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    Frontiers in Offshore Geotechnics: ISFOG 2005 Gourvenec & Cassidy (eds) 2005 Taylor & Francis Group, London, ISBN 0 415 39063 X

    95

    1 INTRODUCTION

    Subsea pipelines are laid on the seabed and may ormay not be buried. Pipeline design needs to accountfor the ways in which the pipeline interacts with thesoil. This paper is about pipeline-soil interaction and, in

    particular, the geotechnical issues that must be facedby pipeline designers if soil-pipe interaction is to becaptured adequately in the design process. Pipelinegeotechnics is an emerging specialty that involvesapplications of geotechnical theory and practice uniqueto the construction of underwater pipelines.

    For this State-of-the-Art review, the material hasbeen organized broadly in the order in which the phe-nomena arise embedment and lateral friction mobil-isation during pipelay, stability against current andwave action when on the seabed, development of axialfriction as a result of thermal and pressure loading etc.Since embedment during pipeline affects axial and lat-

    eral friction this appears to be a logical progression.Some areas that are not covered include self-burial

    of pipelines as a result of wave/current action andsediment transport, liquefaction around pipelines, andthe whole area of ploughing and trenching (Cathie &Wintgens 2001) which would make the paper tooextensive.

    2 GEOTECHNICAL PARAMETERS

    If soil-pipeline interaction is to be understood andmodelled then it is clear that geotechnical data alongthe pipeline route is required. This basis point is notalways recognized and a group involved in the offshore

    geotechnical industry has put together a guidancedocument to assist non-specialists to understand the

    basic data required from a survey (OSIF 1999). Char-acterisation of the pipeline route with emphasis ontrenching issues is discussed in Cathie (2001).

    Pipeline design and engineering will require geo-technical data if major assumptions are to be avoided.For design, all soils require classification tests (par-ticle size distribution, index tests and basic shearstrength parameters). The in situ density of granularsoils and the undrained shear strength of cohesivesoils should be determined.

    If the soil is very soft it is not sufficient to collectdrop cores as the disturbance can lead to underestimat-ing the soil strength. This could affect axial and lateralfriction assessments, and trenching engineering issues.The box corer coupled with in situ vane testing (OSIF1999) is recommended for obtaining good quality datain very soft soils. A substantial quantity of soil can be

    collected using a box corer and this may be necessary ifriser-soil interaction needs to be investigated in specifictests, or if backfill properties need to be investigated.

    In granular material, the cone penetration test (CPT)remains the most attractive tool to determine the in situdensity (OSIF 1999). CPTs should always be com-

    bined with sampling techniques in order to determinethe physical properties of the sand, particularly thegrain size, which may be important for trenchability.

    3 PIPELINE EMBEDMENT

    Pipeline embedment begins during pipe lay and mayincrease with time due to hydrodynamic forces, pipe

    Pipeline geotechnics state-of-the-art

    D.N. Cathie & C. JaeckCathie Associates SA/NV, Brussels, Belgium

    J.-C. Ballard & J.-F. WintgensFugro Engineers SA/NV, Brussels, Belgium

    ABSTRACT: Pipeline geotechnics deals with soil-pipeline interaction. This covers installation issues (pipelinepenetration and short-term lateral stability), axial and lateral response to loads. It then encompasses pipelinetrenching, backfill engineering and pipeline stability when buried. This review provides an overview of all aspects

    of pipeline geotechnics except trenching. The focus of the paper has been on the mechanics of each problem,explaining the issues with a view to developing understanding, rather than providing ready made solutions. Theinterested reader can make use of the references for going deeper into particular aspects of the subject.

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    where z penetration, D pipeline diameter,su undrained shear strength, soil unit weight,Fz vertical load per unit length on soil.

    Verley & Sotberg (1992) propose equivalent equa-tions for penetration in sand but for cyclic loading.

    Murff et al. (1989) have presented upper and lowerbound plasticity solutions for rough and smooth pipes(full adhesion and no adhesion) in cohesive soils. Thelower bound results are shown on Figure 3 and com-

    pared with experimental data and the Verley & Lundapproaches. Upper bound solutions were only slightlyhigher.

    These solutions are of most importance in verysoft clays where embedment may be significant. Muchof the experimental data shown in Figure 3 is from

    Wagner et al. (1987) and is applicable to remouldedclay with an undrained shear strength around 1 kPa.Some of the scatter may arise from the difficulty ofmeasuring strength and its variation with depth.

    The plasticity solutions do not account for buoy-ancy effects, soil heave, or any increase in strengthwith depth. Murff et al. have shown the potential impactof heave and increase in strength. The effect of buoy-ancy would typically be about 510% depending onthe soil strength but would be more important in fluidmuds. Soil heave is potentially even more important.Soil heave around the pipe increases the bearing area

    by about 20% at z/D 0.2 and increases the contactarea at the same depth by 35%. At a penetration ofabout z/D 0.25 almost the full diameter of the pipeis bearing on the soil when heave is accounted for. Itis difficult to see how laboratory and field data can be

    properly interpreted without taking account of the

    geometric changes that occur with penetration. Thereis scope for further theoretical work in this respect.

    Since pipe penetration involves remoulding thesoil locally to the pipe wall, and since repeated load-ing due to hydrodynamic effects would only accentu-ate this effect, it seems reasonable to assume that

    penetration assessment should consider a zone ofremoulded soil below the pipeline during laying. Thiswould suggest that using a low soil-pipe adhesion forembedment assessment would be appropriate.

    In the view of the authors, the issue of initial pene-tration resistance is relatively well understood and

    established. Focus should now be on the nature andmagnitude of the loading applied by a pipeline duringlaying, and the possible magnitude of the cyclic effects.

    3.3 Penetration due to repeated loads

    As discussed above, as a pipe is laid on the seabed themovement of the pipe during the laydown processwill increase the penetration.

    Cyclic vertical loading has been investigated byDunlap et al. (1990), Fontaine et al. (2004) and in the

    STRIDE and CARISIMA joint industry projectsaimed at understanding riser-soil interaction (Bridgeet al. 2004). The latter authors provide a clear explan-ation of repeated penetration and pullout, particularlyfocusing on the suction that develops during pulloutin soft cohesive soils. Dunlap et al. indicate thatlimited cyclic loading without break out resulted inlittle additional burial while larger cyclic loads which

    broke suction and pulled the pipe free resulted in fur-ther penetration. This may be due to the additionalremoulding experienced in the full breakout case.

    Repeated lateral loading induced by current load-ing and dynamic response of the suspended section of

    pipeline can result in a considerable increase in pene-tration (Morris et al. 1988), particularly if the pipe is

    97

    heave

    z

    a)

    b)

    Figure 2. Pipeline embedment (a) initial penetration, (b)lateral movement.

    0

    1

    2

    3

    4

    5

    6

    0.0 0.2 0.4 0.6 0.8 1.0

    Normalised penetration (z/r)

    Normalisedresistance(F/SuD)

    Murff et al - rough Murff et al - smooth

    Verley and LundVerley linear

    Data (Murff et al)

    Figure 3. Lower bound plasticity solutions and empiricalapproaches for pipeline embedment for cohesive soil.

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    not overpenetrated. Consider a pipe penetrated with acertain vertical load and let us assume that the load iskept constant. Any horizontal load will then result in

    both vertical and horizontal movements since the pipeis on the failure envelope. This is discussed further inthe next section.

    Morris et al. (1988) carried out laboratory tests toassess the embedment due to horizontal cyclic load-ing in very soft clay. The additional penetration waslargely dependant on the magnitude of the force, ordisplacement, and the duration during for which itwas applied. Although the rate of penetration wasfound to decrease over the duration of each test, thisrate did not appear to reach a constant value or the

    pipeline to reach a limiting burial depth. As the pipecontinued to sink, the mounds grew in size, attainingheights of up to 0.8D. Morris et al. also consideredthe accumulated burial effect of variable cyclic load-

    ing and found that the effect could be modelled withsets of uniform cycles.

    Verley & Lund (1995) proposed the empiricalequation

    (2)

    where a is the amplitude of horizontal movement, toassess the maximum penetration that can be achievedfor a given amplitude of motion and to assess the

    development of the penetration as a function of thework done by the pipe on the soil. The authors suggestthat the applicability should be limited to z/D of 0.3due to range investigated in their study.

    Lund (2000) performed an investigation on the pene-tration depth achieved for a large diameter pipeline(1.2 m OD). While the calculated embedment wasabout 0.05m, the actual embedment varied between0.250.4m. Much of the route was sand. Lund con-cluded that much of the additional embedment wasdue to lateral oscillation at the touchdown point.

    4 LATERAL AND AXIAL RESISTANCE OFPARTIALLY EMBEDDED PIPELINES

    Lateral and axial resistance of an unburied pipelineneeds to be assessed at the design stage. Lateral resist-ance is important for the design of the pipeline on-

    bottom weight and any weight coating that may berequired. Axial resistance controls pipeline expansionand affects end connections, spool pieces, and upheaval

    buckling. Since environmental conditions (wave/cur-rent) leading to lateral forces and thermal/pressureeffects leading to axial forces are variable and cyclic innature, the subject must be considered for both mono-tonic and cyclic loading. The pipeline outer coating

    properties also take on more importance for lateraland axial resistance.

    4.1 Lateral resistance

    Three different approaches have evolved for assess-ing lateral resistance:

    1 a single friction factor approach where the lat-eral resistance is related to the submerged weightof the pipeline and the soil type;

    2 a two component model consisting of a slidingresistance component and a lateral passive pressurecomponent frictional model supplemented with

    passive resistance of the wedge of soil (Nyman1984, Wagner et al. 1987, Lieng et al. 1988, Verley &Sotberg 1992, Verley & Lund 1995);

    3 a plasticity model approach (Zhang et al. 1999,

    2002, Cassidy 2004).Generally, the two component models are based on

    empirically fitting laboratory test data. A summary ofsome of the proposed equations is given in Table 1.

    While being practically useful, these models do notenlighten the user with the actual mechanics of the prob-lem. Moreover, they become more and more empiricalwhen cyclic behaviour needs to be introduced.

    The plasticity framework outlined by Zhang et al.(1999, 2002) provides a much more fundamental wayof understanding the mechanisms involved as well as

    being much more general. Zhang et al. have developed

    the plasticity model for calcareous sands but theapplication to clays should be straightforward and isin progress by the authors. The concepts are now welldeveloped for surface footings. As described byCassidy (2004) the models contain a yield surface, astrain-hardening expression, elastic behaviour insidethe yield surface, and a flow rule to define the direc-tion of movement during yield.

    Considering pipe behaviour under combined verti-cal (V) and horizontal (H) loads in calcareous sand incentrifuge tests, Zhang et al. have confirmed whatwas known for surface footings that the shape of the

    yield surface is almost parabolic. A simplified repre-sentation was proposed, given by

    (3)

    where is a parameter associated with the frictionbetween pipe and soil for low vertical load, Vmax andVmin are the positive and negative intercepts (H 0)on the vertical load axis, and

    (4)

    where is determined from calibration tests. Thisexpression gives an intercept on the horizontal load

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    axis when vertical load is zero, and represents a pas-sive soil resistance with a magnitude Vmin. Vmax rep-resents the maximum vertical load for a given

    penetration (the preload). Figure 4 shows the nor-malised form of the yield surface for different valuesof . The peak horizontal resistance is achieved atabout 40% of the maximum vertical load.

    The proposed hardening function is based on themonotonic vertical penetration resistance of the pipe(plastic stiffness) and the rebound response gives theelastic stiffness. This enables the increment of vertical

    plastic strain to be defined in terms of the elastic andplastic stiffnesses and the increment in Vmax.

    The plastic potential takes a different but similar formto the yield surface, and was defined by considering thedisplacement increment vectors in different tests:

    (5)

    where t is the shape parameter. The exponent m hasthe effect of adjusting the value of the vertical load atwhich the normal to the plastic potential becomes

    parallel to the H axis. Figure 5 shows the plasticpotential for different values of m. The model predictsupwards pipe movement and strain softening when

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    Table 1. Examples of lateral resistance models.

    Reference Equations Details

    Wagner et al. (1987) Sands: 8.6kN/m3

    Fy (W FL) A Monotonic

    0.6 38

    Fy horizontal resistance 9.6kN/m3

    A 0.5xEmbedded area MonotonicW submerged pipe wt 0.6FL hydrodynamic lift 79

    Cyclic load( static failure)Embedment x 2

    reduced by 50%

    Cyclic load( 5% D)Embedment x 3

    reduced by 8090%

    Clays: MonotonicFy (W FL) Su A/D 0.2

    39.7

    Cyclic( static failure)Embedment x 2

    31.7

    All clayscyclic load( static failure)Embed x 2.5

    15.7

    Lieng et al. (1988) Fy (W FL) FR 0.6(sands) 0.2(clays)FR calculated consideringaccumulated energy

    Verley & Sotberg (1992) Fy Fc FR All sandsFy (W FL) FR 0.6FR D

    2 (4.5 0.11D2/Fc) (z/D)1.25

    Verley & Lund (1995) Fy Fc FR ClaysFy (W FL) FR (Su 70kPa)

    0.2FR 4.13 DSu(Su/(D))

    0.392 (z/D)1.31

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    V/Vmax is between 0 and 0.25 depending on theexponent m.

    For the calcareous sand tested by Zhang et al.Figure 5 makes it clear that unless the pipe is highlyoverpenetrated (i.e. V/Vmax is very low) further pene-

    tration of the pipe can be expected under horizontalloading.The model has been calibrated for the specific cal-

    careous sand. However, it is of interest to see how themodel compares with other methods for calculatinglateral resistance. Care must be taken to correctlyinterpret laboratory tests in terms of whether they aresideswipe tests (horizontal displacement control atconstant vertical displacement) or probe tests (hori-zontal displacement at constant vertical load).

    Figure 6 shows experimental data for monotonictests with an overpenetration ratio of 1 (current verti-cal load is the maximum experienced). Normalisationwith the penetration depth is logical since vertical

    penetration is a key variable but the plot tends to put

    the data scatter in a good light. Presenting data interms of D (Fig. 7) shows the wide range of data.Both plots demonstrate that the Zhang et al. plasticitymodel is a reasonable representation of all the datadespite the fact that it has not been calibrated forother sands.

    100

    -0.2

    0

    0.2

    0.4

    0.6

    0.8

    1

    0.0 0.1 0.2 0.3 0.4 0.5 0.6

    H/Vmax

    V/Vm

    ax

    =0

    =0.1

    =0.2

    =0.7

    Figure 4. Yield envelope for sand (Zhang et al. 1999).

    -0.2

    0

    0.2

    0.4

    0.6

    0.8

    10.0 0.1 0.2 0.3 0.4 0.5 0.6

    H/Vmax

    V/Vmax

    m=0.1

    m=0.2

    m=0.3

    m=0.4

    =0.05

    Figure 5. Plastic potential for sand (Zhang et al. 1999).

    0

    0.1

    0.2

    0.3

    0.4

    0.5

    0 50 100 150 200

    Normalised lateral resistance H/ 'z2

    Embedment(z/D)

    Plasticity model

    Lieng et al (1988)

    Brenodden et al, 1989

    Verley and Sotberg (1994)

    Verley & Sotberg (1994)

    Palmer et al, 1988

    Wagner et al, 1987

    Zhang et al (2001)

    Figure 6. Lateral resistance models for sands (monotonictests).

    00 1 2 3

    0.1

    0.2

    0.3

    0.4

    0.5

    4

    Normalised lateral resistance H/ 'D2

    Embedment(z/D)

    Plasticity model

    Lieng et al (1988)

    Brenodden et al, 1989Verley and Sotberg (1994)

    Verley & Sotberg (1994)

    Palmer et al, 1988

    Wagner et al, 1987Zhang et al (2001)

    Figure 7. Lateral resistance model for sands normalisedwith pipe diameter.

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    In order to relate the models back to the traditionalfriction factor, the horizontal resistance has been nor-malised with the applied vertical load to give thefriction factor (Fig. 8). The low prediction of the

    plasticity model at greater depth reflects the Zhangdata but not the other data.

    The plasticity model is likely to require develop-ment as the model described above is a single surfacestrain hardening model. Other approaches are likely to

    be required for modeling cyclic loading. In particular,geometric changes will need to be considered wherelarge lateral movements of several diameters occur.

    4.2 Axial resistance

    Axial loads normally apply some time after installation.The implications are different for sands and clays. Insands, wave and current action may have inducedminor lateral loading which has in turn induced some

    further embedment or densification of the soil. Therecould be some build up of sediment against the pipe.In clays, set-up following remoulding will have hadtime to develop. This could include components ofstrength increase arising from thixotropy and consoli-dation. It is also likely that good adhesion between the

    pipe and the soil will have developed in soft clays.Although the axial resistance is influenced by

    embedment and time dependent factors such as thosedescribed above, simple Coulomb friction models areoften adopted to evaluate the axial resistance of par-tially embedded pipelines in all soils, given by

    (6)

    Strictly, this is only valid for drained conditions,but this may be a reasonable assumption for bothsands and clays if the loading rate is slow enough. For

    thermal expansion, the temperature increase is likelyto take several hours and this could be taken as justi-fication for a drained analysis.

    Axial friction assessment then reduces to evaluat-ing the friction coefficient . depends on the internalfriction angle of the soil and on the properties of thesoil-pipeline interface. There are various guidelinesused in the offshore industry for both pipelines and

    piles:

    tan( 5) (API RP2A WSD 2000) 2/3 tan() (Bureau Vritas) frtan() (Finch et al. 2000)

    The first two formulations assume that the inter-face is soil-steel. For pipelines this is rarely the casesince the outer coating is generally a corrosion pro-tection such as polypropylene (PP) or a concreteweight coating. PP coatings may be smooth but spe-cial materials can be ribbed to improve friction fortransport and handling. Based on research performedon a range of coatings (Finch 1999), Finch et al.(2000) recommend values of frshown in Table 2 as afunction of coating roughness and soil grain size:

    For fine-grained sediments, and where loading

    may be rapid enough to elicit an undrained responsefrom the soil, the axial resistance would be a functionof the contact area, and of the undrained shearstrength of the soil, expressed as

    (7)

    where adhesion factor and L arc length inembedded soil (including heave).

    Appropriate values of the shear strength and adhe-sion factor will depend on whether the peak or residualaxial resistance is required, and how long the pipelinehas been installed without load. Laboratory sheartests are recommended for the specific soil and coat-ing under consideration. In very soft clays, in the

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    0

    0.1

    0.2

    0.3

    0.4

    0.5

    0.4 0.6 0.8 1 1.2 1.4 1.6 1.8

    Friction Factor

    Embedment(z/D)

    Plasticity model

    Lieng et al (1988)

    Brenodden et al, 1989

    Verley and Sotberg (1994)

    Verley & Sotberg (1994)

    Palmer et al, 1988

    Wagner et al, 1987

    Zhang et al (2001)

    Figure 8. Lateral resistance models for sands in terms offriction factor.

    Table 2. Resistance factor for pipeline axial friction coef-ficient under fully drained conditions (after Finch et al.2000).

    Condition f r

    Granular cohesionless soil and f r 1D50 pipe roughnessGranular cohesionless soil and 0.75 fr 0.9D50 pipe roughnessFine-grained cohesive soil and f r 1D50 pipe roughnessClay and D50 pipe roughness fr 0.6Silt and D50 pipe roughness fr 0.4

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    absence of specific data, and subject to the roughnessof the interface, the adhesion factor may be taken as 1for the peak resistance and related to the soil sensitiv-ity, St, for the residual strength ( 1/St ).

    The research programmes discussed by Finch(1999) have shown that the peak resistance in sands istypically achieved within an axial displacement of2 mm. This value is consistent with the 0.1 (2.54mm)generally considered for the axial friction mobiliza-tion of tz curves of driven piles (API RP2A WSD2000). A linear elastic-perfectly plastic axial resistancemobilization curve (with a peak resistance reached ataround 2mm) is therefore a reasonable assumption ingranular soils.

    4.3 Axial creep or walking

    Observations and analysis have shown that pipelinescan walk or creep axially (Tornes et al. 2000, Carr et al.2003) due to internal heating and cooling. The drivingmechanism is the expansion and contraction of the

    pipeline and whether there is an effective anchorpoint where no movement occurs. The rate of creepwill depend not only on the temperature profiles butalso on the magnitude of the axial resistance, themobilization distance and the degradation to residualconditions.

    5 PROPERTIES OF TRENCH BACKFILLS

    The properties of a trench backfill are necessarily afunction of how the backfill is placed. Four broad cat-egories can be considered:

    natural infill (wave/current induced); mechanical backfilling; backfill following jetting; active jet cutting and collapse.

    5.1 Natural infill

    Wave or current induced natural infill is applicablemainly in sands and in relatively shallow water whereseabed currents are sufficient to induce transport.Soil particles are deposited under a relatively highenergy environment which results in a structure that istypically loose to medium dense. Rates of trenchinfill can be estimated using methods such as Schapp(1982) and Niedoroda & Palmer (1986), or bydirectly modeling the flow regime accounting forspoil heaps using computational fluid dynamics simu-lations. Other relevant information is given in VanRijn (1993) and Fredsoe (1978). Note that naturaldensification can also occur with time as a result ofwave action (Clukey et al. 1989).

    5.2 Mechanical backfilling

    Mechanical backfilling involves scraping the spoil

    (previously removed from the trench) back into thetrench. It is applicable to all types of soil conditions.The backfilling process is discussed in detail in Cathieet al. (1998). Soil in the spoil heaps, and sometimessome of the in situ seabed soil, is mixed and depositedinto the trench very rapidly. Water is believed to beentrained with the backfill and the resulting mass isexpected to have a higher macro water content than inthe spoil heaps, particularly if the soil contains a cohe-sive component.

    Mechanical backfills are not unlike hydraulic fillsand a starting point for considering the properties of

    backfills is to use this work. Whitman (1970) pro-posed a classification system, which provides a start-ing point.

    There is no specific published data on the proper-ties of mechanically backfilled trenches as far as theauthors are aware.

    For sand backfills, some information about in-situdensities after hydraulic filling and slumping is pro-vided by Stoutjesdijk et al. (1998). They consideredthe formation of submarine slopes with hydraulicsand fill and found that very low densities coulddevelop during hydraulic filling (1030%) but that

    liquefaction and flow could increase the relative dens-ity (to 2050%). The rate of f illing was apparently notimportant (Bezuijen & Mastbergen 1988). It seemsreasonable to assume that similar considerations applyto mechanical backfilling of sands. Therefore, sand

    backfills are expected to be in a loose state after back-filling. It is not difficult to demonstrate that drainageof a 1.5m trench should be largely complete in a fewminutes for most sands.

    Clay backfills or mixed sand/clay materials arebelieved to be heterogeneous after backfilling. Stiffclay is ploughed out of the trench into the spoil heapsand then left exposed to free water until backfillingtakes place. The surfaces of the lumps will take in waterand soften. During backfilling, further disturbance

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    Table 3. Classification of hydraulic fills (Whitman 1970).

    Nature of original material Characteristics of fill

    Fairly clean sand Reasonably uniform fill(15% passing No.200 sieve) of moderate density

    Silty or clayey sand Very heterogeneous fillof large void ratioStiff cohesive soil Skeleton of clay balls,

    with matrix of sand andclay

    Soft cohesive soil Laminated normallyconsolidated orunderconsolidated clay

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    and deformation of lumps will occur and the voidsbetween lumps will be filled with water, slurry andmore mobile sand components. Further details anddiscussion can be found in Karthikeyan et al. (2001,2002), Hartlen & Ingers (1981) and Mendoza &Hartlen (1985). Lumpy clay backfills consisting ofstiff clay will generally consolidate much morequickly than would be expected for homogeneous in-situ soil due to the voids and channels available forfree water. A model for consolidation of lumpy clayfills has been proposed by Yang et al. (2002).

    Soft clay backfills created by mechanical trench-ing and backfilling are also believed to be heteroge-neous and consist of softened and remoulded materialclose to the in-situ water content in a slurry of muchhigher water content soil. Their properties are stronglytime dependent as consolidation and thixotropic regaintakes place. Bruton et al. (1998) studied the consoli-

    dation of clay slurry in the laboratory and showed thatcavities and channels formed within the mud as it set-tled. Such channels tend to form even in a homoge-neous slurry as it sediments but are further encouragedif it is heterogeneous, if there is sufficient adhesion inthe clay to permit cracks to remain open, and if thereare silt or sand inclusions. This heterogeneity wasfound to result in much more rapid consolidation thanwould have been measured in an oedometer test onhomogeneous soil. Mechanical backfilling probablydoes not destructure the soil as completely as jettrenching and therefore the quality of the backfill soil

    should also be better. This is an area of active work atthe present time and much is as yet unpublished.

    5.3 Backfilling following jetting

    Lowering or burial of a pipeline by jetting results indestruction of the in situ soil, local movement of thesoil behind the trencher and deposition of the soil par-ticles or lumps back into the trench or just outside.While clays are cut and broken before being trans-

    ported, sands and silts are eroded and transported insuspension.

    Sand backfills. There is almost no data related toactual soil densities in a real backfill post-jetting.However, Kvalstad (1999) does indicate that low coneresistances have been measured even some time aftertrenching indicating the sand to be loose to veryloose. These loose sands had apparently been stable fora long period. We have also seen very loose sand off-shore that have cone resistances equivalent to soft clays.

    As discussed for mechanical backfilling, themechanisms described in Bezuijen and Masterbergen(1988) for example for hydraulic filling are likely to

    be relevant. However, the mechanism of sedimenta-tion of sand in the trench following trenching is quiterapid but may occur in a low energy environment.Rapid sedimentation of sand after trenching is quite

    similar to the method of water pluviation to achieve aminimum density soil (Kolbuszewski 1948) and it isknown that slower rates of deposition/sedimentationlead to higher densities. Resedimentation tests byKvalstad on fine sand have shown that relative dens-ities in the range 1020% have developed. The dens-ity of the backfill is likely to be affected by:

    rate of trenching and sedimentation, use of back-flow jets (rate of deposition and water flow throughthe sand affects final density (Kolbuszewski1948);

    particle angularity, grain size distribution (uniform-ity) and grain size, mineralogy (Youd 1973, Hightet al. 1999);

    depth of backfill (density of very loose sand willprobably increase with stress level due to the highcontractive potential at very loose states);

    time, Pipeline movements, wave action, and ageingwill all tend to cause the loose sand to densifywith time.

    Some experiences with uplift of pipelines in veryloose backfills appears to confirm that the sand can

    be very loose and susceptible to structural collapseand static liquefaction. Laboratory testing on soils atvery low densities requires special techniques (Lade1992). Logically, samples prepared by water pluvia-tion would appear to have the best potential for repre-senting the fabric of a post-jetted sand.

    Clays. A jetted backfill consisting of clay is

    believed to consist of lumps of semi-intact material ina matrix of unconsolidated slurry. While not unlikethe mechanical backfill of the same material, thewater content is probably higher initially and theremaining intact soil reduced to smaller more dis-turbed lumps. Again, very little field data has been

    published but Newson et al. (2004) have demon-strated by in situ CPT and T-bar tests in a high plas-ticity clay in the Nile delta that about 50% of the insitu strength was regained within 3 months and littleadditional strength gain occurred over the next5 months. This appears to confirm the assumption of

    heterogeneous conditions postulated for mechanicalbackfilling which allows drainage and consolidationto occur relatively quickly. Unpublished laboratorytesting on soft blocky clays does confirm the samehypothesis but no testing that really simulates the jetcutting and deposition process has been attempted todate. One important conclusion however from thistesting is that shear strength of blocky soil for upliftresistance may be characterised by a frictional modelas used for sands (Bruton et al. 1998, Bolton &Barefoot 1997) although this should be confirmed bylaboratory testing on project specific soils.

    Finally, the strength regain of a soft clay backfill ismade up of both reconsolidation and thixotropicregain. Laboratory testing is beginning to be used to

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    characterise both thixotropic regain (without changein water content) and consolidation strength regain.Reference is made to Burland (1990), Leroueil &Hight (2003), Leroueil (2003), Locat et al. (2003),Schmertmann (1991), Sills (1995), Silva (1974),Skempton & Northey (1952) and Skempton (1970) forfurther details related to the behaviour and strength ofsoft clay backfills.

    6 PIPELINE STABILITY DURINGTRENCHING AND BACKFILLING

    6.1 Ploughing

    Pipeline stability is generally not an issue duringtrenching by ploughing since the pipe is lowered intothe cut trench as the plough moves forward. However,

    backfill ploughing can lead to pipeline uplift in cer-tain conditions (Cathie et al. 1996, 1998). A pipelinelying in a V-shaped trench may be susceptible to upliftcaused by:

    transverse flow of the soil down the slopes and hightransient hydraulic pressure below the pipe;

    turbulence and in-line water flow driven by thebackfill plough;

    pipe spans and out-of-straightness; low pipe weight; soil liquefaction and slow drainage.

    Pipeline uplift is known to have occurred inpipelines with a specific gravity of between 1.2 and 1.6.As suggested by Cathie et al. (1998), there are usuallyseveral factors that combine to cause uplift during

    backfilling. A backfill plough creates considerable tur-bulence and the effect of the plough (particularly themould boards) and the soil mass advancing rapidlycan cause uplift of a lightweight pipeline in a trench.

    Flow of soil down the slopes of the trench is animportant cause of uplift since high transient upliftforces are developed when the soil impacts the

    pipeline (Powell et al. 2002). Both transient hydraulic

    pressures and possibly wedging effects if the soil hasa shearing resistance would act to destabilize the pipe.The kinetic energy of the soil flow is likely to begreater if the trench is deep and the backfilling israpid. High backfilling speeds, even if short lived,may create conditions in with the uplift initiates. Afterinitiation of an uplift feature, the gap between thetrench base and the pipe is ahead of the plough. Thisin turn permits soil to flow under the pipe and propa-gate the feature in a progressive manner.

    According to Powell et al. (2002), and based onboth model testing and experience, uplift does notoccur if the pipe weight is sufficient. They suggestthat the minimum specific gravity for a mechanic-cally backfilled pipeline should be 1.8. Pipelines of

    lower density may need to be held in position, forexample, by stitch rockdumping.

    6.2 Trenching

    Pipeline lowering and burial by jet trenching requirescutting, erosion and fluidization of the soil by the jetsand lowering of the pipeline into the area cut or flu-idized by the jets. Clays are cut and broken by the jetsand transported behind the trencher. Sometimes educ-tors are used to clean out the trench since the claywalls remain stable for some time. In the view of theauthor, the problem of pipeline lowering and stabilityis mainly dependent on the amount of slurry or lumpsleft in the base of the trench. However, Powell et al.(2002) suggest that the product specific gravityshould be greater than the specific gravity of a lique-fied clay at the onset of effective stress. This may not

    be a useful criterion since it is typically around 1.2 formost muds.

    For sands, the trench walls are not stable and thepipe lowering is dependent on the pipeline beingheavier than the fluidized soil. The sand minimumdensity could be used as a guide to the minimum pipeweight but the author is unaware of work specificallyaimed at defining what this minimum pipe weightwould be. Nevertheless, it is clear that heavier, moreflexible, pipes will be easier to lower than light stiff

    pipes. Flexure of the pipe below the trencher coupledwith maximizing the length of fluidised soil will

    result in the most efficient lowering process.To minimize the risk of floatation in sands (or to

    maximize the lowering efficiency), Powell et al.(2002) suggest that for rapid draining soil a specificgravity of 1.5 is sufficient but that if silt or clay arealso present (lowering the permeability) then the riskof floatation is significant if the product specificgravity is less than 1.7.

    The subject of pipeline stability during trenchingand backfilling is still wide open for further insightsand research. We caution against focusing too muchon the 2D aspects of the problem (cross section) and

    only on soil mechanics. Uplift initiation and propaga-tion during trenching/backfilling is a 3 D problem andinvolves pipeline structural response to bending andhydrodynamic loads, soil transport and deposition,and consolidation. These aspects should be investi-gated together.

    7 AXIAL AND TRANSVERSE RESPONSEOF BURIED PIPELINES

    Axial and transverse response of a buried pipeline isimportant for assessing the behaviour of the pipelineto hydrotest or operational load conditions. Axialloads induced by internal pressures or temperatures

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    render a pipeline susceptible to transverse loads as aresult of its out-of-straightness i.e. imperfections.Generally the most serious imperfections are in thevertical plane and make the pipeline susceptible toupheaval buckling. However, horizontal or inclinedimperfections can also occur due to laying around a

    bend or to the pipeline lying on one side of a slopingtrench.

    Therefore, the pipeline response to both axial andlateral loads is important for assessing buckling

    potential. An early case history presented by Nielsenet al. (1990) focused industry attention on the prob-lem and Pedersen & Jensen (1988) and Nielsen et al.(1988) suggested solution methods. Since then a lotof experimental work has been performed to charac-terize uplift resistance in various soils. We provide asummary in Section 7.2.

    Transverse forces in a pipeline are a result of the

    internal compressive forces which are unable to bereleased due to the axial restraint of the soil sur-rounding the pipe. A model for the axial response istherefore necessary and this is discussed first.

    A buried pipeline may be partially in contact withrelatively undisturbed soil as well as being sur-rounded by backfill. Axial response is reasonablyassessed assuming the pipeline is surrounded by

    backfill but considering the effect of some contactwith firmer/denser soil as part of a parametric analy-sis. Lateral response of a buried pipeline should con-sider in which direction the lateral movement is

    taking places and select soil properties accordingly.For both axial and lateral resistance assessment,

    consideration should be given to whether the backfillwill be in a drained or undrained condition.

    7.1 Axial behaviour

    7.1.1 Ultimate axial resistanceThe ultimate axial resistance of a buried pipeline infreely draining soil may be determined by consider-ing the mean normal (lateral) pressure, on the

    pipeline and the axial friction coefficient, fr (see

    Section 4).Considering the normal stresses on the top, bottomand sides of an equivalent square leads to (Schamineeet al. 1990, Finch et al. 2000):

    (8)

    so that the axial resistance of a buried pipe is:

    (9)

    For cohesive soils, a decision has to be maderegarding drained or undrained behaviour. It is com-

    mon practice to assume that operational loadsdevelop relatively slowly (e.g. over a period of hoursfor temperature increases) and that both sands andclays can be treated as drained for axial loading(Finch et al. 2000). However, this may not necessarily

    be the case and it is prudent to consider both drainedand undrained response in the design analyses unlessit can be demonstrated otherwise.

    For undrained conditions, the undrained shearstrength is the basic parameter, with the resistancegiven by

    (10)

    Finch (1999) suggests that for clays with low shearstrength, values of should be 1.0 for peak resistanceand about 1/St where St is the sensitivity of the soil fora residual strength.

    Using axial resistance based on shear strengthimplies a uniform strength along the length in areas ofuniform backfill. This is not the reality. Seabed fea-tures and out-of-straightness have the effect of increas-ing the apparent frictional resistance due to high

    pressures on supporting areas. Finch recommends toconsider as 1.0 but use of equation 10 can result invalues of greater than 1.

    7.1.2 Axial responseThe response of buried pipelines to axial loads is gen-erally assumed to be similar to partially buried

    pipelines as discussed in Section 4. It is difficult toescape the analogy with axial loading of piles andsimilar approaches could be considered (e.g. Kraft et al.1981, Randolph & Wroth 1978). Differences arise

    because of the infinite length of the pipeline andthe non-axisymmetric nature of the problem. Never-theless, it can be reliably speculated that the soil shearmodulus would fundamentally govern the initial axialresponse. Since the modulus is likely to be related tothe stress level, the axial stiffness of the materialaround the pipe would vary, particularly for shallow

    burial. In practice, the traditional elastic-perfectly

    plastic model is generally used in finite elementanalyses of buried pipes using a mobilization distanceof 23 mm to mobilize ultimate resistance, values thatare justified by some experimental work.

    7.2 Transverse behaviour

    Since the most critical form of transverse response isuplift, this is addressed first. A brief treatment oftransverse resistance in other directions is given later.

    7.2.1 Uplift resistance in uniform soilUplift resistance models were developed initially forthe pull-out capacity of anchors. Rowe & Davis (1982)made theoretical and experimental investigations and

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    provide a good summary of earlier work coveringsands and clays. Subsequently, Murray & Geddes(1987), Dickin (1988) and others developed the under-

    standing mostly concentrating on uplift of anchorplates in sand. Recently, Merifield et al. (2001) haveprovided a rigorous theoretical solution for theundrained anchor problem.

    Work on buried pipes includes Matyas & Davis(1983), Trautmann et al. (1985) and others. As theoffshore industry recognized the upheaval bucklingrisk of buried pipelines, Schamine et al. (1990) pub-lished a fundamental set of experimental data applic-able to subsea pipelines. Further experimental datahas been collected (Ng & Springman 1994, White et al.2001) but much is proprietary and only partially pub-

    lished (Finch 1999, Finch et al. 2000, Fisher et al.2002).

    Uplift resistance models for sand have been devel-oped from the work on anchors in which clearlydefined slip planes are apparent up to relatively largedepths and the focus was on peak uplift resistance. Asshown by White et al. (op cit), a peak resistance may

    be accompanied by an upward sliding block mech-anism (Fig. 9) but as a gap is formed below the pipe the

    boundary conditions of the problem change and themechanism changes to a circulation or flow aroundmechanism. Pedersen & Michelsen (1988) sketched

    the same concept but without elaboration andSchamine et al. (1990) identified it from their testingprogramme.

    In very loose sands which contract when sheared,Vanden Berghe et al. (2005) have demonstrated thiscirculation mechanism numerically (Fig. 10).

    Compression of the very loose sand and upwardmovement makes room for the flow mechanism todevelop. The dilatancy of the soil determines at whichH/D ratio the flow mechanism governs the peakresistance as well as the residual.

    The different mechanisms of failure go some wayto explain why the wedge failure model cannot capturethe whole range of soil density and H/D ratios veryeffectively. It may also explain some of the scatter in

    test results (Schamine et al.) and particularly for theloose soils.

    Wedge failure models can broadly be classified intoa) simple vertical slip model (e.g. Schamine et al. seeFig. 9 with 0) and b) inclined wedge models (e.g.White et al).

    Consider a vertical block mechanism (Fig. 9 with 0) and assume only the soil above the crown of

    the pipe is active (height, H). The uplift resistance, P,per unit length for frictional (drained) behaviour isgiven by:

    (11)

    where f K tan and is known as the uplift factor.Uplift factors indicated by the Schamine tests are

    shown in Table 4 for the limited range of H/D tested(4 for sands).

    A variation of the frictional model (Eqn (11)) is

    attributed to Pedersen in which the whole volume ofsoil above the pipe is involved:

    (12)

    where fp K tan (using the suffix p to differentiatewith the f of Schamine) and 0.1 (D/H) represents theweight of the soil wedges between the pipe centrelineand the crown.

    The equivalent vertical slip model for cohesive(undrained) behaviour leads to:

    (13)

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    H

    D

    Wz

    PP

    Figure 9. Uplift wedge failure mechanism.

    Figure 10. Circulation mechanism in very loose sand(Vanden Berghe 2005).

    Table 4. Uplift factors (Schamine et al).

    Soil Uplift factor, f []

    Very loose sand 0.15

    Loose sand 0.4Gravel/rockfill 0.6

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    Palmer et al. (1990) recommend more conserva-tive factors (0.5 for dense material and 0.1 for loose)and put forward a variation of this equation to addressthe problem of deep or flow failure, expressed as:

    (14)

    For the cohesive model, the more rigorous solu-tions of Merifield et al. (2001) are preferred. Uplift

    bearing capacity factors (P/suD) have been computedusing both upper and lower bound solutions account-ing for shallow and deep failure mechanisms. Thorneet al. (2004) has investigated in detail the issue of suc-tion behind the pipe.

    Bolton & Barefoot (1997) and White et al. (2001)have justified the wedge mechanism shown in Figure9 by demonstrating that the angle corresponds tothe dilation angle of the soil and accounting for anincrease in the vertical stress (and shear resistance) inthe vicinity of the pipe. The shear stress along theslip surface is expressed as:

    (15)

    and thus:

    (16)

    and K0 can be taken as 1 sin critThe dilation angle can be assessed from the rela-

    tive density of the sand, the stress level and the par-ticle characteristics (Bolton 1986). For a range of sands

    between loose and dense White et al. show that thecorrelation is good. In the experience of the authors,this approach may overestimate uplift resistance forvery loose sands where a negative dilation anglewould be applicable and where a circulation flowmechanism occurs.

    Recommended uplift factors are also given byFinch et al. (2000) for different soil conditions.

    In practice, some other issues must be addressedwhen considering uplift resistance:

    pipelines are not always in continuous contact withthe seabed;

    uplift loading is often cyclic since it is associatedwith production cycles; upward movement mayenable a gap to develop and particles to flow underthe pipe, leading to upward ratcheting or creep;

    very soft backfills are heterogeneous and maybehave like a frictional material.

    These topics are discussed below.Discontinuous contact of the pipeline with the

    trench is normal due to irregularities in the trenchlevel. This results in gaps (spans) under the pipewhich may not be fully filled with backfill soil.Fortunately, this is typically associated with sag bendswhich, when loaded axially, will tend to move downinto the soil. However, gaps could exist adjacent to asupported area. A gap below the pipe increases thetendency for a circulation or flow failure and there-fore the uplift resistance could be lower than antici-

    pated for a uniformly bearing pipeline. Designconservatism must be introduced for incomplete con-tact with the seabed.

    Thermal loading is associated with production andwhen production is halted the pipeline cools down.Thus pipelines experience many cycles of uplift load.If the backfill resistance is low and the pipe is able tomove sufficiently to create a gap under the pipeline,soil particles can flow into this gap during a hot

    phase. On cooling the pipe cannot return to its ori-ginal position. Cyclic loading can thus lead to the pipecreeping or ratcheting upwards. With loss of cover andresistance a strain-softening resistance is experiencedand full uplift failure can occur. Nielsen et al. (1990)

    postulate that this was the mechanism that occurred intheir project. Finch et al. (2000) suggest, based on 1 glaboratory testing that ratcheting only occurs in cleansands where the grains are not held by adhesion. Inorder to limit progressive uplift the magnitude ofuplift displacements must be controlled.

    Very soft clay backfills are heterogeneous and arebelieved to drain more rapidly than the in situ soil.Bolton & Barefoot (1997) have shown that for theAtlantic mud tested, both drainage is relatively rapidand the clay dilates when sheared at very low stress

    levels. This results in higher than anticipated upliftresistance. They argue that adopting a frictionalmodel with an uplift factor f of 0.4 can be justified forthe mud tested. The authors have been involved in

    projects involving jet trenching of very soft clays. Anuplift factor of 0.3 was used based on centrifuge testson the soil and no uplift problems have been reported.

    7.2.2 Uplift responseNumerical methods used to design against upheavalbuckling require not only the ultimate capacity butalso the displacement response to mobilize this cap-acity. Results can be sensitive to this stiffness. Muchless interest has been shown on uplift displacements

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    to reach failure but Finch et al. (2000) propose guide-lines based on their experimental program. Trautmannet al. (1985) suggest that the displacement at peakresistance is between 0.51.5% of the pipeline depth.For a 16 (0.4m OD) pipeline buried at a depth of 1 m(H/D 2.5) the displacement at peak would be

    between 26mm which is in agreement with Finchet al. at that depth.

    7.2.3 Lateral resistanceThe lateral resistance of buried pipelines is not gener-ally important for buckling but becomes importantif ground movements occur, such as by faults ormudslides.

    Pipelines buried in sand have been studied byAudibert & Nyman (1977), Nyman (1984) andTrautman & ORourke (1985). The ultimate lateralresistance can be written:

    (17)

    where the dimensionless lateral bearing capacity fac-tor Ny depends on the relative density of the sand andon the embedment of the pipeline. Trautman &ORourke (1985) showed that for loose and mediumdense sands, Ny increases approximately linearly withthe embedment for H/D 8, whereupon Nybecomesconstant, indicative of the transition from shallow todeep soil failure mechanism. For dense to very densesands, the transition was not reached at H/D of 11.

    Trautmann & ORourke also demonstrated that thevalues of Ny defined for the holding capacity ofanchor plates (Rowe & Davis 1982) were in goodagreement with their own data for pipes. Rowe andDavis showed that Ny depends primarily on the fric-tion angle and embedment ratio, and on the roughnessof the embedded structure.

    For homogeneous cohesive soils, the ultimate lat-eral resistance of buried pipelines can be based on thework of Merifield et al. 2001 for plate anchors :

    (18)

    where the dimensionless factor Nyu depends on theembedment of the pipeline and to a lesser extent on itssurface roughness.

    Considering conservatively the results of the lowerbound plasticity analysis quoted by Merifield et al.(2001), the dimensionless factor Nyu can be written:

    (19)

    (20)

    The limiting value of 10.47 reflects the transitionfrom shallow to deep behaviour.

    7.2.4 Lateral responseMoving on to the lateral force-displacement modelsin sands, as described by Trautman & ORourke, ahyperbolic relationship is proposed given by

    (21)

    where Py* Py/(Ny H D) is the normalised force and

    y* y/yu is the normalised displacement; a and b arethe model parameters. Proposed values for a and b byTrautman & ORourke are 0.17 and 0.83, respectively.Slightly different values are proposed by Audibert &

    Nyman (1977). The displacement at the peak resist-ance yu depends on the embedment ratio (H/D) anddecreases with increasing relative density of the sand.

    The hyperbolic force displacement curve can besimplified into a bilinear representation. Trautman &ORourke (1985) suggest an initial stiffness equal tothe secant stiffness at 70% of the ultimate resistance.In that case, the maximum force is reached at a dis-

    placement of 0.4yu. Normalised force-displacementcurves are plotted on Figure 11. Note that the soilsmodelled in this study have effective friction angles

    between 2030 and thus would be in the relativedensity range 020%.

    7.2.5 Resistance to inclined transverse loadsIn a study related to pipelines buried in very loosesand, Vanden Berghe et al. (2005) have shown that thereis very little difference in uplift resistance when the loaddirection is within about 30 of the vertical. Figure 12shows the displacement patterns and Figure 13 depictsthe uplift factor Nz as a function of direction.

    The implication of this finding for upheavalbuckling is that modes of deformation in an inclined

    108

    0

    0.2

    0.4

    0.6

    0.8

    1

    1.2

    0 0.2 0.4 0.6 0.8 1 1.2

    y / yu

    Py

    /P

    yu

    Audibert & Nyman (1977)

    Trautman & ORourke

    (1985)

    Simplified bi-linear

    Figure 11. Lateral force-displacement curves for pipelinesembedded in sand.

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    direction are very likely to occur since imperfectionsare rarely 2-dimensional. This agrees with experience

    of surveyed buckles.

    8 THERMAL PROPERTIES OF SOILS

    Efficient transportation of crude oil in a pipelinerequires a sufficiently high temperature to maintainlow product viscosity and avoid unwanted depositionof wax. Since the temperature of the soil and wateraround the pipeline is lower than the temperature ofthe oil, heat flows from the oil to the environment.Heat lost along the length of the line results in a tem-

    perature drop from inlet to outlet. Pipeline designmust ensure that this temperature drop is withinacceptable limits. The temperature loss along the line

    will depend on the thermal conductivity of the pipelineand its surrounds. If transient solutions are required(for example for the heating up or shutting down of thesystem) the specific heat capacity will also be required.Hence there is a need to know the thermal propertiesof the soil and make use of the low thermal conduct-ivity where possible to provide thermal insulation.

    8.1 Factors affecting thermal properties

    Heat transfer can take place by conduction, convec-tion and radiation. In saturated soils, however, heattransfer is mainly due to conduction through the solidframework and the pore water (Farouki 1986). Con-vection may be more important in coarse grained soilsand rockfill.

    Heat conduction in soil can be described, for theone dimensional case, by the Fourier equation:

    (22)

    where T is the temperature at time t and depth z; k isthe thermal conductivity (amount of heat that flowsthrough unit cross-sectional area under a unit temper-ature gradient (T/z) in unit time); C is the specificheat of soil; and is the density. Since the specificheat of particles of sand and clay is only about 20%that of water (Geological Society of America 1942), itis largely on account of the water content that soil iscapable of storing heat.

    The main factors that affect the thermal conduct-ivity of saturated soil are: mineral composition, particlesize distribution, density/water content, and tempera-

    ture. In general, solids conduct heat better than liquids,and liquids better than gases (Brandon & Mitchell1989).

    Table 5 indicates some typical values for soil com-ponents:

    Mineral composition. Heat conduction through theparticles is an important mechanism of heat transfer.All other factors being equal, sands containing a high

    percentage of quartz will have a higher thermal con-ductivity than those containing a high percentage ofmica (Brandon & Mitchell 1989). Clays with arelatively high content of kaolinite have relatively lowconductivities. Moreover, sands containing high per-centages of silica may exhibit increased thermal con-ductivity with time, possibly due to the formation of

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    a) vertical b) 22.5to the vertical

    c) 45to the vertical d) horizontal

    Figure 12. Failure mechanisms for different displacementdirections (Vanden Berghe et al. 2005).

    0

    2

    4

    6

    8

    0 22.5 45 67.5 90

    Displacement Direction []

    Nv

    [-]

    Phi=20- Psi= - 10

    Phi=25- Psi= - 5

    Phi=30- Psi = 0

    H/D=2

    Figure 13. Comparison between vertical, oblique and lat-eral resistance in contractive soils (Vanden Berghe et al.2005).

    Table 5. Typical thermal conductivities.

    Material Thermal conductivity (W/mK)

    Quartz 4.09.1Water 0.600.67

    Clay (typical) 1.52.9

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    silica precipitants at the contact between grains(Brandon & Mitchell 1989).

    Particle size distribution. Well-graded soils con-duct heat better than poorly graded soils because thesmaller grains can fit in the interstitial voids betweenthe larger grains, thus increasing the density and themineral-to-mineral contact area (Brandon & Mitchell1989). Thermal conductivity in general varies withthe grain size of the soil. At a given density and mois-ture content, the conductivity is relatively high incoarse grained soils such as gravel or sand, somewhatlower in sandy loam soils, and lowest in fine grainedsoils such as silty loam or clay (Kersten 1949).

    Density and water content. Due to the relativelyhigh conductivity of minerals compared to water, theconductivity increases with density (Kersten 1949,Brandon & Mitchell 1989).

    Temperature. Thermal conductivity may also be

    influenced by temperature because each of the con-stituents has temperature dependent thermal conductiv-ities (Brandon & Mitchell 1989). All crystallineminerals in soils show a decrease in thermal conductiv-ity with increasing temperature, except feldspar (clay).The thermal conductivity of water increases with tem-

    perature because of the higher level of molecular move-ment at higher temperatures (heat transfer by collision).

    8.2 Measuring thermal conductivity

    Since thermal conductivity is dependent on the spe-cific mineral constituents of the soil as well as its dens-ity/water content it is preferable to measure theconductivity on soil samples in the laboratory orin-situ. A common method is the thermal probe.

    The thermal probe is a long needle that is inserted inthe soil containing both heating and temperature meas-uring elements. A known amount of current is passedthrough the heater element and the resulting variationof temperature is measured as a function of time. Thethermal conductivity of the soil can be deduced fromthese measurements. The applicable procedure isdescribed by ASTM D5334 (2000). The thermal needle

    probe has been presented by various authors (Hooper &

    Lepper 1950, DeVries 1952, Woodside 1958, Falvey1968, Mitchell & Kao 1978, among others).

    8.3 Empirical methods for determiningconductivity

    In the absence of specific laboratory data, variousempirical equations are available relating the thermalconductivity of the soil to its water content, dry densityand type of soil. Some empirical approaches are givenin Table 6.

    For saturated soils, the relationships presented inTable 6 can be conveniently expressed in terms of thewater content rather than the dry density or porosity.The three approaches presented in Table 6 have been

    used to show variation that may be anticipated(Fig. 14).

    A detailed review is given by Farouki (1986) andRawat et al. 1979, Young et al. (2001) have also con-tributed. Rawat et al. suggested that maximum errorwith the Kersten method was 25%. Farouki recom-mends the Johansen method for coarse to fine sand

    (5% passing 2 microns). This method was found toprovide the best correlation because it takes intoaccount the mineralogy of the sand (which should bedetermined by x-ray diffraction). Based on laboratorydata, Rawat considered that the Makowski andMochlinskis method overestimates the conductivityunless the combined silt and clay fraction was used inthe equations.

    Laboratory tests from high water content deepwa-ter Gulf of Mexico clays, Indonesia and Nigeria(remoulded and undisturbed) were reported by Younget al. (2001). Remoulded data is shown in Figure 15.

    This is of interest because it covers water contentsthat would be typical for trench backfills.The undisturbed soil samples had thermal conduct-

    ivity values ranging from 0.65 to 1.25W/mK whileremoulded values were in the range 0.8 1.05W/mK.

    8.4 Selection of thermal conductivity

    As with all design parameter selection, the use of theparameter should be considered. Finch et al. (2000)suggest the following guidelines: upper bound valuesshould be adopted for thermal insulation design, ashigh thermal conductivity represents high heat loss.Conversely, lower bound values are applicable toupheaval buckling assessments where heat retained in

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    Table 6. Empirical equations for determining soil conduct-ivity.

    Reference Equations (k in W/mK)

    Kersten k 0.144 (0.9 log w0.2) 100.06364d

    (1949) d dry density (in kg/m3

    )for silts and clayey soils

    k 0.144 (0.7 log w 0.4) 100.06364d

    for sands

    Johansen k ksat ks(1n)kw

    n

    (1975) ks kqqk0

    1q

    kw 0.6W/mK; kq 7.7W/mK;k0 2.0W/mKq quartz contentSimplified equation for saturated soils.

    Makowski k (a log(w) b) 10c

    & Mochlinski a 0.1424 0.000465p(1956) b 0.0419 0.000313p

    c 0.00062 dp weight percentage of soil f iner

    than 2m

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    the pipeline will tend to increase the uplift forcesexperienced by a buried pipeline.

    The thermal properties of jetted backf ills may be animportant design component for a deep water system.As discussed in connection with mechanical proper-ties, the changes in water content during jet trenchingare not well known. This introduces significant uncer-tainty when considering the thermal properties of the

    backfill. Deep water sites are commonly associatedwith soft clays and are ideally suited to the use of jettrenching. Even though the properties of the backfillsoil are significantly different from the virgin soil, the

    backfill soil exhibits low values of thermal conductiv-ities and sufficient strength and density to inhibit ther-mal convection currents (Young et al. 2001). In fact,the very soft highly plastic clays encountered at mostdeepwater locations have three characteristics thatmake them a favourable medium for flowline insula-

    tion. First, the clays exhibit cohesion and low permea-bility making them strongly resistant to thermalconvection (water travelling freely through the soil toand from the heat source). Second, saturated clayswith high water contents exhibit low values of thermalconductivity. Third, the soils can be easily jetted to

    produce a trench with steep but stable trench walls.

    9 CONCLUSIONS

    Pipeline geotechnics deals with soil-pipeline inter-action. This covers installation issues (pipeline pene-tration and short-term lateral stability), axial andlateral response to loads. It then encompasses pipelinetrenching, backfill engineering and pipeline stabilitywhen buried. Particularly current topics such as riser-soil interaction, upheaval buckling and lateral buck-ling/ snaking, all of which include many load cycles,are all subject to ongoing investigations and jointindustry projects.

    While performing this review, the authors havebeen made aware again of the very wide range ofissues that must be faced in connection with pipelinedesign. We have found no similar review papers cov-

    ering the subject. Therefore, this paper claims unique-ness and we trust it will provide others with a starting

    point in many of the specific subject areas. It waswritten in the midst of a busy consulting schedule andtherefore does not treat all the subject matter as fullyas we would have liked. There are likely to be someerrors that have crept in. Nevertheless, we trust thatfuture reviews will f ind this a useful starting point.

    REFERENCES

    ASTM D5334, 2000. Standard test method for determina-tion of thermal conductivity of soil and soft rock by ther-mal needle probe procedure.

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    0

    1

    2

    3

    4

    5

    0 20 40 60 80

    0 20 40 60 80

    0 20 40 60 80

    Thermalconduc

    tivity[W/mK]

    0

    1

    2

    3

    4

    5

    Thermalconductivityk[W/mK]

    0

    JohansenQuartz content, %100

    50

    0

    1

    2

    3

    4

    5

    Water content [%]

    Thermalconductiv

    ity[W/mK]

    MakowskiClay content, %

    50

    100

    0

    Kersten

    Sand

    Clay

    Figure 14. Thermal conductivities by various methods forsaturated soils.

    Figure 15. Remoulded conductivities (Young et al 2001).

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