Pipe-soil interaction for submarine pipelines

Transcription

1 RECOMMENDED PRACTICE DNVGL-RP-F114 Edition May 2017 Pipe-soil interaction for submarine pipelines The electronic pdf version of this document, available free of charge from is the officially binding version.

2 FOREWORD DNV GL recommended practices contain sound engineering practice and guidance. May 2017 Any comments may be sent by to This service document has been prepared based on available knowledge, technology and/or information at the time of issuance of this document. 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 responsibility for loss or damages resulting from any use of this document.

3 CHANGES CURRENT This is a new document. Changes - current Recommended practice DNVGL-RP-F114. Edition May 2017 Page 3

4 CONTENTS Changes current... 3 Section 1 Introduction General Objective Scope Application Contributions from joint industry projects Structure of this recommended practice Referenced standards and recommended practices Definitions...9 Contents Section 2 Modelling pipe-soil interaction General Pipe-soil interaction within the design process Establishing a geotechnical model Finite element modelling to assess pipe-soil interaction...20 Section 3 Material properties required for design and assessment General Geotechnical field and laboratory testing Geotechnical properties Pipe properties Section 4 Exposed pipelines General Pipe embedment Axial pipe-soil interaction Lateral pipe-soil interaction Soil stiffness Soil damping...61 Section 5 Buried and covered pipelines General Effect of trenching method Axial pipe-soil interaction Lateral pipe-soil interaction within rock berms Uplift resistance...69 Recommended practice DNVGL-RP-F114. Edition May 2017 Page 4

5 5.6 Rock fill over backfilled clay Downward resistance and stiffness Section 6 Treatment of uncertainties General Considerations for pipeline design and assessment...76 Contents Section 7 Special considerations On-bottom stability Free spanning pipelines Design philosophy of support fills Penetration of falling objects Carbonate soils Section 8 Bibliography Bibliography...90 Appendix A Soil models from DNV-RP-F109 (October 2010)...94 A.1 General A.2 Initial penetration A.3 Passive soil resistance...94 A.4 Nomenclature Changes - historic...96 Recommended practice DNVGL-RP-F114. Edition May 2017 Page 5

6 SECTION 1 INTRODUCTION 1.1 General Pipe-soil interaction is an important aspect of a pipeline system as it may have a large influence on both the structural behaviour and integrity of the pipeline during installation and operation. Knowledge about the soil conditions along the pipeline route is essential to evaluate the pipe-soil interaction, and the planning of soil investigations should be tailor-made for the conditions encountered during the lifetime of the pipeline. Soil variability is inevitable over large distances and is especially the case in the surficial soils. The variation in soil parameters seen in a pipeline development project is thus larger compared to traditional foundation design. During installation of an exposed pipeline, the soil around the pipe will be disturbed, affecting both the strength and stiffness properties as well as the seabed configuration close to the pipe. These installation effects are difficult to predict, as they are highly governed by the pipe motions during laying. For buried pipelines, the state of the backfilled material is challenging to predict. The complexity and uncertainty in pipe-soil interaction are significant, and require simplifications and assumptions in the engineering models. The effort spent on pipe-soil interaction should however reflect the sensitivity to the pipeline design. This recommended practice involves guidance related to pipe-soil interaction for submarine pipelines and supersedes the pipe-soil information given in the following recommended practices: DNVGL-RP-F105 Free spanning pipelines DNVGL-RP-F107 Risk assessment of pipeline protection DNVGL-RP-F109 On-bottom stability design of submarine pipelines DNVGL-RP-F110 Global buckling of submarine pipelines DNVGL-RP-F111 Interference between trawl gear and pipelines DNVGL-RP-F113 Pipeline subsea repair Hence, the pipe-soil interaction parts in the above recommended practices will be removed in next respective revision and a reference to this recommended practice will be included. The pipe-soil interaction guidance for exposed pipelines included in this recommended practice has to a far extent been developed by the SAFEBUCK joint industry project (JIP). Contributions from other JIPs to the recommended practice are included from the PIPESTAB JIP, HOTPIPE JIP and GUDESP JIP. This recommended practice follows the same principles as outlined in ISO and API RP 2GEO, but with more specific guidance. 1.2 Objective The objective of this recommended practice is to provide guidance related to pipe-soil interaction relevant for the various conditions experienced during the lifetime of a pipeline system according to the requirements set out in DNVGL-ST-F Scope This recommended practice gives recommendations on how to evaluate pipe-soil interaction for various design situations or assessments relevant for exposed and buried submarine pipelines. 1.4 Application This recommended practice is written primarily for qualified geotechnical engineers. Hence, basic geotechnical terms are not always explained. It is however important that the geotechnical engineer cooperates with the pipeline engineer to understand the design situations. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 6

7 As far as possible, engineering models based on geotechnical principles should be used. In some cases, it is difficult to establish theoretical models. In such cases, empirical models are necessary. Empirical models should be used with care and the geotechnical engineer should evaluate the validity of the model for the problem at hand, understanding its limitations. Due to the complexity in the pipe-soil interaction assessments, this recommended practice cannot be as prescriptive as other recommended practices. The calculation models presented herein can therefore be considered as examples. In general, more than one model should be evaluated, see Sec.6 for discussion about uncertainties related to pipe-soil interaction assessments. For scenarios involving pipe-soil interaction which are not captured in this recommended practice, specific evaluations by the geotechnical engineer is required. In general, when well-established methods are not available, the consequences of this uncertainty should be evaluated. Alternative design solutions may then be considered. If the consequence of an unfavourable incident do not jeopardize the pipeline integrity, a survey plan in combination with mitigation measures may also be a viable solution. Guidance for exposed pipelines obtained through the SAFEBUCK JIP has been included. Also, calculation models based on geotechnical principles have been included, and where possible, they have been compared with recognized empirical formulations used in the industry. Alternative calculation methods may also be used than those provided in this recommended practice. However, it is recommended that they have a sound theoretical basis, or that they capture better some relevant data or conditions, such as project-specific model testing of PSI resistance, or relevant field observations of existing pipelines from the same region. By use of new calculation methods, its applicability should be documented in a transparent manner which allows for independent verification. The full interaction between the pipe and the soil accounting for the stiffness of the pipe and the loads acting upon the pipe, is not covered in this recommended practice other than as brief discussions where relevant, giving reference to other relevant recommended practices. The recommendations provided in this recommended practice are primarily related to the interaction between the pipeline and the soil per unit length of the pipe. Likewise, the integrated interaction between a pipeline or flowline including spools and connected structures is covered in respective standards or recommended practices for pipelines and structures. The recommendations given herein may be used as input for analysis of such integrated interaction. This recommended practice does not consider the following items that may be included in future editions: geohazards earthquake design and assessment of pipelines riser-pipe interaction pipe-soil-structure interaction. 1.5 Contributions from joint industry projects SAFEBUCK JIP The SAFEBUCK JIP was a joint industry project, which had the aim of developing design methodologies related to high pressure, high temperature (HPHT) pipelines susceptible to lateral buckling. Extensive research with respect to pipe-soil interaction was carried out as a part of the JIP. New calculation models were developed based on small and large scale tests. The JIP mainly focused on exposed pipelines placed on deep water clays, hence the soil conditions that is covered by the SAFEBUCK database is limited accordingly. The database consists of tests primarily carried out on soft West African clays with high plasticity. The vertical embedment model proposed in SAFEBUCK is based on theoretical considerations, but the approach to use remoulded shear strength to account for cyclic effects during laying is highly empirical. The JIP recommended to perform specialized interface tests to measure the axial interface strength directly. The proposed lateral soil resistance models are empirically calibrated towards a limited database of tests. and is therefore expected to show some bias to the underlying database conditions. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 7

8 1.5.2 PIPESTAB JIP The PIPESTAB JIP was a joint industry project which was part of the development of DNVGL-RP-F109. The soil models developed from this work have been reviewed and adopted as found appropriate in this recommended practice HOTPIPE JIP The HOTPIPE JIP was a joint industry project which was part of the development of DNVGL-RP-F110. Pipesoil interaction guidance on uplift resistance of buried pipelines has been reviewed and adopted as found appropriate in this recommended practice GUDESP JIP The GUDESP JIP was a joint industry project which was part of the development of DNVGL-RP-F105. Guidance on simplified soil damping is included in this recommended practice. 1.6 Structure of this recommended practice The recommended practice is structured as follows: 1) Introduction (this section) Presents the overall objective, scope and applicability of the recommended practice, as well as relevant abbreviations and symbols (giving all the symbols in the equations). Referenced standards are listed in this section and referred to by its acronyms while bibliographies and reports are listed in Sec.8 and referenced by reference numbers. 2) Modelling pipe-soil interaction Presents an introduction to pipe-soil interaction and how soil resistance curves may be included in the pipeline analysis 3) Material properties required for design and assessment Provides guidance with respect to soil investigations relevant for pipe-soil interaction and the pipe properties required for pipe-soil assessment 4) Exposed pipelines Provides guidance for evaluating pipe-soil interaction for exposed pipelines 5) Buried pipelines Provides guidance for evaluating pipe-soil interaction for buried pipelines 6) Treatment of uncertainties Discusses in general terms the different sources of uncertainties in geotechnical design, and highlight special considerations related to pipe-soil interaction 7) Special considerations Presents other issues related to pipe-soil interaction which may not naturally be placed in Sec.1 to Sec.6. 8) Bibliography 1.7 Referenced standards and recommended practices General In case of conflict between this recommended practice and referenced DNV GL standards, the standard or recommended practice with the latest edition date shall prevail. The latest valid edition of each of the DNV GL reference documents applies. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 8

9 Referenced relevant standards and recommended practice documents are given in [1.7.2] and [1.7.3] while bibliography is given in Sec DNV GL standards and recommended practices Document code DNVGL-OS-C101 DNVGL-ST-F101 DNVGL-RP-C207 DNVGL-RP-C212 *) DNVGL-RP-F105 DNVGL-RP-F107 DNVGL-RP-F109 DNVGL-RP-F110 DNVGL-RP-F111 DNVGL-RP-F113 Title Design of offshore steel structures, general LRFD method Submarine pipeline systems Statistical representation of soil data Offshore soil mechanics and geotechnical engineering Free spanning pipelines Risk assessment of pipeline protection On-bottom stability design of submarine pipelines Global buckling of submarine pipelines Interference between trawl gear and pipelines Pipeline subsea repair *) DNVGL-RP-C212 will soon replace DNV Classification Notes Other standards and recommended practices Document code ANSI/ASME B46.1 API RP 2GEO Title Surface Texture (Surface Roughness, Waviness and Lay) Geotechnical and Foundation Design Considerations ISO Petroleum and natural gas industries Specific requirements for offshore structures Part 4: Geotechnical and foundation design considerations 1.8 Definitions Abbreviations Abbreviation BE CPT FE FEED HE HPHT JIP Description best estimate cone penetration test finite element front end engineering design high estimate high pressure, high temperature joint industry project Recommended practice DNVGL-RP-F114. Edition May 2017 Page 9

10 Abbreviation LE NC OCR PCPT PIP PLEM PLET PSI SHANSEP TUM Description low estimate normally consolidated overconsolidation ratio cone penetration test with pore pressure measurements pipe-in-pipe pipeline end manifold pipeline end termination pipe-soil interaction stress history and normalized soil engineering properties terrain unit mapping Symbols Greek characters Symbol α γ' γ c γ' fill γ pre γ rate γ' seabed δ f δ peak δ res ε res ζ ζ soil κ a κ p μ A,brk,d μ A,brk,u Description pipe-soil adhesion or roughness factor submerged unit weight cyclic shear strain submerged unit weight of backfilled material consolidation preloading effect rate factor submerged unit weight of seabed soil failure mobilization distance peak interface friction angle residual interface friction angle residual reduction factor wedging factor soil damping ratio active earth pressure coefficient passive earth pressure coefficient axial breakout friction factor in drained conditions axial breakout friction factor in undrained conditions Recommended practice DNVGL-RP-F114. Edition May 2017 Page 10

11 Symbol μ A,res,d μ A,res,u μ fill μ seabed ν ρ ρ s /ρ σ a σ' h σ' s σ' v τ φ φ fill φ peak φ res ω Description axial residual friction factor in drained conditions axial residual friction factor in undrained conditions friction coefficient between pipe and backfilled material friction coefficient between pipe and seabed soil Poisson s ratio gradient of undrained shear strength profile with depth specific mass ratio between the pipe mass and the displaced water atmospheric pressure (100kPa) horizontal effective stress mean effective stress in soil vertical effective stress shear stress drained friction angle friction angle of backfilled material peak friction angle of the soil residual friction angle of the soil angular frequency Symbols Latin characters Symbol a A berm A bm A p A pipe B c C L C V Description horizontal oscillation amplitude prior to lateral breakout displaced soil area creating a berm adjacent to the pipe penetrated cross sectional area of the pipe cross-sectional area of the pipe plugged area of falling pipe pipe-soil contact width viscous damping coefficient lateral dynamic stiffness factor in simplified evaluation for free spanning pipelines vertical dynamic stiffness factor in simplified evaluation for free spanning pipelines Recommended practice DNVGL-RP-F114. Edition May 2017 Page 11

12 Symbol D d ca d q D ref E Dissipated E Elastic E p EI F f F A F A,brk F A,brk,d F A,brk,u F A,deep,d F A,shallow,u F L,brk,d F L,brk,d,fric F L,brk,d,passive F L,brk,u,fric F L,brk,u,remain F L,brk,u F L,res,d F L,res,u Description pipe outer diameter (including coating) depth correction factor for clay depth correction factor for sand reference diameter dissipated energy within one hysteretic loop elastic energy within one hysteretic loop energy absorbed in gravel pipe bending stiffness bearing capacity factor clay (accounting for pipe roughness and soil strength gradient) uplift resistance factor axial resistance axial breakout resistance axial breakout resistance factor in drained conditions axial breakout resistance in undrained conditions axial resistance for buried pipelines in drained conditions (deep failure mode) axial resistance for buried pipelines in undrained conditions (shallow failure mode) lateral breakout resistance in drained conditions frictional part of the lateral breakout resistance in drained conditions part of the lateral breakout resistance in drained conditions which involves passive soil resistance frictional part of the lateral breakout resistance in undrained conditions part of the lateral breakout resistance in undrained conditions which involves active and passive soil resistance and soil weight lateral breakout resistance in undrained conditions lateral residual resistance in drained soil condition lateral residual resistance after breakout in undrained soil conditions F p drained passive resistance as proposed in /22/ F uplift,d F uplift,global,u F uplift,local,u G G max H uplift resistance in drained conditions uplift resistance in undrained conditions (global failure mode) uplift resistance in undrained conditions (local failure mode) shear modulus shear modulus at small strains cover height Recommended practice DNVGL-RP-F114. Edition May 2017 Page 12

13 Symbol I p K k K 0,fill K 0,seabed K L,d K L,s k lay k lay,1, k lay,2 K p K v,d K V,s L L sh m Nγ N c N q p Q v Q v0 Description plasticity index lateral earth pressure coefficient (general) linearized spring stiffness lateral earth pressure coefficient of backfilled material lateral earth pressure coefficient of seabed soil lateral dynamic stiffness lateral static stiffness touchdown lay factor touchdown lay factors (used to determine the initial embedment) passive earth pressure coefficient vertical dynamic stiffness vertical static stiffness span length span support length on one shoulder (for transfer of one-half the weight of the free span) factor that accounts for long term effect of overconsolidation bearing capacity factor for sand Theoretical bearing capacity factor for clay (for constant undrained shear strength) bearing capacity factor for sand pipe-soil contact arc length vertical penetration resistance (including depth effects) vertical penetration resistance (not including depth effects) r roughness parameter according to /21/ R a r pipe-soil S t S u s u,0 s u,1 s u,2 u,active u,backfill,reconsolidated pipe coating roughness pipe-soil roughness factor sensitivity undrained shear strength undrained shear strength at reference level for depth effects average undrained shear strength above the foundation level average undrained shear strength below the foundation level average undrained shear strength within the active failure zone average value of reconsolidated shear strength along the vertical failure plane in the backfilled material Recommended practice DNVGL-RP-F114. Edition May 2017 Page 13

14 Symbol u,bottom,reconsolidated s u,intact u,passive s u,r u,reconsolidated s u,z=0 T 0 V W f W i W op x brk x failure x mob x res y brk y res z z 0 z f z failure z ini z mod z op z su,0 Description average value of reconsolidated shear strength at the bottom half of the pipe intact undrained shear strength average undrained shear strength within the passive failure zone remoulded undrained shear strength average value of reconsolidated shear strength undrained shear strength at seabed (z=0) horizontal effective lay tension in the pipe during installation at touchdown point vertical pipe-soil force submerged pipe weight during hydrotest (flooded/waterfilled condition) submerged pipe weight during installation submerged pipe weight during operation mobilization displacement required to mobilize the axial breakout resistance lateral extent of passive failure surface axial mobilization displacement mobilization displacement required to mobilize the axial residual resistance mobilization displacement required to mobilize the lateral breakout resistance mobilization displacement required to mobilize the lateral residual resistance pipe invert embedment (general) reference level for depth effects in sand pipe invert embedment after flooding vertical extent of passive failure surface initial pipe invert embedment after laying modified height taking into account presence of a berm when calculating lateral breakout pipe invert embedment during operation reference level for depth effects in clay Recommended practice DNVGL-RP-F114. Edition May 2017 Page 14

15 1.8.4 Definitions of verbal forms Term shall should may Definition verbal form used to indicate requirements strictly to be followed in order to conform to the document verbal form used to indicate that among several possibilities one is recommended as particularly suitable, without mentioning or excluding others, or that a certain course of action is preferred but not necessarily required verbal form used to indicate a course of action permissible within the limits of the document Recommended practice DNVGL-RP-F114. Edition May 2017 Page 15

16 SECTION 2 MODELLING PIPE-SOIL INTERACTION 2.1 General Key pipe-soil interaction (PSI) parameters are the soil resistances during vertical, axial and lateral pipe movement. In this recommended practice and in most pipeline engineering practice, the PSI resistances related to exposed pipelines are described in terms of equivalent friction coefficients, defined as the available resistance to axial or lateral movement divided by the current submerged pipe weight. The vertical pipe-soil interaction is particularly important during installation, when the pipe penetrates the seabed. After installation, the vertical pipe-soil interaction is usually less important, but the resulting embedment from the installation phase is important for the subsequent lateral and axial resistances. In reality, the lateral and axial resistances are not solely dependent on pipe weight, but are influenced by pipe embedment, soil type, drainage condition, interface condition and the previous history of loading and pipe movement. It is therefore important to recognize that a pipe-soil equivalent friction factor is not a soil property, but depends on the soil properties, the pipeline properties and the mode and history of loading. PSI may be included in the analysis of a pipeline in various ways, which are listed below in order of increasing complexity: a) As a single limiting value of axial or lateral resistance (or friction factor). b) As force-displacement responses in the axial and/or lateral directions, within a finite element analysis of the pipeline (similar to the t-z and p-y load transfer methods of analysing pile response). Simple forcedisplacement (or friction-displacement) responses are bi-linear (elastic perfectly-plastic), tri-linear (with an initial peak) or piecewise linear. Additional rules may define the cyclic behaviour. c) As a general vertical-lateral response model, based on plasticity theory, implemented within a finite element analysis of the pipeline via a force-resultant macroelement. d) Through explicit modelling of the soil continuum, pipe and pipe-soil interface by using a finite element analysis software. This is computationally expensive given the need to model typically several kilometres of pipeline in a single model. This approach is rarely used except for research. For buried pipelines, the uplift resistance is often modelled as a bi-linear or a tri-linear curve. The adopted modelling approach should reflect the current project requirements, recognizing the project stage, risks and opportunities for optimization. In any design case, the geotechnical engineer needs to consider whether the soil behaviour is drained, undrained or partially drained and select appropriate calculation models. Note that different classification systems exist in different countries with respect to soil characterization, and that the same soil can behave differently for different rate of loading. If there are uncertainties in the soil behaviour, the geotechnical engineer needs to take this uncertainty into account in the further assessments. 2.2 Pipe-soil interaction within the design process During the different stages of a pipeline project, PSI should be addressed with an increasing degree of detail, sufficient to optimize the design through reduced uncertainty. The flowchart in Figure 2-1 illustrates a project PSI workflow for on-bottom pipeline design. Three sets of PSI recommendations are passed through to the pipeline engineering workflow at the different project stages desk study, preliminary and detailed design. As indicated in Figure 2-1, readily available data may be used to obtain preliminary values for the PSI parameters at early stages of the design process. Increasingly detailed estimates may then be obtained through more complex testing and analysis, as described in [3.3.2]. The level of PSI analysis performed in a project should be chosen to suit the project requirements, both to minimize risk and to maximize added value. During a project, as the geotechnical input and pipeline design conditions are refined, more complex PSI analysis becomes possible and the level of uncertainty in that modelling reduces. This potentially leads to more optimized designs. However, the cost benefit of overall optimization should be assessed along with the cost of engineering to try to achieve those optimizations. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 16

17 For re-assessment or modification during operation, field observations from installation and operation may be used to perform back-calculations to update PSI parameters for use in pipeline operational assessments and as feedback to future projects. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 17

18 Figure 2-1 Pipe-soil interaction workflow during a project Recommended practice DNVGL-RP-F114. Edition May 2017 Page 18

19 2.3 Establishing a geotechnical model When assessing any geotechnical problem, the first step is usually to establish a geotechnical model where all parameters and boundary conditions affecting the result are defined. This geotechnical model should contain all assumptions in the geotechnical analysis. This should ideally also be the first step in pipe-soil assessments. In traditional foundation design, usually the intact soil strength and boundary conditions are relatively well defined. However, for a pipeline, given the spacing of soil data along the route and the unknown effects of laying it is challenging, if not impossible to identify all possible scenarios the pipe could experience. As such, when using analytical or empirical methods, it is important to evaluate the assumptions and background for the development of the methods. When a model is based on test results, the model uncertainties are usually related to differences in how the geotechnical model is defined and how the tests are carried out. The main effects to consider when establishing a geotechnical model for pipe-soil interaction for an exposed pipeline are as follows with some of the effects illustrated in Figure 2-2: drainage conditions and loading rate (drained, partially drained or undrained soil behaviour) geometrical boundary conditions (e.g. ideal soil contact or pipe placed inside a trench) shear strength underneath the pipe (may be affected by remoulding during installation and subsequent thixotropy effects and consolidation from pipe weight) shear strength on the outside of the pipe pipe-soil interface roughness durations of and time between pipelay, flooding, hydrotest, dewatering and start-up. Figure 2-2 Geotechnical model for an exposed pipeline For buried pipelines, many of the same items as listed above for exposed pipelines are also important. The soil resistance is extremely dependent on the type and state of the backfill material. The soil properties of natural backfill are uncertain as they could be altered significantly from the in-situ properties. Hypothetically, assuming all aspects of the geotechnical model is well known, which may be the case for other types of foundations, the best way to calculate soil resistances would be to employ finite element analyses. However, as this is never the case for pipe-soil interaction assessment, finite element analyses should not be the only design tool. Finite element analyses may however help to understand the physics involved and is considered to be the best way to perform sensitivity studies, allowing the user to investigate how different scenarios (changes in the geotechnical model) would affect the result. Evaluations using finite element analyses may be relevant for calculating breakout resistances, but not necessarily residual resistances as large displacement analyses are encumbered with more uncertainty. In those cases, empirical based methods are needed. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 19

20 In general, due to the difficulties in establishing the geotechnical model, more than one model should be used in PSI assessments, see also Sec Finite element modelling to assess pipe-soil interaction If pipe-soil interaction is evaluated using finite element (FE) analysis one should thoroughly evaluate possible sources of error and their effect on the results. The following issues are of particular concern in this context: the constitutive soil model should represent the soil behaviour needed for the problem at hand the iteration procedure should not result in an overshoot of failure loads the mesh should be sufficiently fine with proper width/length/height ratios of the elements to ensure a proper load distribution throughout the soil. When establishing an FE model, several assumptions need to be made. The influence of the model assumptions should be investigated and evaluated. The model assumptions include the representation of the pipe, loading conditions, soil behaviour and soil parameters. The model should be able to capture vertical and horizontal strength variations. For pipe-soil assessments, the geotechnical model is never known, and FE analyses will not necessarily give the correct result, but is better suited to evaluate different effects and changes in the boundary conditions for the pipe, compared to analytical or empirical models with prescribed failure modes. When non-linearities associated with large displacements are to be studied by means of an FE model, the non-linearities may be represented explicitly in the FE model or the underlying large displacements may be simulated in a wished-in-place analysis. The modelling limitations for the particular analysis should be assessed. Non-linear large displacement effects are encountered in several situations, including: changes in boundary conditions, such as changes in contact area due to large displacements displacement-dependent loads, such as displacement-dependent changes in load direction large strains. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 20

21 SECTION 3 MATERIAL PROPERTIES REQUIRED FOR DESIGN AND ASSESSMENT 3.1 General This section describes the required material parameters to carry out pipe-soil interaction assessments, both related to geotechnical properties and pipe properties. 3.2 Geotechnical field and laboratory testing For evaluating pipe-soil interaction, the near surface soil properties are critical and require specific pipelinefocused activities in a site investigation and laboratory test program. Uncertain soil conditions will result in a large range between low estimate and high estimate pipe-soil interaction parameters, which will increase the uncertainty and potential mitigation costs. In some cases, it may be difficult or costly to demonstrate a robust design solution. Planning of soil investigations for a pipeline should be performed with focus on the design scenarios of importance for the pipeline. These scenarios include, but are not necessarily limited to: axial expansion and walking lateral and upheaval buckling on-bottom stability and route curve pull-out free spans pipeline supports trenching and back-filling potential external impacts like trawling equipment and anchors potential geohazards like landslides from surrounding areas that could hit the pipeline. The soil influenced by the pipe soil interaction for exposed pipelines is normally within the upper one metre and for many pipelines within a few tens of centimetres. Thus, the sampling and testing should have particular focus on the shallowest soils. Note however that the pipe embedment within a lateral buckle may achieve depths down to two metres due to cyclic movements. Where possible and in particular in soft clay, box coring should be performed obtaining blocks of up to half a meter side dimensions, from which samples can be taken for laboratory testing, or within which small scale in-situ testing may be performed. Some deeper coring should be performed in addition to the box corings. Subsea in-situ testing should be performed in addition to the corings. This could primarily consist of PCPT testing and in clays also T-bar testing, which near the surface can provide more reliable interpretation of undrained shear strength than PCPT testing will allow for. Alternatively, ball penetrometer testing may be performed. Ball penetrometer or T-bar penetrometer cyclic tests provide the fully remoulded soil strength, which is of relevance for the assessment of pipeline embedment and cyclic lateral pipeline response. For analysis of stability of pipeline supports as well as for evaluation of trenching capabilities, soil information to somewhat larger depths, i.e several metres, would be required to capture the soil strength within the predicted failure zone. A geotechnical engineer should in cooperation with the pipeline engineer be involved in the planning of the soil investigations to make sure that the information required in the subsequent design analyses is obtained. For pipelines that are buried by ploughing or jetting, the largest uncertainty is related to how the trenching method has affected the in-situ strength and stiffness parameters of the backfilled soil. The soil investigation program needs to consider both the intact soil conditions and the soil conditions following a trenching/jetting operation. The latter may require the construction of a certain length of dummy trench as part of the soil investigation program. Due to general soil variability, it will be practically impossible to obtain very accurate soil data for each location of interest where the scenarios listed above may be relevant. Thus, a proper strategy for planning the soil investigations would be to identify from geophysical surveys, possibly combined with relevant other Recommended practice DNVGL-RP-F114. Edition May 2017 Page 21

22 information, the various soil units at or very close to the surface along the pipeline route, and to perform soil sampling to identify the range of characteristics for each unit along the route. For a survey particularly aiming to provide a basis for pipeline routing and design, sub-bottom profiling should be included unless available knowledge is such that inhomogeneous conditions of the top soils can be precluded. Note that it is very challenging to characterize the upper soil very close to the seabed and the profiler frequency should be targeted specifically for this purpose. Recognized standards shall be used to carry out laboratory testing. Particular attention should be given to the planning and execution of tests required to determine very low soil and interface strengths corresponding to the very low contact stresses between the pipe and the soil. 3.3 Geotechnical properties General Geotechnical characteristics necessary for evaluations of all relevant loading conditions shall be determined for the soils along the pipeline route, including possible unstable soils in the vicinity of the pipeline. Geotechnical properties may be obtained from generally available geological information, results from geophysical surveys, including seabed topographical surveys and sub-bottom profiling, and from geotechnical in-situ tests and laboratory tests on sampled soil. Supplementary information may be obtained from visual surveys. Soil parameters of main importance for the pipeline response are: shear strength parameters (intact and remoulded undrained shear strength for clay, and angle of friction for sands) deformation characteristics (stress-strain relationships) drainage characteristics (permeability and coefficient of consolidation). These parameters should preferably be determined from adequate laboratory tests or from interpretation of in-situ tests. In addition, classification and index tests should be considered, such as: unit weight water content liquid and plastic limit grain size distribution carbonate content other relevant tests. Such tests are important to assess spatial variability of the soil conditions along the pipeline route and to validate in-situ or laboratory test results from empirical correlations or using such correlations to supplement the laboratory tests. Laboratory tests with specific modifications to suit low stress testing are the preferred method to determine the undrained strength and friction angle of pipe-soil interfaces. Interface materials that are representative of the planned pipeline coating should be used as interface roughness has a strong influence on interface friction. Guidelines regarding specialized interface tests are given in [3.3.2] Specialized pipe-soil interaction testing The following specialist laboratory equipment for low-stress shear testing may be used to derive the soil-pipe interface parameters: tilt table device for drained interface strength /1/ low-stress shear box, for drained and undrained interface strength /3/. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 22

23 For interface tests, the roughness characteristics of the interface should be representative of the planned pipe coating or covering a range of potential coating roughness. Roughness characteristics should be documented. Large and small-scale pipe models may also be used as background for evaluating the vertical, axial and lateral pipe-soil resistance in case the existing engineering models are not considered appropriate or if a certain aspect is particularly important for the project. Four types of tests allow such measurement: small scale centrifuge tests (e.g. /4/) large scale ex-situ pipe-soil tests (e.g. /5/) small scale ex-situ pipe-soil tests large scale in-situ pipe-soil tests (e.g. /6/). This allows for project specific refinement (calibration) of the engineering models and generally leads to a reduction in the model uncertainty and the range of PSI parameters. Tests should cover the range of normal effective stress that can be expected between the pipe and the soil. Testing at normal effective stresses far from the expected value should be avoided whenever possible to avoid extrapolation errors. As far as possible, the tests should be carried out in a way replicating the actual loading history of the soil underneath the pipe, e.g. accounting for penetration arising from motions during pipeline laying and for the pre-consolidation effect from the water filled/flooded condition during the pipeline hydrotest. Laboratories executing such tests should have the correct knowledge about the equipment and procedures to execute and measure the interface strength at low stress levels. 3.4 Pipe properties Typical pipeline properties required for pipe-soil interaction assessments are: submerged pipe weights, W i, W f, W op pipe outer diameter including coating, D pipe coating roughness, R a pipe bending stiffness, EI horizontal effective lay tension in the pipe during installation at touchdown point, T 0. A distinction should be made between the following pipe weights: The submerged pipe weight at installation, W i, which is usually the empty weight. The water flooded submerged pipe weight, W f. The operating submerged pipe weight, W op. A range may be required, considering the range of product density during the operating life and also potential separation on shutdown, e.g. liquid hold up. The roughness (R a ) of the planned pipe coating should be determined or estimated when planning interface testing. To avoid confusion, the pipe coating roughness should be defined as the average deviation from the mean height, in accordance with ANSI/ASME B46.1. Profilometer or laser interferometer devices provide rapid quantification of interface roughness, and may be used to gather data from existing pipe coating samples, and to assess the roughness of interfaces available for laboratory testing. The pipe bending stiffness, EI, should consider the composite behaviour including coatings (if significant), inner and outer pipes for a pipe-in-pipe (PIP) line, etc. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 23

24 SECTION 4 EXPOSED PIPELINES 4.1 General Pipe-soil interaction is a key element in the assessment of exposed pipelines. Typical scenarios involving pipe-soil interaction are lateral buckling, end expansion, pipeline walking, route-curve pullout, flow line anchoring, on-bottom stability, trawl impact and development of free spans. The consideration of low and high estimates of pipe-soil interaction parameters is generally required during design and assessment to satisfy all limit states. If not accounting for a defined range of low and high estimates of resistance, this should be a conscious choice where possible consequences are evaluated and it is possible to rectify unfortunate incidences based on survey and contingency plans. The content of this section is summarized in Table 4-1. In general, the soil behaviour may be drained, partially drained or undrained depending on the loading rate and drainage conditions and should for each scenario be considered when selecting calculation models. Table 4-1 Pipe-soil responses for exposed pipelines Response Embedment (see [4.2]) Description The initial embedment is controlled by the soil conditions and the loads during and following installation. It has a significant influence on the subsequent axial and lateral response. Axial friction (see [4.3]) Lateral resistance (see [4.4]) Soil stiffness (see [4.5]) Axial breakout response Axial residual resistance Cyclic axial response Lateral breakout response Lateral residual resistance Cyclic lateral response Vertical stiffness Lateral stiffness An initial peak in resistance that is mainly relevant to the first load response The large displacement response as the pipe expands or contracts The long term cyclic response under repeated expansion and contraction An initial peak in resistance as the pipe first displaces from the as-installed position The large displacement resistance The long term cyclic response, when the pipe becomes embedded in a trench within a buckled pipe section and soil berms grow causing a rise in lateral resistance Static and dynamic stiffness Static and dynamic stiffness Soil damping (see [4.6]) Soil damping may be introduced in dynamic analyses. Specific guidance regarding cyclic resistances (axial and lateral) are not provided in this recommended practice, as this is an area of ongoing research without clear conclusions. The process is complex, as the development of a trench at a lateral buckle will reduce the vertical reaction in the midspan, transferring vertical reaction forces towards the inflection points. When investigated, it should be followed up during operation through surveys. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 24

25 4.2 Pipe embedment General The pipe embedment is an important factor influencing the pipe-soil interaction as it will determine the boundary conditions around the pipe for the subsequent loading scenarios. The initial pipe embedment is governed by the contact stresses imposed on the soil during laying. The embedment will be influenced by the installation method (e.g. J-lay, S-lay or reeling), the weight and stiffness properties of the pipe and the sea state/vessel motions during laying and lay rate. As such, the sea state is considered the biggest source of uncertainty and a large range in predicted embedment depths is needed to cover possible scenarios. The embedment will be governed by the bearing capacity of the seabed soils, and different scenarios may occur: Pure vertical penetration (laying in calm waters, spools lowered by a crane etc.), see Figure 4-1. Combination of vertical and horizontal motions, reducing the vertical bearing capacity, leading to higher penetrations and also creating more complex boundary conditions, see Figure 4-2. Figure 4-1 Vertical penetration process Figure 4-2 Combined vertical and horizontal penetration process After the pipe is laid on the seabed, usually a pressure test with water-filled pipe is performed. In some cases, the increased weight of the pipe will lead to further pipe penetration. The boundary conditions (pipe penetration, trench development etc.) may vary during the operational life, and regular surveys are recommended to ensure that the boundary conditions around the pipe are covered by the original design assumptions. It should be realized that any embedment model will be a simplification as it is not possible to model the true soil behaviour during an unknown installation scenario including pipe motions in the touchdown zone. In this section, models for calculating the pure vertical penetration resistance only is given. Discussion of the static and cyclic effects of laying is included in [4.2.5]. When regional or local pipe embedment measurements are available from existing pipelines, e.g. post-lay or operational survey data, they can be used to narrow the uncertainty range in the calculation methodology Recommended practice DNVGL-RP-F114. Edition May 2017 Page 25

26 when applied to new pipelines nearby. The selection of key parameters such as undrained shear strength, s u or s u,r, unit weight, γ, and touchdown lay factor, k lay may be refined using these data, before the calculation model is applied to the new pipelines under consideration. Observed embedment data should not be applied directly to new pipelines unless all conditions including pipe characteristics and above mentioned key parameters are closely the same at the two locations. Instead, the calculation method should be calibrated and then reapplied, to scale correctly for the differences in the pipeline and soil characteristics and laying conditions. However, uncertainties related to sea state, type of vessel and the corresponding motions of the pipeline cannot be fully known and back-calculations should be used with care. Changes in seabed conditions, e.g. due to seabed mobility, during the operational phase may also be a source of uncertainty when backcalculating embedment from survey data. In general, due to the uncertainties in the calculation models, more than one model should be evaluated, see Sec.6. The proposed methods in this section can therefore be considered as examples. Other methods may also be relevant, see [1.4]. For re-assessments of existing pipelines, embedment measurements may be used directly in PSI assessments. Embedment measurement error and scatter should then be considered. Back-analysis of embedment can provide a critical review of parameters initially considered in design and help in refining the assessed PSI parameters and the subsequent buckling and walking behaviour. This type of back-analysis is an important aspect of a pipeline integrity management system, especially if the design is sensitive the PSI parameters Definition of pipe embedment Nominal pipe embedment is defined as the depth of penetration of the invert (bottom of pipe) relative to the undisturbed seabed (sometimes termed the far embedment in surveys), see Figure 4-3. Pipeline embedment influences the pipe-soil contact area, which affects the axial and lateral resistance. Heave of soil during penetration increases the local embedment of the pipe (sometimes termed the near embedment in surveys). Data indicates that in cohesionless soil this heave may reduce with time and may then not be reliable in providing additional axial or lateral resistance. The nominal embedment is therefore the conventional embedment definition in design and assessment to define the pipe-soil contact arc length, p. Figure 4-3 Terminology for pipeline embedment The pipe embedment may vary during the lifetime of the pipeline, changing the axial and lateral pipe-soil resistance. In this document, the embedment at various stages of the pipeline life cycle is defined as follows: - Z ini initial pipe invert embedment after laying - Z f pipe invert embedment after flooding Recommended practice DNVGL-RP-F114. Edition May 2017 Page 26

27 - Z op pipe invert embedment during first operation (start-up) - Z pipe embedment (general). Can be replaced in the calculation methods by Z ini, Z f, or Z op depending on the case considered Embedment assessment in undrained conditions General There are different calculation models available for calculating vertical embedment in undrained conditions. For a pipe pushed vertically into the soil, the embedment depth will be the depth where the pipe contact force is in equilibrium with the bearing capacity of the seabed soil. In this section, two approaches are given. Both models give comparable results for normal conditions, but could deviate for special conditions. It is recommended to evaluate various models to assess model uncertainties, see Sec.6. Finite element analyses may also be used as stated in [2.4] Model 1 The penetration resistance may be estimated using the following approach, applying bearing capacity principles from /8/. The vertical force, Q v, required to penetrate the pipe to the embedment, z, assuming linear increase in shear strength with depth, may be calculated as: Q v = Q v0 (1+d ca ) + γ A bm (4.1) where: Qv 0 is the bearing capacity (not including depth effects or soil buoyancy) γ is the soil submerged unit weight d ca is a depth correction factor A bm is the penetrated cross-sectional area of the pipe, see Equation (4.7) Q v0 = F (N c s u,0 + ρ B/4) B (4.2) where: F is a function of pipe roughness and of ρb s u,0 and can be taken from Figure 4-4. It should be noted that the roughness in this respect is related to the degree of mobilized shear stress at the pipe-soil interface. The remoulding process of the soil underneath the pipe during installation will likely be close to a smooth foundation N c is a bearing capacity factor for clay. For pipes considered as smooth, the bearing capacity factor, N c, may be taken as 5.14 for small penetrations, but could reduce to 4 when the pipe embedment is equal to z=d/2 due to the circular arc shaped foundation base. More detailed discussions can be found in /7/ s u,0 is the undrained shear strength at the reference z-level for depth effects, see Figure 4-5 ρ is the gradient of the undrained shear strength with depth B is the pipe-soil contact width. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 27

28 The relation between contact width, B, and the embedment, z, is: for z < D/2 (4.3) B=D for z D/2 where: D z is the pipe outside diameter including coating is the pipe embedment. Figure 4-4 Correction factor according to /8/ It is assumed that there is no depth effect provided that the penetrated pipe is inside the active Rankine zone. The reference z-level for depth effects, z su,0, is taken as the seabed for shallow penetrations. For deeper penetrations, z su,0 is taken as the depth where a tangent to the pipe at 45 intersects the vertical line through the edge of the soil/pipe contact, see Figure 4-5. The reference z-level for depth effects may be expressed as follows: Z su,0 = 0 (4.4) Recommended practice DNVGL-RP-F114. Edition May 2017 Page 28

29 The shear strength s u,0 at the reference z-level for depth effects is taken as: s u,0 = s u,z=0 + ρ z su,0 (4.5) Figure 4-5 Reference level for depth effects in undrained conditions The depth correction factor, d ca, is taken in accordance with DNVGL-RP-C212 as d ca = 0.3 s u,1 s u,2 arctan(z su,0 B) (4.6) where s u,1 = (s u,z=0 + s u,0 ) 2 is the average shear strength above the reference foundation level and s u,2 = Q v0 (B N c ) is the average shear strength below the reference foundation level. The penetrated cross-sectional area of the pipe, A bm, is taken as: A bm = arcsin(b D) D 2 4 B D 4 cos(arcsin(b D)) for z < D/2 A bm = π D D (z D 2) for z D/2 (4.7) Model 2 An alternative model is described in /9/, /10/ and /11/. The vertical force required to penetrate the pipe to the embedment, z, is: (4.8) Recommended practice DNVGL-RP-F114. Edition May 2017 Page 29

30 where: s u is the soil undrained shear strength at pipe invert (and therefore a function of z) D is the pipe outside diameter including coating z is the pipe embedment γ is the soil submerged unit weight A bm is the pipe submerged cross-sectional area (function of z, Equation (4.7)). The first term of Equation (4.8) represents the soil resistance to pipe penetration. The second term accounts for the soil buoyancy which is enhanced by soil heave by a factor of 1.5 (based on a best fit value to numerical analysis, /10/). More information on this parameter can be found in /11/. At very high embedment ratios (z D>0.5) Equation (4.8) may underestimate the penetration resistance and the penetration estimate should be used with caution. Alternative bearing capacity factors may be found in /31/ Embedment assessment in drained conditions There are different calculation models available for calculating vertical embedment in drained conditions. The embedment depth will be the depth where the pipe contact force is in equilibrium with the bearing capacity of the seabed soil. The static penetration resistance may be estimated using the following approach: The relation between the contact width, B, and the embedment, z, can be taken from Equation (4.3). The vertical force, Q v, required to penetrate the pipe to the embedment, z, is calculated as: Q v = 0.5 γ' N γ B 2 + z 0 γ' N q d q B (4.9) where: γ is the submerged unit weight of soil N q is a bearing capacity factor, see Figure 4-6 N γ is a bearing capacity factor /12/, /13/, see Figure 4-6 N γ = 1.5 (N q 1) tanφ N γ = 2 (N q + 1) tanφ φ is the friction angle of the soil B is the pipe-soil contact width d q is a factor accounting for depth effects z is the embedment at pipe invert z 0 is the reference z-level for depth effects. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 30

31 Figure 4-6 Bearing capacity factors for drained conditions It is assumed that there is no embedment effect as long as the penetrated pipe is inside the active Rankine zone, see Figure 4-7. This reference depth, z 0, will be dependent on the friction angle of the soil, and can be found by: z 0 = 0 for z < D 2 [1 cos (π 4 + φ 2)] for z > D 2 [1 cos (π 4 + φ 2)] (4.10) Recommended practice DNVGL-RP-F114. Edition May 2017 Page 31

32 Figure 4-7 Reference level for depth effects in drained conditions The depth factor can be taken as: (4.11) It should be noted that the depth effect is usually not important in sands, because of the large penetration needed before it is relevant. The expression for Q v is based on bearing capacity formulae for ideal 2D foundations. Note that if this model is used to predict the expected penetration z for a given contact force, Q v, it may lead to underestimation of the true penetration due to effects of the pipe laying process, see [4.2.5] Effect of laying process on embedment General Observations show that the as-laid pipeline embedment is typically much greater than would be expected from the static weight alone, due to motions of the pipeline during laying and the interaction between the pipe and the soil in the touch down zone /14/. Vertical and horizontal motions of the pipeline within the touchdown zone may have significant effects on the penetration of the pipe. The penetration of the pipeline is a result of a complex process, the outcome of which depends on intact and remoulded soil properties, the weight and stiffness characteristics of the pipeline, the pipe motions and the mechanisms of the gradual cycle by cycle additional penetration. The dynamic motions of the pipeline owing to vessel motions are dependent on the sea state during laying, the water depth and the lay tension and may thus vary significantly Static touchdown factor A reference static touchdown lay factor, k lay, can be used to account for the increased vertical pipe-soil force. In the absence of project-specific pipe lay analysis k lay may be estimated as /9/: Recommended practice DNVGL-RP-F114. Edition May 2017 Page 32

33 (4.12) where: k lay is the touchdown lay factor (= Q v /W i ) EI is the pipe bending stiffness W i is the submerged pipe weight during installation z ini is the initial pipe embedment after laying T 0 is the horizontal component of the effective lay tension in the pipe at touch-down point during installation. Typical values of k lay lie between 1 and 3 depending on the seabed stiffness, lay tension, departure angle and pipeline bending stiffness. The touchdown factor increases in stiffer soils where the touchdown reaction is concentrated over a shorter length of pipe. In softer soils, where the reaction is spread over a longer length of pipe, the touchdown factor converges towards unity (but can never be less than one). The parameters 0.6 and 0.4 used in Equation (4.12) have been derived from curve fitting to numerical analyses of the catenary response, with a linear variation of seabed resistance with pipe embedment depth. Equation (4.12) applies only for T 0 >[3 (EI) 0.5 W i ] 2 3 /9/. During conceptual design and in the absence of a pipe lay analysis, the horizontal component of lay tension, T 0, can be uncertain. It is then recommended to consider a range of possible lay tension. The embedment is found by first establishing k lay,1 =Q v /W i with penetration depth from [ ] or [ ]. k lay,2 is found by inserting k lay,1 in the right hand side of Equation (4.12): k lay,1 =Q v /W i (4.13) (4.14) The compatible embedment and touchdown factor is found when k lay,1 = k lay,2, as illustrated in Figure 4-8. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 33

34 Figure 4-8 Example of determining touchdown factor and inital embedment Dynamic effects In undrained conditions, /14/ recommends using a vertical penetration model for assessing the initial embedment combined with the remoulded strength to account for dynamic laying effects. In lack of other data this may be a reasonable approach. However, it should be emphasized that this is a simplification, as the pipe will not be pushed vertically into fully remoulded soil, but gradually digs itself down due to a combination of vertical and horizontal pipe movements, see Figure 4-2. In this process the soil is gradually remoulded but engaging new intact soil as the pipe penetrates. The simplification has been proved to provide a reasonable fit for a limited number of installed pipelines, /14/ however there is a need to extend the database with more examples. As-laid surveys will therefore be very valuable in order to increase the confidence in this simplification. For pipe-laying in calm sea states, the use of remoulded strength could overestimate the embedment. As such, when a range in embedment is established based on the above approach, the low estimate embedment should be compared with the static penetration using intact undrained shear strength to represent laying in calm conditions. Note that the approach using the remoulded shear strength is only valid for embedment prediction during installation. When evaluating additional penetration during the hydrotest, a regain of soil strength needs to be considered, and could lead to penetration into intact soil. In drained conditions, the observed embedment of pipelines is often higher than predicted using the submerged pipe weight in a vertical penetration model. The combination of lateral and vertical pipe motions occurring in the touchdown zone during laying may explain the differences. The embedment is therefore strongly related to the sea state during laying. The embedment predictions in drained conditions should be evaluated carefully. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 34

35 4.3 Axial pipe-soil interaction Description of axial response A typical axial pipe-soil response is illustrated schematically in Figure 4-9. The response is described by an equivalent friction factor, F A /V. The soil behaviour during axial pipe movements may be drained or undrained, depending on the rate and duration of pipe movement, the drainage characteristics of the soil and the pipe surface coating. During an axial pipe movement in undrained conditions, particularly in the first cycle, an initial peak is often observed, followed by decay to a steady residual value. During axial pipe movement in drained conditions, the response is generally ductile with no peak. In analyses of global buckling, a breakout peak generally has little influence on long-term cyclic walking. However, it may affect the axial force profile along the pipeline, end expansions during start-up and the possibility of rogue buckle formation, depending on the brittleness of the response and the magnitude of the pipeline displacements involved. The difference between the drained and undrained resistance is due to the generation of excess pore pressure around the pipe in undrained conditions, and therefore is dependent on the soil state, the tendency for contraction or dilation, and the rate of drainage relative to movement. The soil state may change over many cycles of movement, as the soil surrounding the pipe is repeatedly failed and consolidated. This causes a change in the undrained resistance towards the drained value. These mechanisms are discussed in detail in /15/ and /16/. The axial PSI is usually idealized in structural modelling with an elastic-plastic model that consists of two parameters: the limiting (or residual) axial resistance, F A (or equivalent friction, F A /V), and a mobilization distance, x mob. An initial breakout peak can be incorporated using a piecewise linear axial PSI response. Figure 4-9 Illustration of axial pipe-soil interaction response Recommended practice DNVGL-RP-F114. Edition May 2017 Page 35

36 4.3.2 Framework for axial pipe-soil interaction General The model for axial pipe-soil interaction (PSI) is based on the following key concepts, which are summarized in Figure Further background information and research is described in /3/ and /17/. This research has mainly focussed on the residual resistance. However, the same factors are believed to affect the breakout resistance and will therefore fit into the same framework. On fine-grained silty or clayey soils the axial response may be undrained, drained or in the intermediate transitional zone. In this case, design and assessment should as a minimum be based on separate assessments of drained and undrained axial resistance, with a range bounding both cases being adopted. On coarse-grained soils without silt or clay, it is generally only necessary to consider the drained resistance. Figure 4-10 Conceptual model for axial residual resistance Assessment of drained and undrained resistance Assessments of drained and undrained axial resistance should consider the following: The undrained resistance consists of a peak (breakout) and a residual value. A peak may be important in buckling design, and the effect of both including and neglecting a peak should be checked. The drained resistance is usually not significantly affected by large soil displacements/strains. The axial resistance is affected by the pipe-soil interface roughness and the effective stress level (the undrained strength is also affected by any overconsolidation of the soil). The axial resistance may be enhanced by a wedging effect, which causes the integrated normal contact force between the pipe and the soil along the contact area to exceed the pipe weight by a wedging factor denoted by ζ. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 36

37 Assessment of transitional drainage and consolidation behaviour Characteristic values of axial resistance may be refined by considering the level of drainage and consolidation at the pipe-soil interface. Drainage and consolidation cause the undrained resistance to converge towards the drained resistance. Specific guidance on how to quantify axial friction in the drained-undrained transition are not provided in this recommended practice, as this is an area of ongoing research without clear conclusions. When investigated, the undrained-drained transitional axial friction should be bounded by the undrained and drained axial resistance and should be based on project specific pipe-soil interface testing program and followed up during operation through surveys. When it is not possible to define whether the pipe will respond fully drained or fully undrained, the undraineddrained transitional axial friction may be implemented in the design, alternatively both drained and undrained response should be considered Axial breakout resistance Undrained resistance The undrained shear strength underneath the pipe is dependent on the load history the soil has been subjected to. During pipe penetration, the soil underneath the pipe will be at failure until the vertical bearing capacity is high enough to support the contact force during installation. As such, the soil strength at the pipe-soil interface will be degraded towards the remoulded value. Subsequently, the soil will reconsolidate to a higher strength dependent on the soil-pipe contact stresses. It can be assumed that the soil will be reconsolidated in accordance with the normally consolidated strength ratio (s u σ v ) NC. The pipe is usually pressure tested with water prior to operation which can lead to further penetration. If the water-filled period is long enough, the soil at the pipe-soil interface will be consolidated to a higher stress level. This can be treated as an overconsolidation compared to the operational case and the effect on the shear strength can be taken in accordance with the SHANSEP methodology /18/. The lay-induced pipe-soil normal stress should not be considered in the assessment as it is only applied for a short period, preventing full consolidation. In absence of specialized interface testing, see [3.3.2], a method to estimate the breakout axial resistance is given in Equation (4.15), written as an equivalent friction factor. Equation (4.15) contains the factors that are considered to affect the pipe-soil axial resistance in undrained conditions. (4.15) where: α is the pipe-soil adhesion or roughness factor, representing the reduction in soil-interface strength compared to soil-soil strength (S u /σ v ) NC is a ratio for the normally consolidated shear resistance versus the consolidation vertical stress. Note that the ratio is stress dependent /19/, as illustrated in Figure 4-11, and at low stress levels the factor is significantly higher than usually reported in literature for traditional geotechnical design situations. Typical range is from 0.25 to 0.5 for non-carbonate soils. V is the static vertical pipe-soil force for the condition considered, e.g. operation γ pre is the consolidation preloading effect taken as the ratio between the preloading (e.g. water filled condition) and the static pipe-soil force, V, for the condition considered (e.g. operation). This represents the overconsolidation ratio, OCR, of the soil underneath the pipe. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 37

38 m ζ γ rate is a preloading factor that accounts for the long term effect of overconsolidation, being a number less than 1.0. Typical range is from 0.65 to 0.9. is the wedging factor, taken as z>d/2, the wedging factor is constant., where β is the angle defined in Figure For is a rate factor to account for the speed of loading to undrained failure (to be taken as 1.0 for a reference speed of 2 hours to failure, may be increased by 10-15% per log cycle of the rate of loading). Figure 4-11 Stress dependency on shear strength ratio, typical range Recommended practice DNVGL-RP-F114. Edition May 2017 Page 38

39 Figure 4-12 Definition of angle used to determine the wedging factor Note that recommendations for the normalized shear resistance including preloading effects, effect of pipe roughness and the dependency of the effective stress level are not generally available in the literature. Thus, specific soil tests, e.g. special interface tests described in [3.3.2], are recommended. By doing such tests, the stress dependency and preloading effect can be captured. If such tests are performed it is recommended to specify testing procedures representative for the actual conditions the pipe will experience (e.g. consolidate for the stress conditions under the flooded weight and unloading to the operational stress condition prior to shearing). In lack of such tests, each parameter should be assessed based on engineering judgement to establish a conservative range. Such judgement could favourably make use of available tests for similar conditions related to type of clay, preloading effects and roughness of pipe surface. There is a need in the industry to establish a database containing such test results which can be used to assess the above-mentioned factors in routine design. In soft soils, the soil is disturbed and remoulded by the laying process and then reconsolidated by the pipe weight. Both processes alter the soil strength from the in-situ condition and it is not recommended to relate the long-term axial resistance to the intact undrained shear strength simply by using an adhesion factor, α s u,intact. In stronger soils, experience shows that the pipe does not become fully bonded to the seabed and therefore the in-situ soil strength (even after adjustment for interface roughness) cannot be mobilized at the pipesoil interface. The apparent lack of bonding may also be related to the uncertainties in defining the shear strength in the upper soil, and that the real strength underneath the pipe is not captured in the soil investigation. During laying, the soil underneath the pipe will be highly disturbed and any consolidation effects cannot be relied upon. The axial resistance will thus be lower than the subsequent consolidated axial resistance due to excess pore pressure in the soil Drained resistance The drained breakout resistance is governed by the submerged pipe weight, the drained friction angle of the soil and the interface properties, and can be expressed by a friction factor: Recommended practice DNVGL-RP-F114. Edition May 2017 Page 39

40 (4.16) where: δ peak is the peak interface friction angle φ peak is the peak friction angle of the soil r pipe-soil is the pipe-soil roughness factor ζ is the wedging factor, as given in [ ]. Specialized testing (as described in [3.3.2]) for directly assessing the interface friction as a function of stress-level is recommended. For a given soil, the axial friction for a smooth pipe coating may be as low as 30% of the value for a rough coating /20/. For this reason, it is important to use a representative interface material when performing laboratory testing. In case such tests are not available, conservative assumptions should be made for each of the parameters in Equation (4.16) Axial residual resistance Undrained resistance The undrained residual friction is believed to follow the same trends as the breakout resistance with respect to stress-dependency. Specialized testing, as described in [3.3.2], for directly assessing the interface friction as a function of stress-level is recommended. The sensitivity of the clay is the measure of how much the intact shear strength degrades when remoulded. As such, it is believed that the soil sensitivity, S t, is a parameter that affects the residual friction, however not necessarily being a direct correlation. The residual friction can then be expressed as: μ A,res,u = ε res μ A,brk,u = ε res α (s u σ' v ) NC γ pre m ζ γrate (4.17) where: μ A,brk,u is the equivalent peak(breakout) axial friction (from Equation (4.15)) ε res is the residual reduction factor, believed to be in the range 1/St <ε res <1. In order to optimize the outcome of the soil investigations/laboratory tests it is recommended to involve relevant stakeholders in an early phase of a project. This will help defining a scope that accounts for project-specific data, such as type of coating (including surface roughness) and any foreseen level of preconsolidation from the water-filled condition. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 40

41 Guidance note: As part of the SAFEBUCK JIP, several small scale tests were performed within or made available to the JIP focusing on the residual axial friction. Based on these tests, the preloading factor m was found to range between 0.35 and 0.6, which is significantly lower than expected. Experience from SHANSEP tests shows a range from 0.65 to 0.9. The low m values calibrated from the tests are most likely due to an overestimation of the actual OCR. The reported OCR was related to the consolidation pressure applied for preparation of the clay in the test bin before the pipe was pushed in-place. A comment to the obtained low m factors was that they could be explained by the pipe penetration process, i.e. the pipe is erasing some of the in-situ OCR when disturbing the soil during penetration. The SAFEBUCK tests show the importance of trying to replicate the real stress history of the pipe when carrying out model tests/interface tests. The OCR underneath the pipe usually originates from the pipe weight during the water filled period and the tests should replicate this effect as accurate as possible. ---e-n-d---o-f---g-u-i-d-a-n-c-e---n-o-t-e Drained resistance The drained residual resistance is usually similar to the breakout resistance, but could potentially be lower due to a lower residual interface friction angle. Equation (4.16) can be used, replacing the peak friction angle with the residual friction angle, φ res : μ A,res,d = tanδ res ζ = r pipe-soil tanφ res ζ (4.18) Specialized testing, as described in [3.3.2], for directly assessing the interface friction as a function of stresslevel is recommended. Published studies (/1/, /2/, /3/ /17/ and /20/and experience indicate that the drained residual interface friction lies in the range of 0.3<μ res <1.0 (or 15 <δ res <45 ) for non-carbonate soils. Results for carbonate clays and silts extend to a higher upper limit of μ res ~1.4 (δ res ~55 ), see also [7.5]. Other pipeline and soil combinations may lie outside this range Axial mobilization displacements In structural modelling of a pipeline, the axial response is usually modelled using a bi-linear elasticperfectly plastic model, which requires specification of the mobilization displacement, x mob. The mobilization displacement defines the initial stiffness and the unload-reload stiffness. If a breakout peak is modelled, x mob is replaced by two mobilization distances, x brk and x res. The actual response is generally non-linear, with a reduction in tangent stiffness as the axial resistance is mobilized. In assessments of pipe walking, a low value of x mob creates a higher rate of axial walking. To be conservative, a bi-linear fit to the non-linear response should be a tangent fit to the initial part of the axial force-displacement response, which represents the elastic recoverable part, see Figure 4-9. In other situations, such as lateral buckling and feed-in of pipe into the buckle, a higher value of x mob can be more onerous, and the bi-linear fit should be a secant fit to the displacement when F A,brk or F A,res is fully mobilized. In the absence of project-specific assessments, the mobilization displacements can be selected using the advice in Table 4-2. The recommendations may be inaccurate for pipeline properties and soil conditions that are outside of the model test database, see [1.5.1]. For drained conditions, the tabulated high estimates are probably too high. In absence of detailed investigations all values of axial breakout and residual mobilization distance are possible and should be considered in design. Table 4-2 Axial mobilization displacement Axial model Uncertainty case Parameter Typical values Bi-linear (i.e. with no breakout peak) Low estimate, LE 1 Best estimate, BE 1 High estimate, HE 2 x mob Minimum of 1.25 mm and D Minimum of 5 mm and 0.01 D Maximum of 250 mm and 0.5 D Recommended practice DNVGL-RP-F114. Edition May 2017 Page 41

42 Axial model Uncertainty case Parameter Typical values Low estimate, LE 1 Minimum of 1.25 mm and D Best estimate, BE 1 x brk Minimum of 5 mm and 0.01 D Tri-linear (i.e. with breakout peak) Notes: High estimate, HE 2 Low estimate, LE 1 Best estimate, BE 1 High estimate, HE 2 x res Maximum of 50 mm and 0.1 D Minimum of 7.5 mm and D Minimum of 30 mm and 0.06 D Maximum of 250 mm and 0.5 D 1) Represents a tangent fit to the initial part of the axial force-displacement response 2) Represents a secant fit to the displacement when F A,brk or F A,res is fully mobilized 4.4 Lateral pipe-soil interaction Description of lateral response There are two stages of lateral pipe-soil response behaviour: First load displacement (monotonic), characterized by a breakout resistance and a steady residual resistance. Cyclic displacement characterized by the growth of soil berms at the limits of the pipe displacement range, with the pipe descending into a shallow trench. This load scenario may be valid for repeated heating and cooling corresponding to breaks in production at the location of the pipe where lateral buckle displacements occur. The stages above are integrated in the schematic model presented in Figure The red line represents the first load response and the blue line the cyclic response. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 42

43 Figure 4-13 Illustration of lateral pipe-soil response This model is appropriate to light pipes that tend to rise after breaking out from the as-laid position. Heavier pipelines (heavy pipes) can move downward in the soil after the initial breakout resistance is mobilized. This downward movement, coupled with the growth of the soil berm ahead of the pipe, leads to a continuous increase in the lateral resistance, rather than a steady value. The model presented in this recommended practice is not applicable to heavy pipes beyond mobilization of the breakout resistance. Limited data is available on heavy pipe behaviour and no general calculation method to assess their behaviour after breakout is included in this recommended practice. If heavy pipe behaviour is expected, alternative solutions should be considered, such as initiation of buckles on gravel carpets or sleepers. If the pipeline tends to penetrate deeper within a lateral buckle during the first buckle initiation or in subsequent cycles of cooling and heating, such penetration will be counteracted by transfer of vertical reaction forces towards the shoulders of the buckle. Such 3D effects are not accounted for in typical 2D testing conditions. Further collection of relevant survey data would improve the background for developing recommended practice. A particular concern of a high berm resistance relates to a situation where the production pressure or temperature is increased at a late stage of the operation of the pipeline. Regular surveys should focus on monitoring berm developments. In clayey soil, heavy pipe behaviour is observed when the pipe-soil vertical force is more than approximately half of the ultimate vertical bearing capacity, as calculated at a penetration at half a diameter (V>2 s u D). This is however not a strict limit but special attention should be given when approaching this limit. Heavy pipe behaviour is not observed in sandy soil. In some cases, the vertical pipe-soil force may differ from the pipeline weight (e.g. the touchdown area on either side of free spans, distributed buoyancy sections and natural uneven seabed). In these cases, the recommendations provided in this recommended practice can be followed by adjusting the vertical pipe-soil force V. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 43

44 This section discusses the parameters: Breakout resistance: calculation models to assess the breakout resistance are given in [4.4.2]. Residual resistance after breakout: calculation models to assess the residual resistance are given in [4.4.3]. Mobilization displacements: mobilization displacements are discussed in [4.4.4]. In general, due to the uncertainties in the calculation models, more than one model should be evaluated, see Sec.6. The proposed methods in this section can therefore be considered as examples. Other methods may also be relevant, see [1.4] Lateral breakout resistance General The boundary conditions around the pipe are very important when evaluating the lateral breakout resistance. Depending on the pipe motions during laying, different boundary conditions may apply as illustrated in Figure Different models are available for calculation of the breakout resistance. It is appropriate to use a theoretical approach or a semi-theoretical approach with empirical adjustment factors, if the boundary conditions of the model are known, e.g. shear strength profile, penetration, contact area, trench geometry, pipe to soil surface roughness. It is recommended to use a model that is capable of varying the above mentioned effects and assess how they influence on the break-out resistance. As the geometrical boundary conditions (e.g. trench geometry) are not known prior to the pipe installation, this represents a challenge in the design phase so the range in the breakout resistance should include different scenarios. Two models are proposed in this section. It is recommended to evaluate various models to assess model uncertainties, see Sec.6. The use of finite element analysis where all parameters affecting the resistance can be defined is recommended for sensitivity studies, see [2.4]. The challenge however remains to identify the range of assumptions needed to capture the likely conditions the pipeline could experience. The boundary conditions (e.g. pipe penetration, trench development etc.) may vary during the operational life, and regular surveys are necessary to ensure that the boundary conditions around the pipe are covered by the original design assumptions. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 44

45 Pipe in perfect contact with the seabed installed in calm sea state Pipe inside a trench installed in rough sea state Figure 4-14 Effect of boundary conditions and laying effects on the lateral resistance Undrained resistance Model 1 The undrained lateral breakout resistance, F L,brk,u, can be taken as the sum of: a shear resistance on a horizontal surface underneath the pipe, F L,brk,u,fric, similar to the axial friction (from Equation (4.15)) only without the wedging factor a remaining resistance, F L,brk,u,remain, due to mobilizing the soil in front of the pipe and at the rear of the pipe (suction). Suction may be considered if the pipe is installed by vertical penetration, assuring good contact with the soil at both sides of the pipe, or when a high lateral resistance is unfavourable. F L,brk,u = F L,brk,u,fric + F L,brk,u,remain (4.19) F L,brk,u,fric = α (s u σ' v ) NC γ pre m γrate (4.20) (allowing for suction at the rear of the pipe) (4.21) (not allowing for suction at the rear of the pipe) Recommended practice DNVGL-RP-F114. Edition May 2017 Page 45

46 where: κ a, κ p are soil pressure resistance coefficients, in the range from 2 to 2.5 is the average undrained shear strength within the active failure zone, typically at a depth equal to z/2 is the average undrained shear strength within the passive failure zone, typically at a depth equal to z/2. The possibility to use anisotropic shear strength being different in the passive and active zone is included. However, it may often be more realistic to use an equivalent average isotropic strength. If there is no suction at the rear, the resistance due to the weight of the soil is added to the passive resistance. When the active suction is included the soil weight terms from the two zones cancel each other out. A simplification in the model is that the passive resistance is taken in accordance with classical earth pressure theory against a vertical surface in front of the pipe. As such, it does not fully capture the real failure surface being affected by the shape of the pipe. Model 2 An alternative undrained lateral breakout resistance model is as follows: (4.22) where: F L,brk,u s u is the lateral breakout resistance in undrained soil conditions is the soil undrained shear strength at the pipe invert depth D is the pipe outer diameter including coating z is the pipe embedment V is the static vertical pipe-soil force for the condition considered, e.g. operation γ is the soil submerged unit weight at the pipe invert depth. In this equation, the first term reflects the passive resistance related to the shear strength of the soil berm pushed in front of the pipe. The second term is a frictional component and the third term captures the passive self-weight resistance from the soil ahead of the pipe. Equation (4.22) was fitted to a database which comprised of 67 pipe tests on mainly very soft West African clays (s u between 0.4 kpa and 9 kpa at mudline). The pipes were generally embedded by monotonic penetration and then broken out laterally at a relatively rapid rate. The nominal pipe bearing pressure, V/D, was generally in the range between 1 kpa and 7 kpa. Results from 5 tests showing heavy pipe behaviour was included in the calibration of this equation. The model uncertainty factor associated with this model is found to be 1.5, meaning that the LE and HE resistance respectively is found by dividing and multiplying Equation (4.22) by 1.5. Note however, that this model uncertainty is related to scatter in test results. Whether these tests have captured all likely scenarios the pipeline will experience in the field is still uncertain. Uncertainties in soil parameters along the pipeline route should also be considered. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 46

47 Discussion A disadvantage of model 2 is that the stress-dependency of the frictional part is fixed and does not account for pre-consolidation effects. Model 1 encourages the engineer to use a consistent frictional term both in the axial and lateral direction. Also, the passive resistance relating to the shear strength of the pipe is for model 2 using the shear strength at the invert of the pipe thus not reflecting the shear strength variation between the seabed and the pipe invert, which sometimes may be of significance. The two models are compared in Figure 4-15 for one specific case, however leaving out the frictional term for simplicity. In model 1 the effect of allowing for suction at the rear of the pipe can be investigated and this effect is included in the figure. Allowing for suction to develop at the rear of the pipe could be relevant when a high lateral resistance is unfavourable. The comparison is included to illustrate the effect of suction at the rear of the pipe and may to some extent explain the associated model uncertainty with model 2. Figure 4-15 Comparison of model 1 and model 2 for a typical normally consolidated soil Drained resistance Different models exist for evaluating lateral pipe resistance in siliceous sand. Most methods are empirical formulations fitted to a limited number of tests. However, the testing conditions and the validity range for a given empirical based model are often not well described which makes it difficult to judge whether all relevant scenarios are captured by the test program. The empirical equations are based on the breakout resistance, and have therefore taken the effect of the berm into account. Hydrodynamic effects such as sediment transport and scour could easily with time wash away or build up a berm, leading to changed boundary conditions for the pipe. In order to take these changes into account, the methods have to be based on geotechnical theories rather than empirical equations. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 47

48 Model 1 The drained lateral breakout resistance, F L,brk,d, can be taken as the sum of: friction between pipe and soil underneath the pipe passive earth pressure when pushing a wedge of soil laterally the effect of the additional weight of the berm (due to soil heave during penetration). F L,brk,d = F L,brk,d,fric + F L,brk,d,passive (4.23) The passive earth pressure can be taken as: (4.24) The passive earth pressure coefficient, K p, is dependent on the roughness defined as the shear mobilization on the vertical plane in front of the passive wedge, and can be taken according to classical earth pressure theory, see Figure Note that this roughness parameter, r= τ v /(σ h ' tanφ) is different from the pipesoil interface roughness or adhesion factor. It is not a material parameter, but rather a parameter that reflects the direction of the force transferred to the soil in front of the pipe (or in general terms in front of the foundation) and is as such related to the direction of the load acting on the pipe, see Figure The roughness is an uncertain value and requires special attention. For more details regarding how the roughness parameter is defined and how it influences the passive earth pressure coefficient, see /21/. A simplification in the model is that the passive resistance is taken in accordance with classical earth pressure theory against a vertical surface in front of the pipe. As such, it does not fully capture the real failure surface being affected by the shape of the pipe. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 48

49 Figure 4-16 Passive earth pressure coefficient according to classical earth pressure theory Figure 4-17 Definition of roughness parameter When mobilizing the passive resistance, shear stresses in the upward direction acting from the soil in front of the pipe will develop and reduce the effective normal force acting on the sliding plane underneath the pipe. The roughness on the vertical plane in front of the passive wedge is then of importance when determining Recommended practice DNVGL-RP-F114. Edition May 2017 Page 49

50 the effective normal force underneath the pipe. The upward force acting on this plane cannot be greater than the pipe-soil contact force. Hence, both the roughness and the passive resistance are in principle limited by the pipe weight, and will reduce the uncertainty related to the roughness parameter. For light pipes, the roughness will be close to zero. The frictional term can be taken as: F L,brk,d,fric = r pipe-soil tanφ (V r tanφ F L,brk,d,passive ) (4.25) where the term (V r tanφ F L,brk,d,passive ) cannot be negative. If that is the case, the assumed roughness is too high and shall be reduced. Note that r pipe-soil is a roughness factor between the pipe and soil and is different from r. Figure 4-18 Illustration of lateral failure mechanism in drained conditions, including effect of berm next to the pipe In addition, as illustrated in Figure 4-18, the pipe will get additional resistance from any berm that is built up next to the pipe. A berm will form from soil heave during penetration, and possibly also due to oscillations during laying. To account for the additional berm resistance, the displaced area can be idealized as: (4.26) where A bm is the penetrated cross-sectional area of the pipe, see Equation (4.7), and the second term is the displaced soil due to a horizontal amplitude a. The berm area can then for simplicity be evenly distributed over the extent of the failure surface, x failure, leading to a modified height, z mod, which is replacing z in Equation (4.24). Note that the failure surface in accordance with classical earth pressure theory is also related to the roughness parameter, as indicated in Figure How x failure varies with friction angle and roughness is given in Figure Recommended practice DNVGL-RP-F114. Edition May 2017 Page 50

51 Figure 4-19 Failure surfaces for different roughness parameters (example for φ=40 ) Figure 4-20 Normalized failure surface, (x failure /z), as function of friction angle and roughness Model 2 A relationship to assess drained lateral breakout resistance based on calibration to full scale model tests performed on siliceous sands was proposed in /22/. The breakout resistance is divided into a frictional and a passive component that provides increasing resistance with penetration. F L,brk,d = 0.6 V + F p (4.27) where the passive resistance, F p, is defined as: Recommended practice DNVGL-RP-F114. Edition May 2017 Page 51

52 (4.28) and F L,brk,d is the lateral force at breakout in drained soil conditions V is the vertical pipe-soil force γ is the soil submerged unit weight D is the pipe outer diameter including coating z is the pipe embedment. It should be noted that the tests were performed by first oscillating the pipe laterally until a distance of half the diameter was obtained, and by that creating a berm in front of the pipe. Then it was further pushed through the berm and the breakout resistance was recorded. The published method only relates to achieved penetrations without any evaluation of the effect of how the penetration is achieved. It should also be noted that the tests were carried out using steel pipes. Hence, for other types of coating material, the specified friction factor of 0.6 may not be relevant. Discussion In order to illustrate the effect of the oscillations prior to breakout in the tests reported in /22/, the two models are compared in Figure The input parameters are given in Table 4-3. Table 4-3 Input parameters used for comparison of models Pipe outer diameter, D Submerged pipe weight, V 0.5 m 1 kn/m Submerged unit weight of soil, γ 9 kn/m 3 Friction angle of sand, φ 40 Pipe-soil interface roughness factor, r pipe-soil 0.6 As the penetration of the pipe gets deeper, the passive resistance becomes more important. In general, the mobilized roughness on the vertical passive surface has large impact on passive earth pressure, as seen from Figure Similarly, the roughness is very important for large penetrations for the geotechnical model. However, when limiting the possible vertical shear onto the vertical passive surface to the submerged pipe weight, possible roughness values are limited and accordingly the range in passive resistance reduces as shown in the Figure 4-21 below. No effect of a berm in front of the pipe has been accounted for in model 1. It is seen that the model 2 predicts significantly higher resistance. However, by including the effect of a berm due to oscillations prior to breakout (a/d = 0.5), a good fit can be obtained, as shown in Figure The lateral motions imposed to the pipe prior to breakout in /22/ has a large impact. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 52

53 Figure 4-21 Comparison of model 1 and model 2, using compatible roughness Figure 4-22 Comparison of model 1 and model 2, using compatible roughness and accounting for test conditions used in /22/ Recommended practice DNVGL-RP-F114. Edition May 2017 Page 53

54 It is important to be aware of the uncertainty related to the seabed configuration around the pipe in sandy sediments and how this affects the resistance. Where the pipe after laying lays in a bowl shaped trench rather than with perfect soil contact, see [ ], the pipe may have to displace and rise within the bowl before breaking through, in the same way as experienced for the cyclic tests with large amplitude oscillations in /22/. For such conditions the embedment as well as the effect of the berm would have to be adjusted when calculating the breakout resistance. Hydrodynamic effects like scour could change the seabed geometry around pipelines, and these effects will vary both along the pipeline and in time. If the berms created during laying are removed due to scour, the lateral breakout resistance would be significantly reduced. In contrast, in areas of limited water depth, where there is a significant hydrodynamic effect with varying erosion and deposition, the resulting effect with time could be a significant self-burial. In this case the breakout lateral resistance could be significantly higher. For a deterministic design and assessment situation this uncertainty has to be captured by analysing different seabed configurations. Model 1 is as such considered more appropriate to use when evaluating different seabed configurations, but it is based on certain assumptions and simplifications Lateral residual resistance Undrained resistance The undrained lateral residual resistance may be calculated as: (4.29) where: F L,res,u V D z is the lateral residual force after breakout in undrained soil conditions is the vertical pipe-soil force is the pipe outer diameter including coating is the pipe embedment prior to lateral movement. The numerical parameters in this equation were obtained by fitting the formulation to small scale model test data, with some constraints to reflect theoretical considerations. The equation may therefore show some bias to the underlying database conditions. The model uncertainty factor associated with this model is found to be 1.5, meaning that the LE and HE resistance respectively is found by dividing and multiplying Equation (4.29) by 1.5. Note however, that this model uncertainty is related to the scatter in the test results, and whether these tests have captured all likely scenarios the pipeline will experience in the field is still uncertain. Uncertainties in soil parameters along the pipeline route needs also to be considered. Equation (4.29) was fitted to a database which comprised of 67 pipe tests on mainly very soft West African clays (s u between 0.4 kpa and 9 kpa at mudline). The pipes were generally embedded by monotonic penetration and then broken out laterally at a relatively rapid rate. The nominal pipe bearing pressure, V/D, was generally in the range between 1 kpa and 7 kpa. During lateral movement, a light pipe rises from the aslaid position. However, the model was calibrated based on the initial embedment at the start of the lateral movement. As the model does not include any physical parameters as soil strength or weight of the pipeline, it is difficult to assess its applicability outside the conditions of the test database. To increase confidence in the method, there is a need to test the model further against a larger database, possibly also modifying the model to account for other parameters. The method is applicable for pipes where V < 2 s u D, thus excluding heavy pipes on soft soil. Alternative models to calculate the lateral residual resistance may be found in /52/ and /53/. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 54

55 It should be noted that the tests were performed by moving the pipe sideways but allowing the pipe to move vertically. For global buckling, buckles are not necessarily initiated by the pipe first moving laterally. For some buckles, the pipes may rather lift upwards and then move out of the original trench before they transit into a horizontal buckle. In such cases, initial penetration may not be relevant for determining the residual resistance and Equation (4.29) should be used with care. This would also be the case for the resistance at very large displacements Drained resistance Model 1 In case of light pipe behaviour, the residual resistance may be considered using the same formulation as for the breakout resistance, see [ ], but using a shallower penetration depth. For a conservative low estimate, only the frictional component can be used assuming zero embedment. Model 2 The drained lateral residual resistance may be calculated as: (4.30) where: F L,red,d is the lateral residual force in drained soil conditions V is the vertical pipe-soil force γ is the soil submerged unit weight A p is the cross sectional area of the pipe (π 2 D /4) D D ref is the pipe outer diameter including coating is reference diameter taken as 20 inches (508 mm). The formula was obtained by statistical analysis of available laboratory test data for sand of various densities and pipes from 25 mm to 225 mm diameter, and back-calculated resistances from fitting lateral buckles in existing pipelines. Note that Equation (4.30) gives the mean or best estimate value. Low and high estimates can be taken as: (4.31) (4.32) Recommended practice DNVGL-RP-F114. Edition May 2017 Page 55

56 4.4.4 Lateral mobilization displacements Breakout mobilization The mobilization displacement to reach the breakout resistance, y brk, is difficult to predict. Experience shows that y brk primarily depends on embedment, z, apart from for very low values of z/d. Low, best and high estimates of y brk are given in Table 4-4. These estimates are selected to fit a database for clays, see [1.5.1]. In absence of detailed investigations all values of breakout and residual mobilization distance are possible and should be considered in design. The embedment process (static or dynamic) appears to affect y brk. Model tests in which the pipe is monotonically pushed to the initial embedment show much smaller y brk values (and consequently a stiffer lateral response) than model tests in which the dynamic embedment process is simulated. For drained conditions, there are less data available regarding mobilization distances. However, the large ranges given in Table 4-4 are believed also to be representative for drained conditions. Table 4-4 Lateral mobilization distance to breakout resistance Parameter Uncertainty case Typical values 3 Low estimate, LE 1 y brk Best estimate, BE 2 High estimate, HE 2 Notes: 1) The low estimate is a minimum value which considers the model test results from statically embedded pipe data 2) The best and high estimates consider statically and dynamically embedded model test data 3) All values represent a secant fit to the displacement when F L,brk is fully mobilized. No distinction is made between drained and undrained behaviour Residual mobilization The displacement to mobilize the residual resistance, y res, can be estimated using the values in Table 4-5, which are derived from results in /23/ and the database described in [1.5.1]. In absence of detailed investigations all values of breakout and residual mobilization distance are possible and should be considered in design. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 56

57 Table 4-5 Lateral mobilization displacement to residual resistance Parameter Uncertainty case Typical values Low estimate, LE y res Best estimate, BE High estimate, HE 4.5 Soil stiffness General The soil stiffness should be evaluated differently for static and dynamic analyses. Unless the non-linear soil response can be explicitly accounted for in the pipeline analysis and elastic springs are used to represent the soil response, the non-linearity of the soil has to be accounted for. The stiffness should in that case be a secant stiffness representing the expected load level in the pipeline analysis. The static soil response will relate to a static loading situation, e.g. maximum load. Dynamic stiffness will be characterized mainly by the unloading/re-loading situation Static soil stiffness Vertical soil stiffness The static vertical stiffness is a secant stiffness representative for penetration conditions such as during installation and during development of free spans due to erosion. The static vertical stiffness is defined as: K V,s = Q v /z (4.33) where: Q v z is the static vertical soil reaction per unit length of pipe is the vertical penetration of the pipe required to mobilize this reaction. The penetration curve can be established according to [4.2]. Examples of an equivalent secant stiffness at different load levels are illustrated in Figure Recommended practice DNVGL-RP-F114. Edition May 2017 Page 57

58 Figure 4-23 Examples of vertical static secant stiffness for different load levels The penetration curves given in [4.2] are based on a wished-in-place philosophy and an ideally plastic failure situation, which means not considering the additional distance required to mobilize the failure mechanism at any given penetration depth. As such, the stiffness may be adjusted by adding a mobilization distance: K V,s = Q v /(z+ δ f ) (4.34) where: δ f is the failure mobilization distance at any given penetration depth, which may typically be taken as 10% of the pipe-soil contact width, B Lateral soil stiffness The models proposed in [4.4] should be used to establish the lateral resistance curve. The static lateral stiffness, K L,s, should be estimated from these models. Examples of an equivalent static secant stiffness at different load levels relating to a bi-linear lateral resistance model are illustrated in Figure Recommended practice DNVGL-RP-F114. Edition May 2017 Page 58

59 Figure 4-24 Examples of lateral static secant stiffness for different load levels Dynamic soil stiffness General The soil stiffness may be evaluated from an equivalent shear modulus of the soil. The shear modulus G, defined as a secant modulus, is a decreasing function of the shear strain amplitude, γ c, in the soil. The shear modulus, G max, for sands at small strains may be calculated from the following expression /24/: (4.35) where: σ a is the atmospheric pressure, 100 kpa σ s is the mean effective stress in soil e s is the void ratio For clays, the small-strain shear modulus G max may alternatively be calculated from the undrained shear strength, s u, in the following manner, as a best estimate approximation to laboratory test data /19/: Recommended practice DNVGL-RP-F114. Edition May 2017 Page 59

60 (4.36) where I p denotes the plasticity index (in absolute numbers), and OCR is the overconsolidation ratio of the clay. The relation between the secant shear modulus G and the cyclic shear strain amplitude, γ c, is typically expressed as a curve of G/G max versus γ c, with a typical range given in Figure More details may be found in e.g. /25/. Such relations should be used with care in particular at large strains, where it is important to assure that shear stresses exceeding the shear strength are not obtained. Figure 4-25 G/G max as function of cyclic shear strain, typical range Note that other approaches may also be relevant, and the models presented above can be considered as examples. The soil behaviour during dynamic loading is complex and require specific assessments. Further guidance may be found in DNVGL-RP-C212. For free-span specific scenarios, see [7.2] Vertical soil stiffness For determination of the dynamic vertical stiffness, K V,d, the following expression may be applied: (4.37) Recommended practice DNVGL-RP-F114. Edition May 2017 Page 60

61 which is based on elastic half space theory for a rectangular foundation under assumption of a pipe length that equals 10 times the contact width between pipe and soil. The Poisson s ratio for soil, v, can be taken from 0.3 to 0.35 for sand and from 0.45 to 0.5 for clay. More details about elastic soil stiffness can be found in /26/. The main challenge is to estimate an equivalent shear modulus representative for the elastic half space Lateral soil stiffness For determination of the dynamic lateral stiffness, K L,d, the following expression may be applied: K L,d = 0.76 G (1 + ν) (4.38) which is based on elastic half space theory for a rectangular foundation under assumption of a pipe length that equals 10 times the contact width between pipe and soil. The Poisson s ratio of soil, v, can be taken from 0.3 to 0.35 for sand and from 0.45 to 0.5 for clay. More details about elastic soil stiffness can be found in /26/. The main challenge is to estimate an equivalent shear modulus representative for the elastic half space. 4.6 Soil damping The soil damping is generally dependent on the dynamic loads acting on the soil. It can be distinguished between two different types of soil damping mechanisms: radiation damping associated with propagation of elastic waves through the yield zone material damping associated with hysteresis effects taking place close to the yield zone in contact with the pipe. The radiation damping may be evaluated from available solutions for elastic soils using relevant soil modulus reflecting the soil stress (or strain) levels. The radiation damping depends highly on the frequency of the oscillations, and is more important for high frequency oscillations. Soil damping for free spanning pipelines is normally governed by soil material damping. The case-specific modal soil damping ratio, ζ soil, due to pipe-soil interaction may be determined by: (4.39) where the soil damping per unit length, c(s), may be defined on the basis of an energy balance between the maximum elastic energy stored by the soil during an oscillation cycle and the energy dissipated by a viscous damper in the same cycle. The equation may be solved from a finite element analysis of the pipe modelled with discrete soil supports. The viscous damping coefficient c i of support no. i can be calculated from: (4.40) Recommended practice DNVGL-RP-F114. Edition May 2017 Page 61

62 where: k i is the linearized spring stiffness at support no. i ζ soil,i is the damping ratio representing support no. i ω is the angular frequency of the mode considered. Knowing the non-linear hysteretic reaction of a support length the damping ratio representing the support can be calculated as: (4.41) where: E Dissipated is the energy dissipation at support no. i, as illustrated on Figure 4-26 E Elastic is the equivalent elastic energy at support no. i, as illustrated on Figure Figure 4-26 Energy dissipation at soil support, shown in the load-displacement space Because of the soil non-linearity, the equivalent spring stiffness and the damping ratio are dependent on the displacements at the support. For a case-specific determination of the modal soil damping ratio this needs to be taken into account. An iterative solution will be required to assure compatibility between: the dependency of the mode-shape on the equivalent support springs the dependency of the oscillation amplitude on the modal damping the dependency of the equivalent springs and the damping ratio of the discrete soil supports on the cyclic support displacements the dependency of the modal damping ratio on mode-shape and on support springs and damping ratio. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 62

63 As basis for the iterations non-linear relationships for spring stiffness, k i, and damping ratio, ζ i, as function of pipe penetration and cyclic displacements for the relevant soil and pipe diameter are required. Such relationships are qualitatively shown in Figure 4-27, and may be determined either experimentally or analytically. For free-span specific scenarios, see [7.2]. Figure 4-27 Non-linear characteristics of soil stiffness and damping Recommended practice DNVGL-RP-F114. Edition May 2017 Page 63

64 SECTION 5 BURIED AND COVERED PIPELINES 5.1 General This section gives recommendations for design conditions relevant for buried pipelines. Unlike exposed pipelines, buried pipelines are designed to stay in place. Thus, the focus in the pipe-soil interaction assessment should be on the behaviour prior to breakout. The condition of the soil surrounding the pipe is the most important aspect in any assessment of buried pipelines. This will to some extent be influenced by the trenching method. Further, the boundary conditions around the pipe (e.g. depth to intact soil conditions below the pipe in the trench) will add to the uncertainties related to selection of geotechnical parameters. Unexpected and poorly defined ground conditions are generally the most commonly re-occurring causes of delays in project schedule and budget challenges. In many cases marine pipeline projects, including buried pipelines, will be particularly exposed to such risks due to their reliance on satisfactory seafloor and subsurface soil properties for trouble free installation and operation. For pipelines that are buried by ploughing or jetting, the largest uncertainty is related to how the trenching method has affected the in-situ strength and stiffness parameters. The soil investigation program needs to consider both the intact soil conditions and the soil conditions following a trenching/jetting operation. The latter may require the construction of a certain length of dummy trench as part of the soil investigation program. In the calculation models presented in this section, it is assumed infinite width of the backfill material. In case of a rock berm with limited width, the contact stresses around the pipe can be different from the assumption. Finite element analyses are then recommended to assess the normal stresses acting on the pipe surface. In general, due to the uncertainties in the calculation models, more than one model should be evaluated, see Sec.6. The proposed methods in this section can therefore be considered as examples. Other methods may also be relevant, see [1.4]. In case pipelines are covered with rock, the effect of soil settlements should be considered. 5.2 Effect of trenching method Jetting When jetting in soft clay a water-clay suspension is expected to prevail in the trench immediately after trenching. The pipe will be completely surrounded by this material when it is installed in the trench, and the water/clay suspension will gradually settle. Minor penetration of the pipe into the intact material at the bottom of the trench may also be expected. The shear strength of the clay surrounding the pipe will gradually increase from practically zero to that of a normally consolidated clay, depending on the coefficient of consolidation and the thickness of the clay layer. When estimating the reconsolidated strength, uncertainties in both the normally consolidated shear strength ratio, s u σ' v, the unit weight and the degree of backfill should be considered. Following the consolidation process the height of the clay backfill will decrease. With time, the clay in the trench will regain shear strength. The regained shear strength is eventually expected to reach a constant level. The upward displacements required to reach the maximum uplift resistance are also likely to be affected by the method of trenching. Jetting may introduce water filled voids in the soil in the trench, but generally the soil in the trench will form a homogeneous material. It is not recommended to rely on any upheaval resistance in a jetted cohesive soil shortly after installation. The upheaval resistance at time of operation needs to rely on the degree of consolidation of the backfilled material. When trenching by jetting in sandy soil, the sand is fluidized before settling which is similar to the process of preparing soil for determination of minimum density in the laboratory /54/. Thus initially after trenching Recommended practice DNVGL-RP-F114. Edition May 2017 Page 64

65 the sand will have close to zero relative density. Any kind of subsequent loading may densify the sand. One potential source is varying hydrodynamic pressure on the seabed due to waves. This will cause some vertical gradients of the hydrodynamic pore pressures, but more important also impose shear stresses to the soil /55/, which may build up pore pressures in the sand. Such pore pressure build up with subsequent drainage may lead to a densification of the sand, which will be favourable if occurring prior to the operation of the pipeline. However, increased pore pressures will reduce the upheaval resistance compared to that calculated for a fully drained condition. This should be carefully considered if water depth and wave conditions are such that pore pressures may build up during a storm. The shear stress level in the soil may be evaluated from /55/ Ploughing When the pipe is placed in a soft clay trench formed by a ploughing device with subsequent backfill of the trenched material, the water content of the clay will not increase relative to that of the intact material. Thus the remoulded resistance of the clay as established through sensitivity measurements is likely to represent an expected minimum strength. The regained shear strength with time is eventually expected to reach a strength which is proportional to the effective stresses in accordance with theory for normally consolidated clays, or equal to the remoulded shear strength, whichever is the greater. Ploughing is expected to change the macro structure of the clay by introducing cracks and water-filled voids. When ploughing in a stiff clay, the clay is likely to break up forming lumps of clay, and the upheaval resistance would be related to the interface shear resistance between lumps of clay rather than the shear strength in the clay material. Ploughing in sandy conditions will affect the relative density, however less severe compared to jetting. Subsequent densification due to wave loading may increase the relative density with time, but such positive effects are difficult to quantify. 5.3 Axial pipe-soil interaction General The axial resistance of a buried pipeline can be determined by investigating two different failure modes; a deep and a shallow mode (as illustrated in Figure 5-1 where a two-layer model is shown). The difference between the modes is whether the soil above the pipe slides axially together with the pipe, or if the shear resistance of the backfilled material is high enough to prevent soil movement of the soil above the pipe. The shallow mode will be governing for low cover heights. Recommended practice DNVGL-RP-F114. Edition May 2017 Page 65

66 Figure 5-1 Illustration of axial failure modes for buried pipelines For each scenario the engineer needs to assess the material surrounding the pipe, drainage conditions and loading rate to evaluate whether the soil surrounding the pipe will behave drained or undrained Drained resistance A two-layer model, assuming ideal conditions, as given in Figure 5-1, can be used to estimate the axial friction in drained conditions. The equation for the deep mode has been found by integration of contact stresses around the pipe circumference assuming two different layers. The equation for the shallow mode is found by replacing the formula for the upper half of pipe with the vertical shear resistance through the upper layer. The ultimate resistance will be the lowest resistance obtained by the two failure modes. Deep failure mode: (5.1) Shallow failure mode: (5.2) Recommended practice DNVGL-RP-F114. Edition May 2017 Page 66