ASF01-012-SE-REP-0005_revB

53

description

Offshore pipeline buckling report

Transcript of ASF01-012-SE-REP-0005_revB

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    COMMENTS RESPONSE SHEET DOCUMENT REF. : ASF01-012-SE-REP-0005

    No Document / Page Ref.

    Company Comments Initial Contractor Response Initial Remarks

    1

    2

    3

    4

    5

    6

    7

    Page 2 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Where the term HOLD is specified in this document, it signifies that additional engineering or information shall be required to finalise the document.

    Below is a summary of the HOLDS outstanding in this document.

    HOLDS STATUS SHEET

    HOLD NO SECTION PARA NO DESCRIPTION OF HOLD

    1

    2

    3

    4

    5

    6

    7

    Page 3 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    TABLE OF CONTENTS

    SECTIONS 1.0 INTRODUCTION 10

    1.1 Abadi Gas Field Development 10

    1.2 Scope 11

    2.0 SUMMARY, CONCLUSIONS AND RECOMMENDATIONS 12

    2.1 Summary 12

    2.2 Conclusions 12

    2.3 Recommendation 14

    3.0 DESIGN DATA 15

    3.1 Classification of Location and Safety Classes 15

    3.2 Mechanical Properties 15

    3.3 Operational Data 16

    3.4 Material Properties 17

    3.5 Flowline Equivalent Material Data for Bi-Metal 18

    3.6 Steady State Design Profiles 19

    3.7 Transient Profiles 21

    3.8 Environmental Data 24

    3.9 Seabed Profile 24

    3.10 Pipe Soil Properties 25

    4.0 LATERAL BUCKLING ACCEPTANCE CRITERIA 27

    4.1 Allowable Strain Criteria 27

    4.2 Fatigue Limit State 30

    Page 4 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    4.3 Low Cycle Fatigue 31

    4.4 Limit State Check Summary 31

    5.0 BUCKLING DESIGN METHODOLOGY 33

    5.1 General 33

    5.2 Pipeline Effective Axial Force 33

    5.3 Lateral buckling Susceptibility 33

    5.4 Consequence of single isolated buckle 34

    5.5 Lateral Buckling Design Approach 36

    6.0 RESULTS AND DISCUSSIONS 38

    6.1 General 38

    6.2 Effective Axial Force and Expansion Analysis 38

    6.3 Lateral Buckling Susceptibility 39

    6.4 Consequence of Unplanned Buckling 44

    6.5 Fatigue Damage Assessments 45

    6.6 Planned Buckle to Reduce End Expansion 47

    7.0 REFERENCES 50

    APPENDICES APPENDIX A EQUIVALENT PROPERTIES CALCULATION

    Page 5 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Index of Tables

    Table 2.1 Susceptibility to Lateral Buckling Conclusion .................................................................................. 13 Table 2.2 Reduced Expansion for Planned Buckle ......................................................................................... 14 Table 3.1 Safety Class Definition for DNV-OS-F101 ....................................................................................... 15 Table 3.2 Mechanical Properties ..................................................................................................................... 15 Table 3.3 Operational Data ............................................................................................................................. 16 Table 3.4 Flowline Material Properties ............................................................................................................ 17 Table 3.5 Flowline Coating Properties ............................................................................................................. 17 Table 3.6 Operating Pressure and Temperature ............................................................................................. 19 Table 3.7 Environmental Data ......................................................................................................................... 24 Table 3.8 Axial pipe-soil friction coefficients (Operation) ................................................................................ 25 Table 3.9 Lateral pipe-soil friction coefficient monotonic loading (Operation).............................................. 26 Table 4.1 - Design Factor for Strain Criteria ....................................................................................................... 27 Table 4.2 Limit State Summary ....................................................................................................................... 29 Table 4.3 - S-N Curve Parameters in Fatigue Analysis ...................................................................................... 31 Table 6.1 Low Cycle Fatigue Damage at Weld Root (F1 Curve) .................................................................... 46

    Index of Figures

    Figure 1-1 Field Location Map ............................................................................................................................ 10 Figure 3.1 De-rated API 5L X65 Claded Material True Stress Strain Curve ...................................................... 19 Figure 3.2 Steady State Design Pressure and Temperature Profile .................................................................. 20 Figure 3.3 Shut-in Temperature and Pressure Profile Year 30 ....................................................................... 22 Figure 3.4 Start-up Temperature and Pressure Profile Year 30...................................................................... 23 Figure 3.5 Seabed Profile ................................................................................................................................... 24 Figure 3.6 Axial pipe-soil interaction (Operation) ............................................................................................... 25 Figure 3.7 Lateral pipe-soil friction interaction monotonic loading (Operation) ............................................... 26 Figure 5-1 Material Strength Mismatches Modelling .......................................................................................... 34 Figure 5-2 Insulation Coating SCF Value ........................................................................................................... 35 Figure 5-3 Insulation Coating SNCF ................................................................................................................... 36 Figure 6-1 Effective Axial Force (Lower Bound, Mean and Upper Bound Axial) ............................................... 38 Figure 6-2 Axial Expansion (Lower Bound, Mean and Upper Bound Axial) ....................................................... 39 Figure 6-3 Critical Buckling Force due to Seabed Terrain (LB breakout lateral friction) .................................... 40 Figure 6-4 Lateral Buckling Susceptibility ........................................................................................................... 42 Figure 6-5 Probability of Buckling Along the Unmitigated Pipeline .................................................................... 43 Figure 6-6 Feed-In Capacity of Unplanned Buckle on Seabed .......................................................................... 44 Figure 6-7 Effective Axial Force of One (1) Planned Buckle on Seabed ........................................................... 48 Figure 6-8 Effective Axial Force of Two (2) Planned Buckles on Seabed .......................................................... 48

    Page 6 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Abbreviations / Nomenclature

    3D 3 x outside diameter

    5D 5 x outside diameter

    5LPP 5 layer polypropylene

    ALS Accidental limit state

    API American Petroleum Institute

    ASME American Society of Mechanical Engineers

    BE Best estimate

    CITHP Cold in tube head pressure

    CP Cathodic protection

    CS Carbon steel

    CRA Corrosion resistant alloy

    DC Displacement controlled

    DEP Design engineering practices

    DNV Det Norske Veritas

    EAF Effective axial force

    ECA Engineering critical assessment

    FE Finite element

    FEED Front end engineering design

    FLNG Floating liquefied natural gas

    FLS Fatigue limit state

    HAT Highest astronomical tide

    HOOS Horizontal out of straightness

    HT/HP High temperature / high pressure

    ISO International Standard Organization

    KP Kilometre post

    Page 7 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    LAT Lowest astronomical tide

    LB Lower bound

    LC Load controlled

    LNG Liquifed natural gas

    MMscfd Million standard cubic feet per day

    Mtpa Million ton per annum

    MNFL Manifold

    MSL Mean sea level

    Ncr Critical Buckling Force (SAFEBUCK)

    OD Outside diameter

    OOS Out of straightness

    OS Offshore standard

    Pcr Critical buckling force

    PSC Production sharing contract

    RP Return period

    S-N Stress cycle number

    SNCF Strain concentration factor

    SMYS Specified minimum yield strength

    SMTS Specified minimum tensile strength

    SURF Subsea, umbilicals, risers and flowlines

    U Overall thermal coefficient

    UB Upper bound

    UCR Unity check ratio

    UOE U-Ing, O-Ing and Expanding

    ULS Ultimate limit state

    UNS Unified numbering system

    Page 8 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    VOOS Vertical out of straightness

    WD Water depth

    WGK Wood Group Kenny

    WT Wall thickness

    YT Yield to tensile

    Page 9 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    1.0 INTRODUCTION

    1.1 Abadi Gas Field Development

    The Abadi gas field is located in the Masela PSC Block in the Arafura Sea in Maluku Province. The Masela Block is located approximately 800 km East of Kupang, West Timor, Indonesia and approximately 400 km North of Darwin, Northern Territory, Australia. Water depth in the field ranges from 400 m to 800 m. The southern boundary of the block is adjacent to the agreed Indonesia-Australia maritime border as shown in Figure 1-1.

    INPEX Masela Ltd (COMPANY) is considering development of the Abadi gas field with a FLNG concept which has circa 2.5 Mtpa of LNG production capacity. Optimisation of the LNG production capacity will be performed during FEED. The Abadi development plan includes a single subsea drill centre with 5 production wells and a subsea manifold tied back to a FLNG facility via three flowlines and three flexible risers. The subsea wells will be required to produce approximately 467 MMscfd on average for 30 years to meet yearly LNG production targets.

    Figure 1-1 Field Location Map

    Page 10 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    1.2 Scope

    This report establishes the susceptibility of the production flowline to global buckling, the post-buckle response of the pipe and its feed-in capacities. The analysis is based on both thermal expansion calculations and non-linear finite-element model of the flowline.

    Required mitigation measure will be addressed in a separate report, subject to outcome of this report.

    Page 11 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    2.0 SUMMARY, CONCLUSIONS AND RECOMMENDATIONS

    2.1 Summary

    Susceptibility to global lateral buckling for the 12 inch Abadi flowline is accessed based on two methodologies:

    1. Deterministic: comparison of the pipeline effective axial force with the critical buckling force based on Hobbs equation as presented in SAFEBUCK [4] and the critical buckling force due to seabed undulation from FEA model.

    2. Probabilistic: based on Monte Carlo simulation, the probability of rouge buckle formation determined by sampling the critical buckling force along the flowline and compared against the effective axial force. Probability of buckling is quantified by the number of events that trigger a rouge buckle. Log-normal distribution is assumed for all governing parameters used in the simulation [4].

    The overall design methodology broadly follows the SAFEBUCK guidelines while acceptance criteria refers to DNV OS-F101 for displacement controlled condition, i.e. strain based criteria.

    Design temperature of 136oC at inlet and 110oC outlet with constant design pressure of 318 bar are used to access the lateral buckling susceptibility. A model length of 3 km is used to represent the full length of the Abadi flowline.

    Post-buckle capacity of the pipe, in the event of an isolated rouge buckle occurring, is accessed by evaluating the amount of thermal expansion that can be accommodated before the buckle load exceeded the allowable limit. Strain concentration factor (SCNF) arising from insulation coating cutback is also determined in the strain calculation. Fatigue capacity of the buckle due to temperature and pressure cycles are then assessed based on fatigue S/N curve.

    Due to the limitation of tie-in spool expansion offsets at both hot and cold ends, a preliminary expansion calculation is also performed to assess the feasibility of using engineered buckles to control the thermal expansion into the tie-in spool. Both one (1) and two (2) buckle scheme are considered.

    2.2 Conclusions

    2.2.1 Lateral Buckling Susceptibility Assessment

    Deterministic analysis results confirmed that the 12-inch Abadi flowline is not susceptible to lateral buckling along its entire length. This was established based on comparison of unmitigated pipe effective force profile against the critical buckling force from Hobbs equation and from seabed profile.

    Probabilistic calculation based on Monte Carlo simulation also shown that the probability of buckling along the flowline is less than 0.01%, as compare to a threshold value of 5% based on SAFEBUCK [4] guideline.

    Page 12 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Table 2.1 Susceptibility to Lateral Buckling Conclusion

    MANNER PARAMETER VALUE CONCLUSION

    DETERMINISTIC ASSESSMENT

    Max compressive effective axial force

    365 kN Effective axial force < critical buckling force hence rouge

    buckle on seabed is not likely.

    Hobbs critical buckling force 604 kN

    Seabed critical Buckling Force 758 kN

    PROBABILISTIC ASSESSMENT Maximum probability of buckling 0.08%

    < 5% (SAFEBUCK

    threshold value)

    2.2.2 Consequence on Unplanned Buckle

    For a CRA clad pipeline, the design strain is likely to be limited by the ECA requirements, as other acceptance requirements are generally found less stringent. In the present work and from past project experience, a limiting strain of 1% is assumed prior to actual ECA assessment. At 1% strain, the thermal feed-in to the buckle is found to be in excess of 5m, much greater than the end expansion of the pipeline.

    In the unlikely event of an uncontrolled buckle occurring, the maximum feed-in to the buckle is estimated to be about 2.3m which is well within the capacity of the pipe. The corresponding maximum strain at the buckle apex is about 0.2% at best estimate axial and lateral frictions, which is also within the design limit of 1%.

    The accumulated fatigue damage calculation due to operating cycles, assuming 12 cycles per year, gave a total fatigue damage unity check of about 10-3 over the design life of the pipeline. This assumed a shutdown frequency of 12 cycles per years. However, in view of the low fatigue damages, it is expected that considerably more shutdown cycles can be accommodated.

    2.2.3 Reduced Expansions from Planned Buckle

    Expansion calculation shown that the pipe end expansions can be reduced by introducing planned buckles in the pipeline. These buckles shared the thermal expansion loads and hence reduce the end expansions. Preliminary calculations with one (1) buckle located at mid-length and two (2) buckles spaced at 1 km apart, shows able to reduce the hot end expansion from the unmitigated value of 2.52m to 1.88m and 1.63m, respectively, as summarised in Table 2.2 overleaf.

    Page 13 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    It should be emphasized that the end expansion is a function of the buckle locations,

    which in turn, is dictated by the pipe effective force profile to ensure that the planned buckles can be initiated reliably. This therefore limits the minimum effective distance of the buckle from the pipe end.

    Table 2.2 Reduced Expansion for Planned Buckle

    Ends Expansion (m)

    Un-mitigated One (1) Planned Buckle (2) Planned Buckle

    Buckle Location - KP 1.5 KP 1.0 & KP 2.0

    Hot End (MNFL) 2.52 1.88 1.63 Cold End (Riser) 2.26 1.55 1.33

    2.3 Recommendation

    Based on the analysis results, Abadi flowline is not susceptible to uncontrolled global buckling and hence buckle mitigation for the purpose of stress relieving is not deemed necessary. Analysis also confirmed that the chosen pipe section and properties have sufficient capacity to accommodate a rouge buckle on seabed should one is to occur during operation.

    Buckle mitigation is effective in limiting the pipe end expansions due to shared thermal feed-in to the buckles. Preliminary calculation shown that mitigation based on a two (2) buckles is able to reduce the end expansion to a value within the capacity of the tie-in spool. However, more detailed assessments are necessary to confirm the viability of this solution with respect to buckle initiation and post buckle responses. This should include considerations for the actual pipe lay process, buckle initiation method (e.g. ZRB) and, if necessary, increase friction at pipe-soil interface via rough coating of pipe. Various sensitivity assessments are also necessary to quantify the locations of the buckle and their reliability.

    Due to short flowline length and the relatively low axial friction, the flowline may be prone to global axial walking, which can impose significant additional loads onto the tie-in spool. This can also lead to requirements for flowline anchoring, if axial walking persists. This has large impacts on the PLET structure configuration design and, to some extent, the overall field layout. As such, the proposed buckle mitigation analysis should also include axial walking to provide an early indication of walking potential and pipe anchoring loads, if applicable.

    Page 14 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    3.0 DESIGN DATA

    The input data presented in this section is taken from URF and Structure Design Premise [1] unless noted otherwise.

    3.1 Classification of Location and Safety Classes

    The safety class discreteness is presented in Table 3.1 for DNV-OS-F101.

    Table 3.1 Safety Class Definition for DNV-OS-F101

    SECTION FLUID CLASS

    LOCATION CLASSIFICATION

    SAFETY CLASS TEMPORARY

    SAFETY CLASS OPERATION

    Flowline

    E 1 Low Medium

    3.2 Mechanical Properties

    The mechanical properties are summarised in Table 3.2.

    Table 3.2 Mechanical Properties

    PARAMETER UNIT VALUE

    Flowline Material - ISO 3183 L450 / API 5L X65 Manufacturing Method - Seam Welded Clad Pipe Flowline length km ~ 3 Outer Diameter mm 323.9 Wall Thickness (CS) - [2] mm 17.5 Internal Clad Thickness mm 3.0 External Coating - 5LPP mm 5.2

    Wall Thickness Tolerance % -0.5 (Base metal)

    0 (Cladding) Pipe Ovality % 1.5

    Page 15 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    3.3 Operational Data

    Operational data is presented in Table 3.3.

    Table 3.3 Operational Data

    PARAMETER UNIT VALUE

    Product/Content - Gas Design Throughput Capacity each Flowline MMscfd 250 Flowline Design Temperature oC 136 Maximum Operating Temperature oC 131 Normal Operating Inlet Temperature oC 125 Design Pressure at CITHP bar 318 Maximum Operating Pressure bar 219

    Ref. Elevation for Design Pressure m - 609 Hydrotest Pressure @ Seabed Level bar 461.75

    Ref. Elevation for Hydrotest m + 30 Location of Hydrotest - offshore Maximum and Minimum Content Bulk Density 207/86 Assumed Static Lay Tension:

    Tonnes

    - Top 50 - Bottom 25

    Page 16 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    3.4 Material Properties

    The material properties for flowline are summarised in Table 3.4.

    Table 3.4 Flowline Material Properties

    PARAMETER UNIT CS ISO- 3183 L450 / API 5L X65 GRADE

    ALLOY 625 (UNS N06625)

    Density kg/m3 7850 8440 Poisson Ratio - 0.3 0.31 Thermal Expansion Coefficient at 20C

    /oC 11.7 x 10-6 12.8 x 10-6

    Thermal Expansion Coefficient at 130C (1)

    /oC 12.6 x 10-6 13.7 x 10-6

    SMYS at 20C MPa 450 414 SMYS at 136C MPa 405.6 (2) 369.6 (2) SMTS at 20C MPa 535 827 SMTS at 136C MPa 490.6 (2) 782.6 (2) Elastic Modulus at 20C GPa 207 205 Elastic Modulus at 136C (3) GPa 200 (3) 198 (3)

    Notes: 1. Values conservatively adjusted to design temperature based on increments of 8x10-9 per

    C for carbon steel as per SAFEBUCK III [4] 2. De-rated as per DNV-OS-F101 recommendation [8] 3. Taken and de-rated from SAFEBUCK III [4]

    Linear interpolation will be used to establish material de-rating values at intermediate temperatures.

    The flowline coating material property is summarised in Table 3.5.

    Table 3.5 Flowline Coating Properties

    COATING UNIT VALUE

    5LPP Equivalent Density kg/m3 920

    Page 17 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    3.5 Flowline Equivalent Material Data for Bi-Metal

    The properties of CRA and carbon steel, based on specified minimum values, will be combined to obtain an equivalent property for finite element modeling. This is a normal practice for lateral buckling design and has been adopted in past projects in line with SAFEBUCK III [4] recommendations and also other recent industry guidelines [7]. A summary of the important basis in modeling the CRA lining is as follows:

    1. The effective wall thickness is taken as the nominal wall thickness of the pipe plus the cladding.

    2. Equivalent elastic properties are then calculated as the weighted average of the carbon steel and cladding properties as:

    a. cs

    ccsseq tt

    tt+

    +=

    b. where is the property of interest (Elastic Modulus, Poissons ratio, thermal expansion coefficient, etc.), is the nominal wall thickness. The subscript s and c refers to carbon steel and cladding respectively.

    3. The submerged weight of the flowline incorporates the weight of the cladding.

    4. The stress-strain response of the equivalent pipe is obtained through a process of weighted averaging, where the stress value of the given strain is averaged.

    5. For FE analysis, the pipe is modeled as a single pipe with the effective wall thickness of steel plus cladding. The stress-strain curve used is taken as the weighted average of the stress-strain curve for steel and cladding. In addition, the nominal stress is converted into true stress prior to weight averaging.

    However, for characteristic resistance calculation (i.e. local buckling limiting strain), strength contribution from cladding will be excluded.

    The de-rated true stress-strain curve for the steel and cladding at maximum design temperature of 136OC is shown in Figure 3.1 overleaf.

    Page 18 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Figure 3.1 De-rated API 5L X65 Claded Material True Stress Strain Curve

    3.6 Steady State Design Profiles

    Detailed operating pressure and temperature profiles for different possible combinations of flowrate are provided in COMPANYs Flow Assurance report [13] and summarised in Table 3.6.

    Table 3.6 Operating Pressure and Temperature

    FLOW RATE

    PRESSURE / TEMPERATURE

    DOWNSTREAM OF SUBSEA

    CHOKE

    FLET MANIFOLD

    END RISER INLET

    RISER TOP

    1. Base Case 185 MMscfpd x 3 flowlines

    221bara / 130degC

    219bara / 126degC

    215bara / 69degC

    202bara / 63degC

    2. Normal Operation Day 1 185 MMscfpd x 3 flowlines

    204bara / 124degC

    209bara / 125degC

    205bara / 68degC

    192bara / 62degC

    3. Normal Operation Year 15 (202 MMscfpd x 2 flowline) + (150 MMscfpd x 1 flowline)

    179bara / 125degC

    178bara / 122degC

    173bara / 71degC

    160bara / 64degC

    4. Normal Operation Year 30 (222 MMscfpd x 2 flowline) + (111 MMscfpd x 1 flowline)

    113bara / 120degC

    112bara / 119degC

    101bara / 70degC

    91bara / 63degC

    5. Severe Condition 300 MMscfpd x 1 flowline 255 MMscfpd x 1 flowline

    140bara / 125degC

    119bara / 122degC

    99bara / 81degC

    82bara / 73degC

    6. Severe Condition, late production stage, 20% of U- 123bara / 119bara / 110bara / 82bara /

    Page 19 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Page 20 of 50 ASF01-012-SE-REP-0005 Rev. B

    FLOW RATE

    PRESSURE / TEMPERATURE

    DOWNSTREAM OF SUBSEA

    CHOKE

    FLET MANIFOLD

    END RISER INLET

    RISER TOP

    value reduction in jumper, manifold and tie-in spool sections and fully buried flowline (220 MMscfpd x 2 flowlines) + (110 MMscfpd x 1 flowline)

    138degC 131degC 105degC 94degC

    7. Turndown Operation 76 MMscfpd x 3 flowlines

    97bara / 106degC

    97bara / 100degC

    96bara / 24degC

    88bara / 18degC

    From the flow assurance sensitivity evaluations of the above cases, it was identified that Case 6 yields the most critical temperature during late life production when the flowline is assumed fully buried. The expected steady state maximum operating inlet and outlet temperature are 131C and 105C, respectively. A 5C margin is then added to the maximum operating values to give the design profile (136 C inlet) as shown in Figure 3.2 below.

    Figure 3.2 Steady State Design Pressure and Temperature Profile

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    3.7 Transient Profiles

    For cyclic and global axial walking analysis, transient temperature and corresponding pressure profiles representative of the flowline loading under day-to-day operating condition will be considered. The flow transient arises from the operating shutdown and restart cycles.

    The flow assurance study report [13] has considered the full shutdown case whereby after normal production, the well is shut-in and fully cooled down to ambient. The well is then re-opened to the flowline for ramp-up to normal production after two hours. The corresponding transient temperature and pressure profiles along the flowline are shown in following Figure 3.3 and Figure 3.4, respectively.

    Page 21 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Figure 3.3 Shut-in Temperature and Pressure Profile Year 30

    Page 22 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Figure 3.4 Start-up Temperature and Pressure Profile Year 30

    For initial study, the numbers of full shut down cycles assumed are twelve (12) per year throughout the operating life of the flowline.

    Page 23 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    3.8 Environmental Data

    Table 3.7 presents the environmental data.

    Table 3.7 Environmental Data

    DESCRIPTION UNIT VALUE

    Water Depth (MSL) m 607-609 Highest Astronomical Tide (HAT) m 1.27 Mean Sea Level (MSL) m 0.00 Lowest Astronomical Tide (LAT) m -1.62 Seawater density kg/m3 1025 Minimum Seawater temperature oC 5.8 Maximum Wave Height -100yrs RP m 12.49 Marine growth thickness mm 10 Marine Growth Density kg/m3 1175

    3.9 Seabed Profile

    Two types of seabed surface are to be utilised in ABAQUS:-

    1. Flat Seabed this type of seabed is utilised in FE model to evaluate the buckle feed-in and strain.

    2. Actual Seabed this type of seabed is used to determine the critical buckling force for any unwanted/unplanned buckle(s) initiated by vertical seabed imperfection.

    Figure 3.5 shows the seabed profile for the proposed flowline [1].

    Figure 3.5 Seabed Profile

    -609.4-609.3-609.2-609.1-609.0-608.9-608.8-608.7-608.6-608.5-608.4-608.3-608.2-608.1-608.0-607.9-607.8-607.7

    0 500 1000 1500 2000 2500 3000 3500 4000

    Disntance from PLET, m

    Ele

    vatio

    n, m

    PLET Manifold end

    Riser inlet

    Page 24 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    3.10 Pipe Soil Properties

    The pipe soil properties are summarised in Table 3.8 &

    Figure 3.6 and Table 3.9 & Figure 3.7 for axial and lateral soil properties, respectively, based on SURF Pipe Soil Interaction Data technical note [3].

    Following the SAFEBUCK guideline, only the residual axial friction is used in all thermal expansion calculation.

    Table 3.8 Axial pipe-soil friction coefficients (Operation)

    MOBILISATION DISTANCE (m)

    LOWER BOUND

    BEST ESTIMATE

    UPPER BOUND

    0 0 0 0

    0.003 0.38 0.58 0.96

    0.330 0.13 0.19 0.32

    0.659 0.13 0.19 0.32

    Figure 3.6 Axial pipe-soil interaction (Operation)

    Page 25 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Table 3.9 Lateral pipe-soil friction coefficient monotonic loading (Operation)

    MOBILISATION DISTANCE (m)

    LOWER BOUND

    BEST ESTIMATE

    UPPER BOUND

    0 0 0 0

    0.049 0.55 0.92 1.72

    0.989 0.36 0.57 0.90

    1.648 0.36 0.57 0.90

    Figure 3.7 Lateral pipe-soil friction interaction monotonic loading (Operation)

    Page 26 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    4.0 LATERAL BUCKLING ACCEPTANCE CRITERIA

    4.1 Allowable Strain Criteria The allowable strain criteria define the limiting strain capacity of the buckle. Several considerations are to be addresses as summarized below.

    4.1.1 Local Buckling This limit state is to check against the development of pipe wall wrinkle (local buckling) at the buckle apex when subject to axial compressive load. Design standard based on DNV-OS-F101 [8] for local buckling can be used, which provides two levels of checks:

    1. Limiting the bending moment at the buckle apex - a Load Controlled (LC) criterion applicable to all stress based design assessment.

    2. Limiting the compressive strain at the buckle apex - a Displacement Controlled (DC) criterion applicable to strain based design only.

    For lateral buckling assessment, LC criterion is often found to be too restrictive and conservative on pipe bending capacity while DC criterion provides much greater flexibility. In-line with current industry practices for HT/HP flowlines and also many recent project experiences, DC criterion is adopted for this project.

    The appropriate partial safety factors and relevant material parameters for the DC limiting strain are presented in Table 4.1.

    Table 4.1 - Design Factor for Strain Criteria DESCRIPTION DNV-OS-F101 [8] SYMBOL CATEGORY/VALUE

    Hazard Potential Sec.2 C200 - E

    Safety Class Sec.2 C400 - Medium

    Material Strength Factor Sec.5, Table 5.6 U 0.96 Maximum allowed yield to tensile ratio Sec.7, Table 7.5 h 0.90

    Girth weld factor Sec.13, E1000 gw 1.0

    Resistance strain factor Sec.5, Table 5.8 2.5

    Functional Load Effect Factor Sec.4, Table 4.4 F 1.10

    Condition Load Effect Factor Sec.5, Table 4.5 C 1.07

    The calculations of the allowance strain for internal and external overpressure are stated below as per DNV-OS-F101 [8] requirement.

    Page 27 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    175.5101.078.0 5.1min

    +

    gwh

    b

    esd p

    ppDt For Internal Overpressure

    1min

    8.0

    +

    SCm

    C

    e

    E

    C

    SD

    PPP

    For External Overpressure

    Where:

    h Maximum allowed yield to tensile ratio

    gw Girth weld factor

    Resistance factor (strain resistance)

    sd Design axial strain

    minP Minimum internal pressure

    eP External pressure

    bP Containment resistance pressure

    CP Collapse pressure

    t Minimum wall thickness D Outer diameter

    The design compressive strain, sd is given by

    CCFSD =

    Where:

    F is the functional load effect factor (=1.1)

    C is the condition load effect factor (=1.07)

    C is the analysis predicted compressive strain at the buckle.

    It should be stressed that limiting strain for local buckling is dependent on the pipe internal pressure; higher pressure will result in higher resistance and thus larger compressive strain capacity. The anticipated minimum local operating internal pressure therefore shall be used for limiting check.

    In determining the above characteristic strain, strength contribution from the cladding shall be neglected.

    Page 28 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    4.1.2 Material Uniform Strain Capacity

    The strain limit against local buckling defined previously may not be relevant to thick walled pipe for which the design strain could be unrealistically large. For this reason, additional limit on the equivalent strain is necessary to prevent material from necking under tension before local buckling could occur on the compression side.

    Essentially the limit prevents gross plastic deformation of the pipe cross-section. Typically the onset to gross plastic deformation (and prior to failure) under tension is the overall gross reduction in cross sectional area, resulting in a significantly smaller area at the point of impending failure compared to the rest of the section. Hence, forming what is seen as a necking effect as the area continuously reduces under load.

    SAFEBUCK III [4] has proposed a simple expression for equivalent strain limits based on material yield-to-tensile ratio (YT) as below:

    )99.0(25.0 YT

    The allowable equivalent strain limit is therefore 1.5% for present pipe section assuming a maximum YT ratio of 0.93.

    4.1.3 Engineering Critical Assessment Limiting Strain From past projects, it was found that design strain based on local buckling and/or uniform material tensile strength discussed above may result in very stringent Engineering Critical Assessment (ECA) weld requirements, which can be difficult to achieve with respect to acceptable defect flaw size. For this reason, a maximum design strain of 1.0% is imposed at the weld joint.

    4.1.4 ALLOWABLE STRAIN LIMIT STATE SUMMARY The design limiting strain criteria discussed previously are summarised in Table 4.2.

    Table 4.2 Limit State Summary

    PIPE SIZE COMPRESSIVE

    (%) TENSILE

    (%) ECA (%)

    DESIGN STRAIN

    (%)

    323.9mm OD x 17.5mm WT 1.36 1.5 1.0 1.0

    Page 29 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    4.2 Fatigue Limit State

    Fatigue assessments will be conducted using the methodologies outlined in DNV-RP-C203 [9] for weld toe (S/N curve D) and weld root (S/N curve F1). The predicted number of cycles to failure N is expressed in the following general form:

    = k

    refloc t

    tmaN )((.logloglog

    Where:

    alnorloc SCF min =

    And:

    k is the thickness exponent; m is the fatigue exponent; log a is the fatigue strength constant; t is the pipe wall thickness; tref is the reference wall thickness 25mm

    The Stress Concentration Factor (SCF) at the weld toe (curve D) is expressed in equations below.

    )exp(31Dt

    tSCF m +=

    Where:

    min

    minmax

    22

    2/)(DDtt

    MaxOvality

    Thickness

    OvalityThicknessTot

    Totm

    =

    =

    +=

    =

    For simplicity, curves corresponding to seawater with cathodic protection will be used at weld toe (curve D) whereas in-air curve (curve F1) will be used for weld root. The relevant parameters are summarized in Table 4.3.

    Page 30 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Table 4.3 - S-N Curve Parameters in Fatigue Analysis

    Curve F1 at Weld

    Root (In Air)

    Curve D at Weld Toe

    (Seawater with CP) Fatigue exponent, m 3 3

    Fatigue strength constant, log a 11.699 11.764

    SCF 1.0 1.55 (1)

    Thickness exponent, k 0 0.15

    Fatigue knockdown factor 2 (2) 1

    Allowable Fatigue Usage 0.1 (3) Notes: 1. SCF calculated based on the selected pipe section properties. 2. Apply at weld root to account for clad girth weld [7]. 3. A split of 20% and 80% between installation and operational fatigues to be assumed until actual

    flowline installation method is confirmed.

    4.3 Low Cycle Fatigue

    Repeated heat-up and shutdown cycle experienced by the flowline during its design life imposes fatigue damage on the unplanned buckles. The low cycle fatigue is investigated using the transient temperature and pressure profiles determined from flow assurance study [13] based on the startup and shutdown conditions. It is assumed that there will be twelve (12) full shut-down cycles per year, as reflected in the flow assurance study in Section 3.7

    The axial stress range is extracted from the global FE model at pipe mid-wall and used to calculate the low cycle fatigue damages

    4.4 Limit State Check Summary In summary, global buckle design limit state checks are to comprise of the followings:

    i. Ultimate Limit State (ULS)

    Local buckling strain limit at buckle apex of unplanned buckles;

    Material uniform tensile strain limit to ensure non-occurrence of material necking under tensile load;

    Cyclic plasticity limit to ensure non-occurrence of hoop strain ratcheting;

    Page 31 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    ii. Fatigue Limit State (FLS)

    Low cycle fatigue due to operational pressure and temperature fluctuations;

    ECA weld fracture and fatigue crack growth (scope not covered in this document);

    iii. Accidental Limit State (ALS)

    No trawling assumed.

    Page 32 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    5.0 BUCKLING DESIGN METHODOLOGY

    5.1 General

    This section presents the general assessment methodology that will be undertaken for this phase of the FEED work with the objectives of confirming the pipe susceptibility to global buckling and the subsequent post-buckle response and axial working behavior. The methodology discussed here focuses on buckle acceptance criteria, modeling parameters, loading conditions and interactions, limit states and the FE modeling itself.

    5.2 Pipeline Effective Axial Force

    The first step in the buckling assessment is to establish the effective force distribution along the pipeline. The effective force analysis is performed using in-house spreadsheets which compute the effective axial force and its corresponding axial expansion. Lower bound, mean and upper bound sliding axial friction factors shall be considered. It shall be assumed that the pipeline is unrestrained at both ends and is free to expand at the tie-in spools since axial resistance from the vertical jumper spools are expected to be minimal.

    5.3 Lateral buckling Susceptibility

    The next step of the assessment is to determine the susceptibility of the pipeline to lateral buckling. This is carried out by comparing the effective force of the pipeline to the critical buckling force due to vertical out-of-straightness (VOOS) these are imperfection introduced by the undulations of the seabed.

    The susceptibility assessment is carried out in both a deterministic and probabilistic manner. In the former, pipeline susceptibility is assessed through comparison of the pipeline effective axial force with theoretical critical buckling forces from Hobbs and route curvature.

    In the latter, a probabilistic distribution of the VOOS and HOOS is used to generate a random series of potential critical buckling forces which is then compared with the pipeline effective axial force. The probability of lateral buckling is then quantified by the number of events that trigger an unplanned/rogue buckle.

    Pipeline OOS data is unknown prior to installation and there is no existing database of pipeline OOS in this region. Therefore, in lieu of these data, the recommended OOS distribution from [4] will be utilized.

    Page 33 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    5.4 Consequence of single isolated buckle

    If the pipeline is found to be susceptible, the acceptance of an event of a single isolated buckle will be qualified through: -

    1. Local buckling limit state the buckle shall be assessed and shown to fulfil the requirements of Section 4.1;

    2. Fatigue limit state the buckle shall be assessed and shown to fulfil the requirements of Sections 4.2 for low cycle fatigue only.

    5.4.1 Material Strength Mismatches

    To account for possible material strength mismatch at the buckle apex, a weak link concept is normally adopted as per recommendations in SIEP-EP-5154 [11] and [10]. It is assumed that the fully plastic moment in the weak section is 12% below the fully plastic moment of the adjacent pipe and the length of the weak section is set to be 6m in length. In past projects, the weak section property has been based on temperature de-rated SMYS and the mismatch in the adjacent flowline is set to be 12% higher than the de-rated weak section as shown in Figure 5.1 below.

    Figure 5-1 Material Strength Mismatches Modelling

    However, in recent SAFEBUCK III update [4], it was stated that modelling of normal joint to joint strength mismatch on strain localisation is not required, although other forms of SCNF still need to be accounted for. Given the relative small size of the Abadi flowline with no concrete weight coating, buckle strain is expected to be low to moderate, any impact on strain location due to material strength mismatch is expected to be relatively small. For this reason, it is excluded from the weak-link model.

    5.4.2 Wall Thickness Variations

    For pipe manufactured to UOE process with long seam welds, wall thickness tolerances between joints or within joint are generally small as compare to seamless pipe [10]. As such, wall thickness fabrication tolerance is also not included in the weak-link model.

    Page 34 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    5.4.3 Coating Cutbacks

    The change in local bending stiffness due to cutback of external insulation coating at field joint can lead to strain localization. The degree of localization depends on the elastic modulus of the insulating and field joint coating material. If the same high integrity insulation coating material is adopted as field joint infill, stiffness discontinuity will be minimal and strain localization may be neglected.

    However, Abadi flowline coating has yet to be confirmed and hence the SCNF effect from insulation coating cutback will be accounted for by using a methodology presented in [12]. The

    approach is based on conservation of total moment across the steel pipe and the relevant insulation coated sections, as depicted in

    Figure 5-2.

    Figure 5-2 Insulation Coating SCF Value

    Total moment, Mt = Mcc = Mfj, where

    Mcc = Moment by insulation + Moment by steel pipe (nominal moment) and;

    Mfj = Moment by field joint + Moment by steel pipe (peak moment).

    The above two equations can be further expressed in terms of nominal and peak strain respectively using the basic relationship between sectional stress and moment arm. The two equations are then solved for a series of bending moments and the SCNF is simply the ratio of the peak strain at the field joint to the nominal strain at the insulated section.

    Page 35 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Figure 5-3 below shows the insulating SNCF at the unplanned buckle region against the normalised bending strain.

    Figure 5-3 Insulation Coating SNCF

    As can be seen from the figure above, strain concentration factor (SNCF) increases rapidly after the nominal strain exceeded 0.2%. Therefore, the predicted bending strain from finite element (FE) model at unplanned buckle is factored by the SNCF value to give the total expected peak strain.

    5.5 Lateral Buckling Design Approach

    The design process involves a series of steps, as detailed below:-

    Step 1 Define acceptance criteria

    Defining the acceptance criteria (i.e. local buckling limit states, engineering critical assessment constrain and allowable axial displacements).

    Step 2 Perform pipeline effective axial force and end expansion analysis

    Perform a deterministic pipeline effective axial force and its corresponding end expansion calculation. If the end expansion exceeds the allowable limit, expansion control measures are required.

    Step 3 Perform lateral buckling susceptibility assessment

    Page 36 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Once the requirement for expansion control has been defined, the susceptibility of the pipeline to lateral buckling is determined by comparing the effective axial force with the critical buckling force from all possible out-of-straightness considered in this work.

    Step 4 Assess the acceptability of unplanned buckle

    If Step 3 shows that the pipeline is likely to buckle, the next step is to determine if the unplanned/uncontrolled buckling would result in non-compliance with the acceptable limit states. This is done by investigating the post-buckle behaviour of a single isolated buckle on seabed.

    If Step 4 shows that unplanned buckling results in non-compliance of the limit states then buckling mitigation will be developed and carried out in separated document.

    Page 37 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    6.0 RESULTS AND DISCUSSIONS

    6.1 General

    The detailed lateral buckling assessments of the 12inch Abadi flowline are discussed in this section.

    6.2 Effective Axial Force and Expansion Analysis

    The first step in the lateral buckling assessment is to establish the effective axial force distribution along the pipeline.

    The effective force and corresponding axial expansions for lower, mean and upper bound axial frictions are shown in Figure 6-1 and Figure 6-12, respectively.

    Figure 6-1 Effective Axial Force (Lower Bound, Mean and Upper Bound Axial)

    From Figure 6-1, it can be seen that the Abadi flowline can be considered as short pipeline where the axial soil friction is fully mobilised. The peak effective force at UB axial friction is about -365kN while the LB value is only -150kN.

    In Figure 6-2, it is seen that the flowline hot end expansion is about 2.5m for all three axial friction cases considered.

    Page 38 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Figure 6-2 Axial Expansion (Lower Bound, Mean and Upper Bound Axial)

    6.3 Lateral Buckling Susceptibility

    6.3.1 General

    Unplanned buckles can be initiated by OOS incurred during installation and/or by natural occurring seabed terrain. To establish the possible range of buckling force, the followings are considered and analysed in ABAQUS:-

    1. The critical buckling force along the pipeline due to naturally occurring vertical imperfections based on actual seabed profile along the centre line of the straight pipeline route.

    2. The critical buckling force calculated based on Hobbs equation as per SAFEBUCK recommendations.

    6.3.2 Critical Buckling Force Due to Seabed Terrain

    The overall seabed profile along the flowline is generally flat. Nevertheless, there are vertical imperfections at various locations which may act as buckle initiators. Critical buckling forces at these locations need to be established.

    In order to investigate this, an FE analysis has been carried out with the actual seabed profile. The general features of the FE model used in this analysis are:-

    Page 39 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    1. The entire seabed is modelled as a rigid contact surface. Since only the vertical

    profile is available, this profile is extruded to generate a 3D surface with variation in the vertical direction.

    2. The analysis is carried out using lower bound lateral breakout friction in combination with minimum content density to provide the least resistance to unplanned buckle formation.

    The results, together with the corresponding seabed profile are summarised in Figure 6-3 below. It showed that there is one location at approximately KP 0.2km, where the critical buckling force is lowest at about 758 kN as denoted by the solid blue dot. However, unplanned buckle at this location is deemed very unlikely due to its proximity to the pipe end where there is insufficient effective force built-up to initiate the buckle, see Figure 6-1.

    Figure 6-3 Critical Buckling Force due to Seabed Terrain (LB breakout lateral friction)

    .

    Page 40 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    6.3.3 SAFEBUCK Hobbs Equation

    For screening purposes, the critical buckling force, Ncr, for a nominally straight pipeline is given in SAFEBUCK [4] as:

    ),65.0min( crBcr NNN =

    Where

    DWEIN Ls = 86.3 and RWN LscrB =

    EI is pipeline flexural rigidity

    L is minimum lateral break-out friction.

    sW is minimum pipe submerged weight.

    D is steel outer diameter

    R is lateral radius

    Since the initial layout of the flowline route does not have any route curvature, the critical buckling force is solely determined from N to give a value of 604 kN.

    6.3.4 Susceptibility of Lateral Buckling

    6.3.4.1 General The susceptibility of the 12-inch Abadi flowline to lateral buckling is assessed by comparing the critical buckling forces with the pipe effective axial force. Two different approaches are considered in this work.

    6.3.4.2 Deterministic Assessment The lateral buckling susceptibility of the pipeline is determined by directly comparing the critical buckling forces determined in Sections 6.3.2 and 6.3.43 against the pipe effective axial force shown in Figure 6-1. The comparison plot is summarised in Figure 6-4 and it can be clearly seen that the effective force is inadequate to initiate any rouge buckle on seabed for entire length of the flowline. This therefore confirms that the flowline is not susceptible to global lateral buckling.

    Page 41 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Figure 6-4 Lateral Buckling Susceptibility

    6.3.4.3 Probabilistic Assessment Lateral buckling probability is also assessed using a probabilistic model as outlined in SAFEBUCK [4]. The probability assessment is performed using an in-house Mathcad spread sheet based on Monte Carlo simulation with 10,000 sampling points.

    Figure 6-5 overleaf shows the bucking probability distribution along the length for a higher axial friction of 0.39 (design UB is 0.32) after additional soil sensitivity assessments. The maximum buckling probability is less than 0.08% at KP 1.75, which is much lower than the threshold value of 5% for the full pipe length given in the SAFEBUCK guidelines [4]. This therefore further affirms the previous conclusion from the deterministic model that the flowline is not susceptible to global buckling.

    Page 42 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Figure 6-5 Probability of Buckling Along the Unmitigated Pipeline

    6.3.5 Lateral Buckling Susceptibility Summary

    Both the deterministic and probabilistic calculations above have shown that the 12inch Abadi flowline is not susceptible to global lateral buckling.

    Although the buckling likelihood is very low, it is nevertheless useful to find out the capacity of the buckle should one occur on seabed. This is discussed in the next section.

    Page 43 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    6.4 Consequence of Unplanned Buckling

    The consequence of a rouge buckle can be assessed by investigating the post-buckle characteristics of an isolated buckle on seabed from the FE model. This is determined based on the thermal feed-in capacity as a function of buckle strain.

    Once a pipeline has developed compressive buckle globally, there is a release of the axial effective force within the buckle and, as a result, the resultant axial expansion feeds into the buckle. The capacity of a buckle is defined as the amount of thermal expansion which can be accommodated before the buckle is overloaded i.e. exceed the allowable strain.

    A short FE model is utilised to determine the buckle capacity. Some additional attributes are incorporated into the model to capture the stress/strain at buckle apex accurately. These include:

    1. The use of non-linear material properties with de-rated equivalent SMYS as shown in Figure 3.1 previously.

    2. The model is progressively heated up, the post-buckle pipe feed-in and the corresponding compressive strains at the apex are extracted at each step.

    The results for LB, BE and UB lateral fictions are shown in

    Figure 6-6, as buckle strain vs feed-in. At 1% design strain, it can be seen that the buckle has feed-in capacity in excess of 5m, which is greater than the total pipe end expansions. This observation suggested that the current pipe properties have large capacity to accommodate an uncontrolled rouge buckle.

    Figure 6-6 Feed-In Capacity of Unplanned Buckle on Seabed

    Page 44 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Using the same 3km pipeline model, a buckle is artificially initiated with a large horizontal imperfection to determine the actual pipe feed-in and strain at design temperature and pressure. However, due to the relatively low axial friction, the buckle can only be initiated in the FE model with BE axial and lateral frictions. The pipe effective force is insufficient to initiate buckle at LB axial or UB lateral frictions, although this friction combination is more onerous to the buckle.

    The predicted feed-in and corresponding strain is indicated in Figure 6-6 as a blue solid dot. The 2.3m feed-in and 0.2% strain are well below the allowable of 5m and 1%, respectively, and hence confirmed the pipe structural integrity even in the event of an uncontrolled buckle on seabed.

    6.5 Fatigue Damage Assessments

    A preliminary fatigue assessment is performed to determine the damage at the rouge buckle due to cyclic temperature and pressure loads. To this, the, the transient temperature and pressure profiles (Figures 3.3 and 3.4) are used for one (1) full shutdown and restart cycle. This is conservative since the stress at the buckle generally reduces with load cycles due to shakedown effects until the buckle profile has stabilised.

    The 3km length FE model discussed previously with BE frictions is restarted with transient load steps. Maximum longitudinal stresses are extracted at pipe mid-wall and plotted against cyclic load steps as shown in figure 6-7 below. The stress range between full start-up and shut-down cycle, under maximum operating condition, is about 24MPa as highlighted in the figure.

    Figure 6-7 Cyclic Stress Range at Buckle

    24 MPa

    Page 45 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Using the fatigue S/N curve and methodology presented in Section 4.2, fatigue damage calculation is performed for 360 cycles (assume 12 cycles per year for 30-year design life) and the results are summarised in table 6.1 and 6.2 below at weld root and weld toe, respectively. It can be seen that the fatigue damage Unity Checks are very low at both weld locations.

    Table 6.1 Low Cycle Fatigue Damage at Weld Root (F1 Curve)

    No. Cycle/Year

    alnor min alnorloc SCF min =

    Log (N) N Damage Ratio

    12 24 24.00 7.16 1.440E+07 8.33E-07

    Damage per year 8.33E-07

    Damage over 30 years design life 2.50E-05

    Fatigue knockdown factor (1) 2

    Allowable usage factor (2) 0.08

    Fatigue Unity Check 6.25E-04

    Table 6.2 Low Cycle Fatigue Damage at Weld Toe (D Curve)

    No. Cycle/Year

    alnor min alnorloc SCF min =

    Log (N) N Damage Ratio

    12 24 37.20 (3) 6.59 3.87E+06 3.10E-06

    Damage per year 3.10E-06

    Damage over 30 years design life 9.31E-05

    Fatigue knockdown factor 1

    Allowable usage factor (2) 0.08

    Fatigue Unity Check 1.16E-03

    Notes (1): To account for internal cladding [7].

    (2): Assume 20% fatigue budget for installation.

    (3): SCF due to field joint misalignment included at weld toe, refer Table 4.3.

    Page 46 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    6.6 Planned Buckle to Reduce End Expansion

    It has been shown in Section 6.3 and 6.4 that the 12in Abadi flowlines are not susceptible to uncontrolled global buckling and hence buckle mitigation for the purpose of stress relieving is not deemed necessary. However, due to the high operating temperature and pressure, pipe end expansion was found to be in excess of 2.5m, which is beyond the limit of the current tie-in spool capacities. To reduce expansion, a buckle mitigation scheme can be considered, which directly introduces additional virtual anchor points into the pipeline. The distance between the first or last virtual anchor points at the respective pipe ends effectively controlled the end expansions. In other words, the closer the buckle location to the pipe end, the less expansion it will see. However, there is a limit to the buckle position as too close to the end, there may not be sufficient effective force to initiate the buckle.

    To demonstrate this concept of using buckles to limit the pipe end expansion, expansion calculations are performed for one (1) and two (2) buckle cases. For simplicity, the one buckle case assumes the buckle is located at KP 1.5 while the two buckle case at KP 1 and KP 2.

    The post buckle effective forces for the two cases are shown in in Figure 6-8 and Figure 6-8 overleaf. It can be seen that, the virtual anchor length reduces to about 1km and 0.75km for the one and two buckle cases, respectively, at LB axial friction from the original unmitigated value of 1.5km as shown in Figure 6-1.

    The corresponding pipe end expansions are summarised in Table 6.3 overleaf, which shown that one buckle scheme can reduce the maximum end expansion to 1.9m while the two buckle scheme to 1.6m at hot end, from the unmitigated value of about 2.5m. Preliminary spool calculation indicated that 1.6m is close to the anticipated spool capacity and this implies that a two buckle scheme is necessary. The two buckle scheme is also likely to be the maximum that can be accommodated in the present flowline due to low effective axial force to initiate more buckles.

    It should be emphasized that the above results are indicative only and preliminary in nature. The actual location of the buckles and the resultant end expansions have to be quantified with more detailed analyses, incorporating buckle initiators (e.g. ZRB) and pipe lay process into the analysis model. Due to the critical nature of these buckles to the structure integrity of the tie-in spool, sensitivity analysis should also be performed to confirm the buckle reliability.

    Page 47 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Figure 6-7 Effective Axial Force of One (1) Planned Buckle on Seabed

    Figure 6-8 Effective Axial Force of Two (2) Planned Buckles on Seabed

    Page 48 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    Table 6.3 Reduced Expansion due to Planned Buckles

    No of Planned Buckle Ends Expansion (m)

    LB BE UB

    1 (KP 1.5) Hot End (MNFL) 1.88 1.71 1.70

    Cold End (Riser) 1.63 1.48 1.46

    2 (KP 1 and KP 2) Hot End (MNFL) 1.55 1.38 1.21

    Cold End (Riser) 1.33 1.17 1.02

    Page 49 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    7.0 REFERENCES

    [1] WGK Document No. ASF01-012-SE-SBD-0001; URF AND STRUCTURE DESIGN PREMISE; Revision A; January 2013.

    [2] WGK Document No. ASF01-012-SE-REP-0001; TECHNICAL NOTE PRELIMINARY SURF FLOWLINE WALL THICKNESS ASSESSMENT; Revision B; February 2013.

    [3] WGK Document No. ASF01-015-SE-TCN-0002; TECHNICAL NOTE - SURF PIPE SOIL INTERACTION DATA; Revision B; February 2013.

    [4] SAFEBUCK III JIP Document No. 5087471/01/A, Safe Design of Pipeline with Lateral Buckling Design Guideline, 30th August 2011.

    [5] DEP 31.40.00.10-Gen; Pipeline Engineering (Amendments/Supplements to ISO 13626); February 2011.

    [6] DEP 31.40.20.32-Gen; CRA-Clad and CRA-Lined Steel Pipe (Amendments/Supplements to API Spec 5LD); Sept. 2011.

    [7] Technical Report; JIP Lined and Clad Pipeline Materials; Phase 2; Guideline for Design and Construction of Clad and Lined Pipelines; Report No. 2007-0220; Rev 02; DNV; 10th November 2008.

    [8] Det Norske Veritas; DNV-OS-F101; Submarine Pipeline Systems; October 2011.

    [9] Det Norske Veritas; DNV-RP-C203; Fatigue Design of Offshore Structures; April 2008.

    [10] Note: Definition of Weak Section to Assess Strain Localisation in Seamless Pipes; by Ralf Peek (SIEP EPT-PNR) and George W. Brown (SUKEP-EPE-T-PC); 2009.

    [11] SIEPP-EP2005-5154; Design Specification for a Clad Pipelines Subject to Lateral Buckling on a Flat Soft Clay Seabed; April 2005.

    [12] OMAE 2009-79779; Strain Intensification Due To Material Discontinuity At Field Joints Adjacent To Thick Wall Insulation Coating.

    [13] COMPANY Doc. No. A0510-010-FA-TCN-0001, Flow Assurance for Commencement of FEED, Rev 0, December 2012.

    [14] A0510-010-SE-BOD-0001; Abadi Gas Field Development; SURF Basis of Design.

    Page 50 of 50 ASF01-012-SE-REP-0005 Rev. B

  • Report - SURF Flowline Global Buckling Analysis Doc. No: ASF01-012-SE-REP-0005 Rev. B

    APPENDIX A EQUIVALENT PROPERTIES CALCULATION

    ASF01-012-SE-REP-0005 Rev. B

  • INPEX ABADI GAS DEVELOPEMENT PROJECT Equivalent Properties for Lateral Buckling Analysis

    BY:CHK:

    DRST

    Density kg/m3 7850 8440Thickness mm 17.5 3

    Thermal Expansion Coefficient at 20C /

    oC 1.17E-05 1.28E-05

    SMYS at 20C MPa 450 414SMTS at 20C MPa 535 827

    Elastis Modulus 20C GPa 207 205

    Parameter Unit CS ISO- 3183 L450 / API 5L X65 Grade Alloy 625

    MDT 136C:= Input the maximum design temperature !

    ma 8 109

    C 1:=

    C KMDT ma MDT 20C( )+:=

    MDT 12.6 106

    13.7 10 6( )=y SMYS MPa:=

    yMDT y MDT 50Cif

    y MDT 50 C( ) 0.6 MPa C1

    50C MDT< 100Cif

    y 30 MPa MDT 100 C( ) 0.4 MPa C1

    + otherwise

    :=

    yMDT 405.6 369.6( ) MPa=

    T SMTS MPa:=

    TMDT T MDT 50Cif

    T MDT 50 C( ) 0.6 MPa C1

    50C MDT< 100Cif

    T 30 MPa MDT 100 C( ) 0.4 MPa C1

    + otherwise

    :=

    TMDT 490.6 782.6( ) MPa=

    E E GPa:=

    EMDT E MDT 20Cif

    E MDT 20C( ) C 1 0.05 GPa 20C MDT< 100C( )if

    E MDT 20C( ) C 1 0.06 GPa 100C MDT< 200C( )if otherwise

    :=

    EMDT 200.04 198.04( ) GPa=

    1 of 2 ASF01-012-SE-REP-0005

  • For input into ABAQUS or MathCAD Calculation that requires equivalent (Claded X65 Pipe) Properties

    Density kg/m3 7850 8440 7936Thickness mm 17.5 3.0 20.5Thermal

    Expansion Coefficient at

    MDT

    /oC 1.26E-05 1.37E-05 1.28E-05

    SMYS at MDT MPa 405.6 369.6 400.3SMTS at MDT MPa 490.6 782.6 533.3

    Elastis Modulus at MDT GPa

    200 198 199.75

    Parameter Unit CS ISO- 3183 L450 / API 5L Alloy 625

    Claded Pipe

    (equivalent)

    2 of 2 ASF01-012-SE-REP-0005

    Abbreviations / Nomenclature1.0 Introduction1.1 Abadi Gas Field Development1.2 Scope

    2.0 summary, conclusions and recommendations2.1 Summary2.2 Conclusions2.2.1 Lateral Buckling Susceptibility Assessment2.2.2 Consequence on Unplanned Buckle2.2.3 Reduced Expansions from Planned Buckle

    2.3 Recommendation

    3.0 Design Data3.1 Classification of Location and Safety Classes3.2 Mechanical Properties3.3 Operational Data3.4 Material Properties3.5 Flowline Equivalent Material Data for Bi-Metal3.6 Steady State Design Profiles3.7 Transient Profiles3.8 Environmental Data

    Table 3.7 Environmental Data3.9 Seabed Profile3.10 Pipe Soil Properties

    Table 3.8 Axial pipe-soil friction coefficients (Operation)Table 3.9 Lateral pipe-soil friction coefficient monotonic loading (Operation)/4.0 Lateral BUckling ACCEPTANCE CRITERIA4.1 Allowable Strain Criteria4.1.1 Local Buckling4.1.2 Material Uniform Strain Capacity4.1.3 Engineering Critical Assessment Limiting Strain4.1.4 ALLOWABLE STRAIN LIMIT STATE SUMMARY

    4.2 Fatigue Limit State4.3 Low Cycle Fatigue4.4 Limit State Check Summary

    5.0 BUCKLING DESIGN METHODOLOGY5.1 General5.2 Pipeline Effective Axial Force5.3 Lateral buckling Susceptibility5.4 Consequence of single isolated buckle5.4.1 Material Strength Mismatches5.4.2 Wall Thickness Variations5.4.3 Coating Cutbacks

    5.5 Lateral Buckling Design Approach

    6.0 ResultS and discussionS6.1 General6.2 Effective Axial Force and Expansion Analysis6.3 Lateral Buckling Susceptibility6.3.1 General6.3.2 Critical Buckling Force Due to Seabed Terrain6.3.3 SAFEBUCK Hobbs Equation6.3.4 Susceptibility of Lateral Buckling6.3.4.1 General6.3.4.2 Deterministic Assessment6.3.4.3 Probabilistic Assessment6.3.5 Lateral Buckling Susceptibility Summary

    6.4 Consequence of Unplanned Buckling6.5 Fatigue Damage Assessments6.6 Planned Buckle to Reduce End Expansion

    7.0 ReferencesAPPENDIX A eQUIVALENT PROPERTIES CALCULATION