DNVGL-RP-F110 Global buckling of submarine pipelines · Recommended practice — DNVGL-RP-F110....

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RECOMMENDED PRACTICE

DNVGL-RP-F110 Edition April 2018

Global buckling of submarine pipelines

FOREWORD

DNV GL recommended practices contain sound engineering practice and guidance.

© DNV GL AS April 2018

Any comments may be sent by e-mail to [email protected]

This service document has been prepared based on available knowledge, technology and/or information at the time of issuance of thisdocument. The use of this document by others than DNV GL is at the user's sole risk. DNV GL does not accept any liability or responsibilityfor loss or damages resulting from any use of this document.

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Recommended practice — DNVGL-RP-F110. Edition April 2018 Global buckling of submarine pipelines

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CHANGES – CURRENT

This document supersedes the August 2017 edition of DNVGL-RP-F110.Changes in this document are highlighted in red colour. However, if the changes involve a whole chapter,section or sub-section, normally only the title will be in red colour.

Changes April 2018

• Front pageThe title has been changed as the application of this recommended practice has been extended to alsoinclude moderately axially loaded pipelines.

• Sec.1 GeneralMinor modifications and updates have been made.

• Sec.2 Introduction to global bucklingUpdates have been made due to merger with the Safebuck guideline.

• Sec.3 Basis for designMinor modifications and updates have been made and description of uncertainties has been removed.

• Sec.4 Pipe-soil interactionMinor updates have been made and detailed soil description has been moved to DNVGL-RP-F114.

• Sec.5 Load effect calculationUpdates have been made due to merger with the Safebuck guideline.

• Sec.6 Exposed pipelineOld Sec.6 and Sec.7 have been merged and significantly modified due to merger with the Safebuck guideline.

• Sec.7 Buried pipelineMinor modifications and updates have been made.

• Sec.8 Limit state criteriaThis is a new section which is partly replacing old Sec.10.

• Sec.9 Condition load effect factor for exposed pipelinesNo significant changes have been made.

• Sec.10 Operational structural integrityThis is a new section.

• App.A Mitigation measures for exposed pipeline (informative)Updates have been made due to merger with the Safebuck guideline.

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• App.B Probabilistic buckle formation (informative)Old App.B has been removed and replaced with new content.

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• App.C Effect of environment on fatigue and fractureOld App.C has been removed and replaced with new content.

Editorial correctionsIn addition to the above stated changes, editorial corrections may have been made.

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AcknowledgementsThis revision of this recommended practice is developed based on results from the Safebuck joint industryproject (JIP).The following companies, listed in alphabetical order, are acknowledged for their contributions to the JIP.

ABS Allseas BP Boemre

Bureau Veritas Cambridge University Cathie Associates Chevron

ConocoPhillips DNV GL ENI ExxonMobil

Fugro Inpex JFE OTM

Oxford University Petrobras Saipem Shell

Statoil Subsea7 Technip Tenaris

Total TWI Uni of Western Australia Votadini Consultants

Woodside

The following companies, listed in alphabetical order, are acknowledged for their contributions to previousrevisions of this recommended practice.

BP DNV GL ENI Norge Hydro

Norsk Agip Shell Snamprogetti Statoil

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CONTENTS

Changes – current.................................................................................................. 3Acknowledgements.................................................................................6

Section 1 Introduction.......................................................................................... 101.1 General........................................................................................... 101.2 Objective.........................................................................................101.3 Scope.............................................................................................. 111.4 Application...................................................................................... 111.5 Relationship to other standards......................................................111.6 Definitions.......................................................................................131.7 Structure of the document..............................................................21

Section 2 Introduction to global buckling............................................................. 222.1 General........................................................................................... 222.2 Global buckling of exposed pipelines..............................................222.3 Pipeline walking............................................................................. 282.4 Upheavel buckling of buried pipeline..............................................29

Section 3 Basis for design.....................................................................................303.1 General........................................................................................... 303.2 Pipe geometry.................................................................................303.3 Pipeline material.............................................................................313.4 Loads.............................................................................................. 323.5 Time effects.................................................................................... 34

Section 4 Pipe-soil interaction.............................................................................. 364.1 General........................................................................................... 364.2 Exposed pipeline.............................................................................374.3 Buried pipelines.............................................................................. 40

Section 5 Load effect calculation.......................................................................... 415.1 General........................................................................................... 415.2 Analysis loading sequence..............................................................415.3 Analytical methods......................................................................... 425.4 Detailed finite element analysis......................................................43

Section 6 Exposed pipeline................................................................................... 476.1 Objective and applicability..............................................................47

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6.2 Lateral buckling design procedure..................................................476.3 Suspectibility to buckling................................................................506.4 Tolerable virtual anchor spacing calculation...................................546.5 Calculation of characteristic virtual anchor spacing........................ 556.6 Relevant limit states - exposed pipelines....................................... 596.7 Additional considerations for uneven seabed..................................62

Section 7 Buried pipeline...................................................................................... 667.1 General........................................................................................... 667.2 Upheaval buckling design procedure.............................................. 667.3 Effective axial force........................................................................ 687.4 Analytical methods for up-lift assessment......................................697.5 Detailed finite element analyses.....................................................697.6 Upheaval buckling soil limit states................................................. 717.7 Relevant pipeline limit states - buried pipeline...............................76

Section 8 Limit state criteria.................................................................................778.1 General........................................................................................... 778.2 Local buckling limit state - combined loading.................................778.3 Axial loading limit state..................................................................798.4 Uniform strain capacity.................................................................. 798.5 Cyclic plasticity limit state..............................................................808.6 Fatigue and fracture....................................................................... 82

Section 9 Condition load effect factor for exposed pipelines.................................859.1 Basic principles...............................................................................85

9.2 Calculation of CoV(XA) due to axial soil resistance......................... 85

9.3 Calculation of CoV(XL) due to lateral soil friction............................87

9.4 Calculation of CoV(XB) due to stress-strain relationship................. 87

9.5 Calculation of CoV(Xc) due to trawl loads.......................................889.6 CoV for parameters with large variation and non-symmetricalupper and lower estimates...................................................................89

Section 10 Operational structural integrity...........................................................9110.1 General......................................................................................... 9110.2 Surveys of pipeline out-of-straightness and bucklingbehaviour..............................................................................................9110.3 Operating load condition monitoring............................................ 9210.4 Pipeline structural integrity assessment.......................................9210.5 Survey frequency..........................................................................93

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10.6 Re-qualification.............................................................................93

Section 11 Documentation for operation...............................................................9411.1 General......................................................................................... 94

Section 12 References...........................................................................................9512.1 General......................................................................................... 95

Appendix A Mitigation measures for exposed pipelines (informative)...................96A.1 General........................................................................................... 96A.2 Prevent development of buckling................................................... 96A.3 Triggering buckling.........................................................................98A.4 References.................................................................................... 103

Appendix B Probabilistic buckle formation (informative)....................................104B.1 Probabilistic buckle formation model............................................104B.2 Probabilistic results and interpretation........................................ 113B.3 Finite element assessment........................................................... 115B.4 References.................................................................................... 116

Appendix C Effect of environment on fatigue and fracture..................................117C.1 Environment and loading frequency............................................. 117C.2 Corrosion resistance alloys...........................................................118C.3 Fatigue testing in a corrosive environment...................................119C.4 Fracture toughness testing in a sour, corrosive environment........121C.5 References.................................................................................... 123

Appendix D Example of calculation of tolerable and characteristic virtualanchor spacing................................................................................................... 124

D.1 Tolerable virtual anchor spacing.................................................. 124D.2 Characteristic virtual anchor spacing........................................... 126

Changes – historic.............................................................................................. 128


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SECTION 1 INTRODUCTION

1.1 GeneralGlobal buckling of a pipeline implies buckling of a section of the pipeline as a column in compression. Theglobal buckling may appear either downwards in free spans, horizontally as lateral buckling on the seabed orvertically such as upheaval buckling of buried pipelines or upwards on a crest of unburied/exposed pipelinesnormally followed by a lateral turn-down. Local buckling on the other hand is a gross deformation of the pipecross section due to too high combined loading.Global buckling is governed by compressive effective axial force. Pipelines exposed to global buckling areeither those with high effective axial compressive forces, or those with low buckling capacity, typically lightweight pipelines with low lateral pipe-soil resistance.The structural integrity of a pipeline susceptible to global buckling can be assured by two different designconcepts:

— restraining the pipeline with the high compressive forces (buried pipelines), or— allowing the pipeline to buckle globally and thereby releasing the expansion forces (exposed/unburied

pipelines).

The final selection of design concept depends on several factors. This recommended practice will in depthdescribe both design concepts.Another potential global response of pipelines exposed to high pressures and high temperatures is pipelinewalking which is accumulated end expansion with shut-down and start-up cycles. Pipeline walking is apotential challenge for short flowlines and may cause end structures, connected risers and/or spools/jumpersto be over-utilised and fail if not mitigated.Global buckling is a structural response to a high compressive effective axial force and not a failure mode assuch. Global buckling may, however, imply failure modes such as:

— local buckling— fracture and fatigue— excessive displacements.

This recommended practice merges guidance from the HOTPIPE and SAFEBUCK JIPs and presents twoalternative local buckling limit states for unburied pipelines. The intention of the merger of the guidelines hasbeen to reduce cost for the industry and increase predictability without compromising quality and safety.The design procedures and criteria in this recommended practice are based on significant research and a vastnumber of numerical simulations. Despite of this, some non-typical pipeline and loading scenarios may giveunexpected results, and the user should check the results carefully.

1.2 ObjectiveThis recommended practice is to provide design methodology and criteria to fulfil the functional requirementsregarding global buckling in DNVGL-ST-F101. In case of a pipeline laying unburied on the seabed, therecommended practice gives design methodology and criteria to allow the pipeline to buckle in a safe andcontrolled manner. For a buried pipeline, the recommended practice ensures that the pipeline stays in placewithout any upheaval or lateral buckling.

Guidance note:A software, SimBuck, is offered by DNV GL Software (https://www.dnvgl.com/software/), as an optional tool to perform globalbuckling design according to this recommended practice.

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1.3 ScopeThis recommended practice gives guidance on design for two global buckling scenarios for high temperatureand high pressure pipelines:

1) Unburied pipelines on even and uneven seabed where global buckling occurs either fully in the horizontalplane, or first in the vertical plane as feed-in and uplift and subsequently in the horizontal plane.

2) Buried pipelines where global buckling occurs in the vertical plane, so called upheaval buckling. Onlycriteria for avoiding upheaval buckling are given in this recommended practice.

Detailed design procedures and criteria are given for global buckling of unburied pipelines on even anduneven seabed and to avoid upheaval buckling of buried pipelines. The methodology and criteria are valid foroffshore rigid steel pipelines.The recommended practice also contains examples of mitigation measures, some examples on probabilisticbuckle formation calculations and guidance on effect of sour service.

1.4 ApplicationThis recommended practice applies to the structural design of offshore rigid steel pipelines susceptible toglobal buckling, see Table 1-1. This is most relevant for so-called high temperature high pressure pipelinesbut may also be relevant for moderately axially loaded pipelines, such as pipelines with a low submergedweight depending on the global buckling resistance.The trawl interference evaluation in this recommended practice is limited to lateral buckling only. Hence,trawling in free span is not covered by this recommended practice, see DNVGL-RP-F111 .

Table 1-1 Applicability of this recommended practice

Aspect Description

pipeline response laterial buckling, upheaval buckling and walking

pipeline exposure buried and exposed on seabed (unburied)

pipeline size calibrated for 10" og 42" pipelines, also considered valid for 6" and 8" pipelines

materials all kind of rigid steel pipes, parts of this document such as load effect calculations may also beapplied to flexible pipes and plastic/composite pipes while at least failure modes and acceptancecriteria need to be re-defined

buckling initiation initiation due to imperfections both artificial and natural (uneven seabed) and trawl interference

pressure andtemperature

no limitations to pressure or temperature if temperature effects on materials, insulation, thermalconditions and flow assurance are considered in the design (such considerations are outside thescope for this document)

1.5 Relationship to other standardsIn the context of this document, the term standard shall be understood to cover document types such asstandards, guidelines and recommended practices in addition to bona fide standards.This recommended practice complies with DNVGL-ST-F101 and complements the functional requirementson global buckling with specific design procedures and criteria. The design procedures and criteria in thisrecommended practice have been determined by sound engineering judgement in combination with structuralreliability methods and target failure probability in compliance with DNVGL-ST-F101.Global buckling is the structural response to a too high compressive effective axial load. This maysubsequently cause different failures modes. In this recommended practice, only failure modes directly


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arising from global buckling will be considered. For other structural failure modes, the following referencesapply:

— general - DNVGL-ST-F101— vortex induced vibration - DNVGL-RP-F105— on-bottom stability - DNVGL-RP-F109— trawling interference in free spans - DNVGL-RP-F111 (see also [3.4.3]).

Referenced relevant standards are listed in this section, see Table 1-2, while other references are given inSec.12.In case of conflict between requirements of this recommended practice and a referenced DNV GL standard,the requirements of the standard with the latest revision date shall prevail, any conflict is intended to beremoved in next revision of these documents.In case of conflict between requirements of this recommended practice and a non-DNV GL referenceddocument, the requirements of this document shall prevail.For undated references, the edition valid at the time of publishing this document applies.

Table 1-2 Referenced standards

Document code Title

DNV GL service specifications

DNVGL-SE-0475 Verification and certification of submarine pipelines

DNVGL-SE-0476 Offshore riser systems

DNVGL-SE-0160 Technology qualification management and verification

DNV GL standards

DNVGL-ST-F101 Submarine pipeline systems

DNVGL-ST-F201 Dynamic risers

DNV GL recommended practices

DNVGL-RP-A203 Technology qualification

DNVGL-RP-C203 Fatigue design of offshore steel structures

DNVGL-RP-C205 Environmental conditions and environmental loads

DNVGL-RP-F105 Free spanning pipelines

DNVGL-RP-F108 Assessment of flaws in pipeline and riser girth welds

DNVGL-RP-F109 On-bottom stability design of submarine pipelines

DNVGL-RP-F111 Interference between trawl gear and pipelines

DNVGL-RP-F114 Pipe-soil interaction for submarine pipelines

DNVGL-RP-F116 Integrity management of submarine pipeline systems

Other standards

BS 7910 (2013) Guide to methods for assessing the acceptability of flaws in metallic structures


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1.6 DefinitionsFor general definitions, see Table 1-3. Abbreviations used in this document is listed and explained in Table1-4, while Greek and Latin symbols are listed in Table 1-5 and Table 1-6, respectively. For definition of verbalforms, see Table 1-7.

Table 1-3 Definitions of terms

Term Definition

2D analysis pipeline analyses with all degrees of freedom (i.e. 3D) but with the initial pipeline geometrymodelled in one plane only

2½D analysis pipeline analyses with all degrees of freedom (i.e. 3D) both modelled and analysed in threedimensions, i.e. including curves in the vertical horizontal plane, but with the seabed in thelateral direction modelled flat

3D analysis pipeline analysis, modelled and analysed in three dimensions giving possible benefit forcurves in the horizontal plane, and where the seabed is realistically modelled with sidewaysslopes, etc.

anchor length length along the pipeline where axial sliding occurs, the length it takes to build up theeffective axial force from zero to the fully restrained value, depending on axial pipe-soilresistance

associated load used to define other types of loads that are associated with an extreme load in a loadcombination, e.g. for an environmental loading scenario with a 100-year storm, the extremeenvironmental load is not combined with the extreme design temperature and pressure, butwith associated (normal operating) temperature and pressures

anchor point locations along the pipeline route, often in between buckles, where the pipeline does notmove axially (feed-in) towards a buckle or pipeline end

best estimate (BE) the best estimate of a stochastic variable, typically the mean value

characteristic value value of design parameter used in design criterion, typical an upper fractile for load effectsand a mean or lower fractile for resistances

cover added material, e.g. gravel or seabed material, either in trench or un-trenched on flat bottom

design value from a criterion point of view it normally implies that the required partial safety factorsare included, e.g. design load (characteristic loads times the load effect factors), designresistance (characteristic resistance divided by the relevant partial safety factors) et cetera

effective axial force combined axial action of the stress in the pipe wall and force effects of internal and externalpressure (end cap effects)

exposed pipeline pipeline that is resting at the seabed, not trenched and buried, parts may be intervened byspot rock dumping

feed-in axial expansion into an area when the resisting force has been reduced, e.g. the release ofthe stored energy in a pipeline into a buckle

functional load designcase

extreme functional load effects with associated interference and environmental loads effects

global buckling on-set of transverse instability for a significant length of pipeline, either in the vertical planeas upheaval buckling or in the horizontal plane as lateral buckling


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interference load designcase

extreme interference load effects with associated functional and environmental load effects

lateral buckling global buckling in the horizontal plane

limit state state beyond which the structure no longer satisfies the requirements, often referred to asfailure mode

load effect response in the pipeline (axial force, moments, strain, et cetera) due to the applied loads

local maximum designtemperature

the temperature at a specific location along the pipeline corresponding to the maximumdesign temperature profile

long pipeline pipeline longer than the double anchor length, i.e. the fully restrained effective axial force isat least theoretically reached if the pipeline does not buckle

lower estimate (LE) the lower estimate of a stochastic variable, typically the mean minus two standard deviations

pipeline walking the accumulation of axial expansion at pipeline ends and in buckles due to load cycles

post global buckling development of the pipe configuration after the initial buckling

propped shape the configuration of a pipeline given by the relevant self-weight and pipe stiffness when liftedfrom a horizontal plane at a certain height δ

ratcheting normally defined as accumulation of cross-sectional plastic/irreversible deformation or ovalitywith load cycles, here also used in a broader sense about accumulated deformation of endexpansion or lateral displacements in buckles with cyclic loading

re-qualification re-assessment of a design due to modified design premises and/or sustained damage

safety class concept adopted to classify the significance of the pipeline system with respect to theconsequences of failure

short pipeline pipeline shorter than the double anchor length, i.e. the fully restrained effective axial force isnot reached

sleeper simple structure, such as a pipe joint or similar, used to create an imperfection that triggerslateral buckling

slip zone section of the pipeline where axial sliding (feed-in) towards a buckle or pipeline end occurs

S-N approach assessment of fatigue using stress(S)-number of cycles to failure(N) capacity curves

snake-lay pipeline installed with regular lateral imperfections, i.e. route curves, to trigger globalbuckling

snaking the post-buckling configuration of lateral buckling

triggers imperfections that may trigger global buckling, may be natural ones like uneven seabed orengineered ones like rock berms, sleepers, route curves et cetera

upheaval buckling the consecutive deformation in the vertical plane, occurring for buried pipelines, or forunburied pipelines prior to developing lateral buckling

upper estimate (UE) the upper estimate of a stochastic variable, typically the mean plus two standard deviations


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Table 1-4 Abbreviations

Abbreviation Description

ALS accidental limit states

BE best estimate

CoV coefficient of variance for a stochastic variable, CoV(X) = standard deviation(X)/mean (X)

CRA corrosion resistant alloy

CT compact tension

DFF design fatigue factor

ECA engineering criticality assessment (fracture mechanics based assessment of cracks and crack growth)

FCGR fracture crack growth resistance

FE finite element

FEA finite element analysis

FL fusion line

HPHT high pressure and/or high temperature and in this document linked to the susceptibility to globalbuckling rather than to defined limits for the pressure and temperature

KP kilometre post

LE lower estimate

LRFD load and resistance factor design

n/a not applicable

NDT non-destructive testing

OOS out-of-straightness

PDF probability density function

PLEM pipeline end manifold

PLET pipeline end termination

PP parent pipe

PSI pipe soil interaction

ROV remotely operated vehicle

SCR steel catenary riser

SENB single edge notch bending (type of test)

SENT single edge notch tension (type of test)

SLS serviceability limit states

SMTS specified minimum tensile strength at room temperature

SMYS specified minimum yield stress at room temperature

SNCF strain concentration factor


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UC unity check

UE upper estimate

UHB upheaval buckling

ULS ultimate limit states

UOE pipe fabrication process for longitudinally welded pipes, formed as U, then O, welded and thenexpanded

VAS virtual anchor spacing

WM weld metal

TS tensile strength at room temperature

YS yield stress at room temperature

pdf probability density function

Table 1-5 Definition of Latin symbols and characters

Latin symbol Definition

A pipe cross-sectional area

Ai pipe internal bore area

As pipe cross-sectional steel area

D nominal outer diameter of steel pipe

Do nominal outer diameter of steel pipe

E Young's modulus

E[x] expected value for x

FD maximum hydrodynamic drag force per unit length

FL maximum hydrodynamic lift force per unit length

FP trawl pull-over load

FT characteristic trawl pullover load

FTLE/BE/UE lower/best/upper estimate of trawl pullover load

H cover height from top of pipe to soil surface

H1, H2 heights of different soil layers, H = H1 + H2

Hlay residual laying tension

I second moment of area

L buckle length, and length between buckle and virtual anchor point

LUplift length of pipeline uplifted at crest

L0 imperfection length for a propped shape, from apex to touch down

M bending moment


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Mf functional bending moment

MSd design bending moment

MBE/1/2.. bending moments for different sensitivity studies

Mp plastic bending moments capacity

N true axial force (pipe wall force)

R radius of imperfection, and uplift resistance

Rc uplift resistance of cover

RLE/BE/UE radius of imperfection, lower/best/upper estimate

Rmin required resistance on buried pipelines to account for imperfections not detected due to insufficientaccuracy of the survey equipment

Rreq required resistance on buried pipelines

Rspec specific resistance of the cover required to avoid upheaval buckling based on a specific measuredimperfection

S effective axial force

S∞ critical effective axial force for the infinite buckling mode

S0 fully restrained effective axial load

SBH effective (axial) force in the buckle

ŜCR critical buckling force

upper estimate of critical buckling force

ŜCRT critical buckling force associated with trigger

ŜCRU critical buckling force associated with an unplanned buckle

lower estimate of critical buckling force associated with an unplanned buckle

Sinit critical value of effective axial force that triggers global buckling

SLECurve lower estimate critical buckling load in horizontal route curve

Sp plastic effective axial force capacity

SR effective axial force in uplifted free span

SPost post-buckling force (tension is positive)

lower estimate post-buckling force (tension is positive)

upper estimate of the post-buckle force in the planned buckle

SSd design effective axial force

S∞ critical buckling force for infinite mode

T temperature of content


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Td design temperature

T(k) temperature at failure for downward soil stiffness k

Tmax maximum design temperature

Tl,max local maximum design temperature

TRd design (resistance equivalent) failure temperature

TSd design (load) equivalent temperature

VAS virtual anchor spacing

VAScharacteristic characteristic virtual anchor spacing

VASLength virtual anchor spacing limited by pipeline length

VASSo virtual anchor spacing limited by fully restrained effective axial force

VASSharing virtual anchor spacing limited by sharing criterion

VASTolerable tolerable virtual anchor spacing

XA uncertainty in load effect due to uncertainties in axial pipe-soil resistance

XB uncertainty in load effect due to uncertainties in applied stress-strain curve

XC uncertainty in load effect due to uncertainties in applied trawl load

XL uncertainty in load effect due to uncertainties in lateral pipe-soil resistance

XF Uncertainty in load effect due to uncertainties in input parameters

Xtrigger buckle spacing

ai Taylor series expansion coefficient

fALE/BE/UE axial pipe-soil resistance force, lower/best/upper estimate

fLLE/BE/UE lateral pipe-soil resistance force, lower/best/upper estimate

fT annual trawling frequency per relevant pipeline section

fu characteristic tensile strength

fu,temp tensile strength derating factor

fy characteristic yield strength

fy,temp yield strength derating factor

kLE/BE stiffness in clay, downwards, lower/best estimate

n number of independent surveys, and strain hardening factor

p pressure

Pb burst pressure

Pe external pressure

Pli local incidental pressure (internal)


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ri ratio between failure temperatures for best estimate and lower estimate downward stiffness

s bend arc length in kilometres

t pipe nominal wall thickness, and time in years

t1 pipe minimum wall thickness adjusted for relevant corrosion allowance

t2 pipe nominal wall thickness adjusted for relevant corrosion allowance

tcorr corrosion allowance

tfab pipe wall thickness fabrication mill tolerance

w submerged pipe weight

wB submerged pipe weight at a buoyancy trigger

wo submerged pipe weight during installation

wp submerged pipe weight during operation

Xtrigger buckle spacing

xi basic parameters in Taylor's series expansion

Table 1-6 Definition of Greek symbols

Greek symbol Definition

∆pd differential design pressure

∆pi internal pressure difference relative to laying

∆T differential temperature

α thermal expansion coefficient

αB Bauschinger factor

αC hardening effect factor

αp pressure effect factor

αU material strength factor

αLuder material strength factor for Lüder plateau

αh strain hardening material factor

δ imperfection height of a propped shape imperfection, and feed-in to buckle

δe failure displacement

δF propped shape imperfection height for minimum cover

εc characteristic strain capacity

εca characteristic axial strain capacity


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εf functional strain load effect

equivalent plastic strain

longitudinal/hoop/radial plastic strain

εeq equivalent total strain

εf functional bending strain

εSd design bending strain

εUS characteristic uniform strain capacity

ε1 basic strain capacity

axial strain resistance factor

condition load effect factor

load effect factor for environmental loads

load effect factor for functional loads

uniform strain safety factor

strain resistance factor

safety factor on lift force

material resistance factor

safety class resistance factor

load factor for temperature for use in upheaval buckling

uplift resistance safety factor

v Poisson's ratio

рij coefficient of correlation between parameter i and parameter j

σ standard deviation, statistical

σconfiguration standard deviation on the configuration measurement accuracy

σconsol consolidation stress in clay due to a weight of the soil above

σcover standard deviation on the pipeline cover depth measurement accuracy

σeq equivalent stress (von Mises stress)

σh hoop stress

σR axial stress range


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σu ultimate strength, average at room temperature

σx,i standard deviation of parameter xi

σy yield stress, average at room temperature

σeq equivalent von Mises stress

Table 1-7 Definition of verbal forms

Term Definition

shall verbal form used to indicate requirements strictly to be followed to conform to the document

should verbal form used to indicate that among several possibilities one is recommended as particularlysuitable, without mentioning or excluding others, or that a certain course of action is preferred but notnecessarily required

may verbal form used to indicate a course of action permissible within the limits of the document

The use of international system of units (SI) and SI derived units is strongly recommended. However, any setof consistent units may be applied. Exceptions for the uplift resistance factor, , that shall be expressed inmetres. Temperatures shall be expressed in Celsius.

1.7 Structure of the documentThis recommended practice contains the following:

— Sec.2 - description of the different design scenarios, decision flowchart and background— Sec.3 - input parameters like pipe cross-section, material, operational parameters, survey and trawling— Sec.4 - pipe-soil interaction— Sec.5 - analytical equations and general requirements to the response model (FE model)— Sec.6 - detailed procedures and criteria for exposed pipelines - buckling on even and uneven seabed— Sec.7 - detailed procedures and criteria for buried pipelines - upheaval buckling— Sec.8 - complementary limit states to those in DNVGL-ST-F101— Sec.9 - calculation of the conditional load effect factor for exposed pipelines on even/uneven seabed— Sec.10 - structural integrity during the operational phase— Sec.11 - recommendations on documentation for operations— Sec.12 - bibliographic references, while standard references are given in [1.5]— App.A - includes examples of mitigation measures that are applicable to exposed pipelines— App.B - gives detailed guidance on probabilistic buckle formation calculations— App.C - gives guidance on effect from sour service on fatigue and fracture failure modes— App.D - gives example of virtual anchor spacing (VAS).


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SECTION 2 INTRODUCTION TO GLOBAL BUCKLING

2.1 GeneralThis section gives a brief introduction to global buckling of submarine pipelines. Global buckling for exposedpipelines is discussed in [2.2], pipeline walking in [2.3] while upheaval buckling of buried pipelines isdescribed in [2.4].A design based on controlled buckling and thereby release of the expansion forces has some specificchallenges during operation with respect to evaluation of inspection results. Therefore, clear requirementson how to follow this up during operation is found in Sec.10 while requirements on how to document theseaspects are given in Sec.11 .

2.2 Global buckling of exposed pipelines

2.2.1 Design overviewThe global buckling for pipelines exposed on seabed (unburied pipelines) will either initiate on the seabed,i.e. in the horizontal plane, or in the vertical plane. The buckling will be triggered either by natural out-of-straightness (i.e. lateral imperfections from installation or vertical imperfections due to uneven seabed), orby purpose made out-of-straightness (i.e. triggers or buoyancy).For pipelines on uneven seabed, the major difference compared to pipelines on even seabed, is the initiationor triggering of global buckling. For pipelines on uneven seabed the buckling includes the following threephases:

1) expansion into free spans2) lift off at the seabed crests:

— limited lift-off— maximum lift-off.

3) lateral instability, causing the pipeline to buckle sideways.

Design of exposed pipelines shall include the following design steps:

1) Buckling susceptibility: determine the susceptibility to experience lateral buckling, or initial upheavalleading to lateral buckling, due to temperature and pressure.

2) Limit state check for uncontrolled buckling: the post-buckling bending moments and/or longitudinalstrain including any cyclic load effects shall be within acceptable limits. Subsequent over-trawling shallbe considered, if relevant.

3) Development and assessment of buckle mitigation strategy: if uncontrolled buckling is found to giveunacceptable load effects, some form of buckle mitigation strategy is required, see Appendix A. Thiscould either be to prevent development of buckles, or to trigger buckling at known locations, thussharing the expansion between multiple buckles. As for uncontrolled buckling, the limit state checks forthe post-buckled configuration shall be within acceptable limits.

4) Assessment of buckle interaction and walking: the long-term response of the pipeline shall bedocumented to be robust with respect to buckle interaction, e.g. it needs to be considered whether axialfeed-in can localise in one buckle over the design life. Axial displacements due to pipeline walking needto be within allowable limits at all locations such as end structures.

Global buckling design of exposed pipelines is typically performed/checked at two different stages in apipeline project:

— assessment before installation referred to as the pre-installed phase— assessment after installation referred to as the as-installed phase.


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Assessment before installation is to ensure compliance of the pipeline design with requirements, whileassessment in the as-installed phase is to assess deviations from design, if any, and develop remedialmeasures, if needed.

2.2.2 Loading and responseTemperature and pressure effects set up expansion forces and may cause a pipeline to buckle globally. Thedriving force for global buckling of a pipeline is the effective axial force, S, which represents the combinedaction of pipe wall force, N, and internal and external pressures. The effective force for a fully restrainedstraight pipeline, S0, constitutes an upper bound for this axial expansion force.For certain combinations of expansion force and imperfections, the pipeline will experience global buckling.For a partially displacement-controlled condition, this implies that the pipeline will find a new equilibriumposition by moving perpendicular to the pipeline axis and release axial over-length (expansion) by forminga buckle. At the same time as the buckle forms, the pipeline close to the buckle moves axially, feed-in, fromboth sides towards the buckle (expansion). The level of axial force to initiate this global buckling responsedepends on:

— pipe cross-sectional properties— lateral soil resistance— out-of-straightness of the pipeline (geometrical imperfection)— lateral triggering force (e.g. trawling) (load imperfection).

The out-of-straightness may be caused by for example:

— small imperfections on the seabed, e.g. caused by boulders— global imperfections from uneven seabed— curvature in the horizontal plane purposely made or randomly imposed during installation— engineered initiators such as sleepers or increased buoyancy— crossing devices/arrangement.

Figure 2-1 can be used to illustrate lateral buckling of a section in a pipeline. Here it is for simplicity assumeda pipeline with free end expansion, which is a typical assumption but not always true because of endconstraints such as PLETs/PLEMs:

— Prior to applying pressure and temperature, the effective force will be given by the residual lay tension. Incase the ends are more or less free to move, the effective axial force will here be close to zero. Note thatthis step is not shown in Figure 2-1 in order not to make the case too complex.

— When the pipeline is started up and the temperature and pressure increase, the effective force willturn into compression. If the pipeline is totally restrained, i.e. not allowed to move axially, vertically orlaterally, the effective axial force will reach its maximum compressive force of S0 (red dotted line). Thisforce will vary along the pipeline as the temperature and pressure show decreasing profiles along thepipeline. Normally the pipeline ends are free to expand axially, and the effective axial force will thereforeend up at zero at the ends (see the black, solid curve in Figure 2-1).

Note that Figure 2-1 shows the effective axial force prior to any buckling.


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Figure 2-1 Effective force, prior to buckling

Figure 2-2 to Figure 2-6 show a short section of the pipeline illustrated in Figure 2-1 and how the effectiveaxial force reduces after reaching a critical value and triggers global buckling.

Figure 2-2 is split in two, the left part shows the effective axial force, S, over a shorter section of the pipeline.This figure is before the pipeline is started up, so the effective axial force is zero (assuming zero residuallay tension as a simplification). The right part of Figure 2-2 shows the typical development of the effectiveaxial force over time at one location in this pipeline section. For this case, the location is at a significantimperfection where a lateral buckle will be triggered when the effective axial force reach its critical value,Sinit. Note that this critical value, which is sufficient to trigger a buckle, depends on pipeline properties, pipe-soil resistance and the imperfection as discussed previously in this section. This is also illustrated in the leftpart of Figure 2-2 with the upper hatched line giving the varying Sinit along the pipeline section.Figure 2-3 shows the compressive effective axial force, S0, as the temperature and pressure increase duringstart-up. As seen to the right, the force has not reached the critical value, Sinit, and no buckling has beentriggered.In Figure 2-4 the temperature and pressure have further increased, and the effective axial force has reachedthe critical value, Sinit, (point A). The pipeline buckles, and the effective axial force at the apex of thebuckle will drop towards the post-buckling value, Spost (point B). The effective force will gradually build upaway from the apex of the buckle due to the axial pipe-soil restraining force or axial friction, f, giving theslope in the effective force diagram. The axial feed-in to the buckle (indicated by the yellow arrows) will beproportional to the shaded area between the solid line and the potential effective force line.If the pressure or temperature is further increased, a neighbouring imperfection may trigger a buckle andchange the effective axial force diagram as shown in Figure 2-5. From this point, the post-buckling forcein the apex of the buckle (point B) will remain constant, but the axial feed-in to the buckle will increase,proportionally to the shaded area. As this axial feed-in is what drives the lateral buckle, the buckle willincrease in deformation as the feed-in continues.

A typical fully developed force profile for a pipeline with multiple buckles and free ends is shown in Figure2-6. At each buckle, the effective axial force reduces as the adjacent pipeline section feeds in to formthe buckle. The critical buckling force for each buckle varies, depending on the initial out-of-straightness,seabed frictional response and buckle spacing. In the feed-in parts of the buckles, the slope of the effectiveaxial force profile is governed by the axial friction. Between adjacent buckles virtual anchors are formed at


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positions of zero displacement, effectively dividing the system into a series of short pipelines anchored ateach end. The virtual anchor spacing (VAS) associated with each buckle varies. The response of the pipelinebetween virtual anchors is analogous to the response between real anchors, and the loading caused by feed-in in to the buckle is characterised by the VAS.The probability of the initial pipeline configuration to share the available thermal expansion over severalbuckles or to feed into one single or fewer buckles are determined by the initial pipeline configuration and thepipe-soil interaction parameters.

Figure 2-2 General behaviour during global buckling - critical buckling force


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Figure 2-3 General behaviour during global buckling - start-up and increase of effective axialforce

Figure 2-4 General behaviour during global buckling - effective axial force reaches critical valueand a buckle is triggered


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Figure 2-5 General behaviour during global buckling - pressure/temperature increased furtherand second buckle triggered

Figure 2-6 General behaviour during global buckling - final configuration with multiple buckles

The closer the buckles are to each other, the less the axial feed-in to each buckle. The aim of the designapproach for unburied pipelines is to ensure that several buckles form at regular intervals along the pipeline,sharing the expansion potential between the buckles, and leading to limited feed-in and acceptable utilisationof the buckles.


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2.3 Pipeline walkingPipeline walking is a phenomenon in which start-up/shut-down cycles may cause accumulation of axialdisplacement or some sort of ratcheting response over time. Given several load cycles, this ratchetingresponse may lead to very large accumulated axial displacements and potential over-utilisation of jumpers,spools and/or fixed structures. This phenomenon mainly affects short pipelines with high temperature loadingand a temperature transient with a steep gradient during start-up. As lateral buckles essentially divide a longpipeline into a series of short pipelines, pipeline walking may also be relevant to any pipeline that experiencelateral buckling.Pipeline walking can be caused by:

— seabed slopes— thermal transients— steel catenary riser (SCR) tension, and— liquid hold-up.

An example of pipeline walking due to seabed slope is illustrated in Figure 2-7. Here three different cases fora short pipeline are illustrated. One case with no seabed slope (grey lines), one with seabed slope only (blacklines) and one with seabed slope combined with gas/liquid separation on shutdown (red/blue line).For the case with no slope, the loading (start-up) and unloading (shut-down and cooling) curves are bothsymmetrical about the mid-point of the pipeline, the virtual anchor point, meaning that the ends will expandand contract almost equally and no walking will occur. In case of a slope, the expansion in the right end foreach start-up will be larger than the contraction, therefore the pipeline will experience walking to the righteven though the slope is very moderate. A gas/liquid separation during shut-down will further increase thiswalking.

Figure 2-7 Effect of slope and density changes on pipeline walking

In the absence of thermal transients or seabed slopes, the pipeline expands and contracts about a singlevirtual anchor located at the centre of the pipeline and no walking would be expected. If a seabed slopeis present, it is easier for the pipeline to expand or contract down the slope, and thus the locations of thevirtual anchors during operation and shutdown become offset from the centre of the pipeline. A section of the


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pipeline then moves in the same direction during both operation and shutdown, and thus the pipeline walksdown the slope. If gas/liquid separation occurs during shutdown, the axial resistance reduces in the gas filledsection and increases in the liquid filled section, further increasing the distance between virtual anchors andaccelerating the rate of walking.For pipelines subject to thermal transients on start-up, walking is driven by the incremental movement ofthe virtual anchor along the line from the inlet toward the centre as the pipeline heats up. On subsequentrestarts, a second virtual anchor is formed in the downstream section as the force profile movesincrementally from its unload to load form, further increasing walking rates.For pipelines connected to an SCR, walking may also be driven by the tension exerted by the SCR. The SCRtension may dominate the rate of walking, potentially necessitating incorporation of local anchors.Whilst the mechanisms that drive pipeline walking are well understood, it remains difficult to reliably predictthe propensity for and rate of pipeline walking. This is because of the complex interactions of competingmechanisms, uncertainty associated with and sensitivity to pipe-soil interactions, and uncertainty overoperation shutdown and restart cycles.

2.4 Upheavel buckling of buried pipelineIf the pipeline will be covered, the cover or lateral restraint shall be designed to avoid global buckling of thepipeline. This may be done by trenching and covering the pipeline by natural or artificial backfill. Soil nature,pipeline properties and trenching technology influence the evenness of the trench bottom.An un-trenched pipeline may be restrained in its configuration e.g. by covering with continuous graveldumping. This may be the preferred choice in some cases, particularly for soft clay where trenching mayfurther reduce the strength of the clay. Soil nature, pipeline properties and rock dumping technologyinfluence the shape and height of the pipeline cover.All forces in vertical and horizontal directions shall be evaluated when designing the cover because a largehorizontal out-of-straightness may cause the pipeline to break out of its cover in another direction than thevertical one.Upheaval buckling design of buried pipelines is, as global buckling for exposed pipelines, performed/checkedat two different stages in a pipeline project:

— Assessment before installation referred to as the pre-installed phase.— Assessment after installation referred to as the as-installed phase.

Assessment before installation is to ensure compliance of the pipeline design with the requirements and toget a cost and gravel estimate, while assessment in the as-installed phase is to ensure compliance of the as-constructed pipeline including measured vertical imperfections and cover heights.


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SECTION 3 BASIS FOR DESIGN

3.1 GeneralThis section defines the characteristic design parameters and load combinations that shall be used with thedesign procedures and criteria in this recommended practice.All the design parameters such as loads, geometries and material strengths are associated with uncertainties.These uncertainties can be split into four groups:

— natural, physical variability— statistical uncertainty— measurement uncertainty— model uncertainty.

Uncertainty in design parameters may be expressed as a probability density function (e.g. normaldistribution, lognormal distribution).In this recommended practice, lower estimate (LE) and upper estimate (UE) values are used to account foruncertainties.

Guidance note:Natural variability is of random nature and characterised by the fact that more studies will not necessarily reduce the uncertainty.One example is measurement of wall thickness which varies independently of how many measurements are carried out. Note thatlimited measurements (statistical uncertainty) and accuracy of measurements (measurement uncertainty) sometimes are includedin the natural variability.Statistical uncertainty relates to the uncertainty in predicting the statistical variables. Increasing number of samples reduces thestatistical uncertainty.Measurement uncertainty relates to the accuracy in the measurement of each sample.Model uncertainty is characterised by limited knowledge or idealisation of stochastical or physical models. More research willtypically reduce the model uncertainty.


3.2 Pipe geometryLoad effect analyses should be based on the most unfavourable combinations of loads if no project specificloading scenarios are available. The nominal wall thickness should be used for load calculations, while thecorroded section should be used, if relevant, for the resistance.

Table 3-1 Characteristic pipe geometry properties

ResistanceParameters Limit state

Symbol Value

Load effectcalculation

diameter all D nominal D (nominal)

wall thickness all t2 t2 = t-tcorr t (nominal)

Guidance note:Clad or liner thickness shall be included in the wall thickness in the load effect calculations. Any positive effect on the resistanceshould only be included if this effect is documented.



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3.3 Pipeline materialThe material parameters should be based on the nominal values except for the yield stress and ultimatestrength. The stress-strain curve based on yield stress and ultimate strength should be based on thespecified minimum values, fy and fu, as per DNVGL-ST-F101, except when the mean value is explicitlyrequired.It is important to include the temperature effect on the material parameters, not limited to the yield stressand ultimate strength only but also to the temperature expansion coefficient, , and to Young's modulus, E.Note that the thermal expansion coefficient will increase with temperature and neglecting this effect will givenon-conservative results. The characteristic material strength factors are defined Equation (3.1) to Equation(3.4) and a summary of the material parameter definitions are given in Table 3-2.

(3.1)

(3.2)

(3.3)

(3.4)

Where all the parameters are explained in Table 3-2 below.

Table 3-2 Material parameters

Parameter Symbol Value Temperature effect 1)

Young's modulus E nominal/mean above 50oC 2)

temperature expansion coefficient α nominal/mean above 50oC 2)

yield stress, specified minimum SMYS minimum specified 3) at room temperature

yield stress YS mean at room temperature

ultimate strength, specified minimum SMTS minimum specified 4) at room temperature

ultimate strength TS mean at room temperature

Poisson's ratio v mean negligible

reduction in yield stress due to elevatedtemperature 1), 5)

fy,temp mean above 50oC , local designtemperature

reduction in ultimate strength due toelevated temperature 1), 5)

fu,temp mean above 50oC , local designtemperature

supplementary requirement U αu 0.96 or 1.0 none


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Parameter Symbol Value Temperature effect 1)

1) For duplex and super duplex steels, derating shall be considered from 20°C.2) If a constant derating value is used in the load effect calculations, this shall be an equivalent value representing the

total effect at the local design temperature.3) When supplementary requirement U, see DNVGL-ST-F101, is specified, the minimum specified yield stress shall be

at least two standard deviations below the mean value.4) When supplementary requirement U is specified, the minimum specified tensile strength shall be at least three

standard deviations below the mean value.5) For guidance on derating see DNVGL-ST-F101.

If no data exists on material derating effects, values from DNVGL-ST-F101 can be used for both the yieldstress and the ultimate tensile strength.See /3/ for guidance on the temperature dependency of the thermal expansion coefficient for carbon steel.

3.4 Loads

3.4.1 GeneralGlobal buckling shall be checked for the most critical 100-year return period load effect. Various loadcombinations for the 100-year return period may be governing and should, hence, be checked.The following load effect combinations should be checked for exposed pipelines:

— Functional design case - extreme functional load effect (100-year) with associated interference andenvironmental loads effects, see also DNVGL-ST-F101.

— Interference design case - extreme interference load effect with associated functional and environmentalload effects, see also DNVGL-ST-F101.

— Environmental design case - extreme environmental load effect (100-year) with associated functional andinterference load effects, see also DNVGL-ST-F101.

The following load effect combinations should be checked for buried pipelines:

— functional design case - extreme functional load effect (100-year).

The value of the load effects should be in accordance with the relevant load combination, see guidance inTable 3-4.

3.4.2 Operational data (functional loads)The analyses should be performed with relevant operational parameters (pressure and temperature). For thefunctional design case this should represent the 100-year return values, typically the local incidental pressureand local design temperature unless this combination is documented to be unrealistic. For the interferencedesign case, the value of the operational loads will depend on the probability of occurring simultaneouslywith the interference load. This implies that the operational load for the interference (trawl) design case willdepend on the trawling frequency as given in Table 3-4.The relevant (typically the design) temperature profile should be used if available. Otherwise, conservativeassumptions should be applied. The insulation should include conservative assumptions to ensure that theannual probability of exceedance of these temperatures will be equal to or less than 10-2.

3.4.3 Trawling loads and frequencies (interference loads)Evaluation of trawling loads should be based on the principles in DNVGL-RP-F111.

The trawl pull-over load, FT, depends on the trawl frequency, fT, and trawl gear type among otherparameters. If detailed information is not available, the values in Table 3-3 should be applied.


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The trawling frequency is to be considered as the annual frequency per relevant pipeline section. For globalbuckling assessment where the trawl load acts as a triggering mechanism, the section relates to the part ofthe pipeline in a trawl area (order of kilometres) that has a potential for buckling.For trawling assessment in a buckle, i.e. after it has buckled globally, it is conservatively assumed that thetrawl gear will hit the buckle near the apex in an unfavourable manner. The section length refers then tothe length of the buckle or sum of buckles if more than one buckle is anticipated. The length of the relevantsection is typically less than 100 metres per buckle.

Table 3-3 Definition of characteristic trawl pull-over load, FT

Pull-over load fT < 1 10-4 < fT < 1 fT < 10-4

FTUE 1.3 Fp 1.0 Fp NA

FTBE 1.0 Fp 0.8 Fp NA

FTLE 0.4 Fp 0.3 Fp NA

Fp is the trawl pull-over load according to DNVGL-RP-F111 Sec.4.

Table 3-3 should both be used to evaluate the triggering of a buckle, and for the limit state check of abuckled pipeline.

Guidance note:The trawl interference evaluation in this recommended practice is limited to lateral buckling only. Hence, trawling in free span isnot covered by this recommended practice but by DNVGL-RP-F111.


Guidance note:The effect of trawling in this recommended practice is included as a sensitivity study on the overall bending moment response.Since the global buckling moment is mostly displacement controlled, the load-controlled trawl moment will not be added but to alarge extent be replaced by the functional moment from global buckling. If the contribution from trawl loads is dominating, special

evaluations are required in order to determine a higher Yc than given by use of this recommended practice.


3.4.4 Environmental loadsIf no information exists about the most critical 100-year return period condition, the following combinationsare proposed for pipelines on seabed:

— 100-year return period bottom current and 10-year return period wave-induced flow— 10-year return period bottom current and 100-year return period wave-induced flow.

If no information exists about the most critical 1-year return period condition, the following conservativecombinations are proposed:

— 1-year return period bottom current and 1-year return period wave-induced flow.Guidance note:The environmental design case is normally not a governing design case for global buckling. One exception may be triggering oflateral buckles on even seabed.


Guidance note:Environmental loads can be calculated according to DNVGL-RP-C205.



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3.4.5 Load combinationsAll relevant load combinations shall be checked for 100-year return period load effects and the most criticalof the load combinations used. Table 3-4 gives guidance on the governing load combinations.

Table 3-4 Load combinations to be considered in design

Functional loadTrawling

frequency 1) ScenarioPressure load 2) Temperature load

Trawl load 3) Environmentalload

functional design local incidental local design no -

interference design n/afT < 10-4 andburied pipeline

environmental design local operating local operating no 100 year

functional design local incidental local design no -

interference design local operating local operating FTBE = 0.8 Fp -10-4 < fr < 1


functional design n/a

interference design local incidental local design FTBE = 1.0 Fp -fr < 1


1) Trawling frequency is defined in [3.4.3].2) Function loads discussed in [3.4.2].3) In addition to sensitivity analyses as required by the design procedures, see. Sec.6 and Sec.9.

3.5 Time effectsDesign properties may change with time, such as pressure and temperature profiles and wall thickness(corrosion). Further, pipe-soil resistance may vary due to consolidation of soil and build-up of soil berms.Any increase in pressure or temperature beyond design values may give excessive feed-in to a global buckleor excessive uplift forces which may initiate an upheaval buckle. If pressure and temperature are associatedwith decaying trends, the beneficial effects may be accounted for in integrity assessments in the operationalphase. This requires an update of the design values.If the pipeline experiences repeated start-up and shutdown cycles at pressure and temperature levels belowdesign values, this may cause build-up of soil material in the buckles. If the pipeline then is exposed todesign pressure and temperature values, this soil berms may give enhanced lateral resistance at the crown ofthe buckles which may lead to more severe curvature than would occur if the same operating conditions wereimposed on first start-up.Corrosion damage develops over time. Hence at first start-up, the cross-section will be intact along thewhole pipeline, ensuring that the complete cross-section will contribute to expansion forces and expansionto the buckles. If corrosion damage develops almost uniformly along the complete length of the pipeline, theexpansion force will also reduce accordingly. However, the reduction in expansion may be limited.If the corrosion damage is documented to occur at specific locations in the cross-section (i.e. not uniformaround the circumference) an equivalent wall thickness may be established representing the bending capacityof the corroded pipe section. The equivalent wall thickness will be larger than t2 and smaller than thenominal wall thickness t. This wall thickness should only be used for the resistance calculations.


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Corrosion damage will typically not be uniform around the cross-section or in the longitudinal direction ofthe pipeline, i.e. there may be local grooves with different shapes at various intervals. Assuming uniformcorrosion damage along the circumference and in the longitudinal direction may not represent the straincapacity for a pipeline exposed to global bending adequately, and hence it should be confirmed that the straincapacity for such grooves is not giving lower strain capacity than the strain capacity of the uniform corrodedpipe. See DNVGL-ST-F101 Sec.4.

Guidance note:The fact that corrosion develops over time may be accounted for. In this way other time dependent effects such as the beneficialeffects of cyclic loading on buckles can be accounted for as this will to a certain degree balance the reduction in the load carryingcapacity due to corrosion. It is also a fact that corrosion develops over time and that maximum corrosion at the end of the lifetimeis often associated with reduced temperature and/or pressure. However, accounting for this beneficial reduction in functional loadswould normally require re-qualification of the pipeline.


Guidance note:For pipelines susceptibe to corrosion damage it is important to monitor this in the operational phase based on the assumptionsmade in the global buckling design with respect to corrosion.



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SECTION 4 PIPE-SOIL INTERACTION

4.1 GeneralThis section gives some of the fundamental pipe-soil modelling requirements for assessment of pipelineglobal buckling and expansion. Detailed formulations and discussions regarding soil resistance are given inDNVGL-RP-F114.Global buckling behaviour of exposed and buried pipelines is strongly linked to the pipe-soil interaction. Thepipe-soil interaction includes large uncertainties because of variation and uncertainty in characterisationof the soil and is one of the most critical aspects of global buckling or expansion design. The uncertaintiesrelated to pipe-soil interaction are often difficult to quantify and a fair amount of engineering judgement isrequired.A limited amount of survey data should result in larger soil property ranges (upper estimate, best estimateand lower estimate). If the sensitivity to these properties in the pipeline response is large, the designprocedure may require higher safety factors, so more sampling may be an advantage. On the other hand, ifthe response sensitivity is small, further samples may be of limited value.The components of the pipe-soil interaction involved in the potential buckling modes of a pipeline are thefollowing:

— Downward stiffness - downward stiffness is important for smoothing of survey data, and may influence theuplift resistance for upheaval buckling design. Best estimate values should be used as characteristic valuefor exposed pipelines, and best estimate and lower estimate for buried pipelines.

— Lateral resistance - for an exposed pipeline free to buckle laterally, the lateral pipe-soil interaction is oneof the key parameters as it influences both the critical buckling load (break-out resistance) and pipelinepost-buckling configuration (residual soil resistance after break-out). For a buried pipeline, the likelihoodof lateral movement/buckling needs to be assessed.

— Axial resistance - the axial pipe-soil interaction is relevant when any buckling mode is triggered as itaffects the post-buckling configuration. The axial feed-in from the straight sections into the buckled regionis determined by the mobilised axial reaction (of the natural soil and/or of the gravel/rock cover). Theaxial pipe-soil interaction is also important for the axial force build-up, either during buckle development,at the pipeline ends or after a buckle has occurred. The axial pipe-soil interaction also affects buckleinteraction, axial walking and loads on any anchors. For buried pipelines with a high D/t ratio, theaxial restraint provided by the pipe-soil interaction may influence local wrinkling of the pipe and forall D/t ratios influencing upheaval buckling. For the latter, an upper estimate value should be used ascharacteristic value.

— Upward resistance - vertical pipe-soil interaction is of major concern for upheaval buckling, as it affectsthe mobilisation load. A multi-linear interaction model is normally required. A lower estimate value shouldbe used as characteristic value. This is further modified by a resistance safety factor.

The selection of the most suitable pipe-soil interaction formulations and parameters in buckling analysisshould be guided by engineering judgement supported by experience on the specific problem and, wherepossible, by correlation/benchmarking with field measurements. In addition, sensitivity analyses are alwaysrecommended, aimed at determining the criticality of design parameters with respect to parameters andmodelling assumptions.Simplifications of the pipe-soil interaction may be considered in the assessment. Emphasis should then be tomake this simplification representative for the relevant condition, e.g. the model will be different when thebreakout force is estimated as compared to when the post-buckling configuration is determined.Seismic activity causes dynamic excitation (lateral and vertical) and since most upheaval buckling mitigationsare by gravitational rather than mechanical restraints, seismic effects shall be considered when judgedrelevant. Seismic activity may also cause liquefaction of the soil. No specific guidelines are given with respectto how seismic activity influences the pipe-soil interaction.


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4.2 Exposed pipeline

4.2.1 GeneralThe design of exposed pipelines will either aim to ensure that the pipeline will not buckle laterally, or in caseit buckles laterally, to ensure its structural integrity is maintained. The two cases require different pipe-soilresistances, either for small displacements or for large displacements. The appropriate values depend on thebuckling category, see Table 4-1.

Table 4-1 Required friction factors for pipelines left exposed on the seabed

Buckling category Axial Lateral

no global buckling BE LE

maybe global buckling (only relevant for trawling interaction) LE LE BE

global buckling LE/BE/UE 1) LE/BE/UE

1) The "/" means "and", i.e. that all values shall be considered.

— LE - lower estimate— BE - best estimate— UE - upper estimate.

In application of this recommended practice, pipe-soil interaction lower and upper estimates are typicallyassumed to be defined as mean ± two standard deviations for symmetric distributions. For non-symmetricdistributions, it may be more representative to define the upper and lower estimate values as a fractile,where the exceedance probabilities range from 2% to 5%. For a more detailed discussion about theuncertainties, lower, upper and best estimates for pipe-soil parameters, see DNVGL-RP-F114.

4.2.2 Axial pipe-soil resistanceA high axial breakout resistance will influence the global buckling initiation, but to ensure conservatism inthe design process this is generally not accounted for. Therefore, residual axial resistance should be used tomodel thermal expansion, post-buckling behaviour and pipeline walking. However, if there is strong evidenceto support a high breakout resistance then it may be considered in the assessment of buckle formation butonly with great care.A first breakout peak resistance generally has little influence on long-term cyclic walking. If breakoutresistance occurs on all load cycles, it would tend to inhibit pipeline walking under thermal transient loading.As there may be insufficient duration for the pipe-soil interface to drain when the pipeline is stationary duringeither operation or shut-down, incorporation of the axial breakout resistance on all subsequent walking cyclescould be considered but with great care.Sensitivity checks on route curve pull-out should include axial breakout resistance if it may increase the peakaxial tension along the route curve, e.g. if shutdown occurs following a long stable operating period.

4.2.3 Lateral pipe-soil resistanceThe lateral pipe-soil resistance is the key parameter for the lateral buckling on even seabed as it influencesboth mobilisation load (break-out resistance) and pipeline post-buckling configuration. At mobilisation, whenthe pipeline starts to deflect laterally and the displacements are small, the lateral soil resistance is governedby the breakout value. For increasing lateral displacements, the lateral soil resistance may decrease to alower residual value. In the tail of the buckle, where little lateral movement occurs, the lateral resistance willalways be the breakout resistance and this will influence the post-buckling configuration.


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Surface soil, swept ahead of the pipeline on each operational shut-down/restart cycle, can generate soilberms at the extremes of the pipeline lateral displacement. Further cycles of lateral movement lead to asteady increase in the restraint provided by the soil berms. These berms can offer significant resistanceto pipeline movement, preventing growth of buckle amplitude (ratcheting) and wavelength on subsequentcycles so that cyclic displacements remain almost constant. The berms define the shape of the buckle inoperation and prevent excessive reduction in the stress range.Compared to an even seabed, the lateral pipe-soil resistance is less important for buckle initiation on unevenseabed, as the imperfections will generally be larger in the vertical plane, and the initial buckling is likely tooccur vertically before transitioning into a lateral buckle.

4.2.4 Response models for uneven vertical loadingSome consideration is required in the modelling of pipe-soil response in zones where the local vertical loadvaries, for example at the touchdown point to each side of a sleeper or span, or at the ends of a distributedbuoyancy section. The analysis should enable modelling of increased lateral resistance in zones of highcontact stress.The diameter variations within distributed buoyancy sections should also be captured in FE models to ensurecorrect modelling of the pipe-soil contact regions.

4.2.5 Application of pipe-soil resistance in global buckling analysisThe load effects during global buckling should be estimated using the pipe-soil resistance combinations listedin Table 4-2 below.

Table 4-2 Pipe-soil combinations for virtual anchor spacing (VAS) analysis

Failure mode/limit state Axial 1) Lateral 2) Comment

local buckling/load controlled (alternative1 in [6.4] and Table 6-1) BE BE condition load factor γc as per Sec.9 4)

local buckling/load controlled (alternative2 in [6.4] and Table 6-1) BE UE

fatigue 3) BE UE

fracture 3) BE UE

excessive lateral displacement BE LE

1) The axial pipe-soil interaction can be modelled using an equivalent Coulomb friction.2) The non-linear lateral resistance model should include any peak resistance at break-out from the embedded

condition, or increasing resistance following breakout as the soil berm is established. For conceptual studies, anequivalent Coulomb friction is considered sufficient.

3) The full non-linear lateral pipe-soil resistance should be modelled for the first cycle and cyclic resistance forsubsequent cycles, which may increase with each cycle. The lateral pipe-soil model should incorporate the effect ofsoil berm resistance at the extremes of cyclic displacement.

4) For conceptual design, γc = 0.85 may be applied.


The axial and lateral resistance models should both employ the best estimate mobilisation displacements.


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Under certain pipe-soil conditions, the upper estimate of lateral residual resistance may be very high suchthat, if it is assumed that the buckle forms purely in the horizontal plane, a practical design solution for on-bottom buckles (planned or unplanned) becomes difficult to achieve.

Guidance note:If it can be shown that high lateral residual resistance is strongly correlated to high break-out resistance, it may be possible todemonstrate that buckling wholly in the lateral plane is extremely unlikely under high lateral resistance conditions. The design willneed to ensure the most onerous buckle response is identified by considering the conditions under which buckling may occur, thepotential buckle behaviour and the likely pipe-soil response interaction, etc. The design should demonstrate that the approach doesnot compromise the target failure probabilities.


If buckling initially occurs in the vertical plane either due to very high lateral break-out resistance or at theoverbend of a seabed imperfection, it is likely that the uplifted buckle will develop into a lateral buckle.

4.2.6 Application of pipe-soil resistance in buckle interaction and walkinganalysisFor assessment of pipeline walking and buckle interaction the pipe-soil resistance combinations listed in Table4-3 should be applied.

Table 4-3 Pipe-soil combinations for buckle interaction and walking analysis

Design scenario Axial 1) Lateral 2) Comment

buckle interaction and walking LE/BE/UE BE sensitivity assessment should beperformed for the lower and upperestimate axial residual resistance,and the weighting of confidence in thewalking predictions should be basedon the level of confidence in the axialfriction response

1) Axial pipe-soil interaction can be modelled using an equivalent Coulomb friction coefficient. For conceptual studies, itis considered sufficient to apply the best estimate residual resistance only.

2) Lateral pipe-soil interaction should incorporate the effect of soil berm resistance at the extremes of cyclicdisplacement. For conceptual studies, an equivalent Coulomb friction is considered sufficient.


Pipeline walking and buckle interaction are very sensitive to the level of axial restraint and caution is requiredin designing solely to best estimate axial residual friction. Furthermore, walking and buckle interactionbehaviours can be complex with many opposing drivers, and it can be difficult to predict whether the mostonerous behaviour is associated with best, upper or lower estimate or indeed with some intermediateresistance. The design should quantify the sensitivity of the walking susceptibility and rate, and of any anchorloads, to uncertainty in axial pipe-soil resistance.

Guidance note:If walking behaviour is dominated by thermal transients, identification of the most critical axial pipe-soil resistance is moreuncertain and sufficient analyses should be performed to understand the most severe behaviour.If the sensitivity analysis indicates that mitigation may be required, the design may be based on the best estimate values withprovision for mitigation to arrest walking later in the operating life of the pipeline, if operational monitoring indicates that thewalking behaviour is unacceptable.



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For scenarios where walking is dominated by pipeline feeding out of the lateral buckles, lower estimate lateralresistances will cause reduction in buckle amplitude and an increase in pipeline walking. In this case, thelower estimate cyclic lateral resistance can result in the highest rates of walking.The axial and lateral resistance models should both employ the best estimate mobilisation displacements.However, the sensitivity to axial mobilisation displacement should be quantified.

Guidance note:If the best estimate axial mobilisation displacement is small, the sensitivity to a low estimate value is likely to be small, and asensitivity analysis may not be required.


4.3 Buried pipelinesBuried pipelines are covered by soil, either with the seabed soil, trenched and backfilled soil or with additionalcover material (e.g. gravel). Soil properties will be required for each of the three materials as relevant in thedesign:

— in-situ soil conditions— trench material (remoulded/fluidised and reconsolidated)— added cover material.

The in-situ soil conditions are used for the downward resistance of the pipeline. This may either be on theseabed, for pipelines resting on the seabed and covered with gravel, or at the bottom of the trench, for apipeline in a trench.The trench material can be either natural seabed material ensured by ploughing back the material fromthe sides or by jetting where the material has been flushed backwards along the pipeline. When the soil isploughed back, some parts of the soil will maintain its original strength while there may be water pockets inbetween these parts. The jetting trenching will, on the other hand, liquefy the soil, giving a very homogenoussoil, and remould the strength. Attention should be paid to the characterisation of the backfilling material inthe short and long term, as variation in the characteristic soil parameters may be expected.The added cover material will typically be gravel, whose properties will depend on the source for the materialand need to be defined correspondingly.The modelling of pipe-soil interaction generally supplies analytical relationships to describe the ultimate soilcapacity and the relative displacements at mobilisation of ultimate capacity. Proposed pipe-soil interactionformulations are given DNVGL-RP-F114.For assessment of upheaval buckling, the pipe-soil resistance combinations listed in Table 4-4 should beapplied. If horizontal imperfections are present, lateral pipe-soil resistance needs to be included.

Table 4-4 Pipe-soil combinations for upheaval buckling

Failure mode/limit state Axial Vertical

upheaval buckling - upward UE LE

upheaval buckling - downward UE LE/BE



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SECTION 5 LOAD EFFECT CALCULATION

5.1 GeneralThis section contains load sequence description and requirements for design analyses, analytical methodsand detailed FE analysis, valid for exposed pipelines and buried pipelines. Additional requirements for buriedpipelines are presented in Sec.7.A simplified model may be used to calculate the load effects, if it can be documented to give conservativeresults when compared to more advanced methods. In practice, analytical expressions may be used duringconceptual design phases. Such expressions may also be used to assess standard lateral buckling ofexposed pipelines on even seabed and to assess upheaval buckling of buried pipelines. For more complexconfigurations of buckling of buried or exposed pipelines (e.g. where the pipeline may initially bucklevertically at bathymetric features, buckle initiators, pipelines with significant spanning), and during detaileddesign, analyses with advanced finite element methods are normally required.Simplified calculations should be used to confirm the detailed analyses and calculations to avoid gross errors.

5.2 Analysis loading sequenceThe complete loading history of the pipeline should be accounted for in the load effect analysis. This mayinclude the load steps in Table 5-1.

Table 5-1 Finite element analysis load history

Exposed pipeline Buried pipeline

empty empty

water filled water filled

system pressure test system pressure test

dewatering operation 1) dewatering operation

start-up start-up

operation at operational condition operation at design condition

operation at design condition shut-down

trawling (if relevant) repeated start up and shut down (cyclic loading)

shut-down

repeated start up and shut down (cyclic loading)

pack-in and shut-in condition

1) The dewatering operation may in some cases be critical for global buckling with combination of light pipeline andrelatively high pressure.

The cyclic loading analysis should be based on the maximum stress range associated with pressure andtemperature variations. The relevant operational pressure and temperature cycles may be used to calculatethe fatigue loads.


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Guidance note:

For any given cycle of pressure and temperature, the maximum axial stress range generally occurs between the cold pressurisedcondition and the hot depressurised condition. However, the peak stress may occur at an intermediate stage in the loading. Careshould be taken to ensure that the maximum stress range occuring in the entire load cycle is identified.If operating procedures, supported by transient analysis, ensure that neither the cold pressurised nor the hot depressurisedconditions will occur, the stress range may be reduced.


For pipeline walking analyses, appropriate pressure, density and temperature transients need to be included.In addition to the loading history, it is important for the load effect analyse to account for changes in thepipeline boundary conditions and changes in pipe-soil interaction during relevant construction phases andduring different operating conditions. Seabed subsidence, due to for example reservoir depletion, may causevertical and horizontal movement in the effective region. For pipelines with high degree of axial restraint, asis the case for buried pipelines, such horizontal seabed movement can induce large axial forces leading tobuckling and/or overstressing of the pipeline. This should be considered, if relevant.

5.3 Analytical methods

5.3.1 Restrained effective axial forceA pipeline may experience the fully restrained effective force, S0, as given in Equation (5.1), if the locationunder consideration is more than one anchor length from the pipeline end and no buckling has occurred, seealso [2.2] and Figure 2-1.

(5.1)

where:

Hlay = the residual lay tensionΔpi = the difference in internal pressure compared to as-laidAi = the pipe internal bore areaυ = Poisson's ratioAs = the pipe cross-sectional steel areaα = the thermal expansion coefficient for the pipe materialΔT = the difference in temperature relative to the temperature during installation.

Temperature change in the pipeline can be caused by change in the surrounding temperature and/or in theinternal temperature.

Guidance note:

Δpi is the difference in internal pressure between the analysed condition and when laid down on the seabed. Since the internal

pressure during installation normally is zero, Δpi will often be identical to the internal pressure for the analysed condition.


For a more general discussion and application, see /4/ and /5/.

The choice of pipe-soil axial resistance, fa, will affect the load effect calculations, especially the magnitude ofthe effective axial force along the pipeline.

The effective axial force is limited by the fully restrained effective axial force as the theoretical maximumcompressive force. Close to the pipeline ends, the effective axial force reduces from its maximum value, S0,towards zero due to end expansion (assuming the pipeline is not axially fixed at the end). The reduction of


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the axial force along the pipeline is governed by the axial friction between the pipeline and soil. Hence, a highresistance will give higher forces close to the end, and potentially triggering buckles in this area. Restraintsfrom end structures such as spools, PLEM, PLET et cetera should be considered.

The same considerations apply close to buckled sections. A high axial resistance will cause a faster build-up of axial force that may trigger other imperfections closer to the buckled section and development ofadditional buckles.

5.3.2 Global lateral buckling on even seabedGlobal lateral buckling of an exposed pipeline resting on an even seabed may in some cases be analysed withanalytical methods. The analytical method employed should accurately predict both the first buckling loadand the cyclic post-buckle buckling response.A strain concentration factor (SNCF) should be employed to account for so-called global strain concentrationsaccording to DNVGL-ST-F101 Sec.5, (e.g. due to field joints, buckle arrestors et cetera) that may coincidewith a lateral buckle.Various analytical models to assess lateral buckling on even seabed are available in the literature,see /1/, /2/ and /6/. However, analytical methods have several limitations due to the assumptions on whichthey are based:

— linear elastic or simplified elastic-plastic material behaviour— simplified axial and lateral pipe-soil interaction described by a Coulomb-friction— small rotation theory— imposed shape of initial and post-buckling configuration according to assumed buckling mode. For small

initial imperfections, mobilisation load is related to an assumed modal shape that may differ from the realas-laid configuration of the pipeline.

If one or more of the limitations for the analytical methods is not realistic, a more sophisticated analysis isrequired.

5.3.3 Pipeline walkingAnalytic methods can be used to assess the pipeline walking behaviour for a pipeline, see /7/ and /8/.Typically, in these models, individual walking drivers are evaluated separately before combining them into anestimated total response. As for the analytic models for global buckling, knowledge of the limitations of suchmethods should be evaluated before they are employed in design.

5.4 Detailed finite element analysis

5.4.1 Finite element method specificsThe pipeline response during global buckling should be analysed using non-linear finite element methods.The FE analysis should represent the physical phenomena and behaviour of the pipeline adequately.The FE analysis shall consider:

— Non-linear material (steel) behaviour – the material model shall consider the non-linear and bi-dimensional (in the longitudinal and hoop-direction) state of stress by an appropriate yield surface andhardening rule. The choice of true versus engineering stress and strain needs to be consistent with theselected FE program and its element formulation.

— Large rotation theory – this requirement is relevant for pipeline rotation larger than about 0.1 radians.— Element size and type - the pipe element length and number of integration points should be demonstrated

to be adequate to identify the curvatures developed in the buckle and to capture the seabed bathymetry.The element length should typically be in the order of one diameter where the buckle is expected to occurand may be longer in the straight pipeline sections. A strain-based acceptance criterion will be more


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sensitive to the numerical discretisation, i.e. the element size and number of integration points used, asthe strain will be under-predicted in a too coarse element model. Therefore, sensitivity analyses shouldbe performed with different element size and integration points to ensure that strain results that haveconverged.

5.4.2 Finite element modellingThe following requirements apply to the finite element modelling:

— Pipe geometry - for pipe geometry modelling, see [3.2].— Pipeline material - the stress-strain curve should be based on the specified minimum values, fy and fu,

considered being engineering stress values, except for when the mean value is explicitly required bythe procedure. The stress-strain curve should reflect the response of the actual material and shouldrepresent the strain range of interest, e.g. the use of SMYS and SMTS to define the hardening is likely tobe inappropriate. Correct modelling of the stress-strain curve is especially important when a strain basedacceptance criteria is used. The shape of the stress-strain curve should be confirmed against projectspecific test data.

— Initial pipeline configuration - see [5.4.3].— Pipe-soil interaction - see [5.4.4].— Buckle interaction and pipeline walking- see [5.4.5].— Pipe-in-pipe (PIP) systems - For a PIP system, both the inner and outer pipes should be reflected in the FE

modelling. The interaction between the two pipes should be carefully modelled to incorporate all essentialeffects (e.g. insulation stiffness and strength, gaps, centralisers, structural connectivity between pipesand internal friction, field joint stiffness discontinuities). The system modelling should address the stressand strain concentrations that may occur within the pipe-in-pipe system. Installation of a PIP systemby reeling will set up residual curvatures in the outer and inner pipes that needs to be considered in thebuckling design.

— Stress and strain concentration factors - the potential for stress and strain concentrations at the crownof the buckle (e.g. pipeline field joint stiffness discontinuity, buckle arrestor), should be fully quantifiedthrough detailed FE analyses or a test programme. It is not necessary to model the influence of normaljoint-to-joint strength mismatch and misalignment in the global analyses.Guidance note:The manufacturing process affects the stress-strain curve. Seamless pipes tend to have a Lüder plateau, while the stress-straincurve of UOE pipes tend to exhibit behaviour closer to a Ramberg-Osgood plasticity model. Significant plasticity introduced duringinstallation (e.g. reeling) will also affect the stress-strain curve.


Guidance note:The use of SMYS and SMTS to characterise the material stress-strain response may underestimate the forces in a restrainedpipeline, and this should be considered.


5.4.3 Pipeline configurationThe development of global buckles (lateral buckling, upheaval or a combination of the two) is affected by thepipeline as-laid configuration. The pipeline should be modelled stress-free in a straight configuration. Hence,actual (measured) or assumed imperfections for triggering relevant buckling modes should be introducedfrom the initial straight and stress-free pipeline configuration. The pipe-soil should also be stress-free for theas-laid pipeline on the seabed with applied lay tension and relevant imperfections.The design should be performed by analysing the full pipeline route and detailed bathymetry to ensureappropriate modelling of OOS features, spans et cetera and pipeline axial restraints (e.g. anchors, graveldumps).


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Guidance note:To assess the first loading and cyclic loading response within lateral buckles, short FE models (VAS models) of a single isolatedbuckle can be employed, with the length defined by the location of the virtual anchors at both side of the buckle. If manysignificant OOS features are identified along the pipeline route, it may be necessary to undertake separate VAS analyses forseveral or all the features, depending on the variability and severity of the features.


If buckle initiators are employed, the buckle response associated with each type of initiator should beassessed. It may be possible to use a single VAS model to demonstrate acceptable loads for all similar buckleinitiators in a pipeline.For pipeline walking and buckle interaction, full length models are required.For an exposed pipeline on even seabed, imperfections in the horizontal plane will be the governingimperfections. These imperfections may be known (e.g. intentional lateral imperfections or route curves) orunknown (residual out-of-straightness following pipeline installation). A larger imperfection than anticipatedmay be required in the FE model to allow a buckle to form for all sensitivity studies in line with the designprocedure.For exposed pipelines on uneven seabed, imperfections in both the horizontal and vertical planes need tobe considered. These imperfections will normally be modelled both in the design phase and after pipelineinstallation. However, an as-laid survey will reduce the uncertainties in the seabed bathymetry.For an exposed pipeline, the as-laid/as-built OOS should be assessed to confirm the assumptions in design.If the observed imperfections do not comply with the design assumptions, the design analyses and theirresults should be confirmed by analysing the pipeline based on as-built survey data, to account for the actualpipeline position (out-of-straightness, vertical features, free spans and pipe - soil penetration).If deviations from the design assumptions with respect to OOS are identified, and acceptance of the bucklingresponse is sensitive to parameters for which there is a high degree of uncertainty (e.g. an uncontrolledbuckling scheme relying on regular buckles forming due to as-laid out-of-straightness features), the designchecks including analyses should be redone based on the actual configuration.Attention should be paid to the prediction of pipeline as-laid configuration, both in the horizontal and verticalplane, as this determines which buckling modes (horizontal, vertical or a combination of the two) will befirst triggered in operation along the pipeline route. A 3D/2½D seabed description is recommended in theanalysis, but on even/moderately uneven seabed it is acceptable to predict the load effect in buckles by a flatseabed model. Such detailed modelling should be based on accurate processing of survey data.The design analysis may incorporate the residual lay tension which is potentially beneficial to the pipelineresponse. However, care is required in the selection of a suitable design value, particularly if the magnitudeof the residual lay tension is significant compared to the fully constrained force.

5.4.4 Pipe-soil interaction modellingPipe-soil interaction is generally modelled as a contact problem or by using a series of independent non-linearspring-like elements attached to the pipeline. These contact surfaces or springs are characterised by a non-linear force-displacement relationship and represent an integration of the normal and tangential forces actingon the pipeline surface interfacing with the surrounding soil. The effect of peak resistance in the pipe-soilresistance modelling should be carefully evaluated.The axial pipe-soil interaction can be idealised in structural modelling for both conceptual and detailed designby an elastic-plastic model that consists of two parameters: the limiting (or residual) axial resistance, and amobilisation deformation which is elastic and recoverable. If required and if fully justified, an initial breakoutpeak can be incorporated using a piecewise linear axial PSI response.Axial walking behaviour is influenced by the axial pipe-soil resistance and the associated mobilisationdistance. A shorter mobilisation distance yields a higher rate of walking. For this reason, the structuralanalysis should model the elastic mobilisation distance which is the stiff and recoverable part of the axialresponse, rather than the plastic irrecoverable part, to avoid to underestimate the walking.Lateral pipe-soil interaction should include the breakout resistance and incorporate the effect of soil bermresistance where relevant.


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5.4.5 Buckle interaction and pipeline walkingThe cyclic expansion behaviour of the pipeline should be analysed by detailed FE analysis. The assessmentshould evaluate:

— buckle interaction and stability including accumulated lateral displacement due to load cycles— route curve pull-out— pipeline walking.

In addition to the requirements to FE modelling previously outlined in [5.4.1], additional aspects that need tobe included are:

— the full pipeline length, with appropriate end conditions including vertical imperfections/slope— the most likely buckling behaviour (number of buckles and their location) described by the design strategy— number of start-up and shutdown cycles (best estimate number of cycles can be used, which may be less

than the number used in the fatigue assessment)— transient temperature profiles associated with start-up, which shall capture the steepest transients that

occur, with appropriate analysis time steps modelled to adequately capture these within the analysis— phasing of pressure loading relative to thermal loading if it may affect the walking behaviour; this is likely

to be of greatest concern if the pressure loading is significant, and— any significant gas/liquid separation effects on shutdown due to seabed bathymetry which result in

submerged weight variations, potentially exacerbating walking.

The analysis should be repeated until the magnitude of axial displacement for each start-up or shut downcycle has reached a steady state, or until the full expected number of cycles has been analysed shouldthe walking response not stabilise. Sensitivity analyses should be undertaken to evaluate the influence ofrelevant parameters, such as pipe-soil friction and soil berms.For complex operating procedures (e.g. full and partial shutdowns, reverse flows etc.), several different cycledefinitions should be analysed to understand the sensitivity to uncertainties in the operational phase.As the length of the interaction/walking models can be large, the modelling requirements previouslyrecommended for buckle response monitoring may be relaxed. The element discretisation can be coarserthan that employed in the local models, and the pipe-soil interaction need not include the breakoutresponse, but should incorporate the effect of soil berm resistance at the extremes of cyclic displacement, ifappropriate.


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SECTION 6 EXPOSED PIPELINE

6.1 Objective and applicabilityThe objective of this section is to provide design procedures and limit state criteria for exposed pipelines inwhich global buckles may form.A pipeline resting on the seabed and exposed to compressive effective force, may buckle globally. Thissection addresses buckling on both even and uneven seabed. Global buckling may initiate in the horizontalplane or in the vertical plane. When initiating a buckle in the vertical plane, the pipeline will lift off at animperfection before rolling over and developing into a lateral buckle. The details of the analyses required forbuckle initiation in different planes may differ but are all addressed by the design procedure outlined in thissection.

6.2 Lateral buckling design procedureThe design procedure for lateral buckling is illustrated by the flow chart in Figure 6-1.

Figure 6-1 Flow chart for the lateral buckling design

The design procedure for lateral buckling consists of the following steps:

1) Establish of basic design parameters - these will typically include pipe geometry and material, pipe-soilinteraction parameters, operating data (including pre-commissioning) and environmental conditions.


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2) Susceptibility to global buckling - the first decision task is to evaluate the susceptibility to globalbuckling. The resistance towards buckling depends on the trigger e.g. trawl impact, lateral restraint andgeometrical imperfection of the global configuration. If the pipeline is not susceptible to global buckling,only the pipeline walking check remains. In case the pipeline is susceptible to global buckling, it shall bechecked for the post-buckling configuration, see point 5 below.

3) Uncontrolled buckling - the uncontrolled buckling check refers to buckling that occurs where no plannedbuckle initiators are implemented. Although the exact location of unplanned buckles may not beguaranteed, it is often possible to demonstrate that unplanned buckling does not threaten the pipelineintegrity. The task is, hence, to check if an unplanned buckle in a random location can sustain theassociated feed-in. If so, the next task will be to check buckle interaction. In case it cannot sustain thefeed-in, an initiation strategy is required.The general principle of global buckling design is to ensure that the pipeline expansion into one buckleis acceptable. Alternative the design needs to ensure that the feed-in is distributed in a safe mannerallowing design criteria to be fulfilled. This is ensured by the following requirement:

(6.1)

where:

VAScharacteristic = the characteristic virtual anchor spacing, i.e. distance between virtual anchor pointsalong the pipeline length

VAStolerable = the tolerable virtual anchor spacing which exactly match the governing designcriterion.

The relevant pipeline limit states/failure modes within a lateral buckle are (see [6.6] and DNVGL-ST-F101 Sec.5):

— local buckling— fatigue— fracture— uniform strain capacity— cyclic plasticity— lateral pipeline displacements restraints imposed by other pipelines, structure close by etc.

The limit state checks required change from case to case and depend on the susceptibility to buckling.This aspect is covered in more detail in Sec.8. The limit state checks also include an assessment ofbuckle interaction and pipeline walking, to ensure that all relevant design criteria are covered.

For calculations of tolerable VAS and characteristic VAS, see [6.4] and [6.5]. Note that both thecharacteristic VAS and the tolerable VAS may change along the length of the pipeline.

4) Initiation strategy - If uncontrolled buckling is found unacceptable, the next design task is to develop aninitiation strategy that triggers (controlled) buckling (or mitigate buckling). The objective of the initiationstrategy is to ensure that:

— buckle initiation measures introduced in the global buckling design have sufficient reliability to ensurethat the characteristic VAS is less than or equal to the tolerable VAS, and

— the pipeline buckles as anticipated/designed for.

If initiation measures are included, the modelling of the global buckling shall include these measures.Subsequent confirmation that the pipeline behaves as anticipated in operation is a part of ensuring arobust design.A variety of initiation measures are available to control the global buckling behaviour, which generallyfall into one of two categories: measures designed to prevent buckle formation, and measures to triggerbuckle initiation. Examples of measures to prevent buckling are:

— increase soil restraint (e.g. bury the pipeline)


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— reduction of the driving force— change type of pipeline (e.g. bundle, pipe-in-pipe).

Examples of measures to initiate regularly spaced buckles are:

— snake-lay— vertical upset— zero-radius bends— local weight reduction— pre-bent sections— buoyancy modules— rock dumping on critical overbend features or between planned buckle sites.

These initiation measures have different implications for the pipeline design. The selection processdepends on many factors such as severity of operating conditions, environmental conditions, seabedbathymetry, water depth, installation vessels, cost and schedule. The key considerations for eachtechnique are discussed in App.A.

5) Controlled buckling - when a buckle initiation strategy is adopted (as opposed to a buckle preventionstrategy), it is necessary to understand how frequently buckles are required. The trigger spacing shallensure that the characteristic VAS is smaller than or equal to the tolerable VAS, ensuring that the feed-into and combined load effects in the planned and unplanned buckles are acceptable.The lateral buckling analysis should be undertaken in accordance with the modelling requirementsgiven in Sec.5. This will usually involve analyses for a range of VAS to allow the sensitivity of thebuckle spacing to be clearly demonstrated. The tolerable VAS should be calculated for all relevantbuckle initiators, and for unplanned buckles. The tolerable VAS is likely to vary depending on the buckleinitiation method adopted.The location of the triggers may also be influenced by sensitive areas of the pipeline. For example,triggers may be employed to prevent buckling at a pipeline crossing or close to a mid-line tie-in location.For uneven seabed, the vertical imperfections will be associated with uncertainties. For this reason, it isimportant to perform some sensitivity checks of the adopted buckle initiation strategy with respect togetting buckles triggered by adjacent imperfections instead of the planned one.If both the planned buckles, and any potential unplanned buckles, are acceptable, then the lateralbuckling design should be checked for pipeline walking and buckle interaction load effects.

6) Improvement of initiation strategy - if any buckles (either planned or potential unplanned betweenplanned sites) are found unacceptable, then the design is unacceptable. If the initiation strategy canbe improved, the optimum strategy should be identified and the assessment repeated. Improvingthe strategy may involve altering the spacing between triggers, optimising the chosen technique (forexample decreasing the radius of the as-laid snakes), or changing the initiation technique.If the initiation strategy cannot be improved, the basic design parameters should be changed and thedesign process repeated.

7) Pipeline walking and buckle interaction - if the pipeline is susceptible to pipeline walking or buckleinteraction, the load effects may increase over time and needs to be checked for the same limit statesas stated under point 3 above. If all buckles are acceptable, the lateral buckling design is complete.Alternatively, the design solution needs to be revised, e.g. with possible addition of anchors or similar, toconstrain the axial movement of the pipeline.


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6.3 Suspectibility to buckling

6.3.1 GeneralAll relevant imperfections that may introduce global buckling should be considered. Relevant imperfectionswill typically include:

— geometrical imperfections:

— uneven seabed— intentional horizontal curves— initial as-laid lateral imperfections (out-of-straightness).

— external loads:

— hydrodynamic loads— trawl interference loads.

Two buckling response categories are identified by the susceptibility check:

1) buckling2) no buckling.

For pipelines exposed to trawling, a third buckling response category is also relevant:

1) maybe buckling.

The pipeline response is to be categorized as buckling if any of the checks indicate that buckling is possible,or if intentional imperfections (or buckle initiators) are employed.

6.3.2 Global buckling initiated by uneven seabedThe resistance against global buckling initiated in the vertical plane is defined as the lower limit for theeffective axial force to activate buckling. Global buckling assessment can be performed either by a simplified2D analysis or by more complex 2½D or 3D analysis, as required to represent the relevant imperfectionsalong the pipeline route. For pipelines with horizontal curves, these shall be included (i.e. a straight pipelinemodel is insufficient).

In case of using simplified 2D analyses in the vertical plane, one of the following checks in Equation (6.2) and(Equation (6.3) shall be performed.

No buckling will occur, if the effective axial force is limited according to:

(6.2)

Contrary buckling will occur, if the effective axial force exceeds:

(6.3)

where:

SR = the effective axial force in the uplifted span sectionpli = the local incidental pressureTl,max = the local maximum design temperature


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EI = represents the bending stiffness of the pipeLuplift = the length of the pipeline lifted off at the free span crests depending on the effective axial force.

If Equation (6.2) (pinned-pinned Euler buckling length) is fulfilled, the uplifted section remains in the planeand 2D analyses are sufficient. If Equation (6.3) (fixed-fixed Euler buckling length) is fulfilled, the upliftedsection will buckle laterally and 2D analyses with an even seabed or a more optimised design by 2½D or 3Danalyses shall be performed. Equation (6.3) is valid only if the section modulus is the same in all directions,i.e. that yielding does not occur prior to lateral buckling.

If the effective axial force is between the limits given by Equation (6.2) and Equation (6.3), a 2½D or 3Danalysis shall be performed to document the global buckling behaviour.

In case 2½D or 3D analyses are applied, the above checks are not required as the susceptibility to, andrelevant buckling response, will be determined by the global FE analyses.

6.3.3 Buckling initiated by lateral imperfections (even seabed)The resistance against global buckling is also here defined as the lower limit for available effective axialforce to activate lateral buckling. In the absence of trawl loads, it depends on the out-of-straightness of thepipeline and the pipe-soil lateral resistance.

Engineering judgement should be applied to avoid unnecessary conservatism with respect to pipe-soilresistance. The initial buckling may for example be governed by an extreme environmental condition (100-year return period), with a corresponding low lateral resistance (due to lift force and drag force), resulting ina favourable buckle configuration. Further increase in pressure and temperature may, however, be based onhigher lateral resistance representing a more normal environmental condition.

The design check for global buckling triggered by an imperfection of a straight pipeline on even seabed isbased on Hobbs' infinite mode capacity. This capacity depends on the lateral pipe-soil resistance which in turnwill depend on the pipeline weight. The estimated weight shall include the lift effect from current and waves.Two combinations of loads shall be included in this consideration:

1) Lower estimate lateral resistance and extreme bottom flow with associated functional loads (typicallynormal operating temperature and pressure).

2) Lower estimate lateral resistance and extreme functional loads with associated bottom flow (typicallynormal or expected daily flow conditions).

Bottom flow velocity shall include contributions from current and waves. It is then allowed to take benefitof the fact that the extreme wave-induced flow velocity and the extreme current does not necessarily occursimultaneously.

The susceptibility of lateral buckling due to hydrodynamic loads can be evaluated based on the followingcritical level of the effective axial force for the infinite buckling mode, S∞, /9/ based on lower estimate pipe-soil resistance (fL

LE):

(6.4)

(6.5)

where:

EA = the axial stiffness of the pipe


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EI = the bending stiffness of the pipefL = the lateral break-out pipe-soil resistance forcefL

LE = the lower estimate lateral break-out pipe-soil resistance forcefL

BE = the best estimate lateral break-out pipe-soil resistancew = the pipe submerged weightFL = the maximum hydrodynamic lift force per unit lengthFD = the maximum hydrodynamic drag force per unit length.

Hence, fL is taken as the minimum of the lower estimate lateral soil resistance and the effective resistance inthe presence on hydrodynamic loads from wave and current with a specified return period.

No buckling implies that the pipeline will develop marginal deviation from as-laid alignment and fulfil:

(6.6)

where:

S(op) = effective axial force based on operational pressure and temperatureS(des) = effective axial force based on incidental pressure and design temperatureS∞(1yr) = the effective axial force for the infinite buckling mode corresponding to the 1-year return

period environmental condition. The factor of 0.65 ensures that the critical buckling force is areasonable lower estimate to finite mode buckles.

S∞(100yr) = the effective axial force for the infinite buckling mode corresponding to 100-year returnperiod environmental condition.

If there are any route curves in the pipeline on even seabed, a lower estimate capacity corresponding to thisradius shall be used for the global buckling evaluation. This shall include the effect from environmental loadsas given in Equation (6.6), but the factor of 0.65 is not required.

(6.7)

where:

SLECurve = the lower estimate critical level of the effective axial force for the route curveRLE = the expected lower estimate radius of the route curveRBE = the expected best estimate radius of the route curve.

For guidance on the RLE and RBE, see App.B.

The no buckling criterion is met if:

(6.8)

where:

S(op) = effective axial force based on operational pressure and temperatureS(des) = effective axial force based on incidental pressure and design temperature


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SLECurve(1yr) = the effective axial force for the route curve corresponding to the 1-year return periodenvironmental condition

SLECurve(100yr) = the effective axial force for the route curve corresponding to 100-year return periodenvironmental condition.

Should the above criterion fail, the pipeline will be within the buckling response category.

6.3.4 Buckling initiated by trawl interferenceTrawl interference triggering global buckling shall be evaluated for a set of trawl pullover loads and pipe-soilresistances. This evaluation will typically need a set of FE sensitivity analyses. The lateral pipe-soil resistanceand the trawl pull-over load, FT, to be applied in the sensitivity analyses are defined by the soil-trawl matrixin Figure 6-2. The matrix implies a maximum of three FE analyses with different combinations of trawl loadand lateral soil resistance forces. The indices UE and LE indicating upper and lower estimate values specifiedtypically as a mean value ± two standard deviations and index BE indicating a best estimate. For moredescription on upper and lower estimates see DNVGL-RP-F114.

Figure 6-2 Lateral soil resistance and trawl load sensitivity study combinations

The assessment is based on FE analyses of three sensitivity study combinations using (FTUE, fL

LE) denoted#1, (FT

BE, fLLE) denoted #2 and (FT

UE, fLBE) denoted #3, and will categorize the buckling response as

follows:

— No buckling response category if global buckling does not occur for combination #1.— No buckling response category if neither of combinations #2 or #3 cause global buckling.— Maybe buckling (SLS/ALS) response category if either combination #2 and #3 experience global buckling

and a single post-global buckling check is required.— Buckling (ULS) response category if both combination #2 and #3 cause global buckling, then a post-

global buckling check with a full soil matrix is required.


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The FE analyses are performed with a straight pipeline and pressure and temperature corresponding to thetrawling frequency defined in Table 3-4. Trawl loads should be included as per Table 3-3 and as per DNVGL-RP-F111. Hydrodynamic forces need not be included in these calculations as they may reduce the lateralpipe-soil resistance due to lift effects.

6.4 Tolerable virtual anchor spacing calculationThe tolerable VAS is defined by the pipeline limit states, and may be governed by local buckling, fracture,fatigue or lateral displacement of the pipeline. To establish the tolerable VAS, a suitable modelling methodshould be adopted (see Sec.5). The tolerable VAS can either be calculated using an iterative process using aseries of VAS models (first load and cyclic limit states), or using a single feed-in model (first load limit statesonly) which is then used to establish the tolerable VAS.The iterative process, valid for all limit states, is shown in Figure 6-3. As seen in the figure there is a checkon whether the tolerable VAS can be increased if the unity check (UC) is below 1.0. However, there is anupper limit on the tolerable VAS, for long pipelines this is normally related to the fully constrained effectiveaxial force, while for short lines it is related to the pipeline length (normally half the pipeline length). If allunity checks are acceptable at this upper limit, then buckling is acceptable, and the tolerable VAS needs notbeing calculated.

Figure 6-3 Design flow for identification of tolerable VAS

The calculations should be performed at enough locations along the pipeline to develop the variation oftolerable VAS with location (KP), although for short pipelines or highly insulated pipelines, the tolerable VASmay not vary significantly, and a single calculation at inlet conditions may be adequate.Two alternative sets of combined loading - local buckling acceptance criteria are provided to assess thetolerable VAS in Table 6-1 below.


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Table 6-1 Two alternative combined loading criteria

Alternative Local buckling criteria

1 1) LCC - load controlled local buckling check for stress-strain curve based on fy and fu, best estimate valuesof pipe-soil resistance, best estimate trawl loads (if relevant) and a condition load effect factor adjustedfor the tolerable VAS.

DCC - displacement controlled local buckling check for worst case (of the analysed scenarios used toadjust the condition load effect factor).

2 2)DCC - displacement controlled local buckling check for stress-strain curve based on fy and fu, bestestimate axial pipe-soil resistance, upper estimate lateral pipe-soil resistance for the tolerable VAS.

1) Both criteria (LCC and DCC) need to be fulfilled.2) For even and moderately uneven seabed, no trawling interference.

A seabed, or part of a seabed, can be defined as moderately uneven if the load controlled combined loading(LCC) according to DNVGL-ST-F101 Sec.5, for the vertical plane is less than or equal to the following twocriteria:

— (DNVGL-ST-F101 safety class low) in water filled condition, and— (DNVGL-ST-F101 safety class medium) under maximum compression effective axial force prior to

initiation of global buckles.

The condition load effect factor of γc of 1.07 shall be used in these local buckling calculations. If any of thesecriteria are violated, the seabed is defined as uneven.

Guidance note:The criteria for defining a moderate uneven seabed have been established based on evaluations of pipelines with different D/tratios and diameters, for seabed varying from fairly flat to (extremely) uneven. When the criteria for moderately uneven seabedare violated, it is due to presence of significant free spans and/or significant vertical imperfections. The criteria for definingmoderately uneven seabed is for seabed and not for vertical triggers.


An example of calculation of tolerable VAS is given in App.D.

6.5 Calculation of characteristic virtual anchor spacing

6.5.1 Overview of methodologyThe initiation of buckles is sensitive to the imperfections. As the initial out-of-straightness (OOS) of theplanned pipeline is not known prior to pipeline installation, there will always be an inherent uncertaintywith respect to the buckling response of the pipeline. This uncertainty should be quantified and if foundunacceptable, it is either reduced to an acceptable level for the design to proceed, or a design solution whichis more robust with respect to this uncertainty should be developed.As the distance between planned buckle locations is reduced, the likelihood of triggering all planned bucklesalso reduces. This may cause that buckling is not occurring at some of the planned locations. This againmeans that the robustness of the solution reduces, and the distance between buckles which do develop willbe longer than designed for. This aspect needs to be considered in the overall buckling design. The designmethodology addresses this aspect through the characteristic VAS, which is the design value of VAS thatcannot be exceeded to reach an acceptable level of reliability for the buckling design.The following methods are available for calculating the characteristic VAS:

— deterministic definition [6.5.2]

— buckling limited by pipeline strength


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— buckling limited by fully restrained effective axial force— sharing criterion

— probabilistic definition [6.5.3].

If an uncontrolled buckling strategy is proposed based on a probabilistic definition of characteristicVAS, deterministic analyses to define the characteristic VAS should also be performed to obtain a betterunderstanding of the uncertainties, including the pipe-soil interaction and out-of-straightness distribution(and thus the critical buckling force distribution).

6.5.2 Characteristic virtual anchor spacing - deterministicThe deterministic characteristic virtual anchor spacing, VASCharacteristic, is given by:

(6.9)

where:

VASLength = the characteristic VAS limited by the pipeline lengthVASSO = the characteristic VAS limited by the effective axial force envelopeVASSharing = the characteristic VAS limited by sharing of feed-in by neighbouring buckles.

Whether the characteristic VAS is limited by the pipeline length or not depends on the end conditions. For apipeline with ends free to expand, the characteristic VAS is limited to a maximum of half the pipeline length,while for pipeline with fully fixed ends, it is limited by the pipeline length.

The characteristic VAS limited by the effective axial force envelope can be calculated from:

(6.10)

where:

SO = the fully restrained effective axial force= the lower estimate of the post-buckling effective axial force in the buckle

= the lower estimate of the axial pipe-soil resistance.

The characteristic VAS limited by sharing of expansion with adjacent buckles is calculated from:

(6.11)

where:

= the upper estimate of the critical buckling force

= the lower estimate of the post-buckling effective axial force in the buckle.


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The upper estimate of the critical buckling force should be calculated for both planned and unplanned bucklesand is the maximum of:

(6.12)

where:

SCR() = indicates the calculation of the critical buckling force associated with different imperfection radii andlateral soil resistance

RBE = the best estimate imperfection radiusRUE = the upper estimate imperfection radius

= the upper estimate imperfection of the lateral pipe-soil resistance

= the best estimate imperfection of the lateral pipe-soil resistance.

For guidance on imperfections for straight lines and for critical buckling forces, see App.B.

Guidance note:No analytical expression for the post-buckle force is given. Therefore this force need to be assessed by FE analysis or based on aconservative estimate from similar cases.


If a planned buckling strategy is adopted, the following inequality should prevent unplanned bucklingbetween triggers:

(6.13)

where:

Xtrigger = the buckle spacing= the lower estimate of the critical buckling force for an unplanned buckle

= the upper estimate of the post-buckle force in the planned buckle

= the upper estimate of the axial pipe-soil resistance.

Guidance note:Failure to meet the inequality in Equation (6.13) for trigger spacing does not necessarily imply that unplanned buckling will occur.Furthermore, unplanned buckling between planned triggers may not prevent the formation of planned triggers. However, theuncertainties involved in such a scenario cannot be rigorously considered deterministically. The presence, location and post-buckling parameters will vary with pipe-soil interaction and residual out-of-straightness. Therefore, no deterministic criteria isincluded to assess this. Probabilistic methodologies may provide additional confidence in the reliability of the design strategy.The trigger spacing should be no more than the minimum of the planned buckle tolerable VAS or twice the unplanned buckletolerable VAS. This assumes that all the triggers initiate buckling such that the maximum feasible VAS for an unplanned buckleis half the distance between triggers. This strategy will be adequate if the probability of some of the triggers failing to initiatebuckling is low.If the initiation technique reduces the loading within the buckle, for example by distributed buoyancy, the tolerable VAS at thetrigger may be much greater than that for unplanned buckling. The definition of trigger spacing may then be driven by the capacityof unplanned buckles.Guidance on developing critical buckling force distributions for planned triggers is given in App.B.



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6.5.3 Characteristic virtual anchor spacing - probabilisticThe probabilistic characteristic VAS is the VAS whose probability of exceedance is as defined in Table 6-2.

Table 6-2 Allowable exceedance probabilities for characteristic VAS

Unplanned buckles

Full pipeline length 1) Per kilometer of pipelinePlanned buckles

uncontrolled buckling strategy 10% 1% n/a

controlled buckling strategy 10% 1% 10%

1) If there are no engineered triggers, the full pipeline length shall be taken as the total pipeline length. If there areengineered triggers, the full pipeline length shall be taken as the total length of pipeline between initiators.

2) Note that these values are strictly speaking only valid for pipelines with internal overpressure as the structuralreliability analyses have not included external overpressure. In case of external overpressure special considerationsneed to be done.

Guidance note:Planned buckles in this context refer to those forming at locations where specific buckle initiation measures are adopted (e.g.vertical upsets, additonal buoyancy, snake-lay). The uncertainties associated with buckle formation on such engineered buckletriggers are generally much lower than those associated with unplanned buckles, and as such the exceedance probability definingthe characteristic VAS is greater, whilst still achieving target failure probabilities. However, for a well-designed buckle initiationstrategy where all buckle initiators are demonstrably reliable, there is unlikely to be a significant difference between the VASassociated with exceedance probabilities of 1% or 10% (in this case the characteristic VAS will typically be close to the triggerspacing).For an uneven seabed, it may be possible to rely on the formation of buckles at natural seabed features rather than by installingother engineered buckle initiators. In order to employ a probabilistic criterion for characteristic VAS, the buckle formation analysisshould incorporate all relevant effects of the uneven seabed, including the potential variation in imperfection magnitude, spans etc.within the pipeline lay corridor. It is recommended that a 1% exceedance criteria is used, albeit for a significant feature there maybe little difference between the 1% and 10% exceedance probability.


Guidance note:The pipeline is not susceptible to buckling in any region if the probability of buckle formation is less than 1%/kilometre pipeline(i.e. less than 1% in any given kilometre of the pipeline). This can be used to define the section where the pipeline is susceptible tobuckling, e.g. for pipelines in which the temperature and/or pressure reduces significantly along its length.The methodology does not include any specific requirement for the reliability of planned buckling at the engineered initiators.However, high levels of reliability of buckle initiation are generally required to achieve a system that meets the requirements of thedesign method. If either the characteristic VAS of planned initiators is greatly in excess of the initiator spacing or the characteristicVAS of unplanned buckles is large, consideration should be given to improving the reliability of the engineered buckle initiators.


When no buckles occur at a given location (either planned or unplanned) for a given simulation in theprobabilistic assessment, the VAS is zero to allow the probability of no buckling to be identified.Unplanned buckles can be defined as either an exceedance probability for the full pipeline length or perkilometre pipeline. The latter generally allows a more refined and optimised solution.To define, and subsequently use, a probabilistic definition of characteristic VAS, a structural reliability analysisof buckle formation shall be performed. Guidance on developing a probabilistic model of buckle formation,and critical buckling force distributions for unplanned and planned buckles, are given in App.B. Given theuncertainties in the actual critical buckling force distributions, sensitivity analyses should be performedto understand the criticality of these distributions. The chosen distributions for vertical imperfections andengineered buckle triggers should be assessed during a detail design phase to confirm that they are realisticfor the specific case.


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6.6 Relevant limit states - exposed pipelines

6.6.1 GeneralGlobal buckling is not a failure mode as such, but may cause other failure modes to develop such as localbuckling, fracture and fatigue, and excessive displacements. Hence, after the global buckling categorisation(see [6.3]) and combined loading check, the pipeline should be checked for other failure modes. The checksto be performed depend on the design methodology and the buckling response categorisation of the pipeline.The checks to be performed are depending on the buckling response categorisation:

— no buckling/maybe buckling - ALS check in the buckled (post-buckling) configuration— buckling - ULS check in the buckled (post-buckling) configuration.

All limit states should be checked for pressure and temperature values corresponding to the trawlinterference frequency as per Table 3-4.An overview of the required pipeline limit state checks is given in Table 6-3. See also [6.7.3] for additionalrequirements for pipelines on uneven seabed. Guidance on each of the limit state checks can be found inSec.8 (Table 6-3 presents an overview of the sensitivity cases for exposed pipelines), or in the relevantdesign standards (e.g. DNVGL-ST-F101, DNVGL-RP-F111, DNVGL-RP-F105, et cetera).The limit state checks given in Table 6-3 represent typical cross-sectional failure modes. In addition excessivedisplacement, both laterally (contact with other structures/pipeline/objects) and axially (walking and endexpansion), needs to be checked.

Table 6-3 Overview of the relevant pipeline limit states for the different buckling responsecategories - exposed pipelines

Local bucklingBucklingresponsecategory Load controlled Displacement

controlled

Uniform straincapacity

Ratcheting/cyclic plasticity Fatigue Fracture

see [8.2]/DNVGL-ST-F101

see [8.2]/DNVGL-ST-F101

Equation (8.12)Equation(8.17)/ DNVGL-ST-F101

DNVGL-ST-F101, DNVGL-RP-C203DNVGL-RP-F105

DNVGL-ST-F101

no buckling/maybe buckling(post-bucklingcondition)

MBE with γc =0.851

n/a n/a n/a M BE M BE

buckling(alternative 1)

MBE with

calibrated γc 2)

all sensitivitiesincluded in theγc calibration tobe checked, M1… M7

all sensitivitiesincluded in theγc calibration tobe checked, M1… M7

M 3 M 3 M 3

buckling(alternative 2)3)

n/a M 3 M 3 M 3 M 3 M 3


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1) Recommended to be analysed as an accidental limit state. For an explanation MBE see Table 6-4.2) Typically ranging from 0.80 to 1.0.3) Note that alternative 2 (local buckling check) is only valid for cases on even and moderately uneven seabed with no

trawling interference, see Table 6-1.

This table does not rule out other checks required by DNVGL-ST-F101, DNVGL-RP-F111, DNVGL-RP-F105 etc.

Guidance note:For pipelines on even and moderate uneven seabed both alternative 1 and alternative 2 may be used to check local buckling aslong as there is no trawling intereference. These two alternatives may give different tolerable VAS. It is therefore possible to checkboth for optimisation of design.


The condition prior to any buckling shall be checked against the axial capacity (ULS).The no buckling response category shall be checked as a standard pipeline not undergoing any globalbuckling, including the axial wrinkling failure mode, as per Table 6-3.It is recommended to evaluate the post-buckling condition of the pipeline as an ALS case as per Table 6-3 forthe no buckling response category to evaluate criticality.

Guidance note:By modelling realistic imperfections and increasing their size, buckling of the no buckling/maybe buckling response categories maybe triggered. In case the pipeline design criteria are exceeded in the post-buckling condition for such artificially triggered buckles,additional measures to ensure a robust design may be required. This may be most relevant for pipelines with high D/t ratio.


The post-buckling condition for the maybe buckling response category (trawling) shall be checked as an ALScondition based on best estimate lateral resistance, best estimate axial resistance and best estimate trawlloads. A simplified ALS check can be carried out assuming yc equal to 0.85 and yF = 1.0 in line with DNVGL-ST-F101. As an alternative, the design methodology of the buckling response category given in Table 6-3 canbe applied.Table 6-4 presents an overview of the sensitivity cases for exposed pipelines referred to in Table 6-3 andTable 6-5 and in addition used in Sec.9 to calculate the condition load effect factor γc.

Table 6-4 Overview of sensitive cases for exposed pipelines

Case ID Pressure Tempe-rature

Lateralresistance

Axialresistance

Materialstrength

Trawlload Comment

MBE P T fy FT for load controlled local buckling limit state

M1 P T fy FT for calibration of condition load effect factor, γc


M3 P T fy FTfor displacement controlled local buckling limitstate, and fatigue and fatigue limit states


M5 P T σy FT for calibration of condition load effect factor, γc

M6 P T fy for calibration of condition load effect factor, γc


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Case ID Pressure Tempe-rature

Lateralresistance

Axialresistance

Materialstrength

Trawlload Comment

M7 P T fy for calibration of condition load effect factor, γc

6.6.2 Local buckling - post buckling checkA set of non-linear FE analyses is required to document the pipeline integrity in the post-buckling condition.By varying the basic design parameters, a set of possible responses in terms of bending moments and strainsis found.

For the load controlled assessment in alternative 1, the purpose of the required analyses for the bucklingconditions is therefore to determine a specific γc in line with the Sec.9 giving a γc larger than 0.80. A γc of1.0 will correspond to a fully load controlled condition for best estimate values which may not give sufficientsafety. In case of a γc above 1.0, it is recommended to re-assess the design solution.

Guidance note:

A γc above 1.0 means that the design is very sensitive to some of the input parameters, typically the pipe-soil interaction of trawlgear interaction. The design solution should be re-assessed considering for example more buckle triggers and/or rock berms orsimply burial of the pipeline.


The pipeline design check for local buckling, see Figure 6-2, is satisfied if one of the alternative sets ofcriteria, alternative 1 or 2 (see Table 6-1) is fulfilled.

The design bending moment, MSd, is defined as:

(6.14)

where:

Mf(fLBE, fABE, fy, FT

BE) = the functional bending moment based on given pipe-soil parameters, yield stressand trawl load

γf = the functional load effect factor = 1.1γc = the condition load effect factor as per Sec.9.

The design bending strain, εSd, is defined as:

(6.15)

where:

εf(fLUE, fA

BE, fy) = the functional strain inclusive of any relevant SNCF [5.4.2] and [5.5.2]γf = the functional load effect factor = 1.1γc = the condition load effect factor = 1.0.


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For alternative 1, the characteristic functional bending strain is the worst case analysed in the determinationof γc . The worst case for trawling (which shall be checked in addition to the load controlled local buckling)with respect to the design strain is normally one out of:

(6.16)

6.6.3 Pipeline walkingAs for global buckling, pipeline walking is not a failure mode in itself. However, significant axial displacementscan lead to overstressing and subsequent failure of pipeline risers, jumpers or spools. The incremental axialdisplacement associated with each start-up cycle should be evaluated. The total axial displacement should becalculated based on the best estimate, rather than the upper estimate, of the number of start-up/shut-downcycles anticipated over the design life. This number of cycles does not need to be the same as that assumedfor low cycle fatigue design. Historical information from similar projects can be used to assist in quantifyingthe appropriate number of cycles.The behaviour of the pipeline is acceptable if:

— The total axial displacement over the design life of the pipeline is within the design capacity of anypipeline connections, such as jumpers or spools, or to the connection to a SCR, and any associatedmechanical connectors. In addition, it needs to be within the design displacement limit of any sliding in-line structure or end structures such as PLETs/PLEMs.

— The load effects within any lateral buckle remain acceptable.— Any anchors are loaded within their acceptable limits.— Route curve pull-out does not occur or any movement is shown to be acceptable so the walking response

is tolerable, and the route curve stabilises within a reasonable number of cycles.

If the pipeline connections or lateral buckles are not able to tolerate the total axial displacement developedover the design life of the pipeline, remedial measures should be developed to control the walking response,for example anchoring.There is significant uncertainty associated with the pipeline walking estimates. Consequently, an acceptableapproach is to recognise the potential for walking, but to delay significant remedial measures until the actualbehaviour of the pipeline has been identified. This approach shall be integrated into the pipeline integritymonitoring system (see [10.4]).Given the model significant uncertainty associated with pipeline walking predictions, there is low confidencein the predication at low rates of walking. If many cycles combined with a very low walking rate percycle leads to a prediction of a very significant movement over the design lifetime, it may be particularlyappropriate to implement monitoring procedures rather than pre-invest in expensive mitigation techniques.The monitoring procedures should be quantifying the pipeline behaviour in a way that allows mitigationmeasures to be implemented, as necessary. In addition, it should be identified during the pipeline designphase how any operational mitigation could be implemented, for example, suitable anchor attachment pointsmay be incorporated into the pipeline design to ease axial anchoring in later phases.

6.7 Additional considerations for uneven seabed

6.7.1 Design aspectsA pipeline on an uneven seabed may have a more defined imperfection, in the vertical plane, than a pipelineon an even seabed (where the imperfections will be the residual lateral out-of-straightness following pipelineinstallation).


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On uneven seabed, pipeline expansion can be absorbed by vertical deflection in spans. As such, a pipelinethat has been designed for an even seabed will generally be acceptable also on an uneven seabed withoutfurther evaluations of the global and local buckling aspects if:

— the moment contribution from the vertical plane due to the unevenness is negligible— the buckle mode shape is unaffected by the vertical pipeline shape and— the feed-in length to each buckle should be equal to or less than that of an even seabed case.

The presence of the uneven seabed should be considered for the buckle formation assessment (seefurther discussion in App.B). It is particularly important to check the sensitivity of the buckle formation toneighbouring vertical imperfections, with and without buckle mitigation measures. Configuration uncertaintyshould be included. Note that even if a global buckling design on even seabed includes mitigation measures,the need for these will be reduced or even eliminated on uneven seabed.The design procedure in this section [6.7] gives further guidance with respect to expansion capability.Standard design checks with respect to free spans and other failure modes shall be carried out on the actualtopography, as per guidance in relevant standards such as DNVGL-ST-F101, DNVGL-RP-F105 and DNVGL-RP-F111.

6.7.2 Global bucklingA pipeline resting on an uneven seabed may experience the following three types of global response:Type 1: The pipeline in free spans deflect and may touch the seabed.Type 2: Uplift at crests

— The pipeline expansion may lift the pipeline off at a few crests in a limited way: typically less than 50% ofthe pipe diameter over a length of less than 50 pipe diameters, causing the lifted off section to be stableagainst buckling in the lateral direction.

— Even further expansion of the pipeline may increase the upward displacements at crests, it buckleslaterally and develops lateral buckle in the horizontal plane.

Type 3: Further expansion will then increase the bending of the buckle.An important practical task in the design is to limit the complexity of the analyses, by use of 2D, 2 ½ D or3D models when possible. Specific design checks may therefore be required to prove the relevancy of thesimplified models.

6.7.3 Relevant limit states - pipelines on uneven seabedThe limit state checks in the post-buckling condition for a pipeline on uneven seabed follow the sameprinciples as for even seabed in [6.6].Table 6-5 groups the limit states corresponding to the three types of global response listed above [6.7.2]:

— Type 1 - traditional design where the limit state checks relevant for free spans and pipeline in contact withthe sea bottom not subjected to uplift, lateral turn down and subsequent lateral buckling are applied.

— Type 2 - uplifted pipeline without buckling, still the traditional limit state checks relevant for upliftedpipeline sections prior to lateral buckling are applied.

— Type 3 - laterally buckled pipeline sections. Here, the relevant limit state checks are for pipeline sectionsin contact with the sea bottom, subjected to lateral buckling after uplift and lateral turn down.


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Table 6-5 Required pipeline limit state checks - uneven seabed

Local bucklingType ofpipelineresponse Load controlled

Displace-mentcontrolled

Uniformstraincapacity

Ratcheting/cyclicplasticity

Fatigue Fracture Free span

see[8.2]/DNVGL-ST-F101


Equation(8.12)

Equation(8.17)/DNVGL-ST-F101

see[8.6]/DNVGL-ST-F101,DNVGL-RP-C203


DNVGL-RP-F105

type 1 - nouplift M1

1) M1 2) M1 M1 M1 M1 M1

type 2 -uplift

n/a n/a n/a n/a n/a n/a n/a

no lateralturn down M1

3) n/a M1 M1 M1 M1 M1

prior tolateral turndown

M1 4) n/a n/a n/a n/a n/a

type 3- lateralbuckling

MBE withcalibrated γc 5)

allsensitivitiesincludedin the γccalibrationto bechecked M1... M7

M3 M3 M3 M3 M3

1) The load effect factors are depending on loading scenario, i.e. in a free span or on the seabed, see DNVGL-ST-F101.For an explanation of M1 see Table 6-4.

2) The strain corresponding to M1 is applied in the displacement controlled combined loading – local buckling designcriterion.

3) Normal design applies, i.e. with γc = 1.07.4) The cross-section should not have experienced any yielding prior to buckling laterally, i.e. σeq < fy.5) > 0.80, see Sec.9.

In addition to the above limit states, pipeline walking shall be considered.For the alternative 1 combined loading criterion, three different levels of analyses may be applied forcalibration of the condition load effect factor, γc, to minimise the effort required whilst still enabling designoptimisation (only one level necessary):

— Level 1The pipeline limit state check is based on an analysis of the pipeline resting on the seabed for bestestimate pipe-soil properties, and a condition factor γc, Flat for the pipeline on a flat seabed.

— Level 2The pipeline limit state check is based on an analysis of the pipeline resting on the seabed for bestestimate pipe-soil properties and a condition factor for the pipeline on a flat seabed, adjusted for the axialpipe-soil resistance sensitivity (XA).


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Calculation of new load condition factors for the uneven seabed scenarios, γc, Uneven, Lat, are based oncorrecting the load condition factor for the flat seabed, γc, Flat, in order to consider the effect of free spansin the near-buckling conditions adjacent to the buckle. Note that for level 2, the tolerable VAS can beadjusted compared to the flat seabed scenario. The calculated condition factor is in principle only validfor the buckle location where it is calculated, i.e. new updated expressions for CoV(XA) for each bucklelocation should be calculated.Two additional 3D FE analyses of the pipeline laid on the uneven seabed, with relevant mitigationmeasures, are required to calculate the effect of the free spans adjacent to the buckle. These analysesare used to calculate a new coefficient of variance for the equivalent axial resistance factor (CoV(XA)),as specified in Sec.9. The condition factor for uneven seabed, γc, Uneven, Lat , is found by replacing theexpression of CoV(XA), in the calculation of, γc, Flat.

— Level 3The pipeline limit state check is based on both sensitivity analyses (to determine γc) and pipelineresponse analysis of the pipeline on the uneven seabed.

A new load condition factor, γc, Uneven, Lat , is calculated using a full 2½D or 3D FE analysis (consideringand not considering , respectively, the real 3D roughness of the sea bottom). The calculated conditionfactor is in principle only valid for the buckle location where it is calculated, i.e. new condition factorsshould be calculated for each buckle location.The effect of the soil resistance matrix and free spans are analysed using 2½D or 3D FE models. Level 3 isrecommended in cases where the pipeline configuration in the lateral plane is affected by the sea bottomroughness transverse to the pipeline route. This level is also relevant where pressure loads and thermalexpansion loads are released by a combination of pipeline uplift/turn-down/lateral buckling at the crests ofthe most pronounced undulations. It is also relevant for lateral buckling developing at the built-in curvesor along the transversal down slope.

The formal design criterion then becomes identical to that of an even seabed, applying the calculated γcfor the best estimate parameters.


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SECTION 7 BURIED PIPELINE

7.1 GeneralA buried, axially restrained pipeline exposed to pressure and temperature develops a compressive effectiveforce. An out-of-straightness imBurieperfection along the pipeline length exposed to this compressiveeffective axial force may result in global buckling in the vertical plane or the horizontal plane, if the soilsurrounding the pipeline has insufficient resistance to resist this upward or lateral driving force. The verticaltype of global buckling of buried pipelines is referred to as upheaval buckling (UHB).The objective of this section is to provide design procedures and criteria for upheaval buckling design ofburied pipelines. The criteria ensure that upheaval buckling is avoided by calculating sufficient soil cover tokeep the pipeline in place (UHB soil limit state) and that the pipeline itself also remains intact (pipeline limitstates).

7.2 Upheaval buckling design procedureUpheaval buckling design is first performed in the pre-lay phase, before pipeline installation and burial whenthe actual profile of the pipeline is not known, and later the final design check is performed during the post-lay phase using the actual surveyed pipeline profile and cover height.The design processes for upheaval buckling in both the pre-lay and the post-lay phases are outlined in theflow charts shown in Figure 7-1. Note that the two-coloured boxes in the detailed UHB design are detailed inFigure 7-2.

Figure 7-1 Flow chart for pre-lay and post-lay UHB design


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For both the pre-lay and post-lay assessment, the upheaval buckling design is performed for two differentassumed imperfections:

— minimum cover - based on assumed propped imperfections not detected due to the inaccuracy of thesurvey, and

— specific cover - based on the assumed set of propped imperfections (pre-lay) and the surveyed profile(post-lay).

Flow charts illustrating the main steps in determining the minimum cover and the specific cover, respectively,are shown in Figure 7-2.

Figure 7-2 Design process - minimum and specific covers

Survey methods used to determine the configuration of pipelines including the imperfections are associatedwith uncertainties. A minimum cover resistance shall therefore be determined by analyses of out-of-straightness imperfections which are normally not detected by the configuration survey. Required soil/rockcover resistance should be calculated using FE analysis for these imperfections.The design process to determine specific design cover requirements consists of three major parts:

1) an out-of-straightness (OOS) analysis shall be performed to obtain the curvature along the pipeline onthe seabed

2) propped shape analysis shall be conducted with imperfections covering the range of curvatures obtainedfrom the OOS analysis, and

3) based on a comparison of the curvature along the pipeline obtained from the OOS analysis to thecurvature of the analysed propped shapes, proposal of soil/rock cover can be assigned along thepipeline.


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To ensure that this assigned cover provides sufficient capacity towards UHB, an analysis shall be conductedwith the assigned cover where the temperature is increased until the soil fails. If the failure temperatureis less than the design temperature (see Equation (7.2)), the cover capacity at the failure location shallbe increased. The analysis shall then be re-run with the updated cover properties. This iterative process iscontinued until the failure temperature exceeds the design temperature.The final required cover height shall be the maximum of the calculated specific cover height and theminimum cover height.The methods and objectives of the analyses for the different phases are summarised in Table 7-1.

Table 7-1 Purpose of the different assessments described in this section for buried pipelines

Phase Analysis method Objective Limit states Geometry

pre-lay analytical formula(based on proppedshape)

to estimate anindicative cover heightand total volume ofcover in early phases

pre-lay FE analysis (based onpropped shape)

to estimate a moreaccurate cover heightand total volume ofcover

soil limit states 1) propped shape heightequivalent to assumedsurvey accuracystandard deviation 2)

assumed proppedshape modelimperfection heightsand distribution 3)

to estimate therequired cover heightthat will vary alongthe length of thepipeline

soil limit statespre-lay (no trenchscenario)

FE analysis (based onpre-lay survey data)

to ensure sufficientwall thickness andsteel grade

pipeline limit states

propped shape heightequivalent to assumedsurvey accuracystandard deviation 2)

pipeline route pre-laysurvey data 3)

post-lay FE analysis (based onas-laid/as-trenchedconfiguration)

to determine therequired cover heightthat will vary alongthe length of thepipeline

soil limit statespipeline limit states

propped shape heightequivalent to surveyaccuracy standarddeviation 2)

pipeline as-laid surveydata 3)

1) For soil limit states see [7.6].2) For minimum cover assessment.3) For specific cover assessment.

7.3 Effective axial forceSimplified calculations should always be used to confirm the detailed calculations and to avoid gross errors.Within the anchor zone close to the pipeline end, the effective axial force is reduced from maximum, thetotal restrained axial force, So, due to end expansion. The reduction of the axial force along the pipeline isgoverned by the axial friction between the pipeline and the soil. Hence, a high resistance will give higherforces close to the end. An upper estimate axial friction shall be applied for buried pipelines.

For the upheaval buckling analyses, the axial friction in the anchor zone shall be increased with γUF. This willresult in a shorter effective anchor length.


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7.4 Analytical methods for up-lift assessmentThe soil limit state is based on determination of a sufficient cover height to prevent upheaval buckling and tokeep the pipeline in its original position.

As for lateral buckling, there are also analytical methods available in the literature for upheaval bucklingcalculation, see /11/:

(7.1)

where:

So = the fully restrained effective axial forceδ = the (propped shape) imperfectionwp = the submerged weight during operationEI = the bending stiffness.

The above equation along with other analytical methods have their limitations, and should preferably be usedonly in the conceptual design phase. Typical limitations are:

— only linear elastic material behaviour— difficult to describe an arbitrary imperfection shape— soil upheaval resistance is assumed along the entire imperfection wave length. This is not the case in sag

bend regions, where the pipeline will tend to move downwards, resulting in no soil contribution to upliftresistance.

— does not take the vertical soil resistance force-displacement curve into account— cannot account for cyclic loading and possible creep.

7.5 Detailed finite element analyses

7.5.1 GeneralDetailed UHB pipeline response should be analysed using non-linear finite element methods. The FE analysisshall describe the physical phenomena and behaviour adequately. The purpose of the FE analysis is to find atwhich temperature of the content the soil around the pipeline fails.This section presents additional requirements for UHB to those presented in [5.5]:

— pipe-soil interaction - see [4.3]— initial pipeline profile - The development of the upheaval buckling mode is affected by the pipeline as-

laid vertical profile or assumed propped shape size. The pipeline should be stress free in a straightconfiguration. The actual survey profile, representing the as-laid configuration, shall be used as outlined in[7.5.2].

— effective axial force build-up - see [7.3]— upheaval buckling analysis - requires finding an applied (TRd) temperature where the pipe-soil interaction

fails. The pipeline fails for UHB when the axial loading cannot be increased or the mobilisation of the soilexceeds the failure displacement δF.

7.5.2 Modelling propped shape configurationAnalysis of a propped shape imperfection serves three different purposes:

— pre-lay assessment of gravel volume:


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— The propped shape heights used for the preliminary gravel volume estimate will typically be 0.1, 0.2and 0.3 m. Some additional imperfections may be required based on project specific needs.

— minimum cover design:

— The imperfection height shall be taken as δf = σconfiguration where σconfiguration is one standarddeviation of the configuration survey accuracy. This is identical in the pre-lay and post-lay scenario.

— initial estimate of cover height with the actual pipeline profile from post-lay survey:

— The purpose of this analysis is to use a simple model to determine the cover height where curvaturesof the real profile are compared to the curvature of propped shape imperfections. A range ofimperfections heights covering the relevant curvature range is required.

For all the propped shape imperfections, the pipeline section is first laid on a flat seabed, the initial load stepsuntil flooding are applied, and then the pipeline is lifted in the sections centre point to the given proppedimperfection height. An assumption regarding size and number of imperfections should be made to estimaterequired cover volume for the pre-lay assessment since no survey data of the seabed is available.

Guidance note:Unless other information exists, the following can be used:

— 4 propped shapes of 0.1 m height per kilometre

— 4 propped shapes of 0.2 m height per kilometre

— 2 propped shapes of 0.3 m height per kilometre.


It should be noted that larger imperfections and higher number of imperfections can be revealed whensurvey data are available. Therefore, attention should be made to ensure that all imperfections are coveredso that these do not impose risk of failing required uplift resistance when the pipeline is put in operation.

7.5.3 Modelling as-laid/as-trenched configurationA central part of upheaval buckling design for the installed and buried pipeline is measurement of the pipelineprofile, the actual cover height and the mean seabed profile and trench geometry, if relevant.

For buried pipelines, the vertical out-of-straightness is the most important. During the design phase, theimperfections will be unknown, whereas after pipeline installation, the out-of-straightness can be measuredand hence, the imperfections will be known with a certain accuracy.

The survey uncertainty may be significant. This means that the vertical position of the pipeline is associatedwith a measurement error typically in the order of 0.1 - 0.3 m (±2 σconfiguration) with a certain spatialcorrelation. An estimate of the survey accuracy should therefore be made for a set of independentmeasurements, by calculating the standard deviation for each measurement. This is normally not performed,but based on experience and engineering judgement the accuracy for ROV based surveys the following canbe used:

σconfiguration = 0.05 - 0.15 m for top of pipeσcover = 0.10 - 0.15 m for cover height, H

Guidance note:Pipelines subjected to low temperatures, low pressures and requiring a minimum cover height for protection against e.g. 3rd partyloads may not be susceptible to upheval buckling. The requirements for survey accuracy for such pipelines can be re-assessed.


Survey data is normally given as data listings related to position along the pipeline (KP). Average spacingbetween each measurement should be in the order of a couple of diameters but not larger than one meter.


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Guidance note:For large diameter pipelines longer distance between measurements than 1.0m can be reasonable, however, it is judged that thereis limited benefit if increasing the distance between survey points.


Post processing of configuration survey data to transform the measurement to data listing should beperformed, but no smoothing of data should be carried out by the contractor performing the survey. Thedesigner should carry out the smoothing, ensuring a consistent treatment of the raw data, by using thefollowing procedure:

1) Clearly un-physical data shall be removed. This is single points clearly deviating in an unphysical mannerfrom the other data.

2) Representative soil stiffness is applied to the remaining survey points.3) A FE analysis with a straight stress-free pipeline in the as-laid or trenched conditions (i.e. normally laid

empty pipeline on seabed, changed to water filled) is lowered down on the soil springs (as a contactproblem) without any axial friction applied. Alternative ways of simulating this is allowed and maydepend on the FE software’s ability to model the contact problem.

4) In this configuration, the soil springs not in contact with the pipeline are connected in a zero-stressstate simulating soil surrounding the pipeline. Note that this should be performed with great care asactual gaps between the pipeline and seabed will be removed by this process. Actual gaps should notbe removed as this will lead to underestimation of the required uplift resistance. Engineering judgementshould be applied.

5) Engineering judgement should be applied to the obtained configuration of the pipeline versus the surveydata, and comparisons towards a mathematical smoothened profile should be performed to confirmadequate imperfection sizes/pipeline curvature.

7.6 Upheaval buckling soil limit states

7.6.1 Methodology and acceptance criteriaUpheaval buckling may be caused by local soil failure and/or global soil failure. Local failure is only relevantfor cohesive soils (clay). The failure modes are identical for both minimum cover design and specific coverdesign, although the applied safety factors are different. All relevant soil failure modes shall be assessed, andthe worst one considered unless the safety against failure can be documented otherwise. This is for examplerelevant when drained and undrained soil behaviour can be of concern.

The safety is ensured by a load effect factor on the effective axial load, γUF . To account for the pressurecontribution to the effective axial load without influencing on the hoop stress, the design load equivalenttemperature, TSd , includes a pressure term. Further, conservative soil resistance combined with safetyfactors are applied to ensure a safe margin from failure.

The local soil failure mode is given by DNVGL-RP-F114.

The global soil failure is avoided when the design load equivalent temperature (TSd) is smaller than thedesign resistance equivalent failure temperature (TRd):

(7.2)

The design (load) equivalent temperature (TSd) is given by:

(7.3)


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where:

= the load effect factor on the effective axial load, = 1.0, 1.15 or 1.30 for safety class low, mediumand high, respectively

= the difference in internal pressure compared to as-laid

Ai = the pipe bore area= Poisson's ratio

As = the pipe cross-sectional steel areaE = Young's modulus

= the thermal expansion coefficient for the pipe material.

Figure 7-3 illustrates the design principles with respect to upheaval buckling. The curve shows the soilresponse, the maximum uplift resistance as a function of the design load, the equivalent temperature. Thehorizontal axis shows the soil resistance, starting with the expected value, leading to the characteristic valueas a lower fractile and the ending up with the design resistance accounting for an uplift resistance safetyfactor, γUF . The equivalent temperature is increased until failure, the pipeline lift off through the cover andexperience upheaval buckling, for two different downward soil stiffnesses, kBE and kLE, respectively.

The design (resistance equivalent) failure temperature (TRd) is given by:

(7.4)

(7.5)

(7.6)

where:

T(KLE) = the failure temperature for the lower estimate of the downward soil stiffness, see Equation (7.5)T(KBE) = the failure temperature for the best estimate of the downward soil stiffness, see Equation (7.6)KLE = the lower estimate of the downward soil stiffnessKBE = the best estimate of the downward soil stiffnesspli = the local incidental pressureRc = the uplift resistance of the cover.

And the uplift resistance factor, γUR related to the vertical upward soil resistance is found for or cohesive soils(clay) factor, γUF, shall be taken as:

(7.7)

and non-cohesive soils (sand and rock) as:

(7.8)


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The safety factor of 0.85 for the minimum cover design of non-cohesive soils is related to the uplift resistanceaccording to the expressions in DNVGL-RP-F114 [5.5] and DNVGL-RP-F114 [5.6]. If other methods are usedto establish the uplift resistance, the safety factor needs to be reassessed to ensure a safe design againstupheaval buckling.

The values relevant for the specific cover design applies also to initial propped shape analyses of the as-trenched configuration.

The pipe-soil interaction model shall be applied on the propped shape configuration as described in [7.5.2].

Close to the end, the build-up of axial force requires special attention, see [7.3] .

Figure 7-3 Illustration of the design principles for upheaval buckling

The design process to determine the failure temperature is outlined as follows:

1) The uplift resistance in the load response model is represented by a lower estimate characteristic value(from expected resistance (E(R)) to lower estimate (Rc)).

2) This uplift resistance (Rc) is reduced by the safety factor γUR. This safety factor is again dependent onthe configuration survey accuracy.

3) Downward soil stiffness equal to a best estimate is applied (kBE).4) All loads are applied in the model and the temperature is increased until failure occurs at T(kBE) in the

soil, Equation (7.5).5) A corresponding failure temperature is calculated for a lower estimate downward stiffness (kLE),

resulting in a temperature T(kLE), Equation (7.6).6) If T(kLE) is close to T(kBE), this implies that the pipeline will fail upwards. Failure upwards implies that it

is limited by the cover uplift resistance and that it is located on the dashed line in Figure 7-3.7) If T(kLE) is different from T(kBE), this implies that the initial soil failure is downward eventually causing

the pipeline upward penetration. Initial failure downwards implies that it is limited by the downwardstiffness and that it is located on the solid, more horizontal lines in Figure 7-3.


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8) If T(kLE) is higher than T(kBE), then the proposed design approach above needs to be reconsidered forthis case to ensure the safety.

9) Calculate the design resistance equivalent failure temperature, TRd, Equation (7.4).10) Calculate the design load equivalent temperature, TSd, giving axial effective load factor of γUF, Equation

(7.3).11) Confirm that the design load equivalent temperature is less than the design resistance equivalent failure

temperature, Equation (7.2), with the corresponding safety factors, Equation (7.7) and Equation (7.8).

If the analyses fail to fulfil the criterion in Equation (7.2), the cover shall be modified or the configurationchanged (re-trenched) and the analyses re-performed.

The design failure temperature TRd is a local parameter that will vary along the whole pipeline due todifferences in imperfections and soil resistance. The following simplified approach is recommended to avoidthis:

1) For a range of propped shape configurations, the required soil resistance/soil cover height (includingsafety factors) shall be estimated based on the best estimate downward stiffness, kBE, i.e. such that:

(7.9)

The height of the imperfections shall represent the relevant range of out-of-straightness/curvature of thepipeline.

2) Based on the soil resistance/cover height found during this design process, the failure temperatureT(kBE) shall be estimated with the lower estimate downward soil stiffness kLE (including safety factors).For each propped shape, the ratio between the temperatures at failure for best estimate and lowerbound downward stiffness shall be calculated as:

(7.10)

where i indicates propped shape imperfection no i.

3) If the deviation between temperature at failure for the best estimated and lower estimate downwardstiffness is within 5%,i.e. max (ri) < 1.05 mean (ri) the FE analyses on the actual seabed profile may be carried out usingthe best estimate downward soil stiffness. If the deviation exceeds 5%, modification in the safety factorshall be applied according to Equation (7.11), where γUF, OLD is as given in Equation (7.3). The modifiedsafety factor only needs to be applied at the propped shapes exceeding 5%.

(7.11)

For further details on estimation of local and global failure, see DNVGL-RP-F114.

7.6.2 Minimum cover designThe minimum soil uplift resistance, Rmin, along the pipeline length shall be determined for a propped shapeconfiguration. The height of the propped shape, δf , shall be taken as:

(7.12)


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where:

σconfiguration = is one standard deviation of the configuration survey accuracy given in metres.

The requirement of Equation (7.12) shall be regarded as a minimum requirement, previous projectexperience and engineering judgement may call for an increased propped shape height to determine theminimum cover.

7.6.3 Specified cover designThe purpose of the analysis of the pipeline in the as-laid/as-trenched configuration is initially to extract thepipeline curvatures from the configuration after laydown on the actual seabed. Then, these curvatures arecompared with the similar curvatures of the different propped shapes analysed. In this way, it is possible toestimate the necessary cover heights along the pipeline.Then, the as-laid/as-trenched model is analysed with soil springs representing the different cover heights.The temperature is gradually increased until the soil resistance springs fails to keep the pipeline in place.The failure temperature is then compared with the design temperature. If the failure temperature is belowthe design temperature, the uplift resistance shall be increased at locations where the cover initially failed.After applying additional cover at these initial failure locations, the analysis is re-run. This iterative process isrepeated until the failure temperature exceeds the design temperature.The final design shall be based on as-laid/as-trenched data.

TRd is calculated according to Equation (7.4) using T(kBE) and T(kLE) from the case with highest r.

7.6.4 Required coverThe required soil resistance, R, gives the following required cover:

(7.13)

An additional safety margin shall be added to account for uncertainties when documenting this cover heightin the final assessment (i.e. it needs to ensure that the required cover height or soil resistance is met alongthe complete length of the pipeline). The required safety margin depends on the survey accuracy of the coverheight.

7.6.5 Two or more independent surveysA survey will be associated with errors and inaccuracies depending on the survey tool, environmentalconditions, and obstacles on the seabed. The design approach assumes that a single survey represents theaverage profile with an upper and lower estimate represented by a factor on the survey accuracy. This is apragmatic approach to avoid artificial modification to the measured profile such a smoothening or definitionof local imperfections.

In case more than one independent survey is performed an increased confidence in the estimated coverheight is obtained. Each survey of the pipeline configuration can be considered as an average survey.Required specific cover/soil resistance, R1(KP) and R2(KP) to Rn(KP) can be estimated for each of thesurveys, but now calculated with reduced safety factors. Each of these cover heights shall be confirmedacceptable independently of each other.

The general soil resistance/cover height based on n surveys can now be taken as the average of the surveys.

(7.14)


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When calculating the average required cover height, engineering judgement has to be taken into accountbecause maximum required soil resistance may not be at the exact same locations on two surveys.

The safety factors will for a combination of n independent surveys become:

Non-cohesive soil (sand and rock):

(7.15)

Cohesive soil (clay):

(7.16)

where:

σconfiguration = one standard deviation of the configuration survey accuracy given in metres.

The same TRd applies for all n surveys of the pipeline.

The minimum soil resistance/cover height can be calculated for an imperfection of:

(7.17)

The final cover resistance/cover height for n surveys shall be taken as the maximum of the specific and theminimum cover height.

(7.18)

7.7 Relevant pipeline limit states - buried pipelineTable 7-2 shows the pipeline limit state checks normally governing for a buried pipeline. See Sec.9 for furtherdetails.

Table 7-2 Required and normally governing pipeline limit state checks - buried pipeline

Local buckling

Case Pressure cont. Load controlled1)

Displacementcontrolled

Axial Ratcheting Fatigue/failure

reference DNVGL-ST-F101



Equation (8.9) DNVGL-ST-F101


buried pipe X 2) X X X X X

1) Temporary phases (e.g. installation, as-laid empty and waterfilled).2) The X just means that this limit state needs to be checked.

A buried pipeline typically fails due to upheaval buckling when the axial load capacity is exceeded or whenthe mobilization of the soil exceeds it capacity expressed by the soil limit states, see [7.6]. In addition to theabove limit states, pipeline walking shall be considered.


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SECTION 8 LIMIT STATE CRITERIA

8.1 GeneralThe risk principles and limit state methodology in this recommended practice complies with DNVGL-ST-F101.This section describes design limit states based on:

— Definition of characteristic loads (e.g. pressures, temperatures, environmental loads and trawling) whichare described in Sec.3.

— Requirements to analysis of load effects given in general in Sec.5.— Calculation of the characteristic load effects for exposed pipelines, based on a strain based formulation or

a moment based formulation, according to Sec.6.— Calculation of the characteristic load effects for buried pipelines as given in Sec.7.

Overviews of the required design criteria are given in the subsections [6.6], [6.7.3] and [7.6] for exposedpipelines at flat seabed, exposed pipelines at uneven seabed and buried pipelines, respectively.

8.2 Local buckling limit state - combined loading

8.2.1 Moment based formulationThe bending moment for internal overpressure shall fulfil the following criterion (see also DNVGL-ST-F101):

(8.1)

15≤D/t2≤45, Pi>Pe, (SSd/Sp)<0.4

(8.2)

where:

γm = the material safety factor = 1.15, see DNVGL-ST-F101 Sec.5γSC = the safety class safety factor, see DNVGL-ST-F101 Sec.5γc = the condition load effect factor, Sec.9γf = the functional effect factor, see DNVGL-ST-F101 Sec.5MSD = the design bending momentSSD = the design effective axial forcepi = internal pressurepe = external pressureMp = the plastic bending moment capacity , see DNVGL-ST-F101 Sec.5Sp = the plastic axial force capacity , see DNVGL-ST-F101 Sec.5Pb = the burst pressure , see DNVGL-ST-F101 Sec.5αc = a hardening effect factor , see DNVGL-ST-F101 Sec.5αp = a pressure effect factor, see DNVGL-ST-F101 Sec.5D = the pipe outer diametert = the pipe wall thicknessMf = the characteristic functional moment.


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For further definitions of the load controlled local buckling criterion, see DNVGL-ST-F101.

The moment based formulation shall also be checked for strain capacity in [8.2.2] for all sensitivity analysesin Sec.9.

The bending moment for loading cases with external overpressure shall fulfil the criterion as per DNVGL-ST-F101.

8.2.2 Strain based formulationThe bending strain for internal overpressure shall fulfil the following criterion:

(8.3)

(8.4)

where:

εc = the characteristic strain capacity, given by Equation (8.5)εSd = the design strainεf = the functional strainγc = the condition load effect factor, Sec.9γf = the functional effect factor, see DNVGL-ST-F101 Sec.5.

(8.5)

(8.6)

(8.7)

(8.8)

where:

ε1 = the basic strain capacity, given by Equation (8.6)αLuder = the Lüder effect factorDo = the pipe outer diametert = the pipe wall thicknessαh = the maximum specified yield to tensile ratio of the pipeline steel, see DNVGL-ST-F101σh = the hoop stress based on the minimum level of internal overpressure which can co-exist with the

imposed temperature during loadingfy = the material yield stress.

Note that ε1 above is applicable for compressive bending strain only.


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For further definitions of the displacement controlled local buckling criterion, see DNVGL-ST-F101.Guidance note:

The correction factor for Lüder plateau is only likely to apply for seamless pipes with D/t > 20.


Bending in presence of external overpressure shall be checked as per DNVGL-ST-F101.

Corrosion damage will typically not be uniform around the cross-section or in the longitudinal direction ofthe pipeline, i.e. there may be local grooves with different shapes at various intervals. Assuming uniformcorrosion damage along the circumference and in the longitudinal direction may not represent the straincapacity for a pipeline exposed to global bending adequately, and hence it should be confirmed that the straincapacity for such grooves is not giving lower strain capacity than the strain capacity of the uniform corrodedpipe. See DNVGL-ST-F101 Sec.4.

8.3 Axial loading limit stateThe limit state for the compressive axial force may be governing for buried pipelines. Such pipelines shallfulfil the following criterion:

(8.9)

(8.10)

where:

εca = the characteristic axial strain capacity, given by Equation (8.10)εSd = the design strainD = the pipe outer diametert = the pipe wall thicknessγax = safety factor for the axial loading, = 3.5n = is the hardening factor in the Ramberg-Osgood curve for the steel material given as:

(8.11)

For pipelines showing a Lüder plateau, it shall be ensured that the axial strain is below onset of yield with asufficient safety margin. A sufficiently high number of the hardening parameter, n, will ensure this. However,such a criterion may be excessively over-conservative. It should also be recognized that installation byreeling will to large degree remove the Lüder plateau due to the significant plastic deformation. Specialconsiderations may be needed in such cases.

8.4 Uniform strain capacityThe longitudinal strain should not approach the uniform strain capacity of the material, i.e. the straincorresponding to the ultimate tensile strength of the material. The maximum equivalent strain developed inthe buckle should therefore be limited to:

(8.12)


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(8.13)

(8.14)

(8.15)

(8.16)

where:

εeq = the equivalent strainεus = the equivalent strain capacityαh = the maximum specified yield to tensile ratio of the pipeline steel, see DNVGL-ST-F101γus = the uniform strain safety factor

= the equivalent plastic strain

σeq = the equivalent, von Mises stressE = Young's modulusσh = the hoop stress

= the longitudinal plastic strain

= the hoop plastic strain

= the radial plastic strain.

Guidance note:This uniform strain criterion is based on tensile test data at ambient temperature. The applicability of the limit at the maximum

operating temperature should be confirmed during detailed design through mechanical testing. For very high αh the model will

give low allowable equivalent strains which may approach the yield strain. However, such αh values are likely to exceed therequirements of DNVGL-ST-F101.


8.5 Cyclic plasticity limit stateMaterials tend to exhibit a reduced yield stress during cyclic loading compared to the first load response. Atypical cyclic response, for a material which exhibits a Lüder plateau on first load, is shown in Figure 8-1.


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Figure 8-1 Typical cyclic stress-strain response with Bauschinger effect

Cyclic plasticity should be avoided. The maximum axial stress range, σR , both for internal and externaloverpressure should comply with:

(8.17)

where:

fy = the yield stressαB = the Bauschinger factor. It defines the magnitude of the Bauschinger effect and is the ratio for the

cyclic yield stress to the monotonic yield stress (or the ratio of the elastic range to twice yield inFigure 8-1. Recommended values of αB are presented in Table 8-1.

σh = the maximum absolute value of hoop stress that could occur during operation.

Table 8-1 Recommended values for αB the Bauschinger factor

Pipe type αB

seamless carbon steel 0.8

UOE carbon steel 0.7

CRA 0.7


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Guidance note:Table 8-1 should be confirmed through cyclic tensile testing in advance of detailed design.


8.6 Fatigue and fracture

8.6.1 GeneralLateral buckling results in both high stress ranges and high tensile stresses. Consequently, both the fractureand fatigue limit states need to be considered. The fatigue and fracture limit states are related as both dealwith cracks, as illustrated in Figure 8-2. The left part of this figure shows the typical work process assessingfatigue and fracture limit states in a lateral buckling design, and the iterations in case the fatigue and/orfracture design criteria are not fulfilled. The right part shows what parameters that govern the fatigue andfracture limit states. As seen many of the parameters are identical since these limit states are closely related,the blue cells show parameters not relevant for fatigue. A fracture mechanics based fatigue and fractureassessment is often referred to as engineering critical assessment (ECA).

Figure 8-2 Fatigue and fracture limit states


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A fracture assessment in accordance with DNVGL-ST-F101 is required for any pipeline that is designedto buckle laterally in case the longitudinal strain exceeds 0.4%. The primary objective of the fractureassessment is to establish the flaw acceptance criteria for girth welds.The fatigue limit state check should first be addressed using the S-N approach and proved to be acceptable.The fracture assessment check should then be undertaken to prove that the acceptable weld flaw size hassufficient fatigue and fracture capacity throughout the design life of the pipeline.All static and cyclic load cases, from installation to the end of the design life of the pipeline, shall beconsidered. The effect of the loading history on the material properties should also be considered. In apipeline designed to experience global buckling, the most significant source of fatigue loading will typically bethat arising from the deformation of the buckles due to start-up and shut-down cycles. The maximum staticand cyclic loads may also change as successive load cycles are applied. Therefore, it may be necessary toconsider several load cycles to identify the bounding case(s).Pipeline components which can lead to significant stress concentrations and thereby reduced fatigue life (e.g.J-lay collars, buckle arrestors or anode attachment pads) should, where practical, be avoided in sectionswhich are designed to buckle.S-N curves and fatigue crack growth laws should be appropriate to the environment, temperature andloading frequency. Similarly, fracture toughness should take account of the effect of environment andtemperature. Further guidance on environmental and temperature effects is given in App.C.

Guidance note:Information required to address the fatigue and fracture limit states may not be available at the conceptual or front endengineering design phases. Conservative assumptions should be made and sensitivity calculations should be conducted todetermine the significance of the various assumptions, and, hence, whether more detailed calculations and/or project-specifictesting are required in detailed design.


8.6.2 FatigueThe assessment of S-N fatigue should be undertaken in accordance with DNVGL-RP-C203. The S-Ncalculations should consider fatigue cracking from the weld cap and the weld root, and other locations offatigue cracking as appropriate. Allowable design fatigue factors are given in DNVGL-ST-F101 Sec.5.Project-specific testing may be required to determine the endurance curves. Recommendations for projectspecific testing are summarised in App.C.

8.6.3 Fracture assessmentThe fracture assessment shall determine the tolerable flaw size of the structural welds subject to the loadsimposed by the lateral buckling design. The acceptable flaw size should be compared with the workmanshipacceptance levels or feature sizes that could realistically escape detection during NDT of welds. If theacceptable flaw size is smaller than typical workmanship acceptance levels, the constructability of thepipeline may be compromised. The flaw size acceptance criteria should be based on the results of thefracture assessment, but they should also account for the characteristics of the welding process(es), theaccuracy and reliability of the inspection method(s) and good workmanship practices.The fracture assessment should be conducted in accordance with DNVGL-ST-F101. A lateral buckle is subjectto the combination of internal overpressure and longitudinal loading. The influence of internal overpressureshould be considered when determining the loading in the buckle.The fracture mechanics based fatigue crack growth assessments should be conducted with a DFF (designfatigue factor) equal to half the value specified in DNVGL-ST-F101.The fracture assessment should consider internal and external surface flaws, and embedded flaws. In clador lined pipelines, it is common for the acceptance criteria to specify no flaws at the weld root or in the zonebounded by the thickness of the CRA. The fracture assessment should nevertheless consider hypotheticalsurface and embedded flaws in or adjacent to this zone for comparison with the detection limit of theinspection method(s).


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The tensile stress-strain curve used in the analysis of lateral buckles and in the fracture assessment shouldbe identical. If this is not the case, it should be shown that the differences will lead to conservative results.Project-specific testing may be required to determine the tensile properties, the fracture toughness, and thefatigue crack growth law to be used in the fracture assessment. The effects of temperature on the materialproperties, and the implications of temperature changes through the load cycle should be considered toensure that appropriate bounding properties are used in the assessment. Recommendations for projectspecific testing are summarised in App.C.The fracture assessment requires a large amount of detailed information, e.g. geometry, material properties(including fracture toughness), installation and operational loads, welding procedures, environmental effects,etc. Not all this information will be available in the early stages of design. The level of complexity of thefracture assessment (e.g. option 1 or 2 of BS 7910, or 3D FE analysis) should be appropriate to the quality ofthe available data and the severity of the loading. A phased approach is recommended:

— a preliminary fracture assessment during conceptual or front-end engineering design, and— a detailed fracture assessment during detailed design.

In the conceptual design stage, the objective of the fracture assessment is to determine whether typicalworkmanship acceptance levels are acceptable or not under the loading imposed by lateral buckling.Sensitivity calculations should be conducted to determine the importance and implications of any necessaryassumptions or simplifications, and to reduce the number of cases to be considered in detailed design.Information from previous projects could be used to supplement the limited information available inconceptual design. The fracture assessment may indicate that the acceptable flaw size is smaller than typicalworkmanship acceptance levels. It is important to be aware of such possible constraints on the fabrication ofwelds early in the design cycle. The results of the preliminary fracture assessment will indicate whether to:

— conduct more sophisticated analyses to reduce the conservatism in the calculations— undertake project-specific testing, and/or— modify the design.

In the detailed design stage, the objective of the fracture assessment is to determine the acceptable flawsizes and therefore, to establish flaw acceptance criteria.


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SECTION 9 CONDITION LOAD EFFECT FACTOR FOR EXPOSEDPIPELINES

9.1 Basic principlesThis section describes the calculation procedure for the condition load effect factor, γc , to be applied in thecombined loading - moment based local buckling criterion for pipelines which experience global buckling.

The condition load effect factor, yc, is based on the prevailing uncertainty in the response bending momentand given by:

(9.1)

Note that a γc less than unity calculated in this section shall not be applied to the effective axial load in thisrecommended practice.

CoV(XF(p,T,Fc)) is the coefficient of variation of the resulting bending moment in the buckle based oncharacteristic pressure(p), temperature(T) and trawl load (FT). The uncertainty in the bending momentresponse from the global FE analysis is assumed to arise from:

— uncertainty in the axial pipe-soil resistance, XA— uncertainty in the lateral pipe-soil resistance, XL— uncertainty in the applied stress-strain curve, XB— uncertainty in the applied trawl load, XC (for annual trawl frequency larger than 10-4 only).

The uncertainty in the bending moment response may be estimated from:

(9.2)

The terms CoV(XA), CoV(XL), CoV(XB) and CoV(XC) reflects the impact on the resulting bending momentresponse uncertainty originating from the pipe-soil resistance (axial and lateral), the material behaviour(stress-strain curve) and the applied trawl pull-over load, respectively. The condition load effect factor willalso implicitly represent the degree of displacement control in the buckle.

In addition, a model uncertainty may be present. CoV(XF) shall not to be taken less than 5%. This isaccounted for by the minimum value of γc of 0.80 in Equation (9.1).In case global buckling does not initiate for certain pipe-soil combinations, the imperfections should bemodified to allow buckling for all sensitivity analyses.

9.2 Calculation of CoV(XA) due to axial soil resistanceThe required set of non-linear FE analyses used to establish CoV(XA) in the expression for the conditionfactor, γc, is indicated in Figure 9-1. XA and XL is representing the stochastic variables for the axial andlateral pipe-soil resistance, respectively. To avoid a full structural reliability analysis, the condition load effectfactor, γc, is established according to Equation (9.1) and a limited number of sensitivity analyses. Thesesensitivity analyses are numbered 1 to 4 and BE corresponding to the soil matrix in Equation (9.3) andEquation (9.6) and represents the uncertainties in the soil resistance parameters. The triggering mechanismshall be established assuming an imperfection triggering the highest capacity combinations of lateralresistance. The moment responses M1, M2 and MBE shall be taken at the final equilibrium state.


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Figure 9-1 Required soil property combinations to be assessed in sensitivity analyses

The moment responses corresponding to soil matrix become:

(9.3)

where the indices UE and LE indicate upper and lower estimate resistance and fy indicate analyses with astress-strain curve defined from specified minimum values fy and fu.The resulting uncertainty contribution from the uncertainty in the axial soil resistance becomes:

(9.4)

Where nA accounts for the distance between upper and lower estimate values in terms of standarddeviations, i.e.:

(9.5)


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If the upper and lower estimate values are specified as mean value ± two standard deviations (in accordancewith normal interpretation), nA = 4 applies.Note that soil resistance coefficients with upper indices LE, BE and UE represent the complete soil (force-deformation) model rather than a single value. The main objective is to define the deviation of the lower andupper estimate models from the best estimate model in terms of standard deviations.

9.3 Calculation of CoV(XL) due to lateral soil frictionThe required set of non-linear FE analyses used to establish CoV(XL) in the expression for the condition loadeffect factor γ2 is also indicated in Figure 9-1. This corresponds to the soil matrix in Equation (9.6). Themoment responses M3, M4 and MBE shall be taken at the final equilibrium state.The moment responses corresponding to soil matrix become:

(9.6)

where the indices UE and LE indicate upper and lower estimate value and fy indicates analyses with astress-strain curve defined from specified minimum values fy and fu. A total of two additional FE analyses isrequired.The resulting uncertainty contribution from the uncertainty in lateral soil resistance becomes:

(9.7)

where nL accounts for the distance between upper and lower estimate values in terms of number of standarddeviations, i.e.:

(9.8)

9.4 Calculation of CoV(XB) due to stress-strain relationshipUncertainties in the load effect calculations CoV(XB), which are related to the resulting moment-strain curvefrom geometry and material properties are assessed from the base case and one additional response (FE)analysis M5.

(9.9)

Where σy indicates analyses with stress-strain curve defined from mean values of yield and ultimate stress.CoV(XL) is given as follows:

(9.10)


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Where ny is the number of standard deviations between the mean yield strength and minimum specified yieldstrength, typically 2.

λ is the ratio between the moment capacity using σy corresponding to M5 and moment capacity using fycorresponding to MBE. This ensures consistency between global response analyses and capacity. λ is givenby:

(9.11)

where:

σy = mean yield stressfy = the material yield stressαp = a pressure effect factor, see DNVGL-ST-F101 Sec.5qh = the nomalised pressure utilisation with fy.

qh and αp are related to fy.

A CoV (XB) of 5% may be used as a maximum value.

9.5 Calculation of CoV(Xc) due to trawl loadsIn case the buckled section is potentially exposed to trawl pull-over loads, the uncertainty from the trawlloading shall be accounted for by performing two additional analyses:

(9.12)

where UE and LE indicate upper and lower estimate value for the trawl load applied in the apex of thebuckled section.

The moment responses, M6 and M7, shall be taken as the maximum values including the transient values ifthese are the largest. This is normally a conservative approach.The resulting uncertainty becomes:

(9.13)

where nF accounts for the distance between upper and lower estimate values in terms of standard deviations,i.e.:

(9.14)

The trawl load is defined in [3.4.3]. nF is typically to be taken between 2 and 4.


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Guidance note:The intention of the trawl sensitivity study is to include this effect on the overall uncertainty of the resulting bending moment.Since the global buckling moment is mostly displacement controlled, the load controlled trawl moment will not be added but to alarge extent replace the functional moment from global buckling. If the contribution from the trawl is dominating the uncertainty,

special evaluations are required to determine a higher y c than resulting from the above procedure.


9.6 CoV for parameters with large variation and non-symmetricalupper and lower estimatesIt is assumed that the resulting bending moment response can be described by a linear Taylor seriesexpansion as follows:

(9.15)

where:

a1 = the Taylor series expansion coefficient around the mean values/base case value for the parameter,a0is a constant

x1 = the basic parameters (axial and lateral soil friction, stress-strain curve, trawl load).

The standard deviation for M, σM becomes:

(9.16)

where:

ai = the Taylor series expansion coefficient around the mean values/base case value for the parameterρij = the coefficient of correlation between parameter i and parameter jσx,i = the standard deviation of parameter xi.

For independent basic parameters (ρx,i = 1) the coefficient of variation becomes:

(9.17)

The expansion coefficient ai shall be established from parametric studies around the mean value/base casevalue. The recommended procedure is to establish ai based on a 3-point polynomial approximation to M(xi).Using bending moment point values (M1, xi,1) and (M2, xi,2) symmetric around the mean value (MBE, xi,bc)the explicit expressions below appear.

(9.18)


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If the bending moment relationship M(xi) is known to be linear, CoV(XF) may be established by only twopoints. In that case:

(9.19)

See also /10/ for further discussions.


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SECTION 10 OPERATIONAL STRUCTURAL INTEGRITY

10.1 GeneralPrior to start-up, an inherent uncertainty with respect to the buckling and walking response of the systemprobably exists. To ensure the long-term structural integrity of the pipeline due to expansion loads,appropriate operational monitoring shall be undertaken.The pipeline integrity management system developed in the design stage shall include requirements formonitoring the formation of lateral buckles shortly after start-up and subsequent behaviour in operation.Further, it is important to monitor the associated load history, defined by pipeline operating pressures andtemperatures. In addition, monitoring requirements for pipeline walking shall be defined. For further details,see DNVGL-ST-F101 and DNVGL-RP-F116.Operating surveys shall be used to:

— confirm that the buckling strategy has been reliable and that all buckles have formed as planned— identify the number and location of any unplanned buckles— determine the buckle shapes and curvatures— monitor axial displacement at pipeline end and at in-line structures— monitor pipeline/SCR holdback anchors to ensure that pipeline walking has been arrested— monitor route curves to ensure that no lateral displacement or pull-out has occurred, and— identify the requirement for any mitigation measures to ensure pipeline structural integrity.

Positional surveys are required to assess pipeline buckling and curve stability. In case pipeline walking isexpected, visual inspections are required and some additional means to assess the end expansions.Clear, easy-to-read markers should be provided at each end of the pipeline and at in-line connections formonitoring axial displacement and potential pipeline walking. If sleepers or other fixed mitigation measuresare employed, these should include clear, easy-to-read position markings along their length for monitoring ofbuckle amplitudes. In all cases, markings should be on a vertical face to avoid obstruction from soil or debrisdeposition. A visual inspection of these markings should be carried out after installation and before hydrotestto establish the datum for all future measurements.Marking is considered as a good direct measurement of buckle amplitude or axial displacement for deepwater projects. For shallow water application where marine growth is expected, marking on triggers or PLET/PLEM will not be as easy to use as for deep-water pipelines and alternative methods should be considered.

10.2 Surveys of pipeline out-of-straightness and buckling behaviourDNVGL-ST-F101 requires that as-laid, as-built and operational surveys shall be conducted.

Guidance note:As-laid and as-built surveys can be important in the assessment of expansion and global buckling behaviour of pipelines.Experience has shown that for pipelines where these surveys are not available, the integrity assessment of the global buckling andexpansion design can be more challenging.


The as-laid pipeline survey shall be used to establish the pipeline profile along the route to confirm theassumptions made in design. The as-laid survey will also provide a baseline against which unexpectedoperational performance can be checked.A full-length pipeline position survey shall be undertaken after steady state operation is achieved to establishthe location of each buckle (planned an unplanned) and to confirm system performance. Measurement of onebuckle in isolation does not provide sufficient information on virtual anchor locations, feed-in to the buckle, orthe loads within the buckle.Further operational surveys should be performed, as required, to monitor the amplitude and shape of thebuckles to ensure that the range of curvature in the buckle is not excessive and does not exceed designlimitations. Ideally, the survey should cover the full pipeline length, however, once the buckles are fully


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established it may be acceptable to limit the survey only to the areas associated with displacements causedby global buckling.

Guidance note:In some cases, the requirements for surveys after the first operational survey may be relaxed, e.g. when the loads in the bucklesare confirmed to be low.Long-distance transportation pipelines are expected only to be susceptible to global buckling and expansion in the warm end whilemost of the pipeline length is operating at ambient temperature with limited buckling susceptibility. Hence, for further operationalsurveys the survey length can be considerably shorter than the pipeline length for global buckling assessments.


Pipeline surveys should be carefully planned and executed to ensure that data obtained meets accuracyrequirements specified for subsequent assessments. The required accuracy should be sufficient to capturethe expected highest levels of curvatures along the pipeline. The survey should give the absolute and relativeposition at least every metre along the pipeline in X, Y and Z directions. The survey contractor should processbut not smooth the data. Processing here means to calculate the pipeline profile and cover height based onthe measured data and to remove the effects of waves and tide. No further smoothening of the data shouldbe done by the survey contractor.

Guidance note:The objective of these surveys is to measure local out-of-straightness over distances ranging from 10 m to 500 m, and theabsolute (global) accuracy is generally significantly less important than the relative (local) accuracy. However, by improving theabsolute accuracy, it can be inferred that the relative accuracy will also be improved.


Calibration of accuracy including calibration of the measuring devices and repeatability of local out-of-straightness measurements should be undertaken prior to commencing the survey. This can be carried out bysurveying the same section of line a few times in different directions, preferably along a section of line with aknown out-of-straightness feature such as a buckle of a route curve.Pipeline embedment should be measured at the pipe centreline and at least two locations (close and far) toeach side of the pipeline to enable assessment of embedment and local berm features along the length of thepipeline. Cross profiles should be provided at specified regular intervals along the pipeline, with refined dataspacing within each lateral buckle region.The operational data (pressure and temperature) at the time of the survey shall be recorded. Surveys shouldbe performed during steady state conditions (operating or shutdown) rather than during transient flow.

10.3 Operating load condition monitoringTemperature and pressure monitoring at the inlet and outlet is required to establish local operating conditionsalong the pipeline with sufficient accuracy to evaluate functional loading in the operation phase. It is alsorecommended to monitor slugging for buckle design where slugging loads have shown to give significanteffects. To ensure that significant fluctuations are registered, data should be recorded at approximately oneminute intervals. Lower or higher recording frequencies may be proposed, depending on the anticipated flowdynamics. Details of measurement frequency, recording methods and the selection of instrument locationsshould be agreed with operations personnel to ensure that all data relevant to pipeline operating conditionsare recorded.

10.4 Pipeline structural integrity assessmentThe pipeline integrity management system should recognise the potentially severe loading associated withthe global buckling design strategy. Key design parameters should be clearly outlined (for example maximumallowable temperature and pressure profiles) so that operational monitoring is against known limits.All data from inspections, surveys and operating load condition monitoring should be recorded and readilyaccessible. The data should be assessed at regular intervals to ensure the pipeline is operated in linewith design intentions, to confirm pipeline structural integrity and to plan the scope and timing of futuresurveys, or mitigation measures in the unlikely event that loading exceeds design criteria. This may require


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engineering assessment including recalibration of design FE models to more accurately quantify operationalloading.

10.5 Survey frequencyIn the early stages of the design life, monitoring allows comparison between operational behaviour anddesign predictions. The monitoring also enable the operator to identify potential new challenges that maydevelop in the longer term and which may be prevented with appropriate remedial actions. The frequencyof operational pipeline surveys may be reduced as confidence grows in system performance, based onsystematic observation of in-place performance.

10.6 Re-qualificationGuidance on re-qualification of pipelines can be found in DNVGL-ST-F101 and DNVGL-RP-F116. Re-qualification is a re-assessment of the design typically due to modified design premises. Most relevant withrespect to global buckling is changes in the design criteria applied (more recent and updated ones), damagesand/or changes in design conditions, e.g. flow conditions like type of fluid, design pressure and temperatureet cetera.As for the structural integrity assessment mentioned above, all available inspection and survey data shouldbe used to improve the knowledge about the typical behaviour of the pipeline. Indirectly survey data aboutthe configuration may be combined with operational data and applied in the global FE model to validate themodel and to improve the estimates of the pipe-soil resistances. In this way, the uncertainties associatedwith the results which governs the pipeline integrity can be reduced.


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SECTION 11 DOCUMENTATION FOR OPERATION

11.1 GeneralA pipeline that is expected to buckle will have some challenges with respect to interpretation of surveyresults. The documentation should be as per standard documentation requirements (see DNVGL-ST-F101)but with special focus on the following issues that should be documented as function of position along thepipeline (KP) for the as-built pipeline (e.g. included in alignment sheet where possible):Pipeline data:

— outer diameter— pipe wall thickness— coating type and coating properties such as thickness, thermal insulation and density— content density— material data.

Loading parameters:

— temperature profile along the pipeline, both design and normal operating— pressure and pressure profile along the pipeline— lay tension— trawl data (frequency, loads).

Design aspects:

— safety class and any trawl free zones— buckling design (how and where the buckling is expected to be triggered)— key points to check, e.g. places with high utilization or uncertainty— friction models applied including expected penetration— effective axial force along the pipeline— interventions and protection requirements— allowable free spans.

As-built configuration:

— as-built profiles and interventions— location of crossings, inline tees, nearby pipelines etc.

The documentation shall be given in such format that it can be used to:

— updated inspection plans— evaluate inspection results.


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SECTION 12 REFERENCES

12.1 General/1/ Hobbs R E, In-Service buckling of Heated Pipelines, ASCE Journal of Transportation Engineering,

1984, vol. 110, pp 175-189.

/2/ Taylor N and Gan A B, Submarine Pipeline Buckling-Imperfection Studies, Journal of Thin WalledStructures, no. 4, Elsevier Applied Science Publishers Ltd., England.

/3/ Walker A, Spence M and Reynolds D, Use of CRA lined pipe in high temperature systems, OffshorePipeline Technology 2000, Oslo.

/4/ Fyrileiv O and Collberg L, Influence of pressure in pipeline design - effective axial force, Proceedingsof OMAE2005, 2005, OMAE2005-67502.

/5/ Sparks C P, The Influence of Tension, Pressure and Weight on Pipe and Riser Deformations andStresses , ASME transaction vol 106, pp. 46-54, 1984.

/6/ Spinazzè M et al., The HOTPIPE Project - Use of Analytical Models/formulae in Prediction of LateralBuckling and Interacting Buckles, Proc. 9th Int. Offshore and Polar Engineering Conference, 1999,Brest, France.

/7/ Carr M, Sinclair f and Bruton D, Pipeline Walking - Understanding the Field Layout Challenges andAnalytical Solutions developed for the SAFEBUCK JIP , OTC-17945, Proceedings of OTC 2006.

/8/ Bruton D, Sinclair F and Carr M, Lessons Learned from Observing Walking of Pipelines with LateralBuckles, Including New Driving Mechanisms and Updated Analysis Models, OTC 20750, 2010.

/9/ Peek R, Correction to the Infinite Mode for Lateral Buckling of Pipelines, ASCE Journal of TransportEngineering, September 2009.

/10/ Røneid S and Collberg L, Global Buckling Design of Submarine Pipelines - Design Performed Basedupon DNVGL-RP-F110 , Proceedings of OMAE2008, 2008, OMAE2008-57953.

/11/ Palmer A C, Ellinas C P, Richards D M and Fuijit J, Design of Submarine Pipelines Against UpheavalBuckling, OTC 6335, 1990.

/12/ Suzuki N and Toyoda M, Seismic Loadings on Buried Pipelines and Deformability of High Strength LinePipes Proceedings of the Pipe Dreamer's Conference, 2002, Yokohama, Japan.


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APPENDIX A MITIGATION MEASURES FOR EXPOSED PIPELINES(INFORMATIVE)

A.1 GeneralIf an unplanned buckling strategy is unacceptable, mitigation measures should be developed, includingmeans to ensure controlled buckling at regular distances. Examples of different types of mitigation measuresare provided in Table A-1.

Table A-1 Examples of buckle mitigation measures

Prevent buckling Trigger buckling

increase lateral soil restraint intermittent rock dumping

reduce driving force snake-lay

prevent unwanted uplift vertical upset

local weight reduction

zero-radius bends

pre-bend sections

The selection of mitigation measure depends on several factors, including the severity of operatingconditions, environmental conditions, seabed bathymetry, water depth, installation vessel, cost, schedule etcetera.

A.2 Prevent development of buckling

A.2.1 Increase lateral soil restraintThe simplest and most straightforward way of preventing a pipeline to buckle may seem to bury it withsufficient cover material. However, this option is often difficult and expensive, particularly for deep waterand/or high pressure high temperature pipelines, and several alternative design strategies have beenexamined, see /A1/.Another alternative could be to increase the submerged weight of the pipeline. It is normally not practicableto resolve the buckling challenge by an increase in submerged weight, because the increase in weight willnormally be substantial and exceed practical limits.

A.2.2 Reduction of the driving forceAnother obvious method of preventing buckling will be to reduce the driving force by reducing the designoperating temperature and pressure. A reduction in operating temperature could be accomplished byadding a heat exchanger to the system. Reducing the wall thickness of the line will reduce the temperaturecomponent of the effective axial force, which is proportional to the wall thickness (and usually the dominatingcomponent). A reduction in wall thickness can be achieved by increasing the grade of steel, by utilising acorrosion resistant alloy or by utilising a pressure protection system.A second method to reduce the driving force is to increase the lay tension. Residual lay tension balancespart of the compressive force induced by operation, and therefore reduces the resultant force. One difficultyis that the residual tension cannot be measured directly, but will be calculated from the lay conditionsand depends on pipeline route curves and uneven seabed. Furthermore, for pipelines with high operatingtemperature and pressure, the required lay tension to prevent buckling may exceed any practical limit and/orthe tension capacity of the installation vessel.


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A.2.3 Prevent unwanted uplift in pipelineA pipeline on an uneven seabed will tend to raise and lift off at seabed crests in case a given combination ofaxial force and axial feed-in is present. This uplift will act as a triggering mechanism for global buckling, butthere may be several reasons why it is not acceptable:

— predicted loading within the resulting buckles are unacceptable— restraints on the pipeline such as end terminations or in-line flanges are normally not designed for vertical

or lateral movements— unwanted interaction with other installations close to the pipeline— uplift from pipeline supports.

Figure A-1 shows an example of a pipeline (blue line) installed at an uneven seabed (thin black line) whereintermittent rock berms at the seabed crests have been used to prevent the pipeline to lift off at the crestsduring operation.

Figure A-1 Uplift restraints at seabed crests by intermittent rock dumping

A rock berm on top of the pipeline is a common mitigation measure to prevent uplift of a pipeline. However,confident identification of these overbends is a demanding survey task in deep water, or when the criticalimperfection amplitude is small.

A method to improve cover efficiency is to place a geotextile or concrete mattress over the pipeline beforethe rock is placed. However, the use of a geotextile in a subsea environment will require a comprehensiveinvestigation of its long-term stability against creep and structural deterioration.

The uplift restraints shall be modelled with the appropriate force/displacement relation. Uplift forces on therestraints Fr shall be estimated for the worst combination of functional loads. The pipe-soil interaction forlateral and axial resistance can be taken as best estimates.


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The restraints preventing the uplift shall be designed according the following principal equation:

(A.1)

where:

Rr = uplift resistanceFr = uplift force

= safety factor.

The resistance in a rock berm can be estimated according to equations in DNVGL-RP-F114. The safety factorshall be taken as = 2.0.

A.3 Triggering buckling

A.3.1 GeneralMany different techniques can be used to trigger buckling and thereby increase the reliability of buckleformation and sharing of the expansion potential into multiple buckle sites. In addition, some buckle initiatorscan produce more benign buckles compared to buckles that are not triggered by engineered initiators, andin this way, further improve the design. Whilst each method is described separately in the following sections,it may be possible to combine the individual techniques to further improve buckle formation reliability (e.g.vertical upset in combination with local weight reduction) (see /A2/, /A3/, /A4/ and /A5/).

A.3.2 Intermittent rock dumpingStabilisation with intermittent rock dumping can be used to prevent buckling at specific locations.Intermittent rock dumping may also be used to limit the axial feed-in to sections of pipeline susceptible tolateral buckling whether this is due to natural imperfections or due to trawl gear interference. In the firstcase, the rock is placed on critical overbends whilst in the second case the rock is placed between bucklelocations (see /A6/).

A.3.3 Snake-layIn snake-lay, the pipeline is laid in a series of gentle curves, either with straight sections between (asillustrated in Figure A-2, or as a continuously curved route such that the exit tangent to one curve is theentry tangent to the next (see /A7/ and /A8/).


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Figure A-2 Typical snake-lay configuration (exaggerated lateral scale)

The key parameters in the snake-lay configuration are the snake pitch, the offset and the average or targetcurve radius. The offset and curve radius define the arc length which typically ranges from 100 m to 300 m.The aim of the snake-lay strategy is to initiate a true lateral buckle (as opposed to benign expansion of thesnake crown) at some point on the curve. The propensity for buckling is primarily controlled by the curveradius but is also influenced by the snake pitch and arc length. Decreasing the curve radius tends to increasethe likelihood of buckling, although this is ultimately limited by the minimum lay radius capability of the layvessel for the given water depth and seabed soil.The frequency of buckling can be increased by reducing the snake pitch, subject to consideration of theavailable driving force to ensure successful buckle initiation at each snake curve, but buckle interactionshould be avoided.The additional length of pipeline required for a snake-lay strategy is generally not significant, being in theorder of a few pipe joints.Snake-lay can be an attractive buckle initiation technique in shallower water depths as it is not susceptible tohydrodynamic instability, it does not create pipeline spans which may be prone to vortex induced vibrations,and it does not incorporate any features which are a hazard to fishing activities.However, as the pipeline is laid on the seabed, the success of the buckle formation is still highly dependenton the pipe-soil interaction. If the lateral breakout friction is underestimated, or the axial friction isoverestimated, there may be insufficient driving force to initiate buckling at every snake curve.The presence of inherent out-of-straightness features in the straight sections between curves (either lateralor vertical) can also lead to preferential buckling away from snake curves, and may produce a buckle whichabsorbs greater feed-in than the design assumptions (based on snake pitch alone). It may be difficult toguarantee that the out-of-straightness imposed by the snake will always be more severe than inherent out-of-straightness features.Defining a target curve radius which is at the limits of what the lay vessel can achieve may be counter-productive, as the curve radius achieved may be greater than the target, thus reducing the probability ofbuckle formation at each site and leading to a less robust solution.As the pitch between buckle initiators is reduced, the probability of buckle formation generally reduces. Ifthe pitch is reduced too much in order to meet the design criteria, buckling may not occur at some of theintended sites and the robustness of the solution decreases.The effect of these undesirable behaviours is incorporated into the calculation of the characteristic VASwhen using the probabilistic buckle formation method. In general, the characteristic VAS will exceed thesnake pitch, and the design is not compromised so long as the unity checks at the characteristic VAS areacceptable.


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An additional concern is the stability of the curves under tension during shutdown, which may increasewith cycles, particularly if pipeline walking occurs. If the tension becomes sufficient to overcome the lateralpipe-soil resistance, the curves may pull out. This can increase the expansion at the ends of the pipeline, oroverload the adjacent lateral buckles (if they absorb the additional pipe).

A.3.4 Vertical upsetThis approach seeks to induce vertical movement which will then develop into a lateral buckle by deliberatelyintroducing significant vertical out-of-straightness at several points along the pipeline (see /A9/). Bothsleepers and gravel dump berms have been employed to date.In addition to the vertical out-of-straightness, the lateral resistance is also reduced relative to a pipelineresting on the seabed as spans are created at either side of the upset, and the friction between the pipelineand the upset can generally be modified such that it is less than that associated with the pipe-soil response.The simplest sleeper design may comprise of several large diameter pipe joints. The diameter of the sleeperpipe should be sufficient to provide the desired vertical imperfection height following any settlement. Therequired sleeper length depends on the anticipated buckle amplitude, typically, two or three pipe joints havebeen used. The sleeper is coated to provide wear resistance and to control the pipe-sleeper friction.A mud mat is frequently incorporated into the sleeper design to reduce the uncertainty associated withsettlement and allow larger vertical imperfections to be achieved. This may also allow the incorporation of avertical reaction post on the sleeper to prevent pipeline movement beyond a certain point.Steel sleepers should be fitted with independent cathodic protection, as required, and coating integritybetween pipeline and sleepers is required to avoid potential corrosion issues. Some projects have undertakensubmerged seawater tests of coating integrity using the design contact loads and repeated cyclic loading.Sleepers should be preinstalled prior to (or during) pipe lay and accuracy of installation is important. The layvessel will then lay the pipeline over the centre of the sleepers with sufficient accuracy to allow for bucklingin operation. This will require touch-down monitoring, and possibly special lay procedures to ensure that thepipeline is laid gently onto the sleepers to minimise sleeper embedment.Whilst the influence of the pipe-soil interaction on buckle formation is reduced (relative to a pipeline on theseabed), the pipe-soil interaction response at the touch-down locations may be a significant influence on thebuckle behaviour and loading.The vertical upset may result in significant spans which are susceptible to vortex induced vibrations, and maybe a snagging (fishing) hazard. This can limit the maximum allowable height of the sleeper, and may preventusage in shallow water.Vertical upsets may induce or exacerbate slugging in multiphase lines. Passage of the heavier slugs over thevertical upset and the associated pipeline spans can result in vibration at the frequency of the slugging. Anyproject in which slugging is anticipated should fully assess the implications of cyclic loading (fatigue) beforeemploying vertical upsets as the buckle initiation technique. This may be of concern for lighter pipelines,for which the variation in internal contents density due to slugging have a more significant influence on theoverall submerged weight.Buckles forming using the vertical upset method are generally symmetric (mode I or III, see /A9/), butseveral asymmetric mode II buckles have been observed. The buckle mode shape is generally driven by thelocal imperfections induced during pipe lay. Mode II buckles can develop higher curvatures, and cyclic stressranges, than mode I and mode III buckles as the location of peak bending may be on the seabed. The designshould therefore consider the uncertainty in mode shape, or include provision to intervene if an undesirablebuckle mode occurs.In order to improve buckle reliability, vertical upsets may be combined with a local reduction in the pipesubmerged weight over the trigger, either permanently (by removal of concrete coating, use of low densitycoating, application of buoyancy) or temporarily (by use of buoyancy bags).Some sleeper designs have incorporate a push mechanism which is activated prior to start-up to induce alateral deflection in the pipeline which further reduces the critical buckling force, and increases the certaintyof the buckle mode shape. The sleeper structure and mud-mats shall resist the lateral force generated duringactivation of the push mechanism. If the pipeline experiences high tensions on unload and the pipeline pullsagainst the push mechanism, high lateral loads may be generated.


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Reduced length sleepers can also be used. The objective is then to initiate the buckle on the sleeper, but forthe pipeline to fall onto the seabed as the buckle develops, thus avoiding long term vortex induced vibrationloading during operation in shallow water depths. However, if a mode II buckle develops

A.3.5 Local weight reductionTo trigger buckling, the submerged weight of the pipeline can be reduced in some manner. Two methodshave been employed to date, removal of concrete coating, or addition of buoyancy (see /A11/ and /A12/). Inboth cases, the reduced submerged weight will lead to reduced lateral resistance, and thus reduced criticalbuckling force and reduced severity of bending (for a given feed-in). In case buoyancy is added, a verticalout-of-straightness will generally be created, further promoting buckle formation.For concrete coated lines, short sections of pipeline have been installed without concrete coating, orwith reduced weight concrete coating. The no concrete approach removes the strain concentration factorassociated with the concrete coating field joint, which can significantly improve the loading in the buckle.A section of reduced concrete thickness, or reduced concrete density, will tend to reduce the magnitude ofthe strain concentration factor at field joints. However, as concrete coating is generally applied to ensurehydrodynamic stability and/or to provide impact protection, removal of concrete from buckle regions mayeither lead to local instability or vulnerability, which should be addressed in the design.The additional buoyancy is chosen so that the operational submerged weight of the pipeline is a small fractionof the normal pipe submerged weight. The length and uplift provided by the buoyancy modules can beadjusted to modify the shape of and loading in the buckle based on project specific design criteria. Theminimum length of additional buoyancy would generally be equal to the length of the central lobe of thebuckle. In some cases, two lengths of distributed buoyancy have been installed with a gap between to aidinstallation, with mode II buckles in operation predicted.Additional buoyancy has been added either as an increased thickness of the insulation coating (providedthe installation vessel can handle the diameter change during lay), or by offshore installation of discretebuoyancy half-shells which can pass through the lay ramp or lay vessel stinger. A related approach is to usediscrete buoyancy bags to aid buckle initiation, which are then removed after start-up once buckle formationhas occurred.Unless it is used as a temporary measure for buckle initiation only, any additional buoyancy devices shouldremain on the pipeline in operation, and should be designed for the life of the pipeline, including cyclicinteraction with the seabed. The reduction in buoyancy uplift associated with compression and/or wateringress should be considered.When the submerged weight is reduced significantly, local weight reduction tends to reduce the importanceof the pipe-soil interaction on the buckle response. However, consideration of the pipe-soil response at thetransition between normal pipeline sections and more buoyancy sections is important to predict the bucklebehaviour.There are practical limits to the length and uplift of buoyancy that can be applied to avoid excessivefree spans or hydrodynamic instability. If the buckle is unstable, particularly in shallow water, it may besusceptible to increased movements and may even flip from one side of the pipeline centreline to the otherunder severe loading, potentially compromising fatigue performance.

A.3.6 Zero-radius bendsZero-radius bends (ZRB) (see /A13/) consist of a vertical upset trigger with a vertical support at one end,and typically include some sort of mud-mat, see also Figure A-3 which shows the plan view during pipelineinstallation.The vertical trigger is preinstalled on the seabed and the pipeline is initially laid straight towards it until thepipeline is firmly seated on the trigger. At this point the lay vessel moves laterally and changes heading bya small angle, before continuing pipe lay in a straight line. In this manner, a bend is created in the pipelineroute, concentrated within the span created by the trigger. Although the bend radius is non-zero due to thepipeline flexural rigidity, it is much tighter than that which could be implemented in a conventional routecurve, and thus the critical buckling load is significantly reduced.


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Figure A-3 Zero-radius bend structure - plane view

As for sleepers, accuracy of the installation of the zero-radius bend structure and the subsequent pipe lay isimportant, and may require touch-down monitoring and potentially special lay procedures.Zero-radius bends tend to produce low critical buckling forces which may allow planned buckles to be locatedcloser together, or may reduce the probability of unplanned buckles occurring. Furthermore, symmetric mode(I or III) buckles are much more likely to occur when compared to conventional sleepers.The zero-radius bend structure shall be designed to resist sliding and overturning under the action ofimposed lateral loads on the trigger both during, and after, installation. This can lead to large mud-matsand structures. If the pipeline experiences large tensions during unloading (e.g. due to pipeline walking,anchoring or tie-in to a steel catenary riser) and the unloaded buckle pulls against the trigger post, highlateral loads can be generated.

A.3.7 Pre-bent sectionsLike a snake-lay solution, the pipeline can be installed with pre-bends at predefined locations along thepipeline route, designed to impose engineered out-of-straightness features to initiate regular buckling on theseabed, see /A14/and /A15/. These bends could be imposed during pipe-lay by for example partly omittingthe straightening of reeled pipelines.The method has been implemented in many recent projects in the North Sea


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A.4 References/A1/ Guijt J, Upheaval Buckling of Offshore Pipelines: Overview and Introduction, Paper no. OTC

6487, 22nd Offshore Technology Conference, Houston, Texas. Locke R. B. and Sheen R.The Tern and Eider Pipelines, Proc. European Seminar on Offshore Pipeline Technology,Amsterdam, NL. 1989.

/A2/ Nystrøm, P. R., Tørnes, K., Karlsen, J. S., Endal, G. and Levold, E., Design of the ÅsgardTransport Gas Trunkline for Thermal Buckling, Proc. of the 11th International Offshore andPolar Engineering Conference, ISOPE-2001, Stavanger, Norway, 2001.

/A3/ Matheson, I, Carr M, Peek R, Sanders P and George N, Penguins Flowline Lateral BuckleFormation Analysis and Verification, OMAE 2004, 23rd International Conference on OffshoreMechanics and Arctic Engineering, 2004, Vancouver, Canada.

/A4/ Harrison G E, Brunner M S and Bruton D, King Flowlines - Thermal Expansion Design andImplementation, Proceedings of the Annual Offshore Technology Conference. OTC 15310,2003.

/A5/ Jayson D, Delaporte P, Albert J, Prevost M, Bruton D and Sinclair, F, Greater Plutonio Project -Subsea Flowline Design and Performance, Offshore Pipeline Technology Conference, 2008.

/A6/ Ellinas C P et al., Prevention of Upheaval Buckling of Hot Submarine Pipelines by Means ofIntermittent Rock-dumping, Proc. Of 22nd Offshore Technology Conference, 1990, OTC 6332,Houston, Texas, USA.

/A7/ Sævik S and Levold E, High Temperature Snaking Behaviour of Pipelines, Proc. InternationalOffshore and Polar Engineering Conference, 1995, Vol.2, The Netherlands.

/A8/ Sævik S, Levold E, Johnsen O K, Breivik J and Hansen W, Lateral Instability of HighTemperature Pipelines, The 20" Sleipner Vest Pipeline, Proc. of Offshore Mechanics and ArcticEngineering, 1996 Vol. V, Italy.

/A9/ Nystrøm P R et al., Design of the Åsgard Transport Gas Trunkline for Thermal Buckling, Proc.of the Eleventh International Offshore and Polar Engineering Conference, 2001, Stavanger,Norway, June 17-22.

/A10/ Hobbs R E, In-Service buckling of Heated Pipelines, ASCE Journal of TransportationEngineering, 1984, vol. 110, pp 175-189.

/A11/ Peek R and Yun H, Flotation to Trigger Lateral Buckles in Pipelines on a Flat Seabed, Journalof Engineering Mechanics, ASCE, April 2007.

/A12/ Anderson M, Bruton, D, Carr M, The Influence of Pipeline Insulation on InstallationTemperature, Effective Force and Pipeline Buckling, 26th International Conference onOffshore Mechanics and Arctic Engineering, 2007.

/A13/ Peek R, Kristiansen N, Zero Radius Bend Method to Trigger Lateral Buckles,OMAE2008-58046, 2008, Proceedings 27th International Conference on Offshore Mechanicsand Arctic Engineering.

/A14/ Endal G and Egeli H, Reel-lay method to control global pipeline buckling under operatingloads, 2014 OPT Conf, Amsterdam.

/A15/ Endal G and Nystrøm P, Benefits of generating pipeline local residual curvature during reel-and S-lay installation, 2015 OPT Conf, Amsterdam.


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APPENDIX B PROBABILISTIC BUCKLE FORMATION (INFORMATIVE)

B.1 Probabilistic buckle formation model

B.1.1 GeneralWhen the characteristic VAS is defined in terms of an exceedance probability, a structural reliability analysisof buckle formation is required. This appendix presents a probabilistic methodology that can be used toestimate the characteristic VAS. However, any documented method which is capable of evaluating thecharacteristic VAS in a realistic way can be used. Project specific work should be performed in detailed designto confirm the chosen distributions (in terms of probability density functions, PDFs) for vertical imperfectionsand engineered triggers .

Guidance note:Experience has shown the critical buckling load on vertical imperfection as suggested this Appendix may have a bias towards thecritical buckling load established by FE.For route bends and horizontal out-of-straightness it is expected that limited benefit from FE is obtained. For route bends evensmall vertical imperfections will reduce the critical buckling load, and hence FE may be used to compare critical buckling loads onflat seabed and actual seabed to evaluate this effect.


The general steps required in the structural reliability process are:

1) Define a quantitative model of the buckle formation process (see for example [B.1.2]) and develop theunderlying models.

2) Establish probability distributions which describe the uncertainty in the various key input parameters(stochastic variables). The stochastic variables will depend on the buckle formation model adopted.However, any model will require distributions for:

— axial pipe-soil resistance— lateral pipe-soil resistance— pipeline imperfection data.

3) Undertake a probabilistic simulation of the pipeline buckle formation process which can deliver thecharacteristic VAS e.g. a Monte Carlo simulation.

4) Extract probabilistic results. The results of the analysis should be processed to:

— identify the overall probability of lateral buckling for the pipeline— identify the VAS distribution for each potential buckle location and establish the characteristic VAS

along the route of the pipeline.

Since frequent buckle formation is generally desirable, it is important that the simulation does notoverestimate the likelihood of buckling. Specific issues which can compromise buckle formation are:

— Higher than anticipated residual lay tension - this will reduce the driving force for buckling and, forexample, may occur due to residual thermal loads, installation of in-line structures or flooded pipelineinstallation.

— Greater than expected pipeline embedment or significant self-burial, especially if there is an extendedduration between installation and start-up, thus compromising the buckle formation on an even seabed. Asimilar concern may arise if the pipeline initially operates at a lower temperature, and the lateral restraintincreases to be greater than anticipated once full operating conditions are imposed.

These areas of concern can be evaluated using the buckle formation model in [B.1.2].


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B.1.2 Buckle formation modelThe buckle formation model can be formulated in terms of an effective axial force. Buckling will occur at anypoint in the pipeline at which the driving force equals to or exceeds the critical buckling force, i.e.:

(B.1)

This can be reformulated as a load resistance ratio:

(B.2)

The probability of buckling is then given by:

(B.3)

Therefore, to find the probabilistic definition of characteristic VAS, the problem becomes one of determiningthe driving force and critical buckling force in the pipeline.

B.1.3 Driving force modelThe buckle formation assessment should evaluate the pipeline behaviour as the effective force increasesduring hydrotest and then during operational start-up. The buckle formation process during hydrotest isessentially the same as outlined in this subsection for first load and therefore is not explicitly discussed.Buckling during hydrotest may allow the success of the design buckling strategy to be evaluated prior tooperational start-up.The driving force profile depends on the changing operating conditions, and the location and form of anyprior buckles, as well as the axial pipe-soil resistance, any pipeline end reactions, pipeline routing and anyfree spans.The pipeline behaviour during three initial stages during start-up is illustrated in Figure B-1. The dashed blacklines show the development of the fully constrained force as the start-up develops. The shape of the forceprofile during start-up differs from the steady state force profile during operation. The first force profile (red)illustrates the situation at the point just prior to any buckle initiation. Initiation will occur at the first locationalong the pipeline that achieves the critical combination of axial force and out-of-straightness (i.e. the drivingforce exceeds the critical buckling force).The critical buckling force distribution along the pipeline can cause buckling at any point downstream of thetemporary hot anchor point, but in the absence of engineered triggers or significant bathymetric OOS, theeffective axial force profile cause that buckle formation will most likely initiate near the hot end.As the start-up conditions increase, the buckle develops. The force within the buckle drops as the pipelinefeeds-in to the buckle, the slip zones become more developed and the locations of the anchor points change.Prior to buckling the pipeline is expanding from the hot anchor point towards the end of the pipeline (i.e. tothe left in Figure B-1. When a buckle appears, the first virtual anchor forms at a location that has alreadytranslated away from the buckle site, this affects the feed-in to the buckle, but can be conservativelyneglected.The green profile in the figure is just before a second buckle initiates. Here the force is higher than thatattained in the original pre-buckled profile. Again, this may not be the case because whether a second buckleforms at a higher or lower force than the first buckle is depending on the out-of-straightness and the seabedfrictional resistance.


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Figure B-1 Buckle development during start-up - stage 1

Further progress of the start-up process is illustrated in Figure B-2. As the second buckle forms, the forceat this location begins to fall. This gives rise to feed-in to the second buckle and the development of ananchor point between buckle #1 and buckle #2. The force within the first buckle continues to drop as itsdevelopment continues, but the rate of fall slows in-line with the response illustrated in Figure B-2. As forstage 1, the positions of the anchor points vary as the start-up progresses. Formation of further buckles isillustrated in Figure B-3.A third and fourth buckle form. The forces in the other buckles continue to drop leading to the final forceprofile shown in green. This fully developed force profile can occur before the system start-up is complete.As the frictional slopes now govern the force in the system, any further increase in operating conditions willnot modify the force profile but the potentially high fully constrained force will be dissipated into further endexpansion and feed-in to each buckle.


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Once the buckle formation process is complete, the VAS can be identified, as illustrated in Figure B-4.


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Figure B-4 VAS for final buckle configuration

As indicated above, the final VAS for each buckle may be less than the transient VAS during the formationprocess. The load effects within a lateral buckle depend on both VAS and pressure and temperatureconditions (amongst other factors), and will generally be greater at the design pressure and temperaturethan during buckle formation. This buckle formation methodology can be applied to both transient and fullydeveloped VAS, but the use of the fully developed VAS is simpler and, with rare exceptions, conservative. Ifthere is any concern over a proposed buckle formation strategy, the load effects throughout the formationprocess can be readily assessed.

B.1.4 Axial and lateral pipe-soil resistance modelsAxial pipe-soil resistance and lateral breakout pipe-soil resistances should be modelled as independentvariables.

Guidance note:Lateral resistance is only fully mobilized over short distances (at buckle sites) whereas the mobilized axial resistance controls theaxial force profile along the pipeline and is not significantly affected by local variations.


Both distributions can be represented by lognormal distributions fitted to the best, lower and upper estimatevalues. The standard deviations should reflect the confidence in the predictions of friction and breakoutresistance, and the level and suitability of available test data. The distribution of the pipe-soil resistance isnormally not well known, and the estimates have to rely on engineering judgement. Guidance is given inDNVGL-RP-F114.

B.1.5 Critical buckling force modelsB.1.5.1 GeneralThe critical buckling force varies from point to point along the pipeline. It depends on several factors,including lateral pipe-soil resistance and initial out-of-straightness for an even seabed. This introducea significant amount of uncertainty and are important stochastic variables that should be given careful


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consideration when performing a probabilistic buckle formation analysis. For uneven seabed or use ofbuckling triggers, the critical buckling load is mainly governed by the uncertainty in geometry, and hence theuncertainty in the critical buckling load is often smaller as compared to the even seabed case.If the lateral break-out pipe-soil resistance is high, an assessment that only considers lateral pipelineimperfections may underestimate the probability of buckle formation as there may be greater tendency forbuckling to initially occur in the vertical plane.The appropriate model to estimate the critical buckling force varies depending on the feature considered.The models outlined here are suitable for developing distributions for use during conceptual design. Wherepossible, project specific analyses should be undertaken to develop appropriate and robust distributions foruse during detailed design.

Within this appendix, lognormal distributions are described by the mean (μ) and standard deviation (σ) ofthe stochastic variable, X. The standard deviation of ln(X) is given by:

(B.4)

The mean of ln(X) is given by:

(B.5)

B.1.5.2 As-laid imperfectionsThe initial imperfection or out-of-straightness (OOS) in the pipeline is not known prior to installation. Theinformation provided here is based on OOS data gathered for many pipeline installation projects with flator moderately uneven seabed, and care should be taken in its application. Given uncertainties in the actualOOS distribution, sensitivity analyses should be performed to understand the criticality of OOS distributions.Where possible, project specific work could also be performed to support the chosen distributions.If the design strategy is sensitive to the uncertainty in the OOS distribution, and, particularly if engineeredbuckle initiators are not proposed, it may be necessary to perform a post-installation OOS survey to confirmthe design assumptions prior to start-up.

B.1.5.3 Critical buckling force - nominally straight pipeFor straight laid pipeline (and sections in between curves), imperfections occur because of the lateralmovement of the vessel and pipeline during the lay process. The critical buckling force for a straight laidpipeline is given by:

(B.6)

where μL is the lateral breakout friction factor and

(B.7)

where w is the submerged weight of the pipeline per unit length.

XNH is a normalised critical force parameter which is a measure of the inherent lateral out-of-straightnessin a nominally straight pipeline. If XNH is equal to unity, the critical force is the Hobbs minimum force for aninfinite mode buckle.

Identifying a suitable distribution for the parameter XNH is difficult. It shall represent the type of OOS foundin actual pipelines and the effect of this OOS on the buckling force. The OOS will depend on many factors andconsequently XNH will vary from project to project and sometimes from pipeline to pipeline within a project.


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In the absence of more specific data, the XNH parameter can be described by a lognormal distribution with amean value of 1.26 and a standard deviation of 0.33. This distribution represents a best fit to the results ofa set of 2D and 3D analyses of 434 kilometres of as-laid pipeline data with modest levels of vertical out-of-straightness. Sensitivity analyses should be performed to reflect the variability of the distribution.

Guidance note:A lognormal distribution with a mean of 0.93 and a standard deviation of 0.27 represents a lower bound fit to the data set, whilsta lognormal distribution with a mean of 1.58 and a standard deviation of 0.31 is an upper bound fit. It is possible that individualpipelines may have out-of-straightness exceeding these bounds.

Figure B-5 Nominally straight pipeline - XNH distributions



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The distribution is applicable to a one kilometre length of straight pipeline, i.e. it describes the most severeout-of-straightness feature that can be expected in each kilometre of straight pipeline.

B.1.5.4 Critical buckling force - route curvesThis model of critical buckling force is applicable to large radius route curves and large radius curvesassociated with a snake-lay design strategy. The critical buckling force is given by:

(B.8)

For a given target curve radius R, the actual as-laid shape will contain features that are more severe thanthe target, and the parameter xNB is a measure of this variation from nominal radius. In the absence of morespecific data, XNB can be represented by a lognormal distribution with a mean given by Equation (B.9) and aCoV of 30%. The distribution is developed from analysis of as-laid large radius curves.

(B.9)

where s, the curve arc length, shall be defined in kilometres.The distribution describes the most severe out-of-straightness feature that can be expected in the lesser ofthe curve arc length or one kilometre.If the curve radius is large, the critical buckling force distribution may exceed that for a nominally straightpipeline, and the straight pipeline critical force distribution should be adopted.Currently, there is no numerical equations for estimating the critical buckling load for pre-bent sections, butthese can be estimated through finite element analysis for the relevant variations in pipe geometry, pipe-soilresistance and level of pre-bending.

B.1.5.5 Critical buckling force - vertical triggerVertical imperfections (e.g. sleepers) can be employed to initiate buckling. The critical buckling force from apropped shape type vertical imperfection with height, h, is given by:

(B.10)

where:

(B.11)

The parameter XNV describes both the uncertainty in the model and the effect of the lateral OOS near thevertical imperfection.

Analysis of as-laid data from a single deep water project indicated that XNV followed a lognormal distributionwith a mean value of 0.725 and a standard deviation of 0.14. Care should be taken in the use of thisdistribution.

A value of 1.0 for XNV provides an upper bound to the critical buckling forces of the as-laid data consideredand therefore could be used to give an estimate of the critical buckling force for conceptual design.Projects adopting vertical triggers will need to develop suitable distributions of the critical buckling force.These can be developed using FE analysis models and should consider uncertainties in vertical trigger height,horizontal OOS, pipe-sleeper resistance, and pipe-soil resistance and embedment at touch down points etcetera.

B.1.5.6 Critical buckling force - buoyancy triggerBuoyancy triggers can be employed to initiate buckling. These are likely to involve local OOS because of thechange in submerged weight, and possibly change in outside diameter.


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The critical buckling force at a buoyancy trigger is given by:

(B.12)

where:

(B.13)

ScharB accounts for the reduction in lateral resistance associated with a reduced submerged pipe weight (wB is the submerged weight per unit length of the pipeline with buoyancy).

XNBo is a normalised critical force parameter and describes the OOS at the buoyancy triggers due to thereduced level of embedment of the larger diameter, buoyant pipeline. No data exists to provide guidance onan appropriate distribution for XNBo.Projects adopting buoyancy triggers will need to develop suitable distributions of critical buckling force. Thesecan be developed using FE analysis models and should consider uncertainties in buoyancy uplift, embedmentin normal and buoyancy section, vertical and horizontal OOS, pipe-soil resistance et cetera.Alternatively, temporary buoyancy can be used to make the pipeline positively buoyant at the trigger duringstart-up. For subsequent operation, the buoyancy will be removed. Analytical solutions for temporarybuoyancy have been developed (see /B1/), but these should be validated with FE analysis

B.1.5.7 Critical buckling force - zero-radius bendsSuitable distributions of critical buckling force can be developed using FE analysis models, and shouldconsider uncertainties in vertical and horizontal OOS, pipe-sleeper resistance, and pipe-soil resistance andembedment at touch down points et cetera.An analytical solution has been developed for zero-radius bends (see /B2/). This solution should be used withcaution because it is known to underestimate the critical buckling force.

B.1.5.8 Critical buckling force - push sleepersSuitable distributions of critical buckling force can be developed using FE analysis models, and shouldconsider uncertainties in vertical and horizontal OOS, pipe-sleeper resistance, and pipe-soil resistance andembedment at touch down points etc.

B.1.5.9 Critical buckling force - uneven seabedThe potential for seabed OOS-induced buckling should be considered. There may be increased tendency forbuckling to be initiated from vertical seabed OOS if either significant spanning can occur or if lateral breakoutpipe-soil resistance is high.It is not possible to provide generic data for vertical seabed OOS as this is entirely site specific. If goodquality bathymetric data is available, then the vertical OOS can be analysed within FE analysis modelsto provide an estimate of the severity of vertical imperfections. These analyses should identify specificsignificant features which should be incorporated into the buckle formation analysis in the same way asengineered triggers. This would be appropriate where many distinct features are present, each of which canbe treated individually. A few sensitivity cases should be undertaken to assess the variation in buckling forcewith key parameters (e.g. variation in imperfection profile within the lay corridor, vertical seabed stiffness,survey data error, lateral resistance) and to define a suitable distribution of critical buckling force for eachfeature.

Guidance note:One method to do assess the critical buckling force is to deform an initially straight pipeline to the seabed bathymetry data andincrease the operating load in the pipeline until bucking occurs. This will identify the most severe OOS features. Further, forevaluating the less severe OOS features, buckling may be prevented in the FE analyses at the OOS features where buckling isalready captured and evaluated, to identify the critical buckling loads for the remaining features.



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The FE analysis should also identify the influence of any free spans on the force profile in the pipeline. If thebathymetry leads to pipeline free spans, then the expansion behaviour of the pipeline will be modified. Thepipeline will expand into the free spans, absorbing some of the feed-in that would otherwise drive lateralbuckling. Very significant spans will modify the force profile in a similar way to lateral buckling, i.e. theeffective axial force will drop at the span and virtual anchor points may form between spans (or betweenspans and lateral buckles), thus tending to reduce the available driving force for lateral buckling and reducethe feed-in to the buckles. This effect would need to be captured adequately in a Monte Carlo simulation.

Guidance note:It may be time consuming to perform probabilistic buckle formation analyses for uneven seabeds in which there are a significantnumber of vertical features, each of which would require a number of FE analyses to define appropriate critical buckling forcedistributions, and where free spans impact the build-up of effective axial force. In this case, the sharing criterion may be easierto implement (albeit FE analyses will still be required to define suitable upper bound buckling forces associated with the seabedbathymetry).


B.1.5.10 Critical buckling force - pre-bent sectionsIn case pre-bent sections, e.g. by use of the residual curvature method, are used as trigger mechanisms forglobal buckling, these sections may be considered very sharp curves. Due to the high curvature, the bucklingforce will be significantly lower than what may be achieved by for example snake-lay. In addition, the as-laidpre-bent section has an inherent bending moment due to the lay tension trying to straighten out the pre-bentsection, and this bending moment contributes to pushing the pipeline laterally/lateral buckling. Currently,there is no numerical equations for estimating the critical buckling load for pre-bent sections, but these canbe estimated through finite element analysis for the relevant variations in pipe geometry, pipe-soil resistanceand level of pre-bending. Some considerations of residual curvature application for lateral buckling, see /B3/.

B.2 Probabilistic results and interpretation

B.2.1 GeneralThe results of the assessment should be interpreted to identify:

— probability of lateral buckling— VAS exceedance probability distribution and characteristic VAS value for each location along the length of

the pipeline.

B.2.2 Probability of lateral bucklingThe overall probability of lateral buckling should be determined to assess the likelihood of at least one buckledeveloping.The probability distribution for the number of expected lateral buckles within a pipeline may be determined.A typical distribution of the expected number of buckles in a long pipeline with no buckle initiation strategyis illustrated in Figure B-6. The left plot of the figure shows the probability density function (pdf) for thenumber of buckles in total for the whole pipeline, while the right plot shows the associated probability ofexceedance for the number of buckles.In this case, several buckles are likely to occur and the median number of buckles is between 8 and 9. Theprobability that there are no buckles in the pipeline is extremely low. In fact, the probability that there will beless than 5 buckles is very low, the same is the probability that there will be more than 16 buckles.


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Figure B-6 Distribution of number of buckles

B.2.3 Distribution of virtual anchor spacing and characteristic virtual anchorspacingFor a buckle trigger, the distribution of VAS can be obtained by recording the VAS at that trigger in eachsimulation. If no buckle forms, the VAS should be recorded as zero (allowing the probability of trigger failureto be identified). Similarly, for an unplanned buckle, the distribution of VAS can be obtained by recordingthe maximum VAS in each kilometre of the pipeline in each simulation. The VAS cumulative probabilitydistributions can then be plotted as shown in Figure B-7, and the characteristic VAS can be extracted.For the trigger example, Figure B-7a, buckling at the trigger is very reliable. There is a 2% chance that thetrigger fails to initiate a buckle, so the distribution starts with a probability of exceedance very close to one.For a trigger, the characteristic VAS is defined by a 10% chance of being exceeded, which in this examplegives a characteristic VAS of 2 840 metres. Note also that the VAS is almost constant from 10% to 0.1%exceedance probability, implying a robust trigger design.For the unplanned buckle example, Figure B-7b, unplanned buckling is reasonably likely to occur, with a 31%chance of buckling within each kilometre of pipeline. For unplanned buckling, the characteristic VAS is definedby a 1% chance of being exceeded in any kilometre of pipeline, which in this example gives a characteristicVAS of 4 900 metres.


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Figure B-7 Typical distribution of VAS

B.2.4 Characteristic virtual anchor spacing along pipeline lengthThe characteristic VAS along the pipeline length for both planned and unplanned buckles should bedeveloped.

B.3 Finite element assessmentIt is possible to support and confirm, or modify, the formation reliability analysis using an FE model of thepipeline. However, it is important that the results of the FE analysis are robust to the uncertainties withinthe problem. An FE model will buckle from the OOS inherent in the model. However, the existence of OOSwithin the pipeline system cannot be known prior to installation. Hence, an FE model is mostly appropriatefor evaluating the critical buckling forces for vertical imperfections.For example, initial curvature may be introduced into the FE model to initiate buckling in an otherwisestraight model. Although buckling will occur at the imperfection as intended, the assumption that the restof the pipeline is straight is unrealistic. The absence of the real OOS information means that the result of asingle FE analysis gives limited information on the likelihood of buckle formation.This analysis approach may be performed using the bathymetric profile for moderate uneven seabed anduneven seabed. This can provide useful insight into the key vertical OOS features but it is important that theuncertainty in the bathymetry including variations within the lay corridor and other data is fully explored.To demonstrate robust and frequent buckle formation under all conditions, it is necessary to performanalyses to address a significant number of additional issues, for example the effects of:

— global variation in friction, a conservative choice of friction for buckling limit states is not necessarilyconservative for buckle formation

— local variations in friction at both planned and unplanned buckle locations— variations in the critical buckling force of engineered initiators— large OOS away from planned buckle locations— survey errors in vertical OOS— local variations in vertical seabed stiffness.


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B.4 References/B1/ Peek R and Yun H, Flotation to Trigger Lateral Buckles in Pipelines on a Flat Seabed, Journal of

Engineering Mechanics, ASCE, April 2007.

/B2/ Peek R and Kristiansen N, Zero-Radius Bend Method to Trigger Lateral Buckles, OMAE2008 58046.

/B3/ Cooper P, Zhao T and Korteaas F, Residual curvature Method for Lateral buckling of DeepwaterFlowlines, OPT 2017.


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APPENDIX C EFFECT OF ENVIRONMENT ON FATIGUE ANDFRACTURE

C.1 Environment and loading frequencyS-N curves and fatigue crack growth laws should be appropriate to the environment, temperature andloading frequency, and to the type and location of the flaw being assessed. Similarly, fracture toughnessshould take account of the effect of environment and temperature.In corrosive environments, the endurance limit and threshold for the initiation of fatigue crack growth shouldnot be applied unless demonstrated by project-specific testing. Corrosive fluids in a pipeline are characterisedas sweet (i.e. no H2S) or sour. In these environments, the fatigue performance of the weld root and the crackgrowth rate of internal surface flaws will be worse than that in air.The fatigue performance of welds and the crack growth rate of external surface flaws exposed to seawater,even if protected by cathodic protection, will also be worse than in air. The presence of sulphate reducingbacteria in soils may also result in similar effects to those seen in a sour environment. However, there isinsufficient published information to offer quantitative guidance in this case.If a high-integrity field joint coating is used, the weld cap is protected from the external environment andin-air behaviour can be assumed. A high integrity field-joint coating should demonstrably exclude seawaterfrom the weld under all loads that the pipeline will experience from installation to the end of the design life.Project-specific testing may be required to demonstrate the performance of the field joint coating.Corrosion fatigue is sensitive to the loading frequency. Lateral buckling imposes a low-frequency loading(typically less than 10-5 Hz). Per cycle, low frequency loading is more severe than high frequency loading.Indicative knock-down factors (also referred to as fatigue life reduction factors) to be applied to the in-air S-N curves and acceleration factors to be applied to the two-stage in-air fatigue growth laws are given in TableC-1. Knock-down and acceleration factors need not necessarily be equal, but in the case where the fatigueperformance of a welded joint is governed by fatigue crack growth (rather than initiation), they would beexpected to be the same. The factors are based on fatigue tests on girth welds in low carbon-manganeseline pipe steel undertaken as a basis for this recommended practice and literature reviews of a limitednumber of reported tests, see /C1/ and /C2/. The factor for sour conditions is a semi-empirical model thathas been fitted to laboratory test data generated under various environmental conditions. In the absence ofbetter data, these knock-down/acceleration factors are suitable for use in conceptual design. Environmentalparameters such as pH, temperature, and the partial pressures of H2S and CO2 are likely to influence theobserved factors. The values quoted in Table C-1 may be non-conservative in specific cases. Project-specifictesting is recommended where the project-specific environment is significantly different from that covered bythe existing laboratory test data.In non-corrosive environments, in-air behaviour can be assumed.In conceptual design, it is often useful to compare the anticipated environmental conditions within thepipeline with those used in previous testing programmes. However, comparison of different test data is notstraightforward, because differences in cyclic loading frequency, stress range, and in-air fatigue performance,can themselves have an influence on the apparent knock-down, and mask the true environmental effect.Consequently, it is recommended that knock-down factors are determined with respect to an absolutebaseline (e.g. a design S-N curve) rather than the actual in-air performance observed in testing.


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Table C-1 Environmental knock-down/acceleration factors for low carbon manganese steels

Environment Knock-down/acceleration factor

sweet corrosion 1) 3

sour corrosion 2), 3), 4)

sea-water and cathodic protection 5) 9

sulphate reducing bacteria 5) no information

1) It should be noted that it is not very much test data to support the recommendations for sweet corrosion.2) In a nominally dry sour environment, the same knock-down/acceleration factor as for wet sour environment should

be used, unless project-specific tests justify a lower factor.3) A knock-down/acceleration factor of 40 should be assumed when environmental conditions are not well

characterised (corresponding to that derived from tests in a saturated NACE B solution).4) pH2S shall be given in mbar.5) In-air behaviour should be assumed if a high-integrity field joint coating is used and with well documented long-

term properties.

The effect of the local environment and the temperature on the fracture toughness should be considered.In sour conditions, active corrosion results in the adsorption and diffusion of atomic hydrogen, eitherinto the plastic process zone at the crack tip or into the steel. This, in turn, leads to a reduction in thefacture toughness and/or environmental cracking, respectively. Depending on the environment (i.e. pHand the partial pressure of H2S) and the operating conditions (specifically temperature), the material maybe susceptible to hydrogen embrittlement and/or stress corrosion cracking. The presence of corrosioninhibitor may have a significant influence on the rate of hydrogen uptake and therefore the resulting fracturetoughness, but experimental data quantifying the effect is limited.In sour conditions, it is necessary to consider the location of the flaw being assessed when determining theappropriate value of fracture toughness or fatigue crack growth rate. Internal surface breaking flaws will bedirectly exposed to the service environment. Embedded flaws will not be directly exposed, but the crack tipmaterial may still be influenced by hydrogen that has diffused through the steel to that location. Differenttest methods for evaluating material performance under these different conditions are discussed in [C.3] and[C.4] below.Project-specific testing is recommended for sour service, unless clad or lined pipe is used and the weldingconsumable has been shown to be unaffected by the sour environment

C.2 Corrosion resistance alloysCorrosion resistant alloys (CRA) may be used in pipelines subject to high levels of corrosion. The corrosionfatigue behaviour of CRA materials is better than carbon-manganese steels, but there remains significantuncertainty.A suitable CRA material should exhibit little degradation of fatigue performance in a corrosive environmentand, in this case, in-air behaviour can be assumed. However, material selection to avoid general or localisedcorrosion does not necessarily imply immunity to corrosion fatigue. There is some evidence to suggest thatsuper duplex stainless steels suffer a knock-down in a sour environment, see /C3/. Therefore, it is tentativelyrecommended that a knock-down factor of 10 is applied in sour service, where either the base metal or weldmetal are of this alloy type. It is acknowledged that this is based on very few actual test data and that theperformance of the material is also likely to depend on specific environmental conditions. Project-specifictesting is recommended to establish appropriate knock down and acceleration factors.The performance of super duplex material in seawater (with cathodic protection) is generally considered tobe comparable to that in air, although again, there are very few published test data, in particular at very lowcyclic loading frequencies.


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Conventional fatigue design curves are derived from tests measuring the number of cycles to specimenfailure, where failure is represented by crack growth through a significant proportion of the wall thickness.For clad or lined pipeline, the growth of a surface root flaw to the same extent may not acceptable, becausethe corrosive fluids might then come into contact with the backing carbon steel material (depending on thegeometry of the weld and the weld consumables), see DNVGL-RP-F108 for more details.

C.3 Fatigue testing in a corrosive environment

C.3.1 GeneralCorrosion fatigue is sensitive to the material, geometry, load, frequency, waveform, environment andtemperature. The influence of environmental parameters (e.g. H2S and CO2 concentrations, pH, salinity,temperature) and the effect of corrosion inhibitor on corrosion fatigue behaviour is not well understood.Lateral buckling imposes a very low frequency loading (less than 10-5 Hz). Per cycle, low frequency loading ismore severe than high frequency loading. The S-N curves and fatigue crack growth laws quoted in standardsare based on tests conducted at high frequencies (typically 0.1 Hz and above). Consequently, the focus of aproject-specific test programme is most likely to be the effect of frequency on the fatigue behaviour.A significant number of tests are required to produce a new S-N curve or a new fatigue crack growth law.The extent of the required project-specific testing will depend on the similarity of the environment to that inprevious projects or to that in tests in the published literature, and the sensitivity of the design to fatigue.It may be sufficient to conduct a small number of tests to confirm that the assumed S-N curve and fatiguecrack growth law are conservative.A project-specific test programme should consist of tests on project-specific materials and weld proceduresqualified for the project and should be conducted under conditions that are representative of the actualconditions.Temperature varies through the buckling cycle. Consequently, it may be necessary to undertake some initialtesting to identify the most appropriate test temperature. In sour service, the most severe conditions may beat ambient temperature.A test programme should include both endurance tests and fatigue crack growth rate tests because thetwo different types of tests are complementary. The knock-down and acceleration factors relative to in-airbehaviour would be expected to be similar, although this is not always the case.Fatigue testing at low frequencies is time consuming. The implications of the time taken to conduct tests onthe project schedule should be considered.

C.3.2 Fatigue crack growth rate testsThe purpose of fatigue crack growth rate tests is to determine an appropriate fatigue crack growth law oracceleration factor for the environment.Fatigue crack growth rate tests are a relatively quick way of generating relevant corrosion fatigue data andto explore the influences of parameters such as cyclic loading frequency. Individual tests may take severalweeks, but the quantity of data generated by a single test is significant (particularly in comparison with asingle endurance test which may take an equivalent or even longer period of time). Therefore, it is oftenadvantageous to conduct fatigue crack growth rate testing at the start of a project-specific test programme.An outline of a project-specific programme for fatigue crack growth rate testing is given in Table C-2.


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Table C-2 Outline of a programme for fatigue crack growth rate testing

Type of test Notch location Environment Number of tests

increasing or decreasing ΔK testat a high frequency to establish areference point

weld metal (WM)

fusion line (FL)

parent pipe (PP)

air

1

1

1

high ΔK frequency scanning test

WM

FL

PP

corrosive

1

1

1

medium ΔK frequency scanning test WM or FL or PP corrosive 1

low ΔK frequency scanning test WM or FL or PP corrosive 1

increasing or decreasing ΔK test atthe saturation frequency to determinethe acceleration factor

WM or FL or PP corrosive 2

FCGR tests are conducted using standard fracture mechanics test specimens, e.g. single edge notch bend(SENB) or compact tension (CT) specimens. Frequency scanning tests are conducted at a constant stressintensity factor range (ΔK), but at various (low) frequencies. The purpose of the frequency scanning testsis to establish the effect of frequency on the fatigue behaviour, and in particular to determine the frequencybelow which the crack growth rate per cycle remains relatively constant (the saturation frequency). It is notpractical to test at the actual frequency of cycling associated with lateral buckling. The saturation frequencyis typically several orders of magnitude higher than the actual buckle frequency.

The saturation frequency may depend on the value of the applied ΔK. Therefore, tests need to be carried outat a range of different values of ΔK. An ECA can be used to calculate ΔK for a range of different flaw sizes,and hence the range of appropriate values of ΔK to be considered in the test programme.

An increasing or decreasing ΔK test shall be carried out at the observed saturation frequency. Testsconducted at this frequency should give upper bound (saturated) crack growth rate data. However, testingat this frequency at low values of ΔK (at which the crack growth rate will inevitably be lower) may beimpractical. Frequency scanning tests should be conducted over a range of values of ΔK. The project-specificfatigue crack growth law should be determined by taking account of all test data that are demonstrated tocorrespond to a saturated condition.The initial fatigue crack growth rate tests should be carried out with the notch positioned to sample differentareas of the weld, e.g. the weld metal or fusion line. Subsequent tests should then employ the notch positionwhich gave the highest crack growth rates.For the assessment of embedded flaws in sour environment, the potential influence of absorbed hydrogen onthe fatigue crack growth rate should be considered. There is some evidence that rates of fatigue crack growthunder these conditions can be significantly higher than in air, and even comparable to those seen when thecrack is directly exposed to the sour environment. However, there are no standard procedures for conductingthese types of test, so the approach will need to be established on a project-specific basis.

C.3.3 Endurance testingThe purpose of endurance tests is to determine an appropriate S-N curve or knock-down factor for theenvironment.An outline of a project-specific test programme for endurance testing is given in Table C-3.


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Table C-3 Outline of programme for endurance testing

Type of test Environment Number of tests

tests over a range of Δσ, to give a fatigue life of 10 4 to 10 7

cycles, at a high frequency, to establish a baselineair 6

tests over a range of frequencies at high Δσ (from the lateralbuckling design) to establish the saturation frequency 1)

corrosive 4 - 6

tests at the saturation frequency to determine knock-downfactor

corrosive 6

1) To establish background data for this document a limited number of very low frequency endurance tests in sea-water and sour environments were conducted. The results suggest that endurance reaches a saturation (lowerbound) value at approximately 0.1 to 0.01 Hz, albeit the saturation frequency is less well defined in milderenvironments. The effect of frequency observed in frequency scanning tests conducted at high ΔK resembled whatwas observed in endurance tests. The results of frequency scanning tests should be used to select the range offrequencies to be considered in the endurance tests.

Endurance tests are conducted using specimens that contain a section of weld in the central highly stressedregion. For tests in sweet or sour environment, it is failure from the weld root that is of interest, so the capmay be ground flush and/or protected from the environment to ensure failure from the relevant location.In pipelines designed for lateral buckling, it is the fatigue behaviour at high stress ranges that is of primaryinterest. Therefore, the low frequency tests should be focussed on high values of Δσ only (rather than testingover a wide range of Δσ). The slope of the S-N curve should be taken to be equal to that observed in tests inair at high stress ranges (i.e. -3).Knock-down factors should be determined with respect to the weld classification (i.e. the design S-N curve)rather than to the observed in-air behaviour.

C.4 Fracture toughness testing in a sour, corrosive environment

C.4.1 GeneralProject-specific fracture toughness tests may be required if the internal environment is sour. The influence ofenvironmental parameters is not well understood, so whilst tests from previous projects may be informative,they may not be sufficient.Two types of tests are typically required: tests to determine the threshold stress intensity factor forsusceptibility to stress corrosion cracking (KISCC), and tests to determine the degree of hydrogenembrittlement (KIH). KISCC is used in the ECA of an internal surface flaw. KIH is used in the ECA of anembedded or external surface flaw (i.e. a flaw not directly exposed to sour environment), or when the sourenvironment has been displaced, as occurs during cold displacement.

The results of KISCC tests are typically more conservative than KIH tests, so only KISCC tests may besufficient. However, this approach may lead to an overly conservative assessment of embedded and externalsurface flaws for sour service.

KISCC and KIH tests are conducted using standard fracture toughness test specimens, e.g. single edgenotch bend (SENB), single edge notch tension (SENT) or compact tension (CT) specimens. The single edgenotch tension (SENT) specimen will tend to give a higher measured fracture toughness than the SENB orCT specimen when tested in air, but it is less clear whether there is a similar benefit when testing in a sourenvironment. If the results of SENT tests are to be used in the ECA, then the test programme should includecomparative tests conducted using both SENB and SENT specimens.


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Tests should be conducted over a range of temperatures, because it is often not obvious which temperaturewill represent the limiting case (given that both the fracture toughness and the applied load depend on thetemperature), and, in the case of rising load tests (see below), over a range of strain rates.Tests should be carried out with notch locations in the weld metal, fusion line and parent pipe material.Tests should also be conducted in air to provide a reference fracture toughness.

C.4.2 KISSC testingKISCC is a characteristic value of the stress intensity factor measured in an environment under static loadabove which crack extension due to stress corrosion cracking (sulphide stress cracking) may take place.The derived KISCC value may be taken as representative of the toughness of material at an internal surfacebreaking flaw in a pipeline exposed to sour service.

In constant load KISCC tests, the test specimen is immersed in the environment and subject to a constantload. The specimen is left for a long period of time (typically 720 hours), or until it fails. Multiple specimensare tested at different loads to determine the threshold value for crack extension (this is then KISCC).

Alternatives to the constant load test, using a rising displacement, can also be used to estimate KISCC.The rising displacement tests take less time to perform than the constant load tests. However, to obtain asuitable lower bound result, tests shall be conducted at a sufficiently slow displacement rate to ensure thereis sufficient time for hydrogen to diffuse to the fracture process zone at the crack tip.

An outline of a project-specific test programme for KISCC testing is given in Table C-4.

Table C-4 Outline programme for KISCC testing


in-air fracture toughness tests to establish a reference toughness air 3

rising displacement tests over a range of temperatures and strain ratesto determine the worst case corrosive 9 (3x3)

constant load tests at the worst case temperature todetermine the value of KISCC to be used in the ECA;KISCC defined as the threshold for crack extension

6

one of:

rising displacement tests at the worst-case temperatureand strain rate

corrosive

3

C.4.3 KIH testingKIH is a measure of the fracture toughness of a material which contains hydrogen, but is differentiated froma KISCC test by the fact that it is measured in air. The derived KIH value may be taken as representative ofthe toughness of material at an embedded flaw in a pipeline exposed to sour service. As an alternative, KISCCdata may be taken as representative for both internal surface breaking flaws and embedded flaws.The test specimen is immersed in a sour service environment to (fully) charge it with hydrogen, and thenremoved from the environment and tested to failure in air. The crack tip is protected from the environmentduring pre-charging. The loading rate in the test needs to be low enough to allow hydrogen diffusion tothe crack tip, but fast enough to avoid significant hydrogen losses from the specimen during the test. Thehydrogen concentration should be measured before and after testing to ensure that the test conditions areappropriate. Testing fully-charged specimens gives a lower bound value for KIH. Analysis of the thermaltransients in the pipe wall during a shut-down, and the rates of hydrogen diffusion into and through the steel,may show that the use of fully charged specimens to measure KIH is overly conservative.


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An outline of a project-specific test programme for KIH testing is given in Table C-5.

Table C-5 Outline programme of KIH testing


in-air fracture toughness tests to establish a referencetoughness Air 3

rising displacement, single point, tests over a range oftemperatures and strain rates to determine the limiting case corrosive 1) 9 (3x3)

rising displacement, single point, tests at the worst casetemperature and strain rate to determine the value of KIH tobe used in the ECA

corrosive 1) 3

1) The specimen is charged in the corrosive environment and then tested in air.

C.5 References/C1/ Baxter D P and Tubby P, SAFEBUCK JIP: Critical Aspects of the Fatigue Limit State of Pipelines

Designed to Laterally Buckle, OTC21510, 2011.

/C2/ Baxter D P, A Modelling Framework for Describing the Corrosion Fatigue Behaviour of Carbon SteelPipelines and Risers , OMAE 2011-49537, 30th International Conference on Ocean, Offshore andArctic Engineering, June 19-24, 2011, Rotterdam, The Netherlands.

/C3/ Gui F et al., Corrosion Fatigue Performance of Duplex 2507 for Riser Applications, OMAE-20609,Proceedings of OMAE2010, 29th International Conference on Offshore Mechanics and ArcticEngineering, Shanghai, China, June 2010.


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APPENDIX D EXAMPLE OF CALCULATION OF TOLERABLE ANDCHARACTERISTIC VIRTUAL ANCHOR SPACING

D.1 Tolerable virtual anchor spacingA feed-in FE model is indicated in Figure D-1, showing a half buckle, symmetry model to assess feed-in andthe load effects associated with this type of loading.

Figure D-1 Feed-in model

A single (or half) buckle model can be used to obtain the load effects in a buckle as a function of the feed-in for the relevant pipe-soil sensitivity analyses, e.g. see Figure D-2 and Figure D-3 below, where bendingmoment and strain are shown as functions of axial feed-in. The acceptable feed-in for local buckling andfracture limit states are also indicated by the horizontal hatched lines.

Note that a half buckle model with symmetric boundary conditions will not reflect anti-symmetrical bucklemodes e.g. the second mode. If such modes are relevant, for example when a buckle is triggered by asleeper, the FE model needs to reflect this.

Figure D-2 Bending moment response and acceptable feed-in


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Figure D-3 Strain response and acceptable feed-in

The acceptable feed-in can be found when the curvature of pipeline equals the acceptable value for thegoverning limit state. To assess fatigue and fracture, cyclic loading is required at the relevant feed-in level.

The relationship between feed-in and VAS for a buckle located midway between virtual anchor points withlittle or no gradient in the fully constrained force can be estimated from:

(D.1)

where:

Ksym = factor to account for symmetrical model (then = 2, in case of a full model = 1)δ = feed-in to the buckleE = Young' s modulusAs = steel areaSo = maximum theoretical effective axial force (absolute value)SPost = post-buckling effective axial force in buckle apex (absolute value)L = length from buckle to virtual anchor pointf = axial resistance.

For the feed-in corresponding to a limit state unity check of one, the tolerable VAS will then be:

(D.2)

The tolerable VAS will vary along the pipeline for decaying temperature profiles. The tolerable feed-in willbe constant along the length of the pipeline unless design parameters determining the load effects, e.g. soilresistance or effective axial force, and the pipeline capacity, e.g. material characteristics, changes. Variationsof the fully constrained force (So) along the pipeline should also be accounted for.


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D.2 Characteristic virtual anchor spacingThe susceptibility to buckling and the characteristic VAS will vary along the pipeline, as illustrated for a freeended pipeline in Figure D-4, because of changes in operating conditions, routing and possibly bathymetry.The blue and red curves show the characteristic VAS for a pipeline of 20 km length in which the temperatureeither does not fall significantly (blue line) or does fall significantly (red line), respectively. The parameterTout is the temperature at outlet, for the case with the red line the outlet temperature is reduced to theambient one while for the case with the blue line it is still 80% of Tin, the inlet temperature. For the casewhere the temperature drops significantly, the reduction in the characteristic VAS is considerable in thesecond half of the pipeline and buckling is not a concern beyond KP 14 (the probability of buckling in any partbeyond this point is less than 1%).At each end of the pipeline, the characteristic VAS drops to zero due to interaction between the buckle feed-in zones and the end expansion zones.It can also be seen that the maximum value of the characteristic VAS is almost 6 000 metres for both cases.So, the main question in the buckling design is whether this value is acceptable or not. Does this lead toover-utilisation of the pipeline in such buckles? If yes, the buckling design needs to be improved.

Figure D-4 Typical variation in characteristic VAS along a pipeline

When a buckle initiation strategy is adopted, the calculation of characteristic VAS should incorporate theinfluence of the triggers on buckle formation, see Figure D-5. In this example, starting with the blue casefrom Figure D-4, the characteristic VAS found unacceptable and four triggers are implemented in the bucklingdesign.The figure shows the characteristic VAS for each trigger (red squares) and the characteristic VAS perkilometre for unplanned buckling (in between the triggers). The initiation strategy can significantly reducethe characteristic VAS associated with unplanned buckles if the triggers are designed to initiate buckles in areliable way.


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Figure D-5 Influence of triggers on characteristic VAS of unplanned buckles

Cha

nges

– h

isto

ric


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CHANGES – HISTORIC

Changes August 2017 editionContent related to pipe-soil interaction has been moved to DNVGL-RP-F114 Pipe-soil interaction forsubmarine pipelines, and replaced by references to DNVGL-RP-F114.

About DNV GLDNV GL is a global quality assurance and risk management company. Driven by our purpose ofsafeguarding life, property and the environment, we enable our customers to advance the safetyand sustainability of their business. We provide classification, technical assurance, software andindependent expert advisory services to the maritime, oil & gas, power and renewables industries.We also provide certification, supply chain and data management services to customers across awide range of industries. Operating in more than 100 countries, our experts are dedicated to helpingcustomers make the world safer, smarter and greener.

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