HPHT Pipeline Design (DOT 2008, Perth)

1

DESIGN OF HIGH PRESSURE/HIGH TEMPERATURE (HP/HT) PIPELINES AGAINST

LATERAL BUCKLING USING VERTICAL UPSET METHOD

A. Nikkhaah1, Dr. M. Baghernejad1, Shaifuzzaman B. Isa2, B. Baghernejad1, C. Y. Chun2

ABSTRACT

Design of HP/HT pipelines differs significantly from traditional pipeline design. In

HP/HT pipelines, phenomena such as lateral buckling, low cycle fatigue, and

ratcheting is accounted for in addition to addressing the traditional pipeline design

requirements.

In this paper lateral buckling of HP/HT pipelines is discussed by presenting a real case

study, the gas pipeline between PC04 and B11 platforms in Malaysian waters.

To accommodate thermal expansion of the pipeline and to overcome the available soil

uncertainties, a design strategy using strain-based design was adopted; incorporating

mitigation techniques such as a pipeline lay over vertical buckle triggers (sleepers).

The pipeline is designed to buckle laterally on sleepers. Locations of the sleepers are

selected with due consideration for total strain in the pipe wall, pipeline route, and

uncertainties in design input data.

INTRODUCTION

In recent years, demand for high pressure/ temperature pipelines is increasing continuously.

In parallel, some joint industry projects have been developed to address safe design of

HP/HT pipelines.

SLT-Engineering developed a simplified methodology based on requirements of DNV-OS-

F101 code [1], using available published information [2 to 12]. Figure 1 shows the analysis

methodology developed for lateral buckling analysis of the pipeline.

1 SLT-Engineering Sdn Bhd, Asia Pacific Region Office, Kuala Lumpur 2 Petronas Carigali Sdn Bhd, Kuala Lumpur

2

. Figure 1: Design Methodology for Lateral Buckling Analysis.

3

DESIGN REQUIREMENTS

The following design requirements are set for lateral buckling of the pipeline:

1. The buckling force (maximum pipeline axial force capacity) should be considered

equal to minimum of Hobbs mode infinity [2, 3] and out of straightness model.

2. Design feed-in-length (FIL) of each buckle shall be equal to maximum feed-in-

length into the buckle which will not cause pipeline failure under all limiting states.

3. The probability of maximum feed-in-length exceeding the design feed-in-length

shall be less than 10-4 [1].

LATERAL BUCKLING ANALYSIS METHODOLOGY

Pipe-Soil Interaction Model

The soil resistance against pipeline movement is divided into two main sections, breakout,

and residual sections. The lateral resistance model of soil is developed based on the model

and recommendations presented in OTC 17944 [10]. The initial embedment of pipeline is

calculated using Equation (1) [10].

2

45 ⎟⎟⎠

⎞⎜⎜⎝

⎛=

u

t

DSVS

Dz (1)

Where, z is the initial penetration of pipe, D the pipe outside diameter, St the soil sensitivity,

V vertical load on the pipeline inclusive of dynamic load during pipeline installation, and Su

the undrained shear strength at bottom of the pipe.

The lateral breakout and residual resistances of soil then may be calculated using Equations

(2) and (3), respectively [10].

Dz

DSv

DSH

uu

breakout

γ/32.0 += (2)

⎥⎥⎦

⎤

⎢⎢⎣

⎡⎟⎟⎠

⎞⎜⎜⎝

⎛−−−=

DS

vH Duresidual

γ1_

21exp165.01 (3)

Where, γ is the submerged soil unit weight, v the vertical load on pipe, and Su_1D the

undrained shear strength of soil at one pipe diameter below seabed.

The mobilization of breakout resistance is assumed within a pipe movement of less than

half of a diameter, while residual resistance occurs within 3 to 5 diameters [10].

4

For axial resistance of the pipeline the soil model presented in OTC 6846 [20] is used.

According to this model maximum and residual axial resistance of soil may be calculated

using Equations (4) and (5), below.

ucbreakout SAH 05.1= (4)

ucresidual SAH 34.0= (5)

The mobilization of axial breakout resistance is within a pipe movement of less than 0.05 of

a pipe diameter, while residual resistance occurs at 1.2 times the diameter of pipe [20].

The uncertainty of the soil model is treated based on recommendations of OTC 17944 [10].

Screening Criteria for Buckling

Lower bound axial capacity of the pipeline to withstand against lateral buckling is

calculated using Equation (6).

S = min (SH∞, SOOS) (6)

Where,

( )( )

25.0

2

3

...

.29.2

⎟⎟

⎠

⎞

⎜⎜

⎝

⎛=∞

sLB

L

H

AEfIE

IES (7)

( )( )RFFWS DLOOS −−= μ (8)

Where, E is the elastic modulus of the pipeline material, W the submerged weight per unit

length of the pipeline, μ the lateral friction coefficient, I the pipeline moment of inertia, As

the pipeline cross section area, FD the drag force per unit length on the pipeline, FL the lift

force per unit length on the pipeline, and fLLB the lower bound lateral soil resistance.

The pipeline may buckle if maximum axial effective force in the pipeline is higher than

axial capacity of the pipeline.

Calculation of Design Limiting Criteria and Feed-In-Length

To calculate maximum allowable strain in the pipeline the following key failure modes

(except pressure containment and external pressure collapse that were fulfilled during early

stage of pipeline design) are considered:

1. Local buckling

5

2. Fatigue

3. Weld fracture

Local Buckling

As most applied loadings on the pipeline are displacement-controlled, strain based design

criteria based on requirements of DNV-OS-F101 [1] is used for local buckling analysis of

the pipeline, considering the following assumptions [13].

1. Hoop component of stress and resultant strain are kept within the allowable limit

obtained from load controlled criteria

2. Environmental loads are applied to the buckle free span in the zones around the

buckle triggers. Effects of these loads on the resultant strain should be insignificant.

Fatigue

High pressure/ high temperature pipelines generally may be subjected to two main

categories of cyclic loadings, as follows:

1. Low cycle / high amplitude loading, mainly due to pipeline installation and startup-

shutdown of the line

2. High cycle / low amplitude loading, mainly due to environmental load and vortex

induced vibration

For low cycle fatigue analysis of the pipeline, the methodology presented in guidelines of

American Bureau of Shipping [15] may be used. According to this guideline, Equation (9)

may be used for assessment of the fatigue life of the welded structures when the plastic

strain range is significant.

002.0055.0 4.0 ≥Δ=Δ − εε forN (9)

002.0016.0 25.0 <Δ=Δ − εε forN

Where, Δε is the strain range in pipeline.

Investigations [13] show that at least a safety factor of 7.0 is included in Equation (9).

The high cycle fatigue analysis of the pipeline spans aside the buckle triggers is performed

in accordance with DNV-OS-F101 [1] and DNV-RP-C203 [14] under the assumption that

the weld line is in region with highest stress conditions.

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Fracture

Engineering critically assessment is performed based on BS 7910 [16] using

recommendations of DNV-OS-F101 [1]. Based on accuracy of available examination

methods, an undetectable circumferential crack with 25 mm length and 2.5 mm depth may

be considered for fracture analysis.

Design Feed-In-Length

Design feed-in-length of each section of the pipeline is calculated using virtual anchor

model by finite element method considering abovementioned limiting parameters.

Acceptability of Single Buckle in Non-Mitigated Pipeline

In some cases even in pipelines susceptible to lateral buckling, limit state parameters of the

buckled section may be acceptable. In order to assess the acceptability of a non-mitigated

pipeline a single buckle between two virtual anchor points on the pipeline shall be

considered, where the feed-in into the buckle is maximum. The buckle may be triggered by

initial lateral or vertical imperfection, or by trawling load.

Calculation of the Distance between Buckle Triggers

To calculate distance between buckle triggers the following approach may be followed:

1. Maximum allowable expansion of end sections of pipeline shall be calculated based

on expansion spool design capacity and lower bound axial soil resistance. The first

buckle trigger from each side of the pipeline shall be positioned to achieve

maximum allowable expansion at each end of the pipeline.

2. Buckle triggers are positioned in such a way to limit maximum feed-in into the

buckles below the design feed-in-length.

3. If the probability of buckle formation failure at a specific buckle trigger is more than

10-4, the consequences of formation of a buckle at vicinity of the buckle trigger shall

be evaluated.

Calculation of Buckle Formation Probability

In lateral buckling analysis of HP/HT pipelines the key uncertainty is buckle formation at

the expected sites. To calculate the buckle formation probability a simplified version of the

reliability model presented in Carr, et al [5] is used. According to this model, the probability

of buckling can be defined as:

7

]0y[Probabilit ≤= Zp f (10)

Where, Z is the limit state function describes the buckle formation, which is obtained by

recasting the buckling formation criteria. Equation (11) denotes the buckling limit state

function.

ExWRZ alat −−= ..μ (11)

Where, Rlat is lateral resistance against pipeline buckling, μa axial friction factor of soil, W

pipeline weight per unit length, x sleeper distance to the end of line or previous sleeper, and

E is axial force of the pipeline due to spool resistance or residual axial force in previous

buckle.

Acceptability of Mitigated Pipeline

After positioning the buckle triggers along the pipeline route, the mitigated pipeline is

analyzed to check whether the mitigation scheme works appropriately under pipeline heat

up and cool down transients. Under certain circumstances walking of the pipeline section

between two adjacent buckle trigger (towards the cold end of the pipeline) may increase

feed-in into the initiated buckle. This phenomenon cannot be captured by virtual anchor

spacing model; the whole pipeline shall be modeled for finite element analysis.

CASE STUDY

Design Parameters

Pipeline

The pipeline was constructed using 12” API 5L-X65 line pipe, as shown in Table 1. High

density concrete coating was used for achieving adequate stability for the pipeline, and its

effects on strength of the pipeline was ignored.

Design pressure, and temperature distribution along the pipeline were used for analysis.

Design pressure and maximum design temperature were 201.4 barg, and 120 oC,

respectively. Temperature profile along the pipeline, which was used for analysis, is

presented in Figure 2. This temperature profile was calculated based on 50% pipeline

seabed embedment.

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Table 1: Pipeline General Specifications.

KP Zone Length (km)

OD (mm)

WT (mm)

Corrosion Allowance (mm)

0.0 -0.5 Zone II (B11) 0.5 304.8 23.8 6 0.5 – 19.3 Zone I 18.0 304.8 19.5 6 19.3 – 21.8 Zone I 3.0 304.8 23.8 9 21.8 – 22.3 Zone II (PC04) 0.5 304.8 23.8 9

The design life of the pipelines was 6 years. Total number of startups and shutdowns during

lifetime of the pipeline was 24 cycles.

Physical and elastic mechanical properties of the pipeline steel were as given in Table 2.

Table 2: Pipeline Steel Properties.

Property Value

Density [kg/m3] 7850 Young’s Modulus [MPa] 207 x 103 Poisson’s Ratio 0.3 SMYS [MPa] 448 SMTS [MPa] 535 CTOD (Weld at Minimum Temperature) [mm] 0.2 [2]

For global buckling analysis and ratcheting analysis of the pipeline, isotropic strain

nonlinear hardening and simplified linear kinematics strain hardening behavior of X65 were

used, respectively. Figure 3 shows the isotropic and kinematics elastic-plastic models of

X65 at different temperatures.

Seabed profile used for analysis is presented in Figure 4.

0

20

40

60

80

100

120

0 5 10 15 20

KP

Tem

pera

ture

(C)

Figure 2: Temperature Distribution along the Pipeline Considering 50% Pipeline Embedment.

9

0

50

100

150

200

250

300

350

400

450

500

550

0.0% 1.0% 2.0% 3.0% 4.0% 5.0% 6.0% 7.0%

True Strain

True

Str

ess

(MPa

))

X65, 25 C X65, 50 C X65, 120 C

0

50

100

150

200

250

300

350

400

450

500

550

0.0% 1.0% 2.0% 3.0% 4.0% 5.0% 6.0% 7.0%

True Strain

True

Str

ess

(MPa

))

X65, 25 C X65, 50 C X65, 120 C

Figure 3: Nonlinear Isotropic Elastic-Plastic (Left) and Bilinear Kinematics Elastic-Plastic (Right)

Behaviors of API-X65.

-96

-94

-92

-90

-88

-86

-84

-82

-80

-78

-76 0 2000 4000 6000 8000 10000 12000 14000 16000 18000 20000 22000DISTANCE (m)

DEP

TH B

ELO

W M

SL

B11 PLATFORM

PC4 PLATFROM

Figure 4: Seabed Profile along the Pipeline Route.

Pipeline-Soil Interaction

Due to insufficient soil data the following procedure was followed to estimate upper bound,

median, and lower bound soil properties.

1. Investigations were performed on the available soil data in existing routes near the

pipeline proposed route. According to these investigations it was understood that the

undrained shear strength of the soil along the route changed linearly with respect to

the pipeline penetration, for penetration depths less than 1.0 m.

2. Three linear functions were fitted to the available soil data along the pipeline and

defined as upper bound, median, and lower bound undrained shear strength profile

of the soil, as shown in Figure 5.

3. Soil axial and lateral resistances were calculated based on the estimated soil

properties, as shown in Figures 6 and 7.

10

GC/01

GC/04

GC/05

GC/08

GC/09

GC/02

GC/03

GC/06GC/07

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

0 20 40 60 80 100 120 140Undrained Shear Strength (kPa)

Dep

th (m

)

Lower Bound Median Upper Bound

Figure 5: Undrained Shear Strength Profile of Soil.

0

0.2

0.4

0.6

0.8

1

1.2

1.4

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8Axial Displacement (m)

Axi

al F

rictio

n Fa

ctor

Median Upper Bound Lower Bound

0

0.2

0.4

0.6

0.8

1

1.2

1.4

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2Larteral Displacement (m)

Late

ral F

rictio

n Fa

ctor

Median Upper Bound Lower Bound Figure 6: Equivalent Axial (Left) and Lateral (Right) Friction Coefficients of Zone I.

0

0.2

0.4

0.6

0.8

1

1.2

1.4

0 0.2 0.4 0.6 0.8Axial Displacement (m)

Axi

al F

rictio

n Fa

ctor

Median Upper Bound Lower Bound

0

0.2

0.4

0.6

0.8

1

1.2

1.4

0 0.5 1 1.5 2Larteral Displacement (m)

Late

ral F

rictio

n Fa

ctor

Median Upper Bound Lower Bound Figure 7: Equivalent Axial (Left) and Lateral (Right) Friction Coefficients of Zone II.

Description of Finite Element Method

Three types of Abaqus [17] finite element models were used for analysis follows:

11

1. Full finite element model of the pipeline on uneven seabed

This model was used in the first and last stages of pipeline lateral buckling

analysis. In the first stage, this model was used to evaluate if the strain in an

unwanted buckle triggered by the seabed imperfection, lateral imperfection or

trawling was within allowable limit under worst case loading and pipe-soil

interaction conditions. In the last stage of the analysis, this model was used to

check strain increase in the initiated buckles due to hydrodynamic loads and

pipeline walking.

The full finite element model of the pipeline was constructed using full route

geometry of the pipeline and seabed profile. It incorporated both Zone I and

Zone II of the pipeline properties.

Figure 8 (left) shows the full finite element model of the pipeline on uneven

seabed.

2. Virtual anchor spacing model

Two virtual anchor spacing models were used in this analysis, first one, on flat

seabed and second, on flat seabed with effects of sleepers. Figure 8 (right)

shows the second virtual anchor spacing model.

The first model was used for the following purposes:

• Calculation of the feed-in into a buckle rested on seabed for different

pipe-soil interaction conditions

• Calculation of relationship between the feed-in and strain level in

pipeline rested on seabed for different pipe-soil interaction conditions

The second model was used for the following purposes:

• Checking the span length on each side of the sleeper

• Calculation of maximum axial load capacity of the pipeline rested on

sleeper

• Calculation of the natural frequency of the pipeline span due to sleeper

for VIV analysis

3. Three dimensional solid model

This model was used for ratcheting analysis, and obtaining actual stress

distribution in the pipeline for fracture analysis.

An Abaqus UFRIC FORTRAN subroutine was developed to model the cohesive and break

out behaviors of pipe-soil interaction model.

12

For full finite element model of the pipeline on uneven seabed and virtual anchor spacing

models, the pipeline was split into two sections “around imperfection” and “away from

imperfection”, with different element size to achieve reasonable accuracy . Element length

for each section is listed in Table 3 below.

Figure 8: Finite Element Model of the Pipeline on Uneven Seabed (Left), and Virtual Anchor Spacing

Model Rested on Buckle Trigger (Right).

Table 3: Element Length in Line Models.

Model Section Full Model [m] Virtual Anchor Spacing Model [m]

Around Imperfection 0.2 0.1

Away from Imperfection 4 1 Screening Criteria

For out of straightness (OOS) model lay imperfections bend radius of 2500 m along the

Zone I, and 2000 m along the Zone II route profile were assumed. These magnitudes of

imperfections were set as limiting criteria for pipeline laying. Table 4 gives lower bound

axial capacity of different sections of the pipeline.

Table 4: Axial Force Capacity of the Pipeline.

Axial Force Capacity (kN) KP Zone Hobbs Infinite

Mode OOS Model Minimum

0.0 -0.5 Zone II – B11 1235 1651 1235

0.5 – 19.3 Zone I 920 1151 920

19.3 – 21.8 Zone I 1164 1567 1164

21.8 – 22.3 Zone II – PC04 1164 1567 1164

13

Figure 9 shows effective axial force in the pipeline at different operating temperature. As

this figure shows, maximum introduced axial force in the pipeline is about 4250 kN. So, the

pipeline was susceptible to buckle and local buckling limit states should be checked.

-4500

-4000

-3500

-3000

-2500

-2000

-1500

-1000

-500

00 2.5 5 7.5 10 12.5 15 17.5 20 22.5

KP

Effe

ctiv

e A

xial

For

ce (k

N)

Design PressureTmax = 39.1Tmax=59.1Tmax = 79.1Tmax = 99.1Tmax = 119.1Allowable Value

B11 PC04 Figure 9: Effective Axial Force in the Pipeline (Before Mitigation).

Calculation of Allowable Strain in Buckle Apex

Allowable strain of the pipeline was calculated using DNV-OS-F101 [1], and presented in

Table 5. To calculate strain limit of the pipeline, axial force and internal pressure of the

pipeline were set to maximum value obtained from finite element analysis, and 201.4 barg

(design pressure of the pipeline), respectively.

As Table 5 shows, minimum allowable total strain in the pipeline was 2.8% from local

buckling point of view.

Table 5: Pipeline Local Buckling Strain Limit.

Zone Tempe (oC) Corrosion Allowance (mm)

Allowable Strain (%)

100 0 3.75

100 6 2.85

50 0 3.65 Zone I

50 6 2.80

110 0 4.30

110 9 3.12

50 0 4.18 Zone II

50 6 3.45

14

Using Equation (9), maximum allowable strain range for the pipeline was 1.50%.

Calculation of Maximum Strain in an Isolated Buckle

In the absence of mitigation technique, a single buckle between pipeline virtual anchor

points was considered with the following assumptions:

1. An isolated buckle is formed at the early stage of buckling

2. All possible feed-in resulting from soil with upper bound axial friction is fed into the

buckle considering upper bound lateral friction.

3. Buckle can be triggered by both vertical and horizontal imperfections3

Using these assumptions, total strain in the buckle was 2.85%, see Table 6, which was not

acceptable from low cycle fatigue point of view. So, it was decided to mitigate the pipeline

using a reliable buckle initiation strategy.

Table 6: Summary of Results for One Isolated Buckle.

Analysis Limiting Parameter Value Limit

Local Buckling Total Strain [%] 1.90 2.85

Fatigue Damage Ratio 1.75 1.0

Mitigation Method

To overcome the high level of uncertainty in provided soil data, it was decided to use

sleepers to impose vertical out of straightness to the pipeline. Furthermore calculations

showed that applying a vertical out of straightness might not be sufficient in special

circumstances. So, it was decided to add an in-plane imperfection to control lateral

resistance on the pipeline and direction of buckling. Figure 10 shows two different methods

to apply in-plane imperfection to the pipeline.

Figure 10: Methods of Imposing In-Plane Imperfection, Changing in Route Direction (Left), and Post

Pipe Lay Displacement (Right).

3 Buckle initiation by trawling was not considered as the probability of trawling around the pipeline route was very low.

15

Relationship between Feed-in-Length and Pipeline Axial Strain

Relationship between feed-in into a buckle and pipeline axial strain was calculated using

finite element method. Figure 11 shows the introduced plastic strain in a buckle versus

thermal feed-in. As this figure shows maximum strain in a buckle corresponds to maximum

lateral seabed friction coefficient. Figure 12 shows change in buckle shape due to axial

feed-in.

Table 7 gives maximum allowable feed-in into a buckle in different lateral friction

coefficient based on 1.5% total strain range obtained from low cycle fatigue analysis of the

pipeline. As this table shows material property changes due to temperature has minor effect

on the obtained results.

0.00%

0.50%

1.00%

1.50%

2.00%

2.50%

3.00%

3.50%

4.00%

0.00 1.00 2.00 3.00 4.00 5.00 6.00Feed-In (m)

Plas

tic S

trai

n

MuL=0.25BO, T=100C MuL=0.75BO, T=100CMuL=1.25BO, T=100C MUL=1.25BO, T=60C

1.5% Total Strain Limit

Figure 11: Plastic Strain in Buckle due to Thermal Feed-In.

Table 7: Design Feed-in Length in a Buckle.

Lateral Residual Friction Coefficient Temperature [oC] Design Feed-In (m)

0.25 100 5.10

0.75 100 2.16

1.25 100 1.44

1.25 60 1.56

16

-4

-2

0

2

4

6

8

10

12

14

16

-100 -80 -60 -40 -20 0 20 40 60 80 100

X (m)

Am

plitu

de (m

)

FIL = 1.8 m

FIL = 3.5 m

FIL = 5.5 m

FIL = 7.5 m

FIL = 9.5 m

FIL = 10 m

Figure 12: Buckle Shape due to Axial Feed-In.

Calculation of Distance between Buckle Triggers

Distance between two adjacent sleepers was calculated using an iterative process to satisfy

following criteria:

1. Distance of the first and last sleeper to the end points of the pipeline (PC04 and

B11) should be in such a way to maintain maximum end expansion of the pipeline.

These expansions were obtained from pipeline riser design report, and were 0.8 m,

and 1.2 m for PC04 and B11 side of the pipeline, respectively. Besides, maximum

feed-in into the first and last buckles triggered by the sleepers shall be less than

design feed-in length. Considering these criteria maximum distance of the first

sleeper to the end point of the pipeline (PC04 side) was obtained 1.40 m.

2. Other buckle triggers were positioned to maintain maximum feed-in in each buckle

less than design feed-in length presented in Table 7. In order to assist the controlled

buckle formation, the pipeline concrete coating around the buckle triggers was

removed.

Considering the pipeline routing criteria, final positions of the sleepers were calculated and

are given in Table 8.

Buckle Formation Reliability

Important parameters that affect lateral resistance of buckle trigger on the pipeline (axial

force capacity of a pipeline) were as follows:

1. Sleeper height (imperfection in vertical plane)

17

2. Friction coefficient between pipeline and sleeper

3. Friction coefficient between pipeline and seabed

4. Pipeline span length aside buckle triggers

5. Initial imperfection in horizontal plane

6. Submerged weight per unit length of pipeline

Table 8: Buckle Apex Strain due to Thermal Feed-In.

Sleeper No. KP Feed-In (m) Strain (%)

#1 5.927 2.9 1.0

#2 10.427 3.10 0.9

#3 12.927 2.16 0.9

#4 15.127 2.00 0.85

#5 17.127 1.96 0.85

#6 19.127 2.12 0.9

#7 20.927 1.6 0.8 A parametric study was performed to evaluate the effects of above parameters on the axial

force capacity of the pipeline. Obtained results were formulated in such a way to be used in

Equation (10).

The probability analysis was performed based on the limit state condition prescribed in

Equation (10), and by using 108 Monte Carlo simulations. The results indicated in the 7th

buckle trigger (worst case from buckle initiation point of view), the probability of buckle

initiation failure was less than 10-4.

Selecting Sleepers Height

Considering flat rigid seabed, maximum estimated span length at vicinity of the pipeline of

the buckle triggers was estimated using Equation (12).

472

wEIl δ= (12)

Where, E is the pipeline elastic modulus, I the pipeline section moment of inertia, δ sleeper

vertical movement, and w the submerged weight per unit length of the pipeline

18

This estimation was used as a conservative value for vortex induced vibration analysis of

the pipeline.

To calculate the appropriate pipeline span length, the following criteria were used:

1. To obtain appropriate sleeper behavior, minimum span length of the pipeline aside

the buckle triggers considered being at least equal to buckle length aside the sleeper.

2. Pipeline span length aside the buckle trigger shall not cause fatigue damage due to

vortex induced vibration. Vortex induced vibration calculation for this pipeline was

performed based on DNV-RP-F105 [18] using SLTFATFREE in-house software.

Obtained results showed that the sleeper height shall be between 0.3 m and 0.6 m. This limit

was considered for sleeper design.

Analysis of Mitigated Pipeline

Lateral buckling of the mitigated pipeline was performed to check the final condition of the

buckles initiated by the sleepers. To obtain cyclic behavior of the pipeline (pipeline

walking), two startup-shutdown cycles of the pipeline was modeled including the pipeline

heat up and cool down transients, and effects of hydrodynamic forces. Results showed that

changes in pipeline strain due to pipeline walking and hydrodynamic loads were

insignificant.

From finite element analysis, maximum strain in the pipeline after mitigation was about 1%,

which was less 1.5% limit set by low cycle fatigue analysis. Figure 13 shows the effective

axial force in the pipeline after mitigation.

-2250

-2000

-1750

-1500

-1250

-1000

-750

-500

-250

00 2 4 6 8 10 12 14 16 18 20 22

KP

Effe

ctiv

e A

xial

For

ce (k

N)

Tmax = 50.8Tmax = 58Tmax = 65Tmax = 72.1Tmax = 79.2Tmax = 86.3Tmax = 93.4Tmax = 100.5Tmax = 107.6Tmax = 114.7Tmax = 119.1Allowable Value

B11 B11 Figure 13: Effective Axial Force in the Pipeline After Mitigation.

19

CONCLUSIONS

This paper presents a methodology that was successfully used for the design of a high

pressure high temperature pipeline allowing controlled buckles at designated locations

along the pipeline route, using vertical upset method.

ACKNOWLEDGMENTS

The authors would like to express their appreciation to PETRONAS for supporting this

work and permission to use the project data, as its first HP/HT pipeline in the region.

REFERENCES

[1]. Submarine Pipeline Systems, DNV-OS-F101, 2006.

[2]. In-Service Buckling of Heated Pipelines, Hobbs. E., International Journal of

Transportation Engineering, vol 110, No. 2, 1984.

[3]. HOTPIPE JIP, Design Guidelines for HP/HT Pipelines, L. Collberg et al.

Proceeding of OMAE 2005.

[4]. Design Guidelines for HP/HT Pipelines, S. Goplen et al.

[5]. Load and Resistance Modeling of the Penguins Pipe-In-Pipe Flowline Under Lateral

Buckling, M. Carr, et al., Proceeding OMAE 2004.

[6]. Penguins Flow Line Lateral Buckling Formation Analysis and Verification, I.

Matheson, et al., Proceeding OMAE 2004.

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