Confronting New Challenges and Sharing Knowledge, 11–13 .../media/Files/A/Atkins-Middle... · Confronting New Challenges and Sharing Knowledge, 11–13 September 2007, London, UK

Proceedings of the 6th International Offshore Site Investigation and Geotechnics Conference: Confronting New Challenges and Sharing Knowledge, 11–13 September 2007, London, UK

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1. IntroductionSubsea pipelines are increasingly being required to operate at higher temperatures and pressures. The natural tendency of a hot pipeline is to relieve the resulting high axial stress in the pipe wall by buckling. Such uncontrolled buckling can have serious consequences for the integrity of a pipeline. Consequently, the industry has generally sought to restrain pipelines by trenching, burying and rock dumping, or to relieve the stress with inline expansion spools.

A far more elegant and cost effective solution is to work with rather than against the pipeline by controlling the for-mation of lateral buckles along the route. Controlled lateral buckling is an efficient solution for the relief of axial com-pression. Indeed, as operating temperatures and pressures increase further, lateral buckling may be the only economic solution. This challenge has led to a radical advance in pipe-line engineering with a greater need for robust lateral-buck-ling design solutions. The SAFEBUCK JIP1 was initiated to address this challenge and aims to raise confidence in

the lateral-buckling design approach and to improve under-standing of the related phenomenon of pipeline walking.

The pipe-soil force displacement response is the largest uncertainty in the design of such systems. With lateral buckling it is necessary to understand the soil behaviour at large displacements and through many cycles of loading well beyond the point of failure. Such behaviour is outside the bounds of conventional geotechnics or extensive earlier research on pipeline stability. Most previous research into pipe-soil interaction has been related to stability under hy-drodynamic loading, with the aim being to ensure the pipe remains in place. A lateral buckling design requires the pipe to break out from the as-laid position and move across the seabed, typically by several diameters.

The purpose of this paper is to outline the significant influence that pipe-soil interaction has on the pipeline design process and highlight the way in which the inherent uncertainty in pipe-soil resistance severely complicates pipeline design.

THE INFLUENCE OF PIPE-SOIL INTERACTION ON LATERAL BUCKLING AND WALKING OF PIPELINES – THE SAFEBUCK JIP

DAS Bruton and M CarrAtkins Boreas, Aberdeenshire, Scotland

DJ WhiteCentre for Offshore Foundation Systems, The University of Western Australia

AbstractThis paper addresses the influence of pipe-soil interaction on the design of pipelines susceptible to lateral buckling and pipeline walking. The pipe-soil response is the largest uncertainty in the design of such systems, and so generic guidance has been developed to guide the design process.

Force-displacement-response models were developed during Phase I of the SAFEBUCK joint-industry project (JIP) based on large- and small-scale tests carried out by the SAFEBUCK JIP and project-specific test data donated by JIP participants. These models are currently being applied by JIP participants on a number of projects, to quantify the susceptibility to lateral buckling and pipeline walking, and design safe and effective means to control these phenomena. However, of all the design parameters, the soil response causes the greatest uncertainty in design because of the extreme sensitivity of design solu-tions to the axial and lateral resistance imposed by the soil. Improving the understanding of pipe-soil response provides the greatest scope for refining the design of such systems.

The purpose of this paper is to outline the significant influence that pipe-soil interaction has on the pipeline design process and to highlight the way in which the inherent uncertainty in pipe-soil resist-ance severely complicates pipeline design. The paper then describes the research, development and model refinement that is ongoing to reduce the uncertainties.

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Bruton, Carr and White. The Influence of Pipe-Soil Interaction on Lateral Buckling and Walking of Pipelines – the SAFEBUCK JIP

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2. Pipeline Structural Response2.1 Pipeline installationPipe-soil interaction influences the behaviour of the pipe from the moment installation commences. The interaction between the pipeline touchdown loads, combined with the dynamics of the pipe catenary, and the seabed surface soil defines the initial pipeline embedment. The remoulding of the soil that occurs during installation then influences the axial resistance, affecting the conditions at which the pipe becomes constrained2. Meanwhile the lateral resistance af-fects the tightness of route curves that the installation con-tractor can achieve.

Furthermore, the time that passes between installation and operation modifies the pipe-soil response as the soil is con-solidated under the weight of the pipe. This consolidation leads to a small increase in pipe embedment and changes the axial response associated with pipeline start-up. This is particularly relevant for soft clays, where consolidation ef-fects can significantly increase the soil strength close to the pipe, increasing the resistance to first movement.

2.2 First load: general pipeline expansion behaviourAfter installation, some residual effective lay tension remains in a pipeline. This force is usually small in comparison with the forces that develop in operation and is not considered further here. When the internal pressure and temperature increase to operating conditions the pipeline tends to ex-pand, but this expansion is resisted by the axial resistance between the pipe and the seabed. This restraint causes an axial compressive force to develop in the pipeline.

The ends of a pipeline are usually free to expand, so the force at the ends is zero. However, as the cumulative axial resistance increases with distance from the pipe ends, the force can increase to a condition of ‘full-constraint’, as il-lustrated in Figure 1.

The fully constrained effective force presented in Figure 1 is the maximum effective axial force that can occur in a pipeline. This fully constrained force drives the axial expan-sion and structural response. The effective force is made up of the (true) force in the pipe wall and the pressure induced axial force. Since pressure and temperature vary along the pipeline length, the fully constrained force also varies along the length as the pipeline cools (with heat loss to the envi-ronment). This is shown in Figure 1 by the slight fall in the fully restrained force along the pipeline.

The gradual increase in effective axial force, from zero at the free ends to full constraint, is due entirely to the cumulative axial restraint provided by the seabed. The slope of this line is equal to the axial resistance (force) per unit length, which is typically modelled as being ‘frictional’, thus proportional to the pipeline weight, W’µa. This force profile is funda-mental to the pipeline response.

The important influence of a reduced level (or lower bound estimate) of axial friction on the effective force profile is il-lustrated in Figure 2.

In Figure 2 the maximum effective force in the lower bound

axial friction case, µa = 0.10, is reduced by a factor of about 4 and the pipeline does not reach full constraint. This illustrates a real design case with upper and lower bound frictions based on upper and lower bound soil responses that correspond to drained and undrained axial movement, respectively.

The compressive effective axial force in a pipeline therefore depends on the operating condition of the pipeline and the axial friction. If the compressive force is large enough, then the pipeline may be susceptible to lateral buckling. Susceptibility to lateral buckling occurs when the com-pressive force exceeds the ‘critical buckling force’, above which the pipeline becomes unstable laterally. This lateral instability is critically dependent upon the lateral soil re-sistance (specifically the ‘lateral breakout resistance’ which is discussed later). Once lateral buckles have formed, the axial force drops significantly as pipe feeds axially into the buckle, as shown in Figure 3.

In the example shown the pipeline has a totally different behaviour at lower bound axial friction and upper bound axial friction. The lower bound friction, µa = 0.10, means that the pipeline will experience significantly greater end expansion and will be susceptible to pipe-walking. The up-per bound friction, µa = 0.58, means that the pipeline will become fully constrained over some of its length so that this section will not move axially, thus preventing walking. However, the maximum effective force in the pipeline in-creases significantly, making it much more susceptible to lateral buckling.

Low axial friction will increase the end expansion and axial feed-in to lateral buckles, while high levels of axial friction will tend to reduce end expansion and feed-in, as illustrated in Figure 4.

Figure 4 highlights how the different levels of axial resist-ance fundamentally modify the pipeline expansion behav-iour. At low axial friction, the pipe is fully mobilised and expands from its centre, reaching a maximum displacement in excess of 2.5m. At high axial friction, in the absence of buckling, there is no displacement over the central section (the pipe is fully constrained) and the pipe expands only at the ends to reach a maximum displacement of about 1m. With lateral buckling the expansion occurs over the whole length, but the direction changes as feed-in occurs towards the buckles, as well as towards the ends of the pipe, reduc-ing the end expansion to about 0.7m.

For the low axial friction condition, the pipe is not suscep-tible to buckling. The design challenge is then controlling the extreme end expansion and its susceptibility to pipeline walking. For the high axial friction condition, the pipe is susceptible to buckling over most of its length, and the de-sign challenge is controlling the severity of the buckles and pipeline walking. Somewhere between these two extremes, a pipeline design solution must be found.

2.3 Response in a buckle: first loadThe attraction of the lateral buckling design solution is in the reduction in compressive axial force that occurs in the pipeline. However, the buckling displacement results in sig-

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nificant bending moment at the crown of the buckle. This moment can be very high and stresses in excess of yield are normal (although these are relatively limited in axial extent). Thus, under the lateral buckling philosophy, although the axial force is reduced, the bending stresses are significantly increased. This can ultimately lead to a local buckling fail-ure with associated deformation of the pipe cross-section.

At start-up, the soil ahead of the buckle crown is heavily remoulded as a berm of soil builds up in front of the lat-erally sweeping pipe (Figure 5). The lateral displacements involved are very significant; displacements of 10 or 20 di-ameters are typical (in absolute terms, displacement in the range 2 to 10m has been observed in operating pipelines). The length of pipe over which the lateral displacement oc-curs is typically between 100 and 300m (the lateral scale is exaggerated in Figure 5).

The lateral resistance during pipe movement governs the level of curvature and bending stress in the pipe. Lateral soil resistance is usually the largest uncertainty in designing for lateral buckling and unfortunately has a significant influence on the design limit states, as illustrated in Figure 6. This fig-ure shows the effect of the lateral resist-ance, expressed as a friction coefficienta, on the limit state utilisation within the crown of the buckle.

Figure 1: Effective axial force in a straight pipeline

Figure 2: Effective axial force for a range of friction in a straight pipeline

Figure 3: Effective axial force in a short pipeline with lateral buckles

Figure 4: Displacement along a short pipeline with lateral buckles

Figure 5: Side-scan sonar image of a lateral buckle

a The lateral response of a pipe resting on clay soil is not truly ‘frictional’, as discussed later in section 3.1.

Figure 6 shows that low lateral resistance is extremely desir-able since it produces a much lower bending severity in the crown of the buckle. It is also beneficial from a buckle for-mation point of view (as discussed later). However, a quite small increase in lateral resistance can significantly increase the design-limit usage factor, so that uncertainty in lateral resistance is likely to compromise the design (and must not be underestimated).

If the loads within the buckle are too high, then the de-sign must seek to reduce these. The severity of the buck-ling problem is driven by how much pipe feed-in must be absorbed by the buckle. Decreasing the distance between buckles reduces the feed-in and hence controls the severity. The lateral-buckling design solution relies on the regular formation of lateral buckles so that the load in each does not exceed design limits (local buckling, strain capacity or fatigue). This approach defines the maximum allowable

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buckle spacing, or virtual anchor spacing (VAS)b , which is typically thousands of metres.

In Figure 6, it is shown that reducing the distance between buckles (equivalent to a reduction in the VAS from 2.5 to 2.0 km) reduces the loading. A greater reduction in VAS will reduce the loading further, but this can make it dif-ficult to ensure that the buckles form reliably. If the spacing cannot be reduced further, then the only other approach is to reduce the lateral resistance. In Figure 6 this would be required if the lateral resistance is expected to be much higher than an equivalent friction coefficient of 1.0; this would mean adopting alternative buckle control methods to achieve system integrity.

Consequently, the pipeline designer has two ways of con-trolling the design problem:• Decrease the distance between buckles• Decrease the lateral resistance.Both of these solutions lead to the adoption of buckle ini-tiation techniques (Section 2.5).

2.4 Response in a buckle: cyclic behaviour – influence of bermsIn operation, as the pipeline experiences fluctuations in pressure and temperature, the pipe cycles back and forth across the same patch of seabed. Surface soil, swept ahead of the pipe on each cycle, then builds up into berms at the extremes of pipe displacement. Subsequent consolidation increases the strength of the soil berms after disturbance. Pipe feeds axially into and out of the buckle with each cycle. On unloading, the pipeline attempts to return to the as-laid position but is prevented from doing so by both axial and lateral pipe-soil resistance.

Without soil berms, numerical modelling with lateral re-sistance calculated using a single friction coefficient shows that buckles will grow in amplitude with each cycle. In real-ity, soil berms restrict the growth of the buckle so that cy-clic displacements remain almost constant over a number of cycles. This increased restraint in the buckle causes higher stress ranges in operation, leading to a higher level of fa-tigue damage. This effect is illustrated by the example in Figure 7.

This example clearly illustrates the influence of soil berms, although much uncertainty remains in predicting the level of restraint provided by berms. Currently, this uncertainty is addressed in design by assuming a range of berm resist-ance to assess the sensitivity on allowable loading.

2.5 Buckle initiationBuckle initiation is governed by three parameters: (i) the effective compressive force in the pipeline, which is a func-tion of axial resistance; (ii) out-of-straightness (OOS) fea-tures; and (iii) lateral breakout resistance. Although only three parameters are involved, a number of factors feed into

each parameter and there is significant uncertainty over the true magnitude of each. Lateral breakout resistance is gen-erally the largest uncertainty.

If lateral buckles are initiated at regular intervals along the pipeline, the loads are effectively shared between buckle sites. However, the shorter the spacing between buckle ini-tiators, the lower the probability is of buckles forming at each site as desired. Therefore, selecting a suitable spacing is often a difficult design compromise between these compet-ing requirements.

If there is no strategy to form buckles at regular intervals along a pipeline, they will form randomly and (generally) less frequently than if an initiation strategy is employed. This random formation may produce an acceptable design. However, the large uncertainty in pipe-soil resistance means that the pattern of formation is extremely challenging to predict. In all but the most benign cases, it is impossible to demonstrate that inherent buckling will occur to an accept-able level of reliability. In these cases, an engineered buckle initiation technique must be adopted.

2.5.1 Buckle initiation techniquesA number of methods are available to initiate buckling at a controlled spacing; these include:• Snake-lay (Figure 8), where the pipe is laid with regular

tight-radius route curves• Vertical upset (Figure 9), where rock or pipe sleepers are

used to raise the pipe off the seabed• Local weight reduction (Figure 10), where distributed

buoyancy or additional insulation coating is applied, or weight coating is reduced or removed.

For the snake-lay method the lateral restraint is governed by the lateral breakout resistance (discussed later), while the OOS is defined by the tightest curvature for which that installation contractor can install. This curvature is gov-erned by two parameters: lateral breakout resistance and lay tension. In reality, tighter than average design curvatures usually occur due to installation dynamics, but these are difficult to predict in advance of installation.

The weight reduction and vertical upset solutions both provide a vertical OOS to encourage buckle formation, although this is perhaps easier to quantify in the vertical upset solution. Both techniques also benefit from the de-sign by reducing the pipe weight or lifting the pipe off the seabed, reducing contact with the seabed and therefore de-creasing the lateral resistance across the buckle. This reduc-tion in loading means that fewer lateral buckles are needed to share the load. Therefore, buckles can be spaced further apart, which increases the reliability of buckle formation. The benefit that these techniques bring has resulted in their use on a number of challenging deepwater projects (e.g. Harrison et al.3).

2.5.2 Modelling issuesSpecial consideration is required in zones where the local ef-fective pipe weight changes due to loss of contact with the seabed, for example at sleepers or at the start of a distributed

b VAS is the distance between virtual anchors that form to each side of the buckle, usually equivalent to the spacing between lateral buckles.

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Figure 6: Typical design limits for a light pipe with a range of lateral

Figure 7: Displaced shape at the buckle (a) without soil berms and (b) with soil berms

buoyancy section. Such features, introduced deliberately to control lateral buckling, require careful implementation of the pipe soil response in finite element analysis (FEA). A dis-tinction must be made between the pipe effective weight, W’, and the local contact force between the pipe and the soil, V. The same approach is also applicable to spans, which can oc-cur along the line and may also act as buckle initiators.

Particular attention should be given to the definition of soil resistance at: • The span touchdown points (TDP) to each side of a

sleeper or span• The start and end of distributed buoyancy sections, where

there is a sudden change in diameter and pipe weight.The touchdown loads to each side of a sleeper are typi-cally 60% greater than the effective pipe weight. The pipe will experience significant additional lateral restraint in these touchdown regions to each side of the sleeper, which strongly influences the load in the lateral buckle. This effect is shown in Figure 11, which shows the plan, elevation and vertical load across a lateral buckle over a sleeper for a lower and upper bound value of lateral resistance.

Figure 9: Vertical upset buckle initiators using sleepers (typi-cally 2 to 3 joints of large diameter pipe)

Figure 10: Local weight reduction buckle initiators using dis-tributed buoyancy

Figure 8: Typical snake-lay configuration (exaggerated vertical scale)

This example in Figure 11 includes a reduction in pipe weight by the addition of buoyancy between 900 and 1100m, thus combining the vertical upset and local weight reduction techniques. It is shown that the highest vertical load occurs within the lateral buckle and is a significant influence on the its shape. A design approach that provides an increase in lateral resistance in these zones is necessary, usually by defining the lateral resistance as a function of the local contact force between the pipe and the soil. This approach leaves the FEA to take account of seabed stiffness and bending of the pipe in the vertical plane.

A realistic value of seabed stiffness is necessary in design to arrive at realistic contact pressures. A very stiff seabed (hard contact) can overestimate the contact pressures and lateral resistance at the TDP to each side of the sleeper. A realistic seabed stiffness to be adopted in design for soft seabeds is typically in the range of 15 to 50kN/m/m, but default sea-bed stiffness values in FEA are often much higher than this. The variation in diameter that occurs at sections of distrib-uted buoyancy should also be captured in FEA models to ensure correct modelling of the pipe-soil contact regions.

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2.6 Pipeline walking When a pipeline is heated it will expand, but this expansion is opposed by axial resistance. When the pipeline is cooled it contracts, but the axial resistance prevents the pipeline contracting to its original position. Subsequent shutdown and restart cycles are normally accompanied by steady-state expansion and contraction between established pipe-end positions. However, in some cases this cycling can be ac-companied by global axial movement of the pipeline, which is termed ‘pipeline walking’. Over a number of thermal cy-cles, walking can lead to significant global displacement of the pipeline. Walking is not a limit state for the pipeline itself, but without careful consideration can lead to failure at the mid-line or end connections.

The SAFEBUCK JIP investigated the pipeline walking phe-nomenon, which has occurred in a number of pipelines and led to at least one failure. A study4 carried out by the JIP defined the key factors that influence pipeline walking and provided guidance and analytic expressions for assessing the severity of pipe walking.

Figure 11: Influence of touchdown load and lateral

resistance on buckle shape

The main causes of pipeline walking are:• Tension at the end of the flowline, associated with a steel

catenary riser (SCR)• A global seabed slope along the pipeline length• Thermal gradients along the pipeline during changes in

operating conditions.

The key parameters that influence pipeline walking are (i) the axial resistance response and (ii) the pipeline length over which walking occurs. Pipeline walking is traditionally as-sociated with pipelines that are short enough not to reach a condition of full-constraint (see Figure 2); such pipelines are typically less than 5km in length. With the current increase in pipeline operating temperatures, ‘short’ pipelines can be many kilometres in length. The phenomenon can also oc-cur in longer lines where lateral buckling has occurred, as the lateral buckles effectively divide the long pipeline into a series of shorter lines.

The presence of lateral buckles dramatically changes the walking response of a pipeline, and the presence of high

Figure 12: Pipeline walking due to seabed slope for a range of axial friction coefficients

Figure 13: Effective force distribution along a walking pipeline restrained by an anchor

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thermal transients can lead to buckle growth over a number of load cycles. Lateral buckles are a significant influence on the walking response, which highlights the need to under-stand this interaction. Currently the interaction between walking and buckling is assessed on a case-by-case basis by FEA in detailed design.

The example in Figure 12 illustrates the influence of pipe-line length (or buckle spacing) and axial friction, µa, on pipeline walking for a range of seabed slopes. In many cases, as shown in Figure 12, a high level of axial resistance can reduce walking significantly. Pipeline length (or section length) is a significant parameter in the assessment of pipe-line walking, but axial resistance has a greater influence. If the pipe weight is sufficiently high, or the fully constrained force sufficiently low, walking can stop altogether.

2.7 Control of pipe walking When walking is considered unacceptable, it becomes nec-essary to anchor the flowline. Hold-back suction anchors are commonly employed to control pipeline walking. In most cases these anchors allow expansion towards the an-chor but prevent the end of the pipe pulling back much beyond the installed position; this ensures that the effective compression in operation is not increased.

However, as walking commences the anchor loads up, until after some tens of cycles, the anchor experiences the full shut-down tensile load. This restraining anchor load induces ten-sile loads along the length of the pipeline. The force profile at shutdown and the restraining anchor load is then defined by the level of axial resistance, as shown in Figure 13.

In the example shown in Figure 13, walking is caused by the presence of a steel catenary riser with a nominal tension of 1MN at one end of the pipeline (at KP 0). At this tension an anchor is required at the opposite end of the pipeline with a capacity of 3.5MN. The effective tension within the pipeline exceeds 4.5MN at KP 3.0. This is a significant lev-el of tension, which introduces a further design challenge: route curve stability. The concern is that this tension could be sufficient to cause instability of the pipeline – through ratcheting lateral displacement – at the route curves. This instability can pull out the curve, allowing further pipe to walk, until the curvature is small enough to be stable.

The minimum stable radius of curvature is governed by the minimum lateral resistance that the soil can provide. If the lateral resistance is low, then this radius of curvature may be so large as to compromise field architecture. This issue can lead to major field layout changes and should be addressed as early as possible in front-end engineering design. The in-stability of route curves has led in the past to some radical changes to field architecture to eliminate route curves or increase their radius. In some cases, the production facility has been relocated, in others the flowline has been split into two straight sections with a mid-line tie-in, which has seri-ous capital expenditure (CAPEX) implications.

Although a low level of lateral resistance is a disadvantage for stability of route curves, it is an advantage for reliable

buckle formation. This once again illustrates the need to bound soil resistance; there is no conservative extreme.

2.8 Key influences of pipe-soil resistanceThe preceding examples demonstrate the significant influ-ence that pipe-soil interaction has on pipeline design. The axial pipe-soil resistance affects• The maximum axial effective force in the pipeline• The effective force available for reliable buckle initiation• Pipe-end expansion and feed-in to lateral buckles• Pipeline walking, including the rate of walking and the

propensity to walk.

Similarly, lateral pipe-soil resistance affects• The lateral instability that is required for buckle initiation• Route curve stability under axial tension• The lateral buckle bending load at large displacements • The cyclic lateral buckle loading due to soil berms.

The influence of pipe-soil resistance on these various issues is often conflicting. For example, high lateral resistance is beneficial for curve stability, but detrimental to the bending loads developed in a buckle. High axial resistance is benefi-cial for controlling feed-in to a buckle, but increases the sus-ceptibility to buckling in the first place. Consequently, it is not possible to define a safe resistance. Instead, it is necessary to bound the pipe-soil resistance and ensure that the design is acceptable throughout the potential design envelope.

3. Current Approach to Modelling Pipe-Soil Interaction3.1 Soil resistance modelsThe interaction between the pipe and the seabed is incor-porated into the structural analysis of a pipeline – which is usually finite element-based – by attaching ‘spring-slider’ elements at intervals along the pipe. This approach is analo-gous the t-z and p-y load transfer methods of analysing pile response. The most basic pipe-soil elements provide a bilin-ear elastic–perfectly plastic response in the axial and lateral directions. However, in order to capture the more advanced effects of interaction, particularly the large displacement behaviour, it is necessary to introduce subroutines in which the element response is modified to account for brittle brea-kout behaviour and cyclic berm growth.

The most simple bilinear models involve a limiting value of axial or lateral pipe-soil resistance, which is calculated using a simple friction law (Coulomb friction) linking the effective pipe weight to the maximum available resist-ance. A suitable axial and lateral friction coefficient can be used successfully for some flowline design functions (e.g. simple stability calculations or end-expansion) and can be employed in conceptual evaluation of lateral buckling, if treated with care. However, a single friction coefficient is not appropriate in detailed numerical modelling design for lateral buckling, particularly for large-amplitude lateral movement, where a frictional model represents an oversim-plification of the behaviour.

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Nevertheless, it is common to model many tens of kilome-tres of pipeline within a single FE analysis. It is therefore necessary to keep the pipe-soil interaction models relatively simple. It is not feasible to model the soil domain around the pipe along the entire pipeline length, and so it remains necessary to encapsulate the pipe-soil behaviour into the response of a single node. Non-linear force displacement re-sponses are therefore used to represent pipe-soil behaviour. These responses are usually presented as an equivalent fric-tion coefficient – that is the maximum resistance divided by the submerged weight of the pipe. This coefficient is then updated automatically throughout the analysis to simulate the underlying behaviour.

A friction coefficient is an unusual concept to apply to the limiting resistance of a clay, which is usually characterised by an undrained strength. However, this is a convenient way of presenting the data for input into design using ana-lytical or finite element based models and remains compa-rable with historic approaches for pipe-soil interaction. This terminology should not, however, be taken to indicate that lateral pipe behaviour is purely frictional; the limiting lat-eral and axial resistance is not solely dependent on the pipe weight, but is significantly influenced by the embedment and soil conditions.

3.2 Interaction between axial and lateral resistance Axial and lateral resistance should be treated as independ-ent parameters in design for any given set of soil conditions. Since the axial resistance response is global – affecting long sections of pipeline – the total axial resistance is negligibly influenced by the localised mobilisation of lateral resistance.

This approach can provide a low assessment of axial resist-ance where axial and lateral interaction is possible locally, at specific regions of a lateral buckle away from the crown. Axial resistance involves shearing at the pipe-soil surface, whereas lateral resistance involves failure within the soil ahead of the pipe. These mechanisms have minimal inter-action. However, in drained conditions, lateral loading on a pipe will increase the effective stresses between the pipe and the soil, raising the available axial resistance. Therefore, treating axial and lateral resistance as independent quanti-ties is likely to underestimate the axial resistance at specific regions of the lateral buckle where drained axial and lateral displacement occurs simultaneously; this approach is usu-ally conservative.

Caution is necessary with some FEA software that includes built-in soil elements. These often combine axial and lateral resistance using an elliptical combined yield surface. This is non-conservative as it results in reduced lateral resistance when axial resistance is mobilised. In short, any combina-tion of lateral and axial resistance causes a reduction of lim-iting resistance on both axes.

3.3 Pipeline embedment 3.3.1 Definition of pipeline embedmentPipe embedment is defined as the depth of penetration of the invert (bottom of pipe) relative to the undisturbed sea-bed. Pipeline embedment therefore influences the pipe-soil

contact area, which affects the axial resistance, and the pas-sive soil resistance against the pipe, which affects the lateral break-out force. Heave of soil against the shoulders of the pipe raises the contact area above the value for a wished-in-place pipe at the same embedment. Figure 14 illustrates the initial embedment of a pipeline.

3.3.2 Embedment uncertaintiesThe current model and background for prediction of em-bedment has already been published by the SAFEBUCK JIP5. However, pipeline embedment is notoriously difficult to predict. The values of as-laid embedment predicted using this and other models6 is generally found to be substan-tially exceeded in practice. This discrepancy is because the touchdown load is actually cyclic in nature, due to dynamic effects including vessel heave and hydrodynamic loading of the pipe catenary. This increase in embedment is captured in design by multiplying the embedment found from a static embedment calculation by a ‘dynamic embedment factor’, which is a function of the lay vessel dynamics, lay rate, lay tension, pipeline configuration and environmental loading during installation. However, these secondary variables are not known prior to installation. Over-embedment, due to the uncertain and variable influence of the laying process, is the main source of uncertainty in the prediction of pipeline embedment and breakout resistance.

The dynamic embedment factor carries significant uncer-tainty. Comparison between field experience for soft clay seabeds with shear strengths of 2 to 4kPa and predictions from static analysis (based on the as-laid pipe weight mul-tiplied by a suitable stress concentration factor at the TDP) give a typical dynamic embedment factor in the order of 1.0 to 3.0. On stiff clay seabeds with shear strength exceed-ing 100kPa, the factor can be in the order of 5.0 to 8.07. Further work is ongoing in this area, but evaluation has been hampered by the limited availability of suitable data from installed pipelines. Simulation of the lay process is possible in model scale testing, but the input parameters related to the motion of the vessel and hydrodynamic load-ing of the lay catenary are difficult to assess.

A further complication related to pipe embedment is the definition of the soil surface, which forms the datum for observed levels of embedment on installed pipelines. In situ test data of shear strength often does not register an increase in resistance at the observed soil surface level, even when using a sensitive T-bar penetrometer. This can either be be-cause the T-bar has already penetrated the surface during placement of the seabed frame, leading to some depth offset between the visual observations and the in situ test data. Alternatively, the visible surface layer may be so weak that its significance can be overlooked. In deep water, a very weak layer of soil about 70 to 100mm thick is commonly found at the surface and its presence is also commonly ignored. However, any visual assessment of pipeline embedment from post-installation surveys will include this weak layer.

3.4 Axial pipe-soil interaction3.4.1 Definition of axial resistanceAxial pipe-soil resistance is often modelled using a simple

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friction coefficient. However, some axial displacement must occur to reach full axial resistance, so that small axial move-ments can occur at quite low loads. This behaviour influ-ences the initiation of buckling by allowing small OOS features to grow at quite small loads.

In addition, when the pipe is initially loaded, an axial ‘brea-kout’ or peak in resistance can occur that falls away to a residual axial friction after breakout. The displacement at which this peak occurs is defined as the ‘mobilisation dis-placement’, illustrated in Figure 15.

Two stages of axial pipe-soil interaction should be considered in design, (i) breakout resistance and (ii) residual resistance.

Breakout axial resistance: A significant peak in axial resist-ance can occur when the pipe moves axially for the first time, or after some time at rest. This first movement is typi-cally associated with buckle formation (initiation) and with shutdown after some time at rest under steady-state operat-ing conditions. This peak is not expected to occur on restart, unless the pipeline has been shutdown for some time.

Residual axial resistance: Once the pipe has started to move, residual axial friction will control the pipe-end ex-pansion/contraction and axial feed-in/feed-out to each lat-eral buckle. The term ‘residual’ is used by analogy with the residual friction angle that is mobilised within fine-grained soils after continued shearing along a single plane.

3.4.2 Axial resistance uncertaintiesIt is clear that pipeline walking tends to reduce and can stop altogether at high levels of axial resistance. This reaffirms the need to understand and better predict axial resistance in soft clays, in particular the need to predict if the response is likely to be drained or undrained.

Pipe laying is an undrained process for typical rates of lay-ing on typical soft clays, so the effective stress at the pipe-soil interface usually remains close to zero, which was the equilibrium effective stress at the soil surface prior to laying of the pipe. Positive excess pore pressure is generated to bal-ance the pipe weight.

During a period of equalisation, the excess pore pressure dissipates, and the effective stress at the pipe-soil interface increases to balance the applied pipe weight. This develop-ment of effective stress governs the availability of axial pipe-soil resistance. Theoretical solutions indicate that the ma-jority of consolidation typically occurs within one month of laying, even in low permeability natural clays. These so-lutions allow the pipe-soil effective stress at the time when the pipe enters operation to be estimated. At this point,

the undrained strength of the soil beneath the pipe is likely to differ from the in situ value due to two mechanisms: (i) remoulding during pipe embedment, which reduces the strength, and (ii) consolidation and a rise in effective stress after embedment, which increases the strength.

The sum of the normal contact stresses around the pipe periphery exceeds the vertical contact force, V, due to the curved shape of the pipe surface. A simple analysis of this ‘wedging’ effect based on an elastic stress field indicates that the ratio of the vertical to normal contact force increases linearly from 1.0 to 1.25 as the pipe embedment rises from 0 to 0.4 diameters8.

Traditional axial resistance models for ‘cohesive’ soils de-fine the axial resistance of the soil as a product of the shear strength and the contact area between pipe and soil (which is itself governed by the pipe embedment) and a multiplier, α. As the pipe is displaced during breakout, it is assumed that the shear strength reduces to the remoulded shear strength of the soil, defined by the soil sensitivity, giving peak and residual values of resistance (Figure 15).

This model is analogous to the total stress ‘alpha’ method of axial pile shaft capacity. It is described as undrained, meaning that excess pore water pressure is assumed not to dissipate. To use this model, it is necessary to assess a suitable value of α which accounts for (i) any changes in undrained shear stress, su, due to the laying and equalisation processes and (ii) any difference between soil-soil and pipe-soil shearing. The pipe-soil contact perimeter must also be estimated.

Alternatively, axial friction can be defined using a ‘drained’ model – analogous to the ‘beta’ effective stress approach for axial pile shaft capacity. In contrast to pile design, the con-tact stresses between a pipe and the seabed are known, being due to the pipe weight (with a modification for the ‘wedging’ effect described earlier). The drained axial resistance can be calculated based on this effective contact force (noting the possibility that some lay-induced excess pore pressure may remain) multiplied by the soil friction coefficient, tanφ.

While this drained friction coefficient is simpler to define, it is significantly influenced by the level of effective stress, which lies below that considered in conventional geotech-nics. At the effective stress levels of 2 to 10kPa generated by typical pipeline weights, the drained friction angle of soft clays significantly exceeds that measured by traditional test-

Figure 14: Initial embedment

of a pipeline

Figure 15: Axial resistance response with mobilisation dis-placement and breakout

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ing apparatus operating at usual effective stress levels. The dif-ference can be critical to pipeline design, as demonstrated by the preceding examples. Friction coefficients as high as unity (corresponding to friction angles in the range 40–50º) have been measured for soft deepwater clays at low stress levels.

During fast axial pipe movement, the apparent friction coefficient that should be used in the ‘beta’ effective stress model may increase or decrease depending on whether neg-ative or positive excess pore pressure is generated by shear-ing at the pipe-soil interface; this tendency depends on the current over-consolidation ratio (OCR) of the soil.

Since undrained (fast) and drained (slow) responses gen-erally provide very different values of axial resistance, it is important to consider the interaction between the available resistance and the resulting pipe velocity. An analysis based on drained parameters that yields high pipe velocities is meaningless, and vice versa. Limited field data is available on pipe velocities during operation, but one analysis exam-ple is illustrated in Figure 16, which shows some typical pipe velocities during a shutdown and restart cycle.

The velocities shown here are generally quite low and only exceed 0.2mm/s at locations close to the pipeline ends (KP 0 and KP 7.5) and then for quite short durations. The impli-cation is that the axial response is likely to remain drained, but this is by no means certain.

The important transition from undrained to drained condi-tions is not well understood. While dissipation of the excess pore pressure from laying is expected to occur relatively quickly, it does not prevent undrained conditions from be-ing re-established under conditions of rapid or large pipe displacements. However, re-consolidation after axial move-ment is likely to be quicker than after embedment, since only a small region of soil at the pipe surface is sheared during axial motion. This topic is an area of ongoing re-search under Phase II of the SAFEBUCK JIP to improve the fundamental understanding and better define the most appropriate axial resistance models and test methods.

3.5 Lateral pipe-soil interaction3.5.1 Definition of lateral resistanceWhile lateral breakout loads have been the subject of much research and published papers, until recently there has been little guidance on modelling lateral resistance for the large cyclic displacements of typically 10 or 20 diameters experi-enced in lateral buckling.

The current models and background for prediction of brea-kout and residual resistance have already been published by the SAFEBUCK JIP5. The simplified response is presented as an equivalent friction coefficient, H/V, where H and V are the horizontal and vertical load on the pipe per unit length, respectively.

A simplified force-displacement lateral-friction response, based on model tests, is illustrated in Figure 17.

Four stages of lateral pipe-soil interaction should be con-sidered in design, (i) breakout resistance (ii) suction release (iii) residual resistance and (iv) cyclic lateral resistance.

Breakout lateral resistance: A significant peak in lateral re-sistance occurs when the pipe first moves laterally (point 1 in Figure 17). The size of this peak depends on the level of initial embedment (Figure 18). Breakout resistance is criti-cal to the initiation of buckles and is also associated with lateral stability at route curves, when a pipeline can become unstable under tensile load.

Suction release: The breakout force includes the effect of suc-tion between the pipe and soil, which is released (point 2 in Figure 17) as a crack forms between the back of the pipe and the soil. Suction release has little influence on subsequent behaviour. For very slow loading, this crack may form prior to the full mobilisation of breakout resistance. In this case, the breakout load will be smaller and no suction release will be evident. However, the initial breakout of the pipeline is usually a fast event, and some level of tension is sustained at the rear of the pipe, leading to a two-way failure mechanism involving soil ahead of and behind the pipe.

Residual lateral resistance: Once the pipe has started to move, the residual lateral resistance drops or rises to a steady-state value at large displacements (point 3 in Figure 17) as the pipe pushes a berm of surface soil across the sea-bed. The residual resistance during the first loading cycle controls the lateral displacement at which the first buckle stabilises, defining the initial shape of the lateral buckle and the peak bending stress in the pipe.

Cyclic lateral resistance: As the pipe undergoes cycles of shutdown and restart during operation, the pipe cycles back and forth across the same patch of seabed. Surface soil, swept ahead of the pipe on each cycle (points 5 and 8 in Figure 17), builds up into berms at the extremes of the pipe displacement (points 6 and 9 in Figure 17). These berms offer significant resistance to pipe movement and define the shape of the buckle in operation. Suction release also oc-curs as the pipe pulls away from each berm (points 4 and 7 in Figure 17). The release usually occurs as a crack forms between the back of the pipe and the soil, but during steady cycling a lump of soft remoulded soil can adhere to the rear of the pipe, so a more gentle suction release occurs at some distance from the berm. Inclusion of the suction release behaviour in the pipeline analysis has little influence on resulting response. However, the mechanism of crumbling soil from the fixed berms falling back into the trench creates a bowl-shaped trench. The shape of this trench means that the cyclic residual resistance is initially low as the pipe slides down one side of the trench, and then gradually increases as the pipe slides horizontally and then rises up the other side of the trench before meeting the berm.

3.5.2 Lateral resistance uncertainties Many of the tests carried out to date have evaluated the lateral resistance of quite light pipes, representing wet-in-sulated or gas pipelines; uncertainty remains over the long term cyclic response of heavier pipelines, such as water-filled lines or pipe-in-pipe systems.

The key to successful modelling of the cyclic behaviour is to capture the reaction force from the static berms. Once

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Figure 16: Example of axial velocity along a pipeline during shutdown/restart cycle

Figure 17: Schematic of lateral force-displacement response

Figure 18: Breakout response based on embedment level

Figure 19: Typical cross profile at crown of lateral buckle on an operated pipeline

the static berms are formed on the first cycle, they quickly become established and provide a significant resistance to increasing cyclic amplitude of the pipe, although the pipe is likely to encroach into the berm slightly during the first few cycles, until the berm gains sufficient strength.

This behaviour is confirmed by the soil berms established on operating pipelines where lateral buckles have occurred. Figure 19 shows a typical cross profile at the crown of a lateral buckle, where the as-installed and maximum excur-sion positions are inferred from survey data. For compari-son, Figure 20 shows the pair of soil berms established by a laterally sweeping pipe section during a centrifuge model test. This test was part of a programme of lateral pipe-soil interaction tests conducted at The University of Western Australia (UWA) as part of SAFEBUCK Phase II9. By ana-lysing images captured during this test programme using particle image velocimetry (PIV)10, the soil deformation mechanisms during lateral sweeping could be identified. Figure 21a shows the displacement field as a pipe travels horizontally on soft clay. The growing berm ahead of the pipe primarily deforms by shear along the base, with some soil being pushed into the berm close to the pipe invert.

To illustrate the significance of the berm resistance, the berm resistance recorded during these tests after 50 cycles of fixed amplitude was typically five times higher than the residual lateral resistance during the first cycle.

Models for pipe-soil interaction that are used to simulate large-amplitude cyclic motion should therefore capture the effect of soil berms, which tend to constrain the buckle and can increase the loading at the apex of the buckle. The in-

Figure 20:Typical berm geometry during centrifuge modelling of cyclic pipe movement8

Figure 21: Berm growth ahead of a sweeping pipe; (a) velocity field during

centrifuge test and (b) mod-elling idealisation

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crease in lateral resistance at the soil berm is a function of:• The cumulative lateral pipe displacement, defined by the

number and size of previous sweeps, and the vertical em-bedment during these sweeps: these distances govern the volume of soil pushed out from the trench and added to the berms; the size of the berm then governs the resist-ance required to push through it

• The proportion of berm resistance mobilised: as the pipe moves into the berm, the resistance rises approximately linearly until the full berm strength is mobilised

• The rate of displacement into the berm, hence wheth-er the response tends towards undrained or drained behaviour

• The remoulded strength of the clay within the berm, de-fined by repeated cyclic loading and consolidation of the berm.

The amplitude of the cyclic sweeps contributes to how quickly the berm is established during the initial cycles. However, during sweeps of constant amplitude, the berm resistance approaches a limiting value, as after the first few sweeps, any fresh soil added to the berm on a given sweep falls back into the trenchc.

FEA models should include the berm resistance based on the actual vertical load, V, exerted by the pipe caused by the formation of soil berms at lateral buckles. This includes lat-eral buckles that occur on the seabed (on-bottom) or those that occur at sleepers or distributed buoyancy sections. In every case, soil berms provide lateral restraint and strongly influence the cyclic response of the pipe. Berm resistance is generally incorporated into FEA by modifying the equiva-lent lateral coefficient as the pipe approaches the berms, ac-cording to a user routine. In recent projects, the magnitude of the berm strength and the rate at which it is mobilised as the pipe approaches the berm has been estimated from model tests. Currently, a range of values of berm strength is employed in design to address the sensitivity to berm strength during cyclic loading.

Prediction of berm resistance is not well understood and much work is ongoing in this area. To provide a more sys-tematic basis for predicting berm resistance, a methodology that links the erosion of material from the trench base to the growth of the berm, and therefore the berm strength, has been proposed11.

This approach operates by considering conservation of vol-ume between the soil berms and the trench between the berms, which is being ‘ploughed’ by the sweeping pipe. The thickness of material ploughed by the pipe during each sweep, tplough (Figure 21b) is calculated from the pipe weight. From tplough, the growth of the ‘active’ berm ahead of the sweeping pipe can be calculated. The volume of ma-terial ploughed during each cycle is added to the fixed berm at the end of each sweep, and the total berm volume is used to calculate the berm resistance.

c The trench described here is one which is created by cyclic pipe dis-placement and is not the result of a pipeline trenching operation.

This model for berm behaviour mimics the mechanisms seen during experiments in a simple form. The model can be used to assist in the interpretation of physical tests and allow these to be generalised to arbitrary sequences of lateral movement, providing values of berm resistance to be used in FEA. A future development would be to incorporate this model directly into the FEA, allowing berm resistance to evolve automatically during the analysis.

3.6 Advances in pipe-soil modelling methodsThe current pipe-soil response models are characterised by a piecewise linear response defined by equivalent friction coefficients (Figure 17). These provide a practical way to capture the non-linear response demonstrated by pipe-soil interaction testing. However, this approach can be cumber-some, requiring the definition of a non-linear response for each pipeline configuration to be inserted into a user-sub-routine that can update the equivalent friction coefficients for each section of the pipeline, as contact loads and dis-placements change throughout the analysis. These routines must be incorporated into the FEA in such a way that re-versals of direction or non-constant pipe touchdown loads are handled competently.

An alternative to the equivalent friction coefficient ap-proach is to use yield envelopes in vertical and horizontal load space that bound the allowable combinations of load for a given pipe embedment. Yield envelopes are well es-tablished as an approach to describe the bearing capacity of shallow foundations under combined load. By invoking plasticity theory and an appropriate flow rule, yield enve-lopes can be used to create a model for the general load-dis-placement response of a pipe or foundation12, 13, 14, 15. These are known as ‘force-resultant’ plasticity models.

These models are similar to the work-hardening plasticity models used to describe the behaviour of individual ele-ments of soil – such as Cam clay – and are able to capture the brittle behaviour of over-penetrated pipes, but with a smooth response rather than the abrupt changes in stiffness given by the piece-wise linear simplification illustrated in Figure 17.

A force-resultant plasticity model for small displacement be-haviour of a pipe on carbonate sand is described by Zhang et al.13 To capture the behaviour needed for lateral buck-ling design, this type of model must be extended to include the large displacement effects described previously – for example, using the berm modelling approach described in Section 3.5.2, in which the active berm area is an additional hardening parameter.

The SAFEBUCK JIP is supporting ongoing research into force-resultant plasticity models for pipe-soil interaction. The aim is to develop a force-resultant model for pipe-soil interaction that captures the necessary effects for use in lat-eral buckling design. This research is currently focused on identifying the basic mechanisms involved – using model tests and theoretical analysis – then distilling this behaviour into appropriate extensions to existing force-resultant plas-ticity models.

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4. Soil Data and Model Testing4.1 Soil properties required for designThe pipe-soil force-displacement responses used in the de-sign of a flowline should cover the full range of anticipated shear strength and submerged unit weight at the site. The lower bound soil strength profile results in greater em-bedment, which leads to high lateral breakout resistance, whereas the upper bound strength profile results in shallow embedment and a reduced breakout load.

For lateral buckling and pipe walking, it is important to bound behaviour. Upper and lower bound values of soil resistance are both important. It is therefore essential that a detailed soil investigation is conducted, providing detailed measurements of soil strength, particularly in the upper 1m of soil. Uncertain soil conditions results in a large range between upper and lower bound behaviour, which will in-crease the design challenge and potential mitigation costs; in some cases, it may preclude the ability to demonstrate a robust design solution.

It is therefore important to define a range of soil parameters that includes realistic upper and lower bound values with depth. It is also important not to assume a single value or an unrealistically narrow band of response, as this could lead to an unsafe design.

Key soils data required for design includes:• Shear strength profiles measured in situ to about 1.0m

depth, preferably using a T-bar penetrometer• Sensitivity, preferably based on first-in/first-out and cy-

clic T-bar response• Submerged soil unit weight• Laboratory test data using disturbed soil samples and

specialist testing methods, particularly the assessments of pipe-soil interface friction at appropriate low stress levels.

Full-flow penetrometers, such as the T-bar, currently offer the best method of assessing the strength of soft near-sur-face sediments for pipeline design. A T-bar penetrometer comprises a cylindrical bar, typically 40mm in diameter and 250mm long, and may be viewed as a small scale seg-ment of pipe. Measurements of T-bar penetration resistance can be converted to undrained shear strength in a more straightforward manner than CPT tip resistance, and the higher projected area of the T-bar leads to reduced noise on the force measurements16.

To assess axial resistance, it is useful to conduct laboratory tests to assess friction angles and remoulded strengths at the very low effective stress levels relevant to pipelines. Specialist equipment has been designed to measure pipe-soil interface resistance at these low stress levels, including the tilt table device at the University of Texas at Austin17 and the Cam shear device used at the University of Cambridge, discussed further in the next section.

The scopes for specialist laboratory and in situ pipe-soil test-ing are being developed and refined on a project-by-project basis. A specialist pipe-soil testing rig – mounted on a sea-

bed frame – to measure axial and lateral pipe-soil resistance in situ is being developed by Fugro (SMARTPIPE). This programme is ongoing and no test data is currently in the public domain.

4.2 Laboratory model test programmesLateral buckling and pipeline walking behaviour is extreme-ly sensitive to pipe-soil interaction, and there is much uncer-tainty associated with models for pipe-soil interaction in soft clay. In addition, the basic phenomena involved are not fully understood, and detailed numerical modelling is not yet ca-pable of fully representing the response. This has led to the need for project-specific model test programmes. These are often undertaken at full-scale (or half-scale), which requires more time than project time scales normally allow.

The aim of the SAFEBUCK pipe-soil interaction test pro-gramme was, therefore, to provide generic guidance for future projects and improve current understanding. Data from these tests have been augmented by project-specific tests, donated by JIP participants, to create a database that now spans a wide range of pipe diameters, weights and soil conditions; this database is the source on which the SAFEBUCK models published previously5 are based.

However, much uncertainty remains and more testing is ongoing. Many recent projects have carried out small- and large-scale tests in specialist laboratories. These tests have concentrated on the low shear strength and low-permeabil-ity soft clays that dominate the deepwater regions of the world where on-bottom lateral buckling solutions are cur-rently being employed.

Large-scale model tests typically use 3–5m3 of soil col-lected from the field and re-consolidated in a large tank. The model pipe is placed into the soil bed and swept axi-ally and laterally to assess the response. A number of such project-specific tests have been carried out at Cambridge University (UK) and the Norwegian Geotechnical Institute (NGI)18 over recent years. The practical limitations of such tests include the time taken to reconsolidate the soil in the tank and the maximum pipe diameter of typically 12in. Nevertheless, such tests provide detailed measurements of pipe-soil interaction in near full-scale conditions.

Large-scale lateral tests have been supplemented by small-scale centrifuge tests, which were initiated by the SAFEBUCK JIP at Cambridge University19 and have since been carried out by a number of international research fa-cilities. A suite of centrifuge tests for the JIP was recently completed at UWA, including a set of tests that used im-age analysis9 to better understand the failure mechanisms. The advantage of conducting lateral pipe-soil model test-ing in a centrifuge and the reason that this type of testing was initiated by the JIP are that only a small quantity of soil is required and the testing programme is much shorter. Centrifuge model tests, using a suitably advanced actuation system, can also replicate the necessary sequences of pipe laying and lateral motion.

The most advanced centrifuge model testing facilities can replicate the complete load and displacement patterns im-

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posed in real conditions, including a simulation of lay ef-fects, changes in pipe weight due to hydrotesting and op-erating cycles, and any pre-determined sequence of lateral motion (or loading). In the centrifuge, the accelerated rates of consolidation allow long periods of pipeline operation to be simulated within a single continuous centrifuge flight. For example, a centrifuge test lasting 24hr at an accelera-tion of 30g using a 20mm-diameter model pipe can simu-late 3 years of operation of a 0.6m-diameter pipe. In-flight T-bar testing allows the model seabed to be characterised in the same way as the field case. Figure 22 shows a model of a 0.4m diameter pipeline suspended above a soft clay seabed prior to simulation of the lay process, followed by lateral testing.

Large-scale axial model pipe tests have been supplemented by small-scale axial friction tests using the Cam shear de-vice at Cambridge University and the Tilt-table device17 at University of Texas. Both methods provide measure-ments of pipe-soil interface resistance, as well as soil-soil shear resistance, at the low effective stress levels appropriate to pipelines (typically 2 to 10kPa). The Cam shear device measures friction at a range of velocities, while the tilt-table test method is ideally suited to repeatable measurements of drained friction angle.

4.3 Operating data from existing pipelinesObservations from operational pipelines provide a signifi-cant source of data that is contributing to the understand-ing of pipe-soil interaction. Lateral buckles have occurred on many large-diameter operating pipelines in shallow wa-ter, and much has been learned from integrity monitoring assessments of these systems. However, much of this experi-ence is in shallow water on sandy seabeds. Operating data is now coming available from field developments in deep water on soft clay, but currently this data is sparse.

In addition, the level of data available from integrity moni-toring varies significantly by operator and region. To extract reliable pipe-soil interaction data from operational behav-iour, it is important to have recorded all the parameters that

Figure 22: Pipe-soil model testing in the UWA geotechnical beam centrifuge

affect behaviour with sufficient accuracy20. For example, it is not helpful to measure buckle shape accurately if the con-current operating conditions at the buckle location are un-known, since these conditions define the load against which the buckle behaviour is being assessed.

There are three key areas of data collation from operating pipelines, which are outlined in this section.

Embedment: Comparison of actual as-laid embedment with predictions based on laboratory tests will help identify the level of embedment that can be attributed to dynamic load-ing at touchdown during installation. Comparison between predictions and actual embedments has helped to calibrate the dynamic embedment factor, as discussed earlier.

Lateral response: Comparison between observed lateral buckle shape and the shape predicted by FEA allows the analyst to estimate the resistance provided by the soil; such estimates fit well with existing model predictions. Observations of lateral buckle shape from one year to the next have also provided valuable data on the influence of soil berms, as discussed earlier.

Axial response: Comparison of measured and predicted end expansions and pipeline walking predictions over time will improve understanding of axial pipe-soil response. Measurement of displacements at J-lay collars along a pipeline has already provided data that supports higher levels of axial resistance than conventionally assumed in design – supporting the laboratory studies that have re-vealed high values of drained interface friction angle (see Section 3.4.2).

5. ConclusionsThis paper has demonstrated the significant influence that axial and lateral pipe-soil resistance has on several aspects of the design of pipelines laid on the seabed. The areas of influence include the effective force in the pipeline, lateral buckle initiation, lateral buckle loading, end expansion, pipeline walking and route curve stability.

The influence of pipe-soil resistance on these various issues is often conflicting. For example, high lateral resistance is beneficial for curve stability, but detrimental to the bending loads developed in a buckle; high axial resistance is benefi-cial for controlling feed-in to a buckle, but increases the susceptibility to buckling in the first place. Consequently, it is not possible adopt a ‘conservative’ pipe-soil resistance parameter based on the extreme anticipated value. Instead, it is necessary to bound the pipe-soil resistance and ensure that the design is acceptable throughout the potential enve-lope of resistance.

Since the pipe-soil response is the largest uncertainty faced in design, there is a need for better analysis models sup-ported by experimental data and underlain by a theoretical basis. Significant work has already been performed by the SAFEBUCK JIP and by forward thinking projects to im-prove understanding considerably.

However, the observed behaviour is extremely complex.

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Compared to conventional geotechnical problems, lateral buckling design is unusual in that soil failure is required in order for the design to be successful. Also, the operat-ing behaviour involves large amplitude pipe movements, leading to repeated remoulding and consolidation of the seabed and gross changes in geometry due to the formation of trenches and berms. All of this activity occurs at effective stress levels that lie around 1 order of magnitude lower than are conventionally considered in geotechnical engineering. This complexity presents a significant challenge in design, but significant project cost savings can result if an accept-able solution can be reached.

The research described in this paper has identified the key mechanisms that affect axial and lateral pipe-soil interaction on soft clay soils. Each aspect of this interaction includes areas of uncertainty identified in this paper. Current developments aimed at reducing these uncertainties include (i) improved penetration testing techniques for assessing near-surface soil strength; (ii) new laboratory testing techniques for assessing pipe-soil resistance at low stress levels; (iii) improved effec-tive stress-type (‘beta’) methods for assessing axial pipe-soil resistance; and (iv) improved models for assessing lateral re-sistance, incorporating berm effects and building on existing force-resultant plasticity models. SAFEBUCK Phase II has focussed on the latter three aspects. Research into the first aspect is described elsewhere16, 8.

It is anticipated that the current research will lead to an improved understanding of the key parameters and mecha-nisms involved in pipe-soil interaction, providing improved analysis models and raising the capability of the industry to assess lateral buckling and pipeline walking in design.

6. AbbreviationsCAPEX – Capital expenditureFEA – Finite element analysisJIP – Joint industry projectKP – Kilometre point (distance along a pipeline)OOS – Out of straightnessPIV – Particle image velocimetrySCR – Steel catenary riserTDP – Touchdown pointUWA – University of Western AustraliaVAS – Virtual anchor spacing

7. Nomenclaturesu = soil undrained shear strength (kPa)H = horizontal resistance, kN/m V = vertical-unit pipe load, kN/mW’ = submerged pipe weight, kN/mµa = axial friction coefficientα = total stress axial resistance coefficient, α = τ/suδ = pipe-soil interface friction angleφ = soil friction angleτ = shear stress

AcknowledgmentsThe JIP has been very well supported by the offshore in-dustry. BP, ConocoPhillips, ExxonMobil, Petrobras and Shell, as well as the US government through the Minerals Management Service (MMS), participated in Phase I, with installation contractors and suppliers represented by Allseas, JFE-Metal One, Technip, and Tenaris. Additional participants including Acergy, Chevron, Statoil and Saipem have joined Phase II, which will run through 2007.

Many engineers and specialists in the pipeline and geotech-nical community have contributed to the understanding and testing of pipe-soil interaction response outlined in this paper, often through involvement in the SAFEBUCK JIP. In particular, Professor Malcolm Bolton has led the SAFEBUCK pipe-soil interaction work conducted at Cambridge University.

The centrifuge model test shown in Figure 20 was conduct-ed by Helen Dingle.

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20. Baker JHA, Bruton DAS, Matheson IC. (2006). Monitoring and effective integrity management of laterally buckled flowlines in deep water. OTC 17932, Proc. Offshore Technology Conf., Houston, USA.

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