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ASIAN PIPELINE CONFERENCE & EXHIBITION
27th 28th September 2005
Crowne Plaza Mutiara Hotel, Kuala Lumpur
Jointly organized by: ASCOPE Gas Centre, Malaysian Gas
Association and Petromin magazine
DESIGN OF HIGH
TEMPERATURE/HIGH PRESSURE (HT/HP) PIPELINE
AGAINST LATERAL BUCKLING
Mr Lim Kok Kien JP Kenny Wood Group SDN BHD
1
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DESIGN OF HIGH TEMPERATURE/HIGH PRESSURE (HT/HP) PIPELINE
AGAINST LATERAL BUCKLING
Lim Kok Kien*, Dr. Lau Siew Ming* and Dr. Emil Maschner
*JP Kenny Wood Group Sdn. Bhd., Kuala Lumpur, Malaysia.
JP Kenny Engineering Ltd, Staines, London, UK.
ABSTRACT
Recent development in Malaysian waters has involved several high
temperature-high pressure (HT/HP) pipelines. At high temperature
and pressure, significant compressive forces can develop in the
pipeline, due to restricted thermal expansion. As a consequence,
the pipeline tends to release the compressive forces by undergoing
a secondary equilibrium configuration, i.e. by buckling or snaking.
The conditions in which buckling occurs depends on several factors
such as soil friction, pipe weight, pipe sectional properties and
the initial out-of-straightness (OOS) of the pipe.
As a result, the design methodologies for HP/HT pipelines differ
significantly from traditional pipeline design. This paper will
present a case study on a recently completed project by JP Kenny
Wood Group Sdn. Bhd.. It involves a 28-inch gas pipeline on a flat
seabed consisting of soft soil. The design temperature and pressure
is 96C and 157 barg respectively. A design strategy using
strain-base criteria was adopted, incorporating a pipeline lay over
vertical buckle triggers.
In addition, the behaviour of HP/HT pipelines will be discussed
with reference to general theory of effective force and thermal
expansion. Furthermore, design strategy and methodology involving
various finite element modelling techniques will also be discussed,
in particular, the different methods used in determining the
suitable number of buckle triggers and its location.
PIPELINE UNDER TEMPERATURE & PRESSURE Background Pipelines
operating at temperatures and pressures above ambient will tend to
expand, due to thermal and pressure loading. If the pipeline is
constrained, either partially or fully, a compressive axial force
will develop in the pipeline. The magnitude of the compressive
force depends on the extent of constraint applied to oppose the
expansion. For an untrenched subsea pipeline, axial constraint
arises in the form of seabed soil friction and/or the flexibility
at the end tie-in. For a short pipeline, the total axial friction
is insufficient to constrain the pipeline fully. A typical
effective force along the pipeline is as shown in Fig 1. The two
ends of the pipeline will expand, moving in the opposite direction
of each other. Thus, the friction force changes direction at an
equilibrium point where no axial expansion occurs. This is known as
a virtual anchor point. For a sufficiently long pipeline, the build
up in frictional resistance will exceed the axial force required to
fully constrain the pipeline. In such cases, certain portion of the
pipeline will be fully constrained, while the other sections are
free to expand (but still remain in compression due to resistance
from friction). Two virtual anchors will develop in this case as
shown in Fig 2.
Typical effective force for a 'short' pipeline
-3500
-3000
-2500
-2000
-1500
-1000
-500
00 2 4 6 8 10 12
KP distance (km)Fully constrained axial force Friction/effective
force
virtual anchor point
Fig 1 Effective axial force of a short pipeline
Typical effective force for a 'long' pipeline
-2500
-2000
-1500
-1000
-500
00 5 10 15 20 25 30 35
KP distance (km)
Forc
e (k
N)
virtual anchors
friction force
fully constrained axial force
effective force
fully constrained section
Fig 2 Effective axial force of a long pipeline Therefore, on a
flat seabed, a pipeline under temperature and pressure loading will
always be under compression as a result of friction limiting its
expansion. A positive (tensile) effective force, however, can
develop if the seabed is highly irregular or if there is seabed
subsidence. Pipelines, therefore can be divided into two groups: -
Long pipelines which develop the full constrain axial force Short
pipelines that never develop the full constrain force. Response
under compressive load The global response of pipelines under
compression depends on the level of compression developed under the
thermal loading cycle. If the effective compressive force exceeds a
certain threshold, the pipeline will deform (globally) into a new
equilibrium shape in order reduce the compression. This
response
2
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whereby a structure obtain a new equilibrium state by seeking a
large deflection is called buckling and the load necessary to
initiate buckling is the critical buckling load. A buried pipeline
will buckle vertically if the vertical restoring force (pipe and
soil weight) were less than the horizontal (sideways) restoring
load. For an untrenched pipeline on irregular seabed, the tendency
is to initially buckle vertically and subsequently move laterally
on the horizontal plane, as the frictional restoring load in most
cases is less than the weight of the pipeline. The critical
buckling load depends on the pipe properties, weight, friction
factor and initial OSS. The problem of pipeline buckling had been
considered extensively by Hobbs [Ref. 2] using analytical methods.
Experimental work performed as part of his study has found that
pipeline can buckle into different lateral mode shapes; the most
common of which (Modes 1 to 4) are shown in Fig 3.
Fig 3 Possible lateral buckling modes
A buckle region consists of the buckle itself flanked on both
sides by two slipping regions. The slip regions will continue to
expand and feed into the buckle if temperature increases further
after the buckle has developed. The different regions in a buckle
are shown below in Fig 4.
Fig 4 Different regions in a buckle
The length of the slip zone depends on the available frictional
resistance to oppose the feed in. A virtual anchor is developed at
the point where there is sufficient frictional forces to constrain
the slip completely. Upon formation of the buckle, the effective
force (post-buckle) changes to account for the reduced compression
in the buckle and the feed in from the slip zones. The post-buckle
effective force after the development of a single isolated buckle
is depicted below in Fig 5. Further increase in temperature in the
pipeline after post-buckling will increase the slip length. This
causes more pipe length to feed into the buckle and therefore
increases the moment at the buckle. It is possible that more than
one buckle develops along the pipeline. In this case, depending on
the distance between the buckles, the feed in will be shared among
the buckles. If the buckles are spaced such that the distance
between successive buckles is less than the total buckle length (Lo
+ 2Ls) of an isolated buckle, the feed in is shared between the two
buckles. This is illustrated in Fig 6.
Fig 5 Post-buckle effective force of a single buckle
Eff. force
KP distance
Slip zones
Fig 6 Post-buckle effective force with multiple buckles PIPELINE
PARAMETERS Operating conditions Temperature profile. The design
temperature variation along the entire pipeline is shown in Fig
7.
Design temperature profile
0102030405060708090
100
0 5 10 15 20 25 30 35 40 45KP distance (km)
Fig 7: - Design temperature profile
Seabed soil condition. The seabed is flat with soft/very soft
clay type soil. The undrained shear strength ranges from 0.5-6 kPa
for a depth of 0.5m below the mudline. High embedment of the
pipeline is predicted due to the soft nature of the seabed soil,
thus, the main motivation for the lateral buckling design scheme
discussed later in this paper.
Mode 1 Mode 2
Mode 3 Mode 4
Slip length, LS Slip length, LSBuckle length, Lo
Slip zone Slip zone
Buckle amplitude
Post buckle
Critical force
Buckle
Virtual anchors
Critical force
Eff. force
KP distance Buckle spacing, L1
Buckle spacing, L2 Buckle
spacing, L3
3
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Others. In addition, the following design parameters are also
used in this study: - Pipe outside diameter = 727 mm Pipe wall
thickness = 30 mm Concrete coating = 40 80 mm Design pressure = 157
barg Water depth = 75 m Seawater density = 1025 kg/m3Content weight
= 80 125 kg/m3Ambient temperature = 19.1 C Pipe material = API
5L-X65 [Ref. 5] THERMAL BUCKLING Driving Forces The driving force
for buckle initiation (either laterally or vertically) is derived
from the compressive forces in the pipeline as a result of
frictional resistance to thermal and end cap pressure. It is
necessary first, to determine the potential compressive force that
can develop in the pipeline based on the design temperature and
pressure. For a totally restrained pipeline, the effective axial
force is given by [Ref. 1]: - S = H Dpi Ai(1-2n) As EaDT (1) where:
S effective axial force (-ve compression, +ve tensile), H residual
lay tension, DT temperature difference relative to as laid, Dpi
internal pressure difference relative to as laid, Ai internal cross
sectional area of pipe, As cross sectional area of pipe. n Poissons
ratio of pipe The frictional resistance provided by the seabed soil
can be calculated based on the operational submerged weight along
the pipeline length. This is given by: - Ffric = m.Ws (2) where:
Ffric frictional resistance force m coefficient of friction between
pipe and seabed Ws pipe submerged weight The totally restrained
effective force of the pipeline and the available frictional
resistance, based on the design temperature profile and pressure,
is shown in Fig 8. From the plot, it can be seen that this 28
pipeline has two virtual anchor points at KP 8 and KP 39,
exhibiting the characteristic of a long pipeline.
Plot of effective axial force
-12000
-10000
-8000
-6000
-4000
-2000
00 10 20 30 40
KP distance (km)
Forc
e (k
N)
virtual anchors
Fig 8 - Effective axial force of pipeline
Having established the effective force characteristics of the
pipeline, it is then necessary to determine the susceptibility of
it to lateral buckling or snaking. Appropriate buckling mitigation
schemes can then be designed and implemented in areas which are
susceptible to buckling. The resulting global pipeline behaviour
can then be used to define the expected stress/strain level as a
framework for establishing the governing failure modes of the limit
state design. Susceptibility to lateral buckling The condition for
global snaking/lateral buckle to occur depends on and is sensitive
to various factors such as: - - pipe weight; - pipe cross sectional
property; - pipe-soil interaction/frictional resistance; - initial
imperfection introduced during pipe-lay installation; -
imperfection caused by undulating/uneven seabed. Lateral buckling
will happen naturally at intervals along the pipe where the
compressive force is sufficient for the pipe to buckle at some
natural or inherent imperfection. Such imperfections from the
idealised straight pipeline results from the lay process due to
vessel motion, wave or current loading, or seabed variations. This
type of uncontrolled lateral buckling will relieve the axial force
locally within a few hundred metres of the buckle, as pipe feeds
into the buckle from each side. To ascertain the critical
compressive force that will initiate global buckling/snaking, the
above factors have to be taken into consideration in view of
providing a realistic assessment of the susceptibility of the
pipeline to buckling. This is done using FE analysis by considering
an initial lateral lay imperfection of R2000m along the route
profile. This magnitude of imperfection was deemed realistic in
view of the lay corridor tolerance and pipe size. This lay
imperfection is assumed to occur over the lay corridor tolerance of
20m as shown: -
survey corridor
Fig 9 Modelling a naturally occurring imperfection
4
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Based on this, the critical buckling force is then obtained in
the
FE analysis using a short model of 1km, by assuming symmetry.
This is done by gradually increasing the temperature of the
pipeline until buckling or instability occurs. No vertical
imperfection was considered as the seabed was generally flat along
the proposed pipeline route. The critical buckling force obtained
here is then used to determine the area that is susceptible to
global buckling. The post buckle shape and the corresponding
critical buckling for is summarised in Fig 10 and Table 1
respectively.
Buckle shape of 727x30mm pipe (40mm HDCC) with R2000m
100
105
110
115
120
125
130
135
700 750 800 850 900 950 1000 1050Chainage (m)
Late
ral d
ista
nce
(m)
As-laid OOS
Buckle
R2000
Fig 10 Post-buckle of initial imperfection
Pipe size Concrete thickness (mm) Critical force (kN)
40 6062
25 4693 727 x30
55 7686
Table 1 Critical buckling force with imperfection
The critical buckling force from Table 1 includes safety factors
to account for soil friction variations and is used as a limiting
criterion to define the areas that are susceptible to buckling. The
critical areas are those with the effective axial force greater
than the critical buckling force defined by Table 1 and are shown
in Fig 11, which is a plot of Fig 8 with the critical buckling
force superimposed on it.
As can be seen from Fig 11, the buckle prone region extends from
approximately KP 5 to KP 21, resulting in a 16 km length, which has
the potential to buckle. The proposed mitigation schemes focuses on
reducing the critical buckling force in this region to initiate
buckling. This is discussed in the following sections.
Effective force along pipeline
0
2000
4000
6000
8000
10000
12000
0 10 20 30 40
KP (km)
Forc
e (k
N)
Effective force
Critical buckling force
Pcric with initial OOS of R2000
Buckle prone region
Fig 11 Regions susceptible to lateral buckling
LATERAL BUCKLING MITIGATION SCHEMES
Design methodology and objectives
As mentioned earlier, a pipeline can be left on the seabed and
allowed to buckle laterally, naturally. However, if the compressive
force in the pipeline system is sufficiently high (which in this
case, is shown in Fig 11), uncontrolled lateral buckling can lead
to one of the following limit states: - 1. Excessive plastic
deformation of the pipe, possibly leading to
localised buckling collapse; 2. Cyclic fatigue failure in
operation due to continuous heat-up
and cool-down cycles.
Early analysis, similar to those shown in Fig 10, where the
pipeline is allowed to buckle naturally on the seabed, showed that
the subsequent thermal feed-in would result in high plastic
deformation in excess of that allowed in the stipulated code of
practice [Ref. 1]. This is due to the large potential feed-in
lengths from each side of the buckle, if it were to develop within
the buckle prone region shown in Fig 11. As there is likelihood
that approximately 15-16 km of pipeline is susceptible to buckling,
the resulting feed-in could be of unacceptable magnitude. The
resulting buckle will exceed allowable strain limits from such high
feed-in lengths, unless lateral buckling can be controlled.
Buckling can be controlled by sharing the load between buckles,
formed at regular intervals along the route of the pipeline. The
location of these controlled buckles must be selected such that the
resulting thermal feed-ins into each buckle does not lead to
maximum strains level in excess of the allowable limit and the
cyclic strains experienced during heat-up and cool-down gives an
acceptable fatigue life.
Lateral buckles can be triggered at the desired locations by
using buckle initiators. There are several options available in
which these buckle initiators can be designed, constructed and
installed, depending on the seabed soil conditions, practicality
and its effectiveness. These different methods of buckle initiation
are snake-lay, mid-line spools, rock dumping (buckle prevention),
vertical triggers and vertical triggers with lateral pull. They
will be discussed in the following sections.
The following targets/objectives were set in order to qualify
these methods as effective and reliable: -
5
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1. The initiators should ensure that buckles are formed at a low
effective axial force, preferably much lower than the critical
buckling force of the coated pipeline (orange line in Fig 11). This
would ensure early initiation of these planned buckles at a stage
where the overall effective force along the pipeline is still
relatively low. Furthermore, this would also ensure that the
effective force between the buckle sites is sufficiently low so as
not to initiate an unwanted buckle.
2. Due to the soft seabed, high pipe embedment is anticipated
for
the pipe size and submerged weight in this study, generating
high lateral resistance. It is therefore, desirable for the
mitigation scheme to provide a means to reduce the lateral
resistance and allow substantial feed-in without over-straining the
buckle. This can be done either by reducing the pipe weight at the
buckle sites or elevate the pipe from the seafloor, thereby
eliminating contact and any lateral resistance.
3. Practicality, installability and cost effectiveness. The
final
selected mitigation scheme, in addition to satisfying the first
two requirements, must be cheap, relatively easy to manufacture and
can be installed with relative ease using conventional
lay-barge.
4. In addition to the buckle initiators, means of increasing
the
axial resistance need also to be explored, in an effort to
reduce the thermal feed-in into the planned buckle sites. This can
be achieved by re-distributing the concrete weight coating along
the pipeline using non-uniform concrete thicknesses where
appropriate.
Snake-lay
The concept of snake lay is to introduce horizontal
imperfections to the pipeline in the form of curves of given radii
of curvature at predetermined locations. The curves are created by
deviating the lay barge from its nominal route corridor to form a
zigzag or snaky pattern. The crown of the snake then behaves as a
large curvature expansion spool while the pitch (i.e. distance
between two successive crowns) dictates the amount of pipe feed-in
at the crown.
Fig 12 Snake lay configuration
The challenge in snake-lay design is to define a critical
buckle
spacing that will prevent the maximum strain and cyclic strain
range from exceeding acceptable levels, and also ensure that buckle
initiate reliably at each planned initiation site. More
importantly, the major uncertainty for snake lay on soft cohesive
soil is the extent of pipeline embedment and the resultant breakout
force. High embedment results in large pipe breakout force to
initiate early buckle (low temperature). Should buckle initiations
be delayed there is an increased likelihood of
excessive localised growth of only a few of the buckle sites,
which will in-turn compromise the integrity of the pipeline.
Fig 13 Example of buckled pipe section by snake lay Two
dimensional FE analysis has been performed to model the
behaviours of the snake laid pipeline. A series of curvature
radii and the corresponding breakout force has been considered.
Seabed friction sensitivity based on upper bound and the best
friction estimate is also included. The objective is to establish
minimum radius of curvature for a reasonable breakout force. The
results are presented in Figure 14 for breakout force against
radius of curvature and Figure 15 for the corresponding breakout
temperature.
727 OD pipe 30mm Wall thickness 25 concrete
0
1000
2000
3000
4000
5000
6000
7000
8000
500 600 700 800 900 1000 1100 1200 1300 1400 1500 1600 1700 1800
1900 2000
Radius of curvature (m)
Initi
atio
n fo
rce
(kN
)
Best estimate 2400kg/m3 concrete
Upper bound 2400kg/m3 conc
Best estimate 3040kg/m3 concrete
Upper bound 3040kg/m3 conc
Fig 14 Buckle initiation force versus lay radius
Snake crown
Pipeline
Counteracts
Wavelength
6
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Fig 15 Buckle initiation temperature versus lay radius
Fig 16 Conceptual snake lay scheme
From the initial assessment, a conceptual snake-lay scheme for
this pipeline is depicted in Figure 16. The following points can be
made based on these plots: -
The requirement for a confident early buckle initiation
requires the pipeline be laid with a significant bend radius.
Studies presented in Figure 14 show a snake-lay radius of 1000m
will be required to achieve confident buckle initiations between
2200kN and 3800kN effective force levels.
To achieve a tight 1000m-pipelay radius at each apex of the
snake-lay on a soft cohesive seabed will require the use of
some form of pre-installed counteract (clump weights or
similar).
727 OD pipe 30mm Wall thickness 25 concrete
10
15
20
25
30
35
40
45
50
55
60
500 600 700 800 900 1000 1100 1200 1300 1400 1500 1600 1700 1800
1900 2000
Radius of curvature (m)
Initi
atio
n te
mpe
ratu
re (d
eg C
)
Best estimate 2240kg/m3 concrete
Upper bound 2240kg/m3 conc
Best estimate 3040kg/m3 concrete
Upper bound 3040kg/m3 conc
From lateral buckling friction studies undertaken in the high
95C to 55C temperature regions from KP 0 to KP 12 buckle site
spacing will need to be kept at a minimum of 2km to avoid buckle
localisations.
Although the snake-lay method is attractive in view of its
relatively simple execution and no requirement for underwater
activity or specialised equipment, the small radius of 1000m
required for effective buckle initiation is less than the practical
limit of typically 2000-3000m for this pipe size. The high breakout
force due to potentially high embedment could also results in high
post-buckle strain levels in the buckle crown. Therefore, due to
the uncertainty in obtaining the required radius of 1000m and the
high breakout force, an alternative mitigation scheme is
investigated.
Mid-line expansion spools
This option investigates the use of mid-line spool to absorb
the
pipe expansion. The objectives are to identify the preliminary
size and number of spools required for the present pipeline. For
simplicity and ease of installation, a spool size of 30m has been
assumed as shown in Figure 17 below.
~ 30 m
E8DR-A to E11P-B Temperature and snake-lay profile
Fig 17 Typical mid-line expansion spool The spool is modelled in
U configuration with thermal expansion imposed to both ends of the
spool. The equivalent stress is then compared to the maximum
allowable of 72% SMYS and the corresponding end expansion limits
noted. The latter is then used to determine the spacing of the
spool to prevent over-stressing. The maximum allowable thermal
feed-in, based on mean seabed friction, for a typical expansion
spool is 2.2m, which is a considerable amount that can assist in
releasing a significant level of compression.
The required number of expansion spools along the pipeline route
is obtained using standard expansion calculation based on the
temperature profile of Fig 7 and a mean soil friction. The
expansion spool acts as compression relief points, hence changing
the effective axial force along the pipeline to the one as shown in
Fig 18. The resulting effective axial force with the presence of
the thermal expansion spool, is shown by the red line (zig-zag)
line as compression is released through the expansion spool. The
expansion spools are located at KP 6, 12, 18 and 24 at a spacing of
6km apart, with a feed-in from the 3km pipe section from each side
of the spool.
0
10
20
30
40
50
60
70
80
90
100
0 5 10 15 20 25 30 35 40 45
KP (km)
Tem
pera
ture
(deg
C)
-500
-400
-300
-200
-100
0
100
200
300
400
500
Snak
e-la
y of
fset
(m)
2km buckle site spacing with 100m offset from centreline
19 deg C ambient sea temperature
Temperature and snake-lay profile
~ 30m Thermal feed Thermal feed
7
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Effective force along pipeline
0
1
2
3
4
5
6
7
8
9
10
0 10 20 30 40 5KP (km)
0
Pcric with initial OOS of R2000
Fig 18 Modified effective force using mid-line spools The main
disadvantage of mid-line spool is that it introduces potential weak
points into the pipeline system because of the flange end
connection (2 flanges per spool). Hyberbaric welding will eliminate
this concern but has significant cost penalty due to mobilisation
of the specialised equipment. The installation time for the spool
is also relatively high, typically, is about 1 day per spool at the
present water depth. Trench and burial This option considers
burying or rock dumps the buckle prone region as a means to avoid
pipeline buckling, vertically and laterally. Burying by trenching
normally relies on the natural backfill process to cover the
pipeline. For lateral bucking control, a consistent consolidated
backfill is necessary to prevent upheaval buckling and natural
backfill carries considerate uncertainty on this consolidation
processes. Some forms of engineered backfill are normally
preferred. To calculate the required minimum backfill to prevent
upheaval buckling, analysis was carried out using PIPECALC, a JPK
developed software to determine the required overburden in order to
prevent pipe break through the cover. Typical rock parameters used
in the analysis is shown in Table 2 below:
Submerged weight 9 kN/m3
Internal friction angle 40 degrees Shear mobilisation
coefficient 0.6
Table 2 Typical rock properties
The minimum required rock cover over the buckle prone region is
shown in Table 3 below. This includes a safety factor of 1.1
calculated using statistical risk analysis approach.
KP start KP end Rock cover to top of pipe (m) 4 5.5 0.7
5.5 12 0.75 12 18 0.65
Table 3 Typical rock properties
Based on an assumed rockberm width of 3m at the top and side
slope of 1:3, the required gravel is about 35,000 MT per km cover,
assuming wastage of 25% on soft seabed. It may not be necessary to
dump the full length between KP 4 and KP 18,
instead, intermittent dumps at every few km could be feasible.
For this, it is necessary to ensure that if buckle forms between
the intermediate restraints, the feed-in is within the allowable. A
simple hydraulic calculation indicates that gravel size of 50-100mm
is acceptable for stability. Gravel size generally has little
bearing on cost or installation equipment.
The main disadvantage of trenching is uncertainty in the natural
backfill consolidation. Use of engineered backfill or direct gravel
dump, however, is affected by the availability of the specialised
gravel dump vessel and their high mobilisation cost. The
availability of suitably graded rock and gravel in large quantity
is also a concern. The soft seabed also renders gravel dump
inefficient, i.e. large wastage.
Vertical triggers/sleepers This option considers the use of
initial vertical out-of-straightness (OOS) to initiate a lateral
buckle. A sleeper, pre-laid across the route of the pipeline, would
raise and support the pipeline off the seabed. This creates an
out-of-straightness feature, which will initiate buckling. In
addition, pipe at the buckle crown is elevated above the seabed
with the benefit of reduction in lateral frictional resistance and
hence, reduces the uncertainties about lateral pipe-soil
interaction. Sleepers have the advantage of lowering the critical
buckling force, hence, creating a more benign buckle with lower
strain levels in the buckle apex. This allows higher thermal
feed-in capacity into the buckle sites, therefore increasing the
buckle spacing and as a result, reducing the number of required
buckle initiator. Its simple construction, ease of installation and
low fabrication cost makes the trigger option the most viable among
the above-mentioned methods. Vertical sleepers have been
successfully implemented in the King flowlines project [Ref. 6].
Initial evaluation of the sleeper solution showed that the strain
level were acceptable based on a criteria of maximum total thermal
feed-in of 2m. Hence, vertical sleeper was selected as the most
practical option for detailed design and further analysis.
trigger trigger
Fig 19 Vertical sleeper layout
pipeline
Fig 20 3D visualisation of buckle on vertical trigger
8
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SLEEPER OPTION: - DESIGN CRITERIA AND CHALLENGES This section
discusses in detail the various limit state design checks that was
carried out based on the vertical sleeper option. Also highlighted
in this section are some of the challenges encountered during the
design stage.
9
Design criteria The following design checks were made and will
be discussed in the following sections. 1. Local buckling limit
state. This criterion is based on local
wrinkling of the pipe wall as a result of bending/compression.
Initially, formulations based on the moment criterion stipulated in
[Ref. 1] were used. In later stage of the work, it was found that a
strain based criterion (also of [Ref. 1]) was necessary for the
sleeper option to be viable.
2. Fatigue limit state. Fatigue life check was carried out for
low-cycle fatigue due to heat-up/shut-down cycle and high cycle
fatigue due to vortex induced vibration (VIV). The former is based
on [Ref. 3], while the latter on [Ref. 4].
3. Axial creep. The susceptibility of local axial creep/rachet
into the buckle sites were also investigated using the transient
temperature profile.
4. On-bottom stability. The stability of the buckle section on
the trigger against wave and current loads was established using FE
based on limitations outlined in [Ref. 1].
5. Trawl gear interaction. The impact of a static trawl load on
the buckle section was also investigated to assess the magnitude of
additional strain/stress on the buckle crown.
Design challenges Minimising the critical buckling force As
mentioned earlier, one of the main objectives of the mitigation
schemes is to minimise the critical buckling force in order to
initiate an early buckle. The initial target was to reduce the
critical buckling force to at least half the peak effective axial
force shown in Fig 11, which is approximately 9400 kN. This was
dictated by the trigger height. The aim was to use triggers of
sufficient height to ensure buckles initiate preferentially at the
triggers and not at vertical or lateral imperfections along the
route. However, it must be ensured that the free spans on either
side of the triggers are below allowable limits for pipeline
design. Once elevated from the sea floor forming an initial
vertical OOS, lateral buckle is initiated as the compressive force
in the pipe is sufficient to lift-off at the trigger and slide
sideways to form a lateral buckle. This buckling mechanism enables
the critical buckling force to be predicted reasonably well using
analytical upheaval buckling methodology. FE analysis was then used
to confirm the prediction and to investigate post-buckling
behaviour such as feed-in capacity, buckle amplitude and
strain/stress level in the buckle crown. Initial calculations
suggest that a 1m vertical imperfection is required to
significantly reduce the critical buckling force. This, coupled
with a reduced pipe weight section along the buckle will further
reduce the initiation force. It was, therefore, decided to remove
the concrete weight coating for approximately 250m length of the
pipeline, which rests on the vertical sleepers. This has the effect
of reducing the critical buckling force to 3900 kN. A schematic
representation of the light weight pipe section with the vertical
trigger is shown in Fig 21.
The development of the effective axial force within a buckle
section is modelled using 3D finite element using a combination of
beam and contact elements. The changes in the effective axial force
are depicted in Fig 22.
heavy section Light section
Fig 21 Light weight buckle section on trigger
Initially, the pipeline is in tension as a result of the
residual force from installation This is shown by the top red line
corresponding to approximately 450kN. As the pipeline is subjected
to pressure and temperature, the effective force along it changes
to compressive until the critical buckling force is reached. The
critical buckling force in this case is reached at around at
temperature difference of 15C at approximately 3900 kN. As soon as
this level of compression force is reached, the pipeline lifts off
the trigger and buckle laterally, releasing a significant amount of
the compressive force. The post-buckle effective force is shown by
the green line at 16C, reaching a level slightly below 1000 kN.
Subsequent increase in temperature causes further reduction in the
effective compressive force as the compression is being continually
released through the buckle, resulting in growing buckle amplitude,
shown in Fig 23.
Effective force along pipe length
-4500
-4000
-3500
-3000
-2500
-2000
-1500
-1000
-500
0
500
1000
-600 -500 -400 -300 -200 -100 0 100 200 300 400 500 600
Chainage (m)
15 C
16 C
200 C
Fig 22 Effective force development
The total feed-in capacity of the buckle is determined by
computing the resultant moment at the buckle crown and comparing it
against allowable moment capacity, as outlined in the local
buckling limit state based on the moment criteria of [Ref. 1]. The
maximum total allowable feed-in into a single buckle was found to
be 2m. Variation of resultant moment with feed-in is summarised in
Fig 24, along with the limit of allowable moment.
Light weight section
Model length in FE
1m high trigger
-
Lateral displacement along pipe length
-20-18
-16
-14
-12
-10
-8
-6-4
-2
0
2
4
-600 -500 -400 -300 -200 -100 0 100 200 300 400 500 600
Chainage (m)
Fig 23 Buckle amplitude at increasing temperature
10
Resultant bending moment vs feed-in
0
500
1000
1500
2000
2500
3000
3500
4000
4500
5000
0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2 2.2 2.4Total feed-in
(m)
Res
ulta
nt m
omen
t (kN
m)
0
50
100
150
200
250
300
350
400
Stre
ss (M
Pa)
Resultant moment Allowable moment Max Eqv Stress
Allowable 4619 kNm
Fig 24 Variation in moment against feed-in
The above results using moment based criteria were based on mean
pipe-soil friction, both in the axial and lateral direction.
Further investigation of the pipe-soil interaction behaviour
revealed a more complex friction model. It resulted in the need to
utilise the strain based criteria due to higher lateral friction.
This was one of the main challenges of this work and proved to be a
valuable lesson learned which will be discussed next. The need for
strain based criteria pipe-soil interaction The main governing
factor that dictates the curvature of the buckle crown, hence the
strain/stress levels, is the lateral resistance acting against the
direction of the buckle. In the use of vertical triggers, some
portion of the resistance is eliminated. However, the wavelength of
the buckle section will most likely be longer than the free span,
in Fig 23, the wavelength is approximately 400m. Therefore,
inevitably, there is a significant portion of the buckle which
slides laterally on the seabed. Frictional behaviour of pipe on
soft seabed is very complex and without full scale laboratory
testing, a simplified model as shown in Fig 25 is used. The
frictional behaviour on firm soil can be approximated using the
Coulomb model (Curve 1). On soft soil, however, due to high
embedment, the pipe will need to breakout from the soil before
sliding. Hence, there will be a peak resistance, Fp, beyond which
the frictional resistance resembles the Coulomb model again. In
this work, an upper, mean and lower bound frictional
curve was established to cater for the uncertainty in actual
soil behaviour. This is summarised in Table 4 and 5 for lateral and
axial friction respectively.
Force
Curve 1 FP
= friction coefficient Ws = submerged weight
Ws
Slope, KT = tangential stiffness
Displacement
Fig 25 Seabed friction model
Equivalent friction factor Lower bound Mean Upper bound
Break-out 0.5 1.3 2.5 Sliding 0.3 0.85 1.25
Table 4 Lateral pipe-soil friction
Equivalent friction factor
Lower bound Mean Upper bound Break-out 0.8 1.317 1.835 Sliding
0.2 0.329 0.459
Table 5 Axial pipe-soil friction
The various friction coefficient values are used accordingly in
different analysis and sensitivity cases. These are discussed in
the following sections. Improving the effectiveness of the vertical
trigger option In view of the relatively large spread of lateral
friction coefficients (mean, lower and upper bound), it was deemed
necessary to add further certainty that buckle will form at the
designated location. The critical buckling force of 3900 kN
establish using a 1m high trigger was still rather large. There
were concerns that due to the high critical force, the buckle would
not form. Worst still, if it were to form elsewhere, the high
lateral resistance would cause the strain levels to exceed local
buckling limit state. To improve the reliability of the scheme, it
was decided that the vertical trigger option should be combined
with a horizontal imperfection, depicted schematically in Fig 26,
which is a plan view of Fig 21. trigger
pipeline
Fig 26 Introduct vertical stopper
ion of horizontal imperfection during pipe installation
-
The horizontal imperfection is created by pulling the lay barge
sideways during installation when the pipe touches down at the
trigger. A vertical stopper is incorporated into the trigger
structure to prevent the pipe from falling off. The side pull
creates a tight curvature around the stopper, hence creating an OOS
and thus reduces the critical buckling force further. To obtain the
optimum lateral pull angle to achieve the desired critical buckling
force whilst maintaining reasonable stress level, a series of FE
analysis were carried out. In view of the certainty that the
critical buckling force can be reduced significantly, it was
decided that the trigger height be reduced to 0.5m. This improves
the free spans on either side of the trigger, allowing for higher
fatigue life capacity. The following sections present results of
the different limit state checks based on the concept of lateral
pull on trigger. Local buckling limit state strain criteria Based
on the improved scheme of lateral pull over the vertical sleepers,
similar FE analyses were carried out to determine the new critical
buckling force and the feed-in capacity of the post-buckle crown
using the new pipe-soil friction coefficient. In addition, the
sensitivity of the critical buckling force to the lateral pull
angle was also investigated. These checks were carried out in the
strain based criteria. It was found that the critical buckling
force is now reduced significantly, from 3900 kN previously, to
between 1100 to 2000 kN for a 5 degree and 10 degree lateral pull
respectively. Furthermore, it was also noted that the critical
buckling force is relatively insensitive to changes in soil
friction. The results based on a 5 degrees and 10 degrees lateral
pull on various soil friction is summarised in Table 6.
Load case
Angle (degs)
Soil friction
Trigger friction
Pcrit (kN)
1 5 1971 1a 10 0.5 1107 2 5 1977 2a 10 1.3 1325 3 5 2052 3a 10
2.5
0.6
1429
Table 6 Critical buckling force for various friction and pull
angles
To determine the feed-in capacity of the post-buckle crown, the
following considerations have to be taken into account in view of
using the strain based criteria: - - carbon steel strain hardening
characteristics with temperature
de-rating, - material strength mismatch at the pipe joints
(girth weld), - end of life wall thickness based on design
corrosion allowance. Currently available data from pipe
manufacturers have occasionally shown a plateau in the
stress-strain curve of carbon steel. Although such plateaus are
more common in seamless pipes, they have also been observed in
other seam-welded pipes as well. Therefore, to overcome this
uncertainty, it was decided to assume the worst case of no strain
hardening in the carbon steel. Hence, an elastic-perfectly plastic
material stress-strain curve was used in the FE model. Furthermore,
to account for the variation in SMYS with temperature, the SMYS was
de-rated as per outlined in [Ref. 1] resulting in a de-rated SMYS
of 404 MPa.
Strain localisation can occur at any point along the pipeline
where there is a mismatch in material properties or pipe geometry
(wall thickness). To account for this possibility, the concept of a
weak link in the chain of pipeline is used. This weak link is
conservatively placed at the buckle crown of the FE model. In this
weak link, the de-rated SMYS is maintained, however, outside, the
SMYS is modelled with a higher value. To account for corrosion
allowance, an equivalent reduced wall thickness is determined based
on the plastic section modulus of the corroded cross section. This
reduced wall thickness is also incorporated within the weak link
section of the FE model. Based on these considerations, the
effective force development, buckle amplitude and feed-in capacity
is shown in the following plots, based on maximum lateral
frictional resistance.
Effect ive fo rce fo r angle=5degs,727x30mm
-3500
-3000
-2500
-2000
-1500
-1000
-500
0
500
1000
0 100 200 300 400 500 600 700 800 900 1000 1100
C hainag e ( m)
pressurised
14 C
220 C
Fig 27 Effective force based on 5 degree lateral pull
Lateral displcement fo r angle=5degs,727x30mm
-25
-20
-15
-10
-5
0
5
10
15
0 100 200 300 400 500 600 700 800 900 1000 1100
C hainag e ( m)
220 C
Fig 28 Buckle amplitude based on 5 degree lateral pull
Axial feed-in for angle=5degs,727x30mm
-1.5
-1
-0.5
0
0.5
1
1.5
0 100 200 300 400 500 600 700 800 900 1000 1100
C hainage (m)
220 C
Fig 29 Thermal feed-in at various temperatures
11
-
The effectiveness of this scheme in reducing the critical
buckling force is evident from Fig 27, in which the compressive
force is released almost immediately upon application of
temperature. Further increase in temperature will continue to
release the compression (effective force gradually becomes less
compressive) at the buckle section. The high temperature of 220C
was necessary in this short FE model to generate the desired level
of thermal feed-in. The feed-in capacity of the post-buckle crown
was found to be slightly more than 2m (Fig 29) with a corresponding
total mechanical strain level of 0.56%. The allowable design
compressive strain based on [Ref. 1] is 1.2%. This completes one
stage of the design checks for which this scheme is viable. Fatigue
limit state The potential for fatigue damage of the buckle section
arises from two sources: -
- low cycle fatigue due to continuous start-up and shutdown
operation throughout its design life; - vortex induced vibration
(VIV), both in-line and cross flow,
on spans either side of the trigger. For in-line VIV only the
steady current is considered whereas cross-flow accounts for both
steady and wave induced current velocities per [Ref. 4].
The 1km model used in the pre- and post-buckling checks is
used to evaluate the cyclic strain. The model is first heated up
to the point where maximum feed-in occurs (220C) and unloaded to
simulate shutdown. It is then re-heated up to 220C again (so that
maximum possible feed-in occurs) and unloaded, and this is repeated
for 12 cycles, to simulate an annual event (assuming one full
start-up/shutdown a month). The maximum lateral friction and
lateral pull angle of 5 degrees was used.
The cyclic tensile and compressive strain is shown in Fig 30 and
31 respectively. The number of cycles to failure is calculated
based on [Ref. 3]. Based on the maximum strain range of 0.2285%
from the compressive strain (Fig 31), the buckle has a fatigue life
of 2842 cycles with a damage ratio of 0.13 for a design life of 30
years. The allowable damage ratio based on [Ref. 1] is 0.2. More
importantly, yielding/plasticity occurs only during the first
heat-up cycle. No further yielding/cyclic plasticity is observed in
the subsequent heat-up cycles.
Cylic strain variation at 90 degs
-5.00E-04
0.00E+00
5.00E-04
1.00E-03
1.50E-03
2.00E-03
2.50E-03
0 50 100 150 200 250 300 350
Load steps
Axial strain
Elastic axial
Plastic axial
max = 1.96e-3
min=-1.24e-4
plastic=1.81e-3
range=2.084e-3
Fig 30 Mechanical tensile strain variation
For In-line VIV, the fatigue life of free spans on either side
of the trigger is calculated based on DNV-RP-F105 [Ref. 4]. The
free span lengths and the summary of results are tabulated in Table
7 and 8.
Cylic strain variation at 270 degs
-2.50E-03
-2.00E-03
-1.50E-03
-1.00E-03
-5.00E-04
0.00E+00
5.00E-04
1.00E-03
0 50 100 150 200 250 300 350
Load steps
Axial strain
Elastic axial
Plastic axial
max = 5.65e-4
min=-1.72e-3
plastic=-2.04e-3
range=2.285e-3
Fig 31 Mechanical compressive strain variation
Loading condition
Free span length on either side of 0.5m trigger (m)
Empty 72 Content 68
Table 7 Free span lengths
Loading condition In-line VIV fatigue life, years
Empty 7.1 Content(1) 40.5
Table 8 In-line VIV fatigue life
For the case of cross-flow VIV, combined wave and current
velocities were used based on the DNV-RP-F105 [Ref. 4] guideline
and the metcoean data. The calculation is performed for all the
wave height and period combination given in the annual joint
frequency distribution table (ie. scatter diagram). The monthly
maximum steady current has been assumed and this is combined with
the wave velocity for the cross-flow VIV/fatigue calculation.
The result is summarised below: -
Total No. of Wave Per year 8,125,334 Cross-Flow VIV Fatigue
Damage Per Year 1.01338E-6 Pipeline Design Life 30 years
Total Fatigue Damage for 30 year 3.04015E-5 Fatigue Usage Factor
(Safety Class Normal) 0.2 Total Cross-Flow Fatigue Life 197,359
Years From the above, it is concluded that cross-flow VIV fatigue
on
the free span is acceptable. Fatigue due to direct wave load is
assessed by applying a
uniformly distributed load (UDL) on the post-buckle pipeline
model in opposing directions and observing the variation in the
maximum and minimum stress. The metocean data for all year
significant wave height, Hs and peak wave period, Tp was used
12
-
to determine the significant wave-induced current. The drag
forces (UDL) for all the possible Hs, Tp combinations were then
determined using Morrisons equation.
The result is summarised below: - Total no. of wave per year
8,125,334 Direct wave fatigue damage per year 4.165E-9 Pipeline
design life 30 years Total fatigue damage for 30 year 1.249E-7
Fatigue usage factor (Safety Class Normal) 0.2 Total Fatigue Life
> 1 million years From the above, it is concluded that direct
wave fatigue is
acceptable. Confirmatory analysis & axial creep Based on the
allowable feed-in (Fig 29), the number of buckle triggers and the
critical buckle site location/spacing was then determined using the
temperature profile (Fig 7) and analytical expansion calculations.
It resulted in 8 triggers spaced unevenly for the first 20 km, in
which the pipeline is prone to buckling (Fig 11). A full FE
confirmatory analysis was then carried out to verify the total
feed-in into each of these planned sites. The location of theses
planned sites and the resulting feed-in into each site at the
design temperature profile is shown in Fig 32 and 33 respectively.
Fig 33 depicts the axial displacements along the pipeline at the
design temperature profile. Positive values indicate movements to
the right and conversely, negative values signify movements to the
left. Points with zero axial displacements are the virtual anchor
points. In between each buckle sites, there exists a virtual anchor
point, from which the pipeline displaces axially in opposing
direction. Near the buckle apex, the total axial displacement
peaks, contributing the maximum feed-in into the buckle. This is
shown in the plot by the various positive and negative peaks, which
occurs approximately 100m either side of the buckle apex. The
results show acceptable feed-in at all the planned buckle sites
with total of less than 2m. The post-buckle effective force at full
design temperature was also found to be favourable in terms of the
safety margin against unwanted buckle between the planned sites.
This is shown in Fig 34 for both the mean and upper bound axial
friction.
Buckle amplitude with 8 triggers
-300
-250
-200
-150
-100
-50
0
50
100
150
-2 0 2 4 6 8 10 12 14 16 18 20 22 24 26 28 30
KP (km)
As-laid Loaded with design temp
T1-KP2
T2-KP4.5
T3-KP7
T4-KP9.5
T5-KP13
T6-KP16.5
T7-KP20
T8-KP24
Fig 32 Proposed buckle sites
The axial creep behaviour of the pipeline was investigated using
the full FE model by subjecting it to repeated heating and cooling
cycles. Each cycle is model with a set of transient temperature
profile during heat up and cool down. Although this pipeline is
long in terms of its effective force characteristics and
should not pose any global axial creep problems, the potential
of local axial creep into the buckle sites (since the formation of
buckle effectively divides the pipeline into various short
sections) was investigated.
Axial feed-in with 8triggers
-1
-0.8
-0.6
-0.4
-0.2
0
0.2
0.4
0.6
0.8
1
-2 0 2 4 6 8 10 12 14 16 18 20 22 24 26 28 30
KP (km)
T1 T2 T3 T4 T5 T6 T7 T8
Fig 33 Feed-in at proposed buckle sites
Post buckle effective force at design temperature
0
1000
2000
3000
4000
5000
6000
7000
8000
9000
10000
11000
12000
0 2 4 6 8 10 12 14 16 18 20 22 24 26 28 30 32 34 36 38 40 42
44KP (km)
Pcric with init ial OOS of R2000
Safety margin
Fig 34 Post-buckle effective force
The incremental axial displacement of the buckle apex with each
subsequent heat-up/cool-down cycle was then extracted from the FE
model. The results are plotted in Fig 35.
Incremental axial displacements of buckle vs temperature
cycles
-0.8
-0.6
-0.4
-0.2
0
0.2
0.4
0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15
Cycle no.
Rel
ativ
e di
spla
cem
ent (
m)
hot end
T8
Fig 35 Incremental axial displacement into buckle sites
The plot shows that the peak axial displacement in to the hot
end spool and each buckle sites occur only at the first heat-up
13
-
cycle. There is no further net increase in axial displacement in
the next subsequent cycles, indicating no local axial creep is
expected. On-bottom stability The stability of the light buckle
section to wave and current load was investigated for the empty and
operational load case using the full FE model. The minimum content
density of 80kg/m3 and mean lateral sliding friction of 0.85 was
used in the analyses. The drag and lift forces (N/m) used are
tabulated in Table 9.
Item Pipe section Lift
force, FL
Drag force,
FD727 buckle section
on seabed 540 734 Operational 727 spanning buckle section 606
807
Empty Buckle section - 727 OD 241 322
Table 9 Hydrodynamic forces for on-bottom stability check In the
operational load case, the combination of 100-year wave and 10-year
current generated a maximum pipeline lateral displacement of 2.51m
at trigger 8. No vertical uplift was observed. In addition, the
uniform lateral hydrodynamic load tends to reduce the curvature of
the post-buckle at the apex (a greater radius of curvature or more
relaxed curvature). Subsequently, a reduction in the bending stress
at the apex was observed. In conclusion, the buckle section at the
trigger is stable against hydrodynamic load with acceptable level
of displacement and the bending stress at the apex is reduced as a
result of an increase in the radius of curvature due to the action
of wave and current. Pipeline/trawl board interaction A study of
the vessel size around the vicinity of the project indicates the
following possible trawl loads: - Horizontal trawl load, Fp = 35.22
kN Vertical trawl load, Fz = 17.61 kN (downwards) The trawl load
computed above is used in the post-buckle FE model as a
concentrated point force applied at the apex to determine the
increase in strain at the apex. The analysis was carried out using
the mean and maximum sliding lateral friction. In both cases, the
post-buckle shape (prior to application of trawl load) corresponds
to a total feed-in of 2m. The results are summarized in Table 10
and 11 below: -
m=0.85 (mean) Post-buckle With trawl load Increment
Total (mechanical) strain, % -0.2316 -0.2596 0.028
Lateral displacement at apex, m 12.97 13.037 0.067
Table 10 Increase in strain and lateral displacement at apex
due to trawl load for 0.85 friction
m=1.25 (max)
Post-buckle With trawl load Increment
Total (mechanical) strain, %
-0.3535 -0.3779 0.0244
Lateral displacement at apex, m
11.348 11.407 0.059
Table 11 Increase in strain and lateral displacement at apex
due to trawl load for 1.25 friction
The increase in displacement and total strain are small and the
pipeline buckle will survive a trawl load of the above-mentioned
magnitude. CONCLUSIONS This paper has presented various possible
mitigation schemes for use in HP/HT pipeline as means to manage the
thermal expansions. The choice of the optimum scheme depends
largely on the various factors such as cost, practicality, ease of
manufacture, seabed conditions, etc. One of the main obstacles
remains the uncertainty of the nature of pipe-soil interaction
behaviour. In most cases, the most onerous conditions need to be
employed, specific to certain analysis, in order to ensure a robust
design. The mitigation scheme using the combination of vertical
triggers and lateral pull provides a solution which is both
manageable and neat. The elevated pipeline eliminates lateral
resistance and pipe-soil friction uncertainties at the buckle
crown, which helps tremendously in lowering the strain levels in
the buckle. In addition, the lateral pull adds confidence to the
certainty of forming a buckle at the intended location by reducing
the critical buckling force significantly. Both this advantages,
together with its simple and cost-effective fabrication, make it a
first-choice option, especially in cases with soft seabed.
REFERENCES 1. DNV-OS-F101,Submarine Pipeline Systems, 2000. 2.
Hobbs, R.E. In-service buckling of Heated Pipelines, ASCE
Journal of Transportation Engineering, 110(2), 175-189, 1984. 3.
Murphey and Langner, Ultimate Pipe Strength Under
Bending, Collapse and Fatigue, Proc. Of OMAE, 1996. 4. Det
Norske Veritas, Free spanning pipelines, Recommended
practice, DNV-RP-F105. 5. API-5L Linepipe Specification,
American Petroleum Institute,
1992. 6. Harrison G.E.,Brunner M.S and Bruton D.A.S, King
Flowlines Thermal Expansion Design and Implementation, Proc. Of
OTC, 2003.
14
Mr Lim Kok KienJP Kenny Wood Group SDN BHDLim Kok Kien*, Dr. Lau
Siew Ming* and Dr. Emil MaschnerABSTRACT
