR WATER-FILLED PIPES

A comparison of the temperatureat the inside face of the pipe

give by the above theory and that measured n a test for a 50

mm diameter Ameron 2000M pipe filled with water flowing at

18 litres/minute is given in Fig. 5 . A sim ilar comparison for

a pipe tilled with stagnant water is given in Fig. 6. In both

cases, he material characteristics or the pipe which were used

in the analysis were identical to those for dry pipes. The

agreement s good and, evidently, the theory derived for dry

pipes can be extended to water filled pipes with similar

accuracy.

Evidently, the presence of stagnant water leads to a much

better fire performance han that of a dry pipe.

Once the water

in the pipe starts to flow, the situation becomes stable with

negligible further deterioration of the pipe. These conclusions

are born out by the additional test results given in Reference2.

Evidently, only a limited amount of fire protection is required

to the pipes of fire-water systems n order to sustain the pipe

through the initial emergency period until the water starts to

flow. This scenario can be adequatelymodelled, thus reducing

the requirement for fire testing.

ACKNOWLEDGEMENT

The authors wish to acknowledge he financial support given

by the Marinetech Research managed programme The Cost

Effective Use of Fibre-Reinforced CompositesOffshore and ts

twenty five industrial sponsorsand EPSRC.

REFERENCES

Dl

PI

131

141

[51

161

[71

181

191

BS476: Part 20: Fire tests on building materials and

structures, British Standards Institution, 1987 and

AMD 6487, 1990.

Davies, J. M., Dewhurst D. and Wang, H-B.

Characterisation and modelling of the performance

of wet and dry GRE pipes under hydrocarbon tire

conditions, Proc. 7th Int. Conf on Fibre Reinforced

Composites , University of Newcastle upon Tyne,

April 1998.

Bamford, C.H., Crank, J. and Malan, D.H. The

combustion of wood, Part 1 , Cambridge Phil, Sot.

Proc., Vol. 42, pp 166-182, 1946.

Kung, H.C. A mathematical model of wood

pyrolysis, Combustion and Flame, Vol. 18, pp 185-

195, 1972.

Matsumoto, T., Fujiwara, T. and Kondo, J.

Nonsteady thermal decomposition of plastics,

Twelfth Symposium (International) on Combustion,

The Combustion Institute, pp 515-521, 1969.

Kanury, A. M. and Holve, D. J. Transient

conduction with pyrolysis (Approximate solutions for

charring of wood slabs), J. of Heat Transfer, Vol.

104, pp 338-343, 1982.

Fredlund, B. Modelling of heat and mass ransfer in

wood structures during fire, Fire Safety Journal,

Vol. 20, pp 39-69, 1993.

Henderson, J. B. and Wiecek, T. E.

A

mathematicalmodel to predict the thermal responseof

decomposing, expanding polymer composites, J. of

Composite Materials, Vol. 21, pp 373-393, 1987.

Ozisik, M N

Heat transfer, A basic approach,

McGraw-Hill Book Company, 1985.


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1000

Temperature C

600

400

200

0

Cellulosic curve

External pipe temperature

0 2 4 6 6 10 12 U 16 18 20

FIGURE 1: AMERON 2000M 75 MM GR EPOXY PIPE - CELLULOSIC FIRE TEST

alculated i

0 1 2 8 4 6 6 7 8

FIGURE 2: AMERON 2000M 100 MM GR EPOXY PIPE - HYDROCARBON FIRE TEST

012346876

FIGURE 3:

5.6 MM THICK BP PHENOLIC PIPE - HYDROCARBON FIRE TEST


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1200

Temperature

C / I

I

i

1000

I

j


0

2

4 6 8 10

12 U 16

FIGURE 4: 9.5 MM THICK BP PHENOLIC PIPE - HYDROCARBON FIRE TEST

1200

Temperature C

400- . . . . .

. .... I T ..,... .........

Internal pipe temperature

200

: Calculated

Measured

0

25 30 36

Time in minutes


1000

Temperature C


J

,.

00-....

600--

.\ : ...

, .............

>

.......

...

.:.

..........

>

.,,,,_ ......

..;. . ...... .... ...........

FLOWING WATER AT 18 LITRES PER MINUTE

Internal water temperature

Measured

: 1 ; +alculated

I

I

I

0 1

2 3

4 6 6

Time in minutes


STAGNANT WATER

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