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INTERNATIONAL JOURNAL OF MODELING AND SIMULATION FOR THE PETROLEUM INDUSTRY, VOL. 8, NO. 2, DECEMBER 2014 17
Analysis of Oil Tanker Deck Under Hydrocarbon Fire
Manco, M. R.1, Vaz, M.A.
1, Cyrino, J.C.R.
1 and Landesmann, A.
2
[email protected], [email protected], [email protected], [email protected]
1Ocean Engineering Program, COPPE / UFRJ, Rio de Janeiro, Brazil 2Civil Engineering Program, COPPE / UFRJ, Rio de Janeiro, Brazil
Abstract. The present paper presents a numerical study of the
behavior of a deck panel of an oil tanker under fire. The fire
condition assumed in the simulation is that of a hydrocarbon
burning process according to the time-temperature nominal
curve in the part 1.2 of EUROCODE 1 and 3 (EC1 and EC3,
2004), which also specify the variations of mechanical and
thermal properties of the steel with temperature. A finite
element model taking into account the initial imperfections
recommended by the ISSC, 2012 was developed through the
use of ABAQUS , 2011 commercial software. The thermal
and mechanical analyses were uncoupled, so the temperature
transient field caused by the fire condition was analyzed first
and then, this thermal load was applied to the structure to
evaluate its mechanical behavior. Once the fire scenario was
defined, it was possible to assess the development of the
temperature field as a function of time. The induced thermal
loads were considered in the analysis of the structural model
together with the pre-existent operating loads, allowing
assessing the behavior of the panel. Thus, the objective of the
present paper is to present a methodology for assessing the
structural behavior of a steel panel in the event of fire.
Index Terms stiffened panel; fire; finite element
1 INTRODUCTION
The design of steel structures for the offshore exploration and
production of oil and gas provides for the consideration of
different scenarios related to severe accidents, including: large
waves, extreme winds, earthquakes, collision between vessels
and fires. For all these conditions, the integrity of the facility
must be ensured and, in some cases, damages must be limited,
ensuring the maintenance of safe operating characteristics for
a given period of time. Due to the presence of large volumes
of flammable materials (liquid and gas), equipment operating
at high temperatures, active flames (e.g., flares) and human
lives, living and interacting in confined spaces, the manufac-
turers and operators of these facilities are required to follow
strict fire protection criteria (deemed the most severe among
all industrial facilities).
The occurrence of a fire in this type of structures is considered
one of the most unfavorable conditions, for the combustion of
hydrocarbons presents a very high rate of temperature increase
in the initial stage of the fire, causing a very rapid loss in me-
chanical properties of steel, thus generating the possibility of
deaths and economic and environmental damages.
Tragic examples of this kind of accidents are those which took
place in the Piper Alpha platform in the UK, in July 1989,
killing 167 of the 229 occupants in less than 22 minutes and,
most recently, at the Deepwater Horizon platform, in the Gulf
of Mexico in April 2010, where 11 people disappeared and a
large environmental impact was generated.
Temperature changes generate degradation of the mechanical
properties causing effects that significantly change the failure
mode forecasted in the panel design (at room temperature).
Such effects are related to the introduction of axial and shear
forces, as well as bending moments due to the thermal gradi-
ent, causing the local buckling of the web (WB) and the flange
(FB) at the ends of the stiffener and its consequent collapse
because of the emergence of plastic hinges and subsequent
membrane behavior of the panel plate until the complete fail-
ure of the panel (Manco et al., 2013). In addition, the adopted
boundary conditions (described in item 3) facilitate the prob-
lem modeling due to the consideration that the properties in
the restraint do not change with temperature, but generate a
numerical problem in the mechanical analysis which is solved
by considering an additional region called Rigid End (Has-
sanein, 2011).
The numerical analyses are performed using the commercial
code ABAQUS, 2011, according to the finite element meth-
od (FEM), taking into account the structural and thermal ef-
fects resulting from the proposed fire. Variation in thermal and
mechanical properties of materials in case of high temperature
conditions are taken into account in the analysis, in accord-
ance with the applicable standard recommendations, such as is
the case of part 1.2 of EC3, 2004. The fundamentals of the
applied analysis model are described in item 2 below. A case
study is proposed and briefly described in item 3 hereto. A
tanker ship deck panel, with the geometry employed by the
ISSC, 2012 in their benchmarking studies, is submitted to a
fire scenario, caused by the burning of hydrocarbons (Part 1.2
of EC1, 2004), allowing to evaluate the thermo-structural
behavior for different instances of the fire. The main results
obtained with the numerical model are presented and critically
assessed in item 4 of this work, taking into account the ful-
fillment of security requirements. The main conclusions ex-
tracted from the analyses performed are mentioned in item 5,
indicating that this methodology allows assessing qualitatively
and quantitatively the behavior of the panel, and can be used
in the improvement of current regulations related to the safety
of structures in the event of fire.
Manuscript received Jan. 26, 2014. Corresponding Author: Miguel Renato
Manco Rivera (E-mail: [email protected]).
MANCO et al.: ANALYSIS OF OIL TANKER DECK UNDER HYDROCARBON FIRE 18
2 ANALYSIS METHODOLOGY
The analysis is carried out using FEM, the model includes a
direct and rigorous consideration of nonlinear physical and
geometric effects on the numerical formulation, allowing the
estimate of the possible structural collapse modes.
The proposed analysis procedure begins with the review of the
panel layout with the selection of the fire scenario. Then, the
thermal analysis is performed, which purpose is to determine
the variation of the temperature in the elements exposed to
fire. The main numerical formulation aspects of this stage are
addressed in item 2.1. The final stage of the procedure aims at
determining the structural behavior as a function of the
elapsed time of fire, in other words, depending on the thermal
conditions of fire exposure and applied external loads (me-
chanical). Computational characteristics adopted in this final
stage of the numerical simulations are briefly described in
item 2.2 hereto.
2.1 Thermal Analysis
The numerical model used FEM to solve the two-dimensional
transient heat conduction problem, as shown in Cook, 2002,
Skallerud and Amdahl, 2002, Lewis et al., 2004 and Landes-
mann et al., 2010. The DS4 element made with 4 nodes was
used to represent the panels.
The partial differential equation which expresses the tempera-
tures (in degrees Celsius) (x,y,z,t) is shown in Equation (1), subject to a temperature field defined in its contour s, which is represented in this analysis by fire-temperature curves (Part
1.2 of EC1, 2004).
tc
zk
zyk
yxk
x
..... (1)
Where is the specific mass of steel (assumed temperature independent), =7850 kg/m, c is the specific heat and k is the thermal conductivity. In this paper, the thermal properties of
steel, as a function of temperature, are provided by part 1.2 of
EC3, 2004 and shown in Figure 1.
When prescribed temperatures are different from temperatures
on the surface, a heat flux qn with two portions is generated:
(i) one due to convection and another (ii) resulting from radia-
tion, which can be written in a single equation through the
linearization of the radiation portion, as below:
gsqqq .rcrcn (2)
where: gsgs ... 22rrr , r is the resulting emissivi-ty, defined as 0.8 (for steel); r is the Stefan-Boltzmann con-
stant (5.67 .10-8
W/mK4); and c is the convective heat coeffi-
cient adopted as 50 W/mK (part 1.2 EC1, 2004). This lineari-
zation of the portion of radiation is necessary given that the
FEM only solves a system of linear equations. Denoting C as
capacitance matrix, Kl and Kc are conductivity matrices (Kt=
Kl+Kc), fb as vector of nodal flux due to convection, Equation
(1) can be rewritten:
Figure 1. Specific heat and thermal conductivity of carbon steel as a function
of the temperature
)()(.)(
. tftKt
tC bt
(3)
Solution of Equation (3) is based on FEM, being possible to
determine the temperature at time n+1 based on data at time n:
bnbnbnntnt ffftKtCtKC 11 1 (4)
where t is the time interval, is the temporal integration factor (taken as 0.9) and the initial temperature throughout the
solid is assumed to be equal to 20C (o). Analyses presented here use the nominal fire curve corre-
sponding to the burning of hydrocarbons (Part 1.2 of EC1,
2004), as given by Equation (5):
tt eet .5.2.167.0og .675.0.325.01.1080)( (5)
where: t is the elapsed time of fire (in minutes), a is the tem-
perature in the middle (in C) and o is the initial temperature
(equal to 20C).
2.2 Structural Analysis
Since the variation of the temperature field was established in
the previous analysis stage, the finite element mesh used, i.e.,
the nodal coordinates, the elements connectivity and the re-
sults for heat fluxes are used in the simulation of structural
behavior under the postulated fire conditions. The procedure is
initialized by the application of external loads, including the
own structural weight, fluid action and other operational loads.
At this stage, deformations and their respective stresses, corre-
sponding to normal operating conditions of the panel, can be
seen. The variation of the temperature field determined in the
thermal analysis is imposed to the structural model along with
other external loads applied.
In building the mesh of finite elements for the structural anal-
ysis, the S9R5 element is used for the panel simulation. This
element is composed of 9 nodes and 6 degrees of freedom per
node (translations and rotations around global axes X, Y and
INTERNATIONAL JOURNAL OF MODELING AND SIMULATION FOR THE PETROLEUM INDUSTRY, VOL. 8, NO. 2, DECEMBER 2014 19
Z), with capacity for developing nonlinear physical and geo-
metrical analyses. The complete Newton-Raphson solution
process is adopted to update the matrices and the linear solu-
tion of equations. The von Mises criterion is adopted for de-
termining the element plastification criterion.
Apart from the thermal deformation imposed on the structural
model, variations in the mechanical properties of steel as a
result of temperature, as shown in Figure 2, are also taken into
account, including reduction: of yield strength (y,), modulus of longitudinal elasticity (Ea,) and yield point (p,) obtained based on recommendations of part 1.2 of EC3, 2004.
Figure 2. Stress-strain relationship for carbon steel at elevated temperatures.
Defining y,20 as the characteristic yield stress and Ea,20 as the modulus of elasticity, at environment temperature, the degra-
dation factors as well as the coefficient of linear thermal ex-
pansion of steel, as a function of temperature (), are shown in Figure 3.
Figure 3. Reduction factor for the stress-strain relationship and thermal
expansion coefficient for carbon steel at elevated temperatures.
3 CASE STUDY
We studied a stiffened steel panel, which is part of the deck of
an oil tanker, subjected to fire caused by hydrocarbon burning.
The temperature of the hot gases inside the compartment un-
der fire (on the side of the stiffeners) is described by Eq. (5),
while the external temperature was considered constant and
equal to 20 C. In the heat exchange process between the up-
per side of the plate and the environment we considered a heat
exchange coefficient (including radiation and convection) of 9
W/mK according to part 1.2 of EC1, 2004 recommendations
as shown in Figure 4. The geometry of the panel was chosen
similar to the ISSC, 2012 benchmark study and is shown in
Figure 4. The initial geometric imperfections of the panel were
based on the model recommended by ISSC, 2012 according to
Eqs. (7) and (8), where oplv is the imperfection in the plate,
osw is the imperfection in the stiffener and is a parameter
that defines the level of imperfection according to Smith et al.
(1992) (slight =0.00025, average =0.0015 and severe = 0.0046). In the case study we employed the three levels of initial imperfections of the stiffener to assess the influence on
the behavior of the panel under fire. Figure 5, presents the
initial imperfections magnified fivefold.
Figure 4. Case study and initial panel geometry
MANCO et al.: ANALYSIS OF OIL TANKER DECK UNDER HYDROCARBON FIRE 20
puopl
b
y
L
xmv
sinsin1.0 2 (6)
uwuos
L
x
h
zLw
sin (7)
The computational mesh used is composed of three regions.
The first region, in blue in Figure 5, is called Rigid End and
simulates the panel support element (bulkhead or deck trans-
verse) the mechanical properties of which are independent
from temperature. This end is used to avoid numerical errors
caused by the application of boundary conditions in the me-
chanical analysis. The second region, in gray in Figure 5,
covers length uL /12 from the rigid end and has a higher mesh
density to allow assessing the WB and FB of the stiffener that
will occur in that area. Finally, the third region in green in
Figure 5 is considered with a less dense mesh to avoid work-
ing with very large matrices. As boundary conditions we con-
sidered the left cross-section of the Rigid End (see Figure 5) as
restrained, while for the right end of the region with less mesh
density the symmetry condition was considered to work only
with half the panel and reduce the computational cost. Finally,
in the side edges we considered the boundary conditions that
represent the continuity of the panel, as shown in Figure 5.
Other boundary conditions employed in similar problems can
be seen in the studies of Heninisuo & Aalto, 2008 and
Vimonsatit et al., 2005, among others. As loads, we consid-
ered only the structure own-weight and three values for side
pressure of 0.01, 0.02 and 0.03 MPa for each of the initial
imperfection levels considered in the stiffener.
Figure 5. Initial geometric imperfections and boundary conditions
4 RESULTS
In all the cases analyzed the failure on the deck presents the
same behavior, only varying the severity of the stiffener dis-
tortion and the time it takes to occur as described below.
4.1 Temperature Field
As expected, the stiffener web is the element that heats more
quickly, because it has a greater massivity factor, i.e., presents
a larger area exposed to fire with a relatively low volume,
when compared to the plate or the stiffener flange. The flange
has an almost uniform heating as evidenced by the little tem-
perature difference between the points D and E. In the plate,
the temperature difference between points A and B, indicates
the presence of a heat flow towards the stiffener. Note that the
temperature on the plate is always lower than on the stiffener
due to the heat exchange with the environment.
Figure 6. Temperature in points A, B, C, D and E, in view of time
4.2 Stress-strain Field
Since we choose a fire scenario caused by hydrocarbon burn-
ing, considered the most severe possible fire, among the sim-
plified models proposed by the part 1.2 of EC1, 2004, the
temperatures of the different panel components increased very
quickly, originating a pronounced drop of the mechanical
properties of the structural elements. According to Yang and
Gao (2004), those temperature increases, besides changing the
thermal and mechanical properties, generate temperature gra-
dients in the panel constitutive elements, resulting in non-
linear forces and moments which, combined with the increas-
ing of imperfec-tions due to thermal deformation, change the
behavior of the panel. Thus, in all the analyses developed from
the data on panel temperature field variation, we determined
the status of stress and strain at different times of the postulat-
ed fire.
Figure 7, presents the stress fields at the beginning of the heat-
ing, i.e., after considering the effect of gravity and the corre-
spondent lateral pressure L.P. It was observed that the initial imperfection does not significantly affect the results, changing
the stress magnitude by less than 2%. In this figure, the cases
with an average stiffener imperfection for each of the side
pressures considered (0.01, 0.02 or 0.03 MPa) are presented.
During the heating, the panel suffers non-uniform thermal
deformations that change the initial stress field, giving rise to
regions where the plastic regime is reached, generating plastic
INTERNATIONAL JOURNAL OF MODELING AND SIMULATION FOR THE PETROLEUM INDUSTRY, VOL. 8, NO. 2, DECEMBER 2014 21
hinges in the area close to the restraint and in the middle sec-
tion of the panel. Those plastic hinges begin with the stiffener
WB and subsequently with the FB.
Figure 7. Stress fields after the application of gravity and L.P. considered
(panel with average imperfection)
Figure 8 shows the distribution of von Mises stresses ( ,vM ),
axial stresses in the S11 and S22 directions ( ,11 and ,22 ,
respectively) and the shear stress ( ,12 ), normalized in rela-
tion to the yield stress (at temperature determined at the
point at that instant) in the stiffener web. The continuous,
dashed and dotted lines represent the stress distribution for
times of 0 (beginning of heating), 5 and 10 min., respectively
and the colors black, red and blue define the cross-sections
located at 2/uLx , 36/uLx and 72/uLx , respective-
ly. For not having trouble viewing (since the stiffener de-
forms) we used a local coordinate on the vertical axis at the
initial time ( LoY , , the origin of which is the junction of the
plate and the web). In this figure we observe that the von Mis-
es stress has a similar configuration as the axial stress ,11 ,
indicating that it prevails over the others ( ,22 and ,12 )
and that after the appearance of the plastic hinge (graph ,11
black dashed or dotted line) the panel supports only tensile loads and basically shows the panel yield (membrane behav-
ior) until it breaks.
Figure 9 shows the geometric configuration of the panel in the
region close to the restraint ( 36/uLx and 72/uLx ) and
in the middle of the panel ( 2/uLx ) at different times, for
the case of a severe imperfection in the stiffener and a L.P. of
0.03 MPa. This figure evidences WB and FB of the stiffener in
the region close to the restraint (plastic hinge) and maximum
vertical displacement (MVD) as well as maximum transversal
displacement (MTD) in the middle of the panel.
Table 1, presents the analysis times, until panel failure for
each of the cases evaluated. We found that neither the imper-
fection level nor the magnitude of the side pressure (the value
of which was triplicated) significantly changes the analysis
time, varying 5% at the most. This happens because the prob-
lem is governed by temperature and not by the geometric
configuration or the load. MVD and MTD magnitudes are also
presented showing that their values are directly affected by the
S.P., on the other hand it also shows that the level of imperfec-
tion has a very large influence on the value of MTD and on the
stiffener final configuration.
Figure 10 shows the final geometric configurations of the
stiffener web for all the cases assessed, measured in a local
reference system, i.e. in a system with origin in the junction of
the plate with the web in the respective x coordinate. It is
worth emphasizing the important vertical displacements suf-
fered by the panel that surpass 60 cm in the middle section,
anyway, the lateral displacements of the stiffener indicate loss
of its stiffness and the behavior of the plate as a membrane. It
should be mentioned that the MTD occurs in some cases in the
middle of the panel and in other cases in the region near the
restraint.
The same Figure 10 also shows that the magnitude of the
imperfection slightly alters the final geometric configuration.
In this manner a slight imperfection in the stiffener originates
a distortion in the opposite direction to the initial imperfection,
however for average and severe imperfections the slope of the
final configuration accompanies the direction of the stiffener
initial imperfection.
11
22
12
MANCO et al.: ANALYSIS OF OIL TANKER DECK UNDER HYDROCARBON FIRE 22
Figure 8. Stress fields after the application of gravity and L.P.= 0.03 MPa (panel with average imperfection)
Figure 9. WB and FB Figure 9 in the region close to the restraint (L.P.= 0.03 MPa and average imperfections).
Plastic
hinge
INTERNATIONAL JOURNAL OF MODELING AND SIMULATION FOR THE PETROLEUM INDUSTRY, VOL. 8, NO. 2, DECEMBER 2014 23
Figura 10. Final configuration of the stiffener web in all cases analyzed
TABLE 1: ANALYSIS TIMES UNTIL PANEL FAILURE, MAXIMUM VERTICAL DISPLACEMENT (MVD) AND TRANSVERSAL DISPLACEMENT (MTD) VALUES FOR ALL THE
CASES EVALUATED.
Lateral
Pressure (MPa)
Level of imperfection in the stiffener
Slight Average Severe
Time Failure
(min.)
MVD
(mm)
MTD
(mm)
Time Failure
(min.)
MVD
(mm)
MTD
(mm)
Time Failure
(min.)
MVD
(mm)
MTD
(mm)
0.01 39.1 215.4 5.4 39.1 215.2 7.6 39.0 214.9 15.1
0.02 42.5 396.2 6.5 42.2 396.0 11.5 41.2 395.9 25.2
0.03 42.5 574.7 36.3 43.2 574.7 36.3 44.6 621.8 235.0
5 CONCLUSIONS
The numerical-computer methodology for analyzing the be-
havior of steel structures under fire conditions presented in
this paper was applied to evaluate the behavior of a stiffened
panel of the deck of an oil tanker submitted to a fire scenario.
Despite the idealized load conditions the thermo mechanical
behavior could be observed at different times of the postulated
fire, but it should be mentioned that the results obtained are
valid only for the load and boundary conditions taken into
consideration. In a real situation, when the structure suffers the
action of the waves and other loading conditions, a different
behavior can occur. Anyway, as shown by Manco et al., 2013
the influence of the outline on the longitudinal edge affects the
behavior of the stiffener facilitating its distortion. Due to the non-linearity of the distribution of the stresses pre-
sent in the panel during the fire, it is evident that such an anal-
ysis is very complex and cannot be addressed by simple ana-
lytical formulations and thus the use of a numerical method is
necessary. Anyway, we proved that the problem of the ana-
lyzed cases is governed by the temperature and not by the ge-
ometrical configuration or the load.
MANCO et al.: ANALYSIS OF OIL TANKER DECK UNDER HYDROCARBON FIRE 24
From the results we concluded that is necessary to apply ele-
ments of passive protection. It should be highlighted that con-
siderably severe fire conditions were forecasted by assigning
the standardized curve for hydrocarbon burning. More refined
studies can be applied in order to make the simulations of fire
scenarios more realistic, as for instance employing CFD
(Computational Fluid Dynamics) models and, consequently,
the mechanical behavior of the structure can be estimated in a
more reliable way.
The conclusions obtained through the numerical simulations
indicate that the methodology presented in this paper can be
applied to assess the structural behavior of offshore structures,
with different load conditions and different materials for ther-
mal protection and has a potential use in the reduction of
structural passive protection without impairing the levels fore-
casted for global safety. The layer of thermal protection, usu-
ally employed in this type of structure, delays the heating of
the protected element, helping to maintain the mechanical
properties for a longer time, improving the panel structural
behavior.
ACKNOWLEDGMENT
The authors wish to express their gratitude to the National
Petroleum Agency of Brazil (ANP), PETROBRAS and
COPPE-UFRJ for their support for the development of this
work.
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