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946

Williams, E. S. et al. (2013). Gé otechnique 63, No. 11, 946–956 [http://dx.doi.org/10.1680/geot.12.P.067]

Pipe uplift in saturated sand: rate and density effects

E . S . W I L L I A M S , B . W. B Y R N E a n d A . B L A K E B O RO U G H

A key parameter for assessing the unburial behaviour of offshore pipelines, such as during upheaval buckling, is the soil uplift resistance. Despite much research on uplift response, two variables remainrelatively unexplored: the effect of low relative density backfill and partially drained soil response. Toaddress this shortcoming, an investigation into the plane-strain uplift resistance of pipe segments insaturated sand was carried out. The model tests, targeting drained and partially drained soil behaviour,demonstrate that soil state strongly influences the pipe–soil failure mechanism, and hence themobilised uplift resistance. In very loose sand the normalised peak uplift resistance remained constant beyond a critical embedment depth, indicating a possible transition between different failure mechan-isms. Faster rate tests result in positive excess pore pressures, which serve to further reduce the upliftresistance. Tests at varying uplift rates and densities identified a transition relative density between netcontraction and dilation for a given embedment. These findings suggest that a minimum backfilldensity should be ensured prior to pipeline operation, to avoid the negative effects associated withcontractive behaviour of loose soil.

KEYWORDS: liquefaction; model tests; offshore engineering; pipelines; sands; soil/structure interaction

INTRODUCTIONThe exploitation of offshore oil and gas reserves requires avast network of subsea pipelines, such as those used toconnect newly discovered oil fields to existing productionfacilities. These pipes are often buried to some depth of soilcover for thermal insulation and pipeline stability, and to protect them from mechanical damage due to exposureabove the seabed (Schupp, 2009). The required embedmentdepth must be sufficient to prevent pipeline unburial behav-iour, but should be minimised for economy of design

(deeper embedment depths are generally associated withgreater installation costs). One of the main pipeline unburialmechanisms is upheaval buckling; a structural effect caused by high operating temperatures and pressures. A key param-eter in assessing pipeline stability against upheaval bucklingis the vertical resistance provided by the soil to upward movement of the pipe. Previous research on this topic hasfocused on understanding the deformation mechanisms and the force–displacement response during pipe uplift, gener-ally assuming drained conditions. This has led to predictionmodels that form the basis of recommendations for pipelinedesigners, such as in DNV-RP-F110 (DNV, 2007).

Although numerous studies on uplift resistance have beencarried out over the last three decades, there are some issues

that are not yet fully resolved, and therefore are not properlyaccounted for in the industry guidelines: the impact of the backfill relative density, particularly in very loose sand, onthe induced failure mechanisms, and the corresponding upliftresistance. Furthermore, partially drained conditions stem-ming from faster uplift rates or pipe vibration may also berelevant when installation methods such as jet-trenchingresult in very loose, liquefiable backfill soils. In this instanceit is important to account for both density and rate effects,and appropriate provisions for these conditions should beincluded in design guidelines. This paper reports a compre-hensive investigation into these two effects.

BACKGROUND Prediction methods for drained resistance

The basis for a prediction model for peak uplift resistanceis a realistic understanding of the failure mechanism for a pipe moving upwards in the soil. Early work, such as that byTrautmann et al. (1985) or Schaminée et al. (1990), con-cluded that the primary failure mechanism consists of theupward heave of a sliding block of soil bounded by shear planes/zones that, at shallow burial depths, extend to theground surface. More sophisticated imaging techniques used

in recent years have confirmed this assumption, and providea better understanding of its range of applicability (e.g.White et al., 2001; Byrne et al., 2008).

Based on these observations, the simplest model is thevertical slip model (VSM), which assumes that the failure planes of the soil extend vertically upwards from the edgesof the pipe to the ground surface, as shown in Fig. 1(a). Theuplift resistance is composed of the weight of the soil prismabove the centre of the pipe – less the soil displaced by tophalf of the pipe – and the shear resistance of the slipsurfaces caused by the lateral earth pressure. A typicalcalculation is given by:

F ult ¼ 1 D

8 H

þ K tan 9 H

D ª9 HDL (1)

where D is the pipe diameter, H is the embedment depthto the pipe centre (note that this is different from thedepth of cover defined in the industry, which is usuallymeasured to the pipe crown), and 9 and g 9 are theeffective friction angle and unit weight respectively. Typic-ally, K is taken as the at-rest lateral earth pressure coeffi-cient ( K 0 ¼ 1 sin 9); alternatively, K tan9 can bereplaced by an empirical uplift factor f , based on broad soil and density classifications. An adapted version of thismodel (the aVSM) defined by Schupp (2009), which doesnot subtract the displaced soil, is used in the DNV guide-lines (see equations (5) and (6)).

The VSM is appropriate for loose sands at shallow depths,as observed by Schupp (2009) and Wang et al. (2010).However, as the density of the soil or the depth of cover increases, the shear planes appear to deviate from the

Manuscript received 11 May 2012; revised manuscript accepted

22 January 2013. Published online ahead of print 18 April 2013.Discussion on this paper closes on 1 February 2014, for further detailssee p. ii. Department of Engineering Science, University of Oxford, UK.

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vertical. In medium-dense to dense sand conditions theyappear oriented at approximately the angle of dilation fromthe vertical (White et al., 2001; Cheuk et al., 2008). Severalresearchers have proposed theories that adapt the VSM byusing these inclined shear planes. The formulation by Whiteet al. (2001), based on the mechanism in Fig. 1(b), appearsto agree well with data in the literature, particularly for medium to dense soils. This inclined slip surface model(referred to herein as the ISM) assumes that the normalstress acting on the shear planes is equal to the in situ (at-rest) value, meaning that the stress level above the pipeincreases as the uplift resistance is mobilised. Verticalequilibrium of the modified sliding block produces nearlythe same equation as in the VSM, but with K tan9 replaced by f (White et al., 2001, 2008), and is given by

f ¼ tanł þ tan peak

tanł 1 þ K 02

(1 K 0)(cos2ł)

2

(2)

where the peak friction and dilation angles are determined following Bolton (1986).

At greater embedment depths it has been observed thatthe sliding-block mechanism does not develop at smalldisplacements. Tests conducted by Schupp (2009) and Byrneet al. (2008) showed that localised compression, followed bya flow-around mechanism, may occur on initial displacementunder these conditions, as shown in Fig. 1(c). As the pipemoves closer to the surface, the sliding block again developsin a similar manner as described for shallower embedmentdepths. The initiation of this flow-around mechanism has

also been observed in the post-peak response of dense soil(White et al., 2001), which suggests that the transition between flow mechanism and sliding-block mechanism may be related to the state of the soil around the pipe.

Effect of pullout ratePrevious research, as described above, has focused on the

drained behaviour of sand – attained through tests in either dry sand, or saturated sand with an uplift rate sufficientlyslow to allow adequate drainage of the pore water. Plane-strain tests reported by Byrne et al. (2008, 2013), Bransby& Ireland (2009), and Schupp (2009), investigated the partially drained conditions that develop during faster rate

uplift tests, and the consequent effect on the mobilised soilresistance. In very loose sand tests, described by Byrne et al. (2013), as the rate of uplift was increased, the resistancewas found to decrease significantly – almost to zero in some

of the faster rate tests. Pore pressure transducers showed significant positive excess water pressures around the pipecircumference, which resulted in significantly reduced effec-tive stresses acting on the pipe. In some instances the excess pore water pressure was found to increase to the value of the initial vertical effective stress (ue ¼ ª9 H ), suggesting thatstatic liquefaction had occurred. Conversely, in denser sand,the peak uplift resistance was observed to be significantlylarger than predicted values for drained conditions, and thiswas also observed by Bransby & Ireland (2009). This sug-gests that negative excess pore pressures were developed,resulting in an increase in the effective stress operatingaround the pipe.

Current design guidelinesThe Det Norske Veritas recommended practice DNV-RP-

F110 (DNV, 2007; referred to below as DNV) is commonlyused in the offshore pipeline industry. This recommended practice defines global buckling as a load response rather than a failure mode, but one that may lead to an ultimatefailure mode such as local buckling. However, as littleguidance is given on post-buckled pipeline integrity, excessupward displacement is considered to be an ultimate limitstate (ULS) failure for the purpose of buried pipeline design.The recommended design procedure is non-linear finite-element analyses of the buried pipe configuration to deter-mine the operating failure temperature, with incorporated safety factors. An important modelling input is the upliftresistance of the soil, which is represented by a lower-bound characteristic curve. The DNV guidelines propose a tri-linear

force–displacement curve for calculating uplift resistance,and this is shown in Fig. 2. The peak uplift resistance,denoted as Rmax, is usually determined from an aVSM, witha recommended uplift resistance factor, f , depending on thedensity of the sand. There are three categories specified:loose, medium/dense and rock. The parameters used todefine this curve for loose and dense sand are provided inTable 1. Important factors affecting uplift resistance, such asrate effects and liquefaction susceptibility, are not specifi-cally addressed in these guidelines.

EXPERIMENTAL METHODThis paper summarises an experimental research pro-

gramme investigating the plane-strain uplift resistance of buried pipes in saturated sand. The tests were designed todetermine the influence of backfill relative density and upliftrate across a range of embedment ratios up to H / D of 4.5.

Soil surface

(a)

D

z H

(b)

D

z H

(c)

D

z H

W weight

γF friction

φ

ψW weight

γF friction

W weight

γ

Fig. 1. Pipeline geometry and simplified failure mechanisms (after Schupp, 2009; White et al., 2008): (a) vertical slip surface; (b) inclinedslip surface; (c) flow mechanism

PIPE UPLIFT IN SATURATED SAND: RATE AND DENSITY EFFECTS 947

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This involved two series of scale model tests exploringdrained and partially drained behaviour. The overall aims of the investigation were

(a) to improve understanding of the mechanisms involved inmobilising the soil resistance for buried pipes installed inloose sand by trenching

(b) to evaluate prediction models presented in the literature(c) to assess the validity of the recommendations in the

current design guidelines for these conditions.

The experimental rig used for the plane-strain pipe uplifttests is described in detail by Schupp (2009); the mainfeatures are outlined as follows. The rig is shown in Fig. 3.It consists of a testing tank measuring 1 m 3 0.3 m 3 1 m,equipped with a soil fluidisation system by means of anupward hydraulic gradient. This is achieved by cycling water upwards through the soil sample using a pump, which drawsfrom a reservoir mounted to the side of the tank and

distributes water through the base of the tank via a systemof drainage channels. Even pressure distribution and fluidisa-tion of the sample are ensured by the use of a geotextilefilter mounted at the base of the tank.

The tank is fitted with a fully instrumented model pipe,which can be moved vertically at different rates using astepper motor driven actuator. The 100 mm diameter pipeconsists of a 200 mm central ‘active’ test section, and two100 mm ‘dummy’ sections at either end, which sit flush withthe tank sides. The active test section of the pipe is fitted with an internal, moment-compensated load cell, whichallows measurement of the vertical forces exerted on thecentral section of the pipe. Excess pore pressure measure-

0

0·2

0·4

0·6

0·8

1·0

1·2

0 0·2 0·4 0·6 0·8 1·0 1·2 1·4

Normalisedresistance,

/R

R m

a x

Normalised displacement, /d d f

α 0·85

α 0·75

α γ/ UR

1/γUR

Design curve

Characteristic curve

Fig. 2. DNV tri-linear uplift resistance curve (after DNV (2007))

Table 1. DNV suggested soil parameters for tri-linear uplift curve (DNV, 2007)

Category Parameter Range

Loose sand (3.5 < H / D < 7.5) f 0.1–0.3d f 0.5% H c –0.8% H cÆ 0.75–0.85 0.2

Medium/dense (2.0 < H / D < 8.0) f 0.4–0.6d f 0.5% H c –0.8% H cÆ 0.65–0.75 0.2

Post-peak resistance at displacement f r Æf . f Æf 0.65–0.75d fr 3 . d f

Fig. 3. Experimental set-up (from Byrne et al., 2008; courtesy of OTC)

948 WILLIAMS, BYRNE AND BLAKEBOROUGH

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ments are taken using five pore pressure transducers evenlyspaced around the half-circumference of the pipe. A long-range displacement transducer is used to track the pipedisplacement through the height of the tank, and a short-range, high-resolution transducer is used to measure theinitial small displacement response.

The tests described in this paper all follow a procedurethat is predominantly automated, allowing systematic

changes to input variables such as backfill density, embed-ment ratio and uplift rate. All tests were conducted usingRedhill 110, a poorly graded silica sand, in fully saturated conditions. The sand was selected for use as it resembles atypical North Sea backfill sand; relevant soil parameters are provided in Table 2, and the grading curve is shown in Fig.4. To simulate jet-trenching installation, the sand is firstfluidised to a target specific gravity and the pipe is lowered to the required depth. The soil then settles over the pipe,while a load-control routine (mimicking the self-weight of the pipe) is implemented. The soil typically settles at itsloosest state (i.e. zero relative density), corresponding to aneffective unit weight of 7.82 kN/m3: If required, the backfillis densified through vibration of the tank, after which the pipe is pulled upwards at a constant rate. This densification process mimics the possibility that the backfill material can be densified either by pipe vibrations or by wave and currenteffects. The relative density is determined by assuming thatthe settlement of the soil occurs uniformly throughout the

tank, which is not unreasonable for the settlement of veryloose sands.

DRAINED RESPONSEThe first test programme investigated the influence of

backfill relative density, at various embedment depths, onthe peak uplift resistance and the force– displacement re-

sponse under drained conditions. The rate of pipe uplift wasset at 0.002 mm/s, a speed judged to be sufficiently slow toensure adequate drainage of the soil during the test. Thiswas confirmed by the negligible response of the pore pressure transducers on the pipe circumference during uplift.A summary of the tests conducted is provided in Table 3.

Peak uplift resistance/breakout factorsThe results for peak uplift force are presented in Figs 5

and 6 in the form of a soil breakout factor, N ult, whichnormalises the force with respect to the effective unit weightof the soil, the embedment depth and the pipe dimensions.

N ult ¼ F ult

ª9 HDL (3)

Figure 5 presents the peak breakout factors for backfill atvarious initial relative densities across a range of embedmentratios. The prediction models chosen for comparison are theVSM and aVSM (using the critical-state friction angle), aswell as the DNV-recommended ranges (Fig. 5(c)). Thesemodels are normalised in a similar manner as follows.

N ult,VSM ¼ 1 D=8 H ð Þ þ K 0 tan 9 H = Dð Þ½ ª9 HDL

ª9 HDL

¼ 1 D

8 H þ K 0 tan 9

H

D

(4)

N ult,aVSM ¼ 1 þ K 0 tan 9 H

D (5)

N ult,DNV ¼ 1 þ f H

D (6)

where f 2 [0.1, 0.3] for loose sand and [0.4, 0.6] for medium/dense sand.

From Fig. 5(a), it can be seen that the VSM is anexcellent predictor of the uplift resistance for soil at itsloosest state, up to an embedment ratio of approximately 2.Beyond this depth, the breakout factors remain constant,indicating a possible transition to a flow-around failuremechanism. As the relative density increases, the applicabil-ity of both the VSMs extends to greater depths, after whichsimilar plateaus in the normalised resistance are observed

(Fig. 5(b)). The results indicate that, as the relative densityincreases, the transition from sliding-block mechanism toanother mechanism (probably the flow-around mechanism)occurs at greater values of H / D. It would be possible to fit amathematical expression empirically to this plateau behav-iour, but, given only a small number of data points obtained,it is not felt appropriate to do so. For practical purposes the

Table 2. Soil properties for Redhill 110

Property Value Source

Maximum void ratio, emax 1.04 Schupp (2009)Minimum void ratio, emin 0.64 Schupp (2009)Specific gravity, G s 2.63 Schupp (2009)

Critical friction angle, crit 368

Villalobos et al. (2005)

Table 3. Summary of drained uplift tests

Set no. Embedment ratio, H / D Relative density: % Uplift rate: mm/s Number of tests

Drained 1 0.5–4.5 0–20 0.002 62

Drained 2 1.5–4.5 21–60 0.002 13

Total drained 75

0·01 0·1 1

Grain size: mm

100

90

80

70

60

50

40

30

20

10

0

Cumulativepercentagepassing:%

Fig. 4. Grading curve for Redhill 110


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plateau value of N ult could be related to either relativedensity or some estimate of the soil’s friction angle account-ing for stress-level effects.

Figure 5(c) also includes dry sand data obtained bySchupp (2009) using the same rig, which demonstrates thatthe above trends can be observed at higher embedment ratios

and higher stress levels (,

23), as well as for greater-diameter pipes. Given that full-scale pipelines are generallytwo to three times the diameter of the model pipe in thesetests, scaling effects are felt to be negligible. For example,

the difference in stress levels is minimal, and in addition the potential effects of dilation are minimised at the range of relative densities targeted in these tests.

Figure 6 compares the peak breakout factors with predic-tion curves based on the ISM proposed by White et al.(2001). At an embedment ratio of 1.5 the breakout factorsare very consistent with this model, across the entire rangeof densities. However, similar to the VSMs, as the embed-ment depth increases, the ISM increasingly overpredicts the peak force response. The results demonstrate that for veryloose soil conditions an increase in embedment depth, beyond a critical value, does not necessarily result in anysignificant improvement in the breakout factor.

Force– displacement responseExamples of force–displacement curves across the range

of embedment depths tested are shown in Figs 7 and 8,together with VSM and aVSM curves for comparison. InFig. 7, the measured and predicted forces are normalised bythe soil effective unit weight and pipe diameter, as this preserves the shape of the force– displacement curve,through the uplift test. Once again, for the loose sand at 0%relative density, it can be seen that the VSM closely matchesthe force–displacement response of the pipe at low embed-ment ratios, whereas at greater depths it overpredicts theresponse. Tests at an embedment ratio of 4 demonstrate howan increase in relative density to 20% is required for theforce to reach the VSM curves, while the response at RD ¼ 30% is higher than the predicted values, but appears

0·5

1·0

1·5

2·0

2·5

0 0·5 1·0 1·5 2·0 2·5 3·0 3·5 4·0 4·5

Brea

koutfactor,N u

l t

Embedment ratio, /(a)

H D

Sliding-block mechanism

Flow-around mechanism

VSM

aVSM

Measured data – RD 0%

0·5

1·0

1·5

2·0

2·5

0 0·5 1·0 1·5 2·0 2·5 3·0 3·5 4·0 4·5

Breakoutfacto

r,N u

l t

Embedment ratio, /(b)

H D

VSMaVSM

Measured data – RD 0%Measured data – RD 10%Measured data – RD %15Measured data – RD 0%2Flow mechanism approximation

RD 20%

RD 15%

RD 10%

RD 0%

DNV medium/dense

aVSM

DNV loose

0

0·5

1·0

1·5

2·0

2·5

3·0

0 1·0 2·0 3·0 4·0 5·0 6·0 7·0

Breakoutfactor,N u

l t

Embedment ratio, /(c)

H D

VSMMeasured data – RD 25–35%Measured data – RD 40–55%DNV rangesExtended DNV rangesSchupp (2009): Dry, RD 30%, 100 mm D Schupp (2009): Dry, RD 17%, 220 mm D

Fig. 5. Summary of breakout factors compared with vertical slipmodels: (a) very loose sand, relative density 0%; (b) loose sand;relative density 0–20%; (c) loose to dense sand, relative density0–55%

1·0

1·5

2·0

2·5

3·0

3·5

4·0

0 10 20 30 40 50 60

Breakoutfactor,N u

l t

Relative density: %

H D/ 1·5

H D/ 2·5

H D/ 3·5

H D/ 4·0

H D/ 4·5

H D /

4 · 5

H D /

4 · 0

H D /

3 · 5

H D /

2 · 5

H D /

1 · 5

Fig. 6. Breakout factors against relative density, compared withinclined slip surface model by White et al. (2001)

0

200

400

600

800

1000

1200

0·5 1·0 1·5 2·0 2·5 3·0 3·5 4·0 4·5 5·0

Force/ γ

D 3

Embedment ratio, /H D

VSM

aVSM

Flow-around RD 0%Measured data

RD 30%RD 20%

RD 0%

RD 20%

RD 0%

Fig. 7. Representative curves of force plotted against embedmentratio


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to approach these curves after approximately 100 mm of displacement.

In Fig. 8, the initial displacement response is shown for varying initial densities at embedment ratios of 1.5 and 3.5(displacement is equal to the pipe embedment at any giventime, z , subtracted from the original embedment, H ). As therelative density increases from 0%, an increasingly distinct peak response occurs, followed by rapid strain-softening.Estimates of the initial stiffness (taken as a linear least-squares fit of the curve up to 50% F ult) indicate that stiffnessincreases with relative density, but is virtually unaffected byembedment ratio (see Fig. 9). Regardless of the initialrelative density and embedment depth, the tests show that,after the initial displacement, the soil resistance converges toa consistent load response (matched by the VSM at H / D , 2). The displacement required for this convergence isdependent on the embedment depth, where a greater uplift

distance is required for convergence at deeper embedmentdepths.

Mobilisation displacement The measured mobilisation displacements, defined as the

pipe displacement corresponding to peak resistance, arenormalised by both H (Fig. 10(a)) and D (Fig. 10(b)). Bothare presented for comparison with data in the literature;however, owing to the potential for localised failure mechan-isms, it is felt that pipe diameter (a local length scale) is the

more appropriate normalisation. The plotted displacements,d f / H or d f / D, appear to be dependent on both relativedensity and embedment ratio, although the overall variationis small, with values ranging from 0.1% H to 1.7% H . This isconsistent with values found by several researchers (Matyas& Davis, 1983; Trautmann et al., 1985; Bransby et al.,2002; Cheuk et al., 2008), which range from 0.1% H to1.5% H, but is at the lower bound of values reported byThusyanthan et al. (2010) and Wang et al. (2012). Themobilisation displacements tend to be highest in loose testsat low embedment ratios; at deeper burial depths relativedensity has a lesser influence.

The DNV guidelines suggest mobilisation displacementsfor modelling purposes in the range 0.5% H c to 0.8% H c,where H c is measured to the pipe crown. This was chal-lenged in a recent study by Thusyanthan et al. (2010), inwhich it is suggested that instead this value could be related to H / D as follows.

d f

D ¼ 0:02 exp

H

2 D

(7)

Both models are compared with the experimental data inFig. 10, and it can be seen that while the DNV limits appear to match the data at greater embedment ratios, they are not

conservative for loose sand at shallow embedment ratios.The model proposed by Thusyanthan et al. (2010) predictshigher mobilisation displacements than observed in theseexperiments, particularly at higher values of H / D.

0

20

40

60

80

100

120

140

160

180

200

220

0 1 2 3 4 5 6 7 8

Force:N

Displacement, : mmH z

H D/ 1·5Increasing relativedensity (0–55%)

H D/ 3·5Increasing relativedensity (0–55%)

VSM

Measured data

Fig. 8. Initial force– displacement response at selected embed-ment ratios

1000

1100

1200

1300

1400

1500

1600

1700

1800

1900

0 10 20 30 40 50 60

Stiffness:N/mm

Relative density: %

H D/ 2·0

H D/ 2·5

H D/ 3·0

H D/ 3·5

H D/ 4·5

H D/ 1·5

H D/ 1·0

Fig. 9. Initial soil stiffness at various embedment ratios andrelative densities

0

0·5

1·0

1·5

2·0

2·5

3·0

3·5

4·0

0·5 1·0 1·5 2·0 2·5 3·0 3·5 4·0 4·5

Normalisedmo

bilisationdistance

/:%

d

H f

Embedment ratio, /(a)

H D

RD 0%

RD 10%

RD 15%

RD 20%

RD 30–60%

Thusyanthan . (2010)et al

DNV 0·8%H c

DNV 0·5%H c

0

1

2

3

4

5

6

7

0·5 1·0 1·5 2·0 2·5 3·0 3·5 4·0 4·5

Normalisedmobilisation

distance,

/:%

d

D f

Embedment ratio, /(b)

H D

RD 0%

RD 10%

RD 15%

RD 20%

RD 30–60%

Thusyanthan . (2010)et al

DNV 0·8%H cDNV 0·5%H c

Fig. 10. Mobilisation displacements for various relative densities,normalised by: (a) H ; (b) D


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PARTIALLY DRAINED RESPONSEThe second series of tests explored the effect of varying

rates of uplift on soil at different relative densities (see Table4) to determine the effect of soil initial state on piperesponse.

Rate effects

Figure 11 presents the force– displacement curves and corresponding pore pressure response for tests conducted atan embedment ratio of 3.5, with a soil relative density of 0%. It was observed that faster rates of uplift resulted inhigher excess pore pressures around the pipe. At its looseststate, the soil is clearly contractive, creating positive excess pore pressures and a decrease in resistance on shearing. Itcan be seen in Fig. 11(b) that the excess pore pressure risesrapidly on initial displacement of the pipe, followed bygradual drainage. At faster rates, a larger displacement isrequired for the excess pore pressures to dissipate and theuplift resistance to increase to its drained residual value. Thisis consistent with the conclusion by Schupp (2009) that thetime required for dissipation of excess pore pressures inthe loose sand is independent of the rate of uplift. At 5 mm/sthe uplift rate is sufficiently rapid for the initial excess pore pressure to approach the vertical effective stress of the soil,

reducing the resistance temporarily to zero. The rate is alsosuch that the excess pressure does not have sufficient time todissipate over the course of the test, and consequently thedrained resistance is never achieved.

Tests with uplift rates greater than 0.1 mm/s were alsocompared with a model proposed by Schupp (2009) and Byrne et al. (2013) to predict the partially drained upliftresponse. This model is based on the VSM, but where the

ª9

term in the uplift equation is replaced by ª9

(ue/ H ).The excess water pressure is given an initial value equal tothe effective stress, and is assumed to decay based onTerzaghi’s one-dimensional consolidation equation. The re-sults of the analysis plotted in Fig. 11 demonstrate that themodel can successfully capture the pore pressure response infast-rate tests, when the soil is at its loosest state, and for embedment depths for which the VSM is appropriate.

Density effectsTo explore the effect of density on the uplift resistance a

series of tests with an uplift rate of 1 mm/s were carried out.This was chosen as a representative rate at which a partiallydrained response occurred. The force–displacement resultsare shown for a range of relative densities in Fig. 12. Inthese tests, a gradual transition from a contractive to dilative

Table 4. Summary of partially drained uplift tests

Set no. Embedment ratio, H / D Relative density: % Uplift rate: mm/s Number of tests

PD 1 3.5 0 0.01–5 8PD 2 3.5 0–35 0.1 2

PD 3 1.5– 4.5 0–55 1, 5 44

Total partially drained 54

0

20

40

60

80

100

120

0 50 100 150 200 250 300 350

Upliftforce:N

Embedment depth, : mm(a)

z

VSM

aVSM

Schupp model

Measured data

Increasing uplift

rate(0·002–5 mm/s)

Increasing uplift rate(0·002–5 mm/s)

Vertical effectivestress

0·5

0

0·5

1·0

1·5

2·0

2·5

3·0

0 50 100 150 200 250 300 350

A

verageexcessPWP,

:kPa

u e

Embedment depth, : mm(b)

z

Fig. 11. Rate effects at low relative density: (a) force response;(b) pore pressure response

0

20

40

60

80

100

120

140

160

180

0 50 100 150 200 250 300 350

Upliftforce:N


z

Increasing relativedensity (0–35%)

RD 35%

RD 21%

VSM

aVSM

Measured data

0·5

0

0·5

1·0

1·5

2·0

2·5

3·0

0 50 100 150 200 250 300 350

A

verageexcessPWP,

:kPa

u e

Embedment depth, : mm(b)

z

Vertical effectivestress

Increasing relativedensity (0–35%)

Fig. 12. Density effects on fast-rate tests: (a) force response;(b) pore pressure response


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soil response was identified, as evidenced by the excess pore pressure response. For backfill soil with a relative densityranging from 0% to 15%, positive excess pore pressuresdevelop, although the magnitude and effect of these de-creased as the density increased. This is also evident in therecorded force curves. At 21% relative density the pipeexperiences positive excess pore pressures, indicating con-traction, but this is immediately followed by negative excess

pore pressures, indicating dilation. In this instance the forcereaches a higher resistance than the drained peak value atthe same density. Finally, at a relative density of 35%, thereis an immediate large negative pore pressure, indicating adilating soil response, and a single peak resistance muchlarger than the drained value for this density.

COMPARISON BETWEEN DRAINED AND PARTIALLYDRAINED; TRANSITION DENSITIES

Figure 13 shows a comparison between drained and partially drained tests at an uplift rate of 1 mm/s for relativedensities of 0% and 30% respectively. In these two figuresthe average excess pore pressure is converted to an equiva-lent force, based on a half-circumference of the pipe. Thesum of the measured soil force on the pipe and the excess

pore pressure force for partially drained tests is then compared with the drained force–response curves for eachdensity. At a relative density of 0%, it can be seen that thedifference between the drained and partially drained re-sponse is almost exactly equal to the average excess porewater pressure around the pipe (i.e. the ‘sum’ curve and thedrained curve are effectively equivalent). This suggests thatcontraction of soil occurring predominantly around the pipe

is the main cause of decreasing uplift resistance. In contrast,at a relative density of 30%, the difference between thedrained response and partially drained response is greater than the excess pore pressure force. This indicates that, inaddition to dilation around the pipe, dilation elsewhere inthe tank (particularly on the shear planes) is acting toincrease the soil resistance. It is expected that this additionaldilation effect would increase as the relative density and/or embedment ratio further increased.

Trends for peak uplift resistance due to variations in upliftrate and relative density for a given embedment ratio( H / D ¼ 3.5) are summarised in Fig. 14. From both plots,estimates can be made regarding the transition density be-tween net contractive and dilative behaviour, based ondifferences observed between the drained and partiallydrained responses during uplift. Note that the drained resis-

VSM

Measured force: drained

Measured force: 1 mm/sv

Excess pore pressure (force): 1 mm/sv

Pressure measured force: 1 mm/s v

Difference betweendrained and partiallydrained peak resistance

0

50

100

150

200

250 270 290 310 330 350

Force:N


z

0

50

100

150

200

250 270 290 310 330 350

Force:N

Embedment depth, : mm

(b)

z

Increase due to dilationnear pipe

Additional increase

Difference between drained andpartially drained resistance

Fig. 13. Comparison of drained and partially drained response for a fast-rate test: (a) relativedensity 0%; (b) relative density ,30%


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tance is identified by the curve relating tests at a rate of 0.002 mm/s. At this depth the transition range appears tooccur at a relative density between approximately 17% and 21%, above and below which net dilation and contractionoccur respectively. It should also be noted that in drained tests at an embedment ratio of 3.5, RD ¼ 20% is sufficientfor the data to match the aVSM – a similar value to thetransition range identified above.

Figure 15 shows a plot of the ‘transition densities’ deter-

mined for a 100 mm diameter pipe using fast-rate tests for several embedment ratios (as described above), along withthe required relative densities to reach the VSM and aVSMin drained tests. It can be seen that, at embedment ratiosgreater than 2, the transition density between contractionand dilation falls within the range of densities required for the peak resistance to reach the VSM predictions. Further-more, H / D values at which a plateau in the drained breakoutfactors was observed for several relative densities are plotted in the figure, and appear to be consistent with this trend.

DISCUSSIONSoil behaviour

The test results demonstrate that the volumetric behaviour of the backfill soil, which is dependent on its initial state,strongly influences both its drained and its partially drained uplift resistance. In the drained tests, a plateau in the peak

breakout factors observed in very loose soil indicates a possible transition from a sliding-block failure mechanism(which is well captured by the VSM) to a flow-around failure mechanism, where shear planes do not extend to thesurface. The mobilised mechanism appears to be dependenton both relative density and embedment ratio – for example,looser backfills experience a transition at shallower embed-ment ratios – which suggests that this behaviour may bedetermined by a combination of the soil’s critical-state parameters and the burial depth.

Tests conducted at fast rates, inducing partially drained conditions, allowed for relative quantification of dilation(and contraction) through measurement of the pore pressureresponse around the pipe and its contribution to the overalluplift force. Soil that contracts during shear sees the

development of positive excess pore pressure and a subse-quent reduction in peak force, whereas dilation results in theopposite effect. The existence of a localised flow-around mechanism is supported by evidence that, at very low rel-ative densities, contraction appears to occur primarily around the pipe, whereas at higher densities dilation away from pipecontributes to the resistance increase. Accordingly, the den-sity (for a given embedment ratio) at which the transitionfrom net contraction to dilation occurs was determined based on total resistance, rather than on pore pressure measure-ments. This relationship between the volumetric transitiondensity and embedment depth was found to be consistentwith the assumed transition density between failure mechan-isms occurring in the drained tests (as shown in Fig. 15).

Evaluation of prediction modelsFor drained conditions the adapted VSM is an excellent

predictor of the peak resistance for loose sand when thesliding-block failure mechanism is mobilised. This mechan-ism appears to occur for all relative densities up to anembedment ratio of approximately 2, after which an increas-ing density is required to reach the aVSM for increasingembedment depths. The prediction model proposed by Whiteet al. (2001) has applicability for a greater range of densi-ties, as it accounts for dilation along its failure planes.However, as with the aVSM, this model overpredicts theresistance at greater embedment ratios once the flow-around

mechanism is mobilised and the failure planes no longer extend to the surface.

For partially drained conditions, currently the only avail-able prediction model is that proposed by Schupp (2009)

Transition range

0

0·5

1·0

1·5

2·0

2·5

3·0

3·5

4·0

0 10 20 30 40

N ult

Relative density: %(a)

v 0·002 mm/s

v 0·1 mm/s

v 1 mm/s

v 5 mm/s

0

0·5

1·0

1·5

2·0

2·5

3·0

3·5

4·0

102

100

N ult

Uplift speed: mm/s(b)

RD 0%

RD 15%

RD 30%

Fig. 14. Summary of breakout factors with respect to (a) relativedensity and (b) uplift speed ( H / D 3.5)

0

5

10

15

20

25

1·0 1·5 2·0 2·5 3·0 3·5 4·0 4·5 5·0

Relative

density:%

Embedment ratio, /H D

RD to reach aVSM

RD to reach VSM

Transition RD – contraction/dilation

Plateau RD – drained tests

Fig. 15. Comparison of transition densities


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and Byrne et al. (2013), which adds a term to the VSM toaccount for the pore pressure response. This is applicableonly to soil at its loosest state, probably because in this statethe excess pore pressure development appears to occur primarily around the pipe. At higher relative densities the phenomenon becomes more complex, as dilation effectsfurther away from the pipe contribute to the overall upliftresistance.

Implications for design against upheaval buckling The DNV guidelines currently specify two broad ranges

to predict the peak uplift resistance for loose and medium/dense sand in drained conditions, using empirical frictionalfactors (shown in Fig. 5(c)). For the sand tested here, theupper bound of the loose sand region is equivalent to thecalculated aVSM, and virtually all the data for 0–20% RDare contained within this band when it is extended to lower embedment ratios. The data for relative densities of 25– 55% also fit very well within the medium/dense region.Thus, as a broad estimate, the DNV provides very good guidance for the range of breakout factors, although within

these bands more accurate predictions may be required. Thedata from this study also agree well with the DNV-recom-mended range for mobilisation distances to be used infinite-element analyses for upheaval buckling; however, en-gineering judgement should be used when extrapolating thedisplacement values from these model tests to larger-scaleapplications.

Generally, current design practice treats the drained caseas conservative, assuming that if fast rates of uplift were tooccur, the soil would be sufficiently dense to dilate and cause an increase in resistance. The findings of this study,however, show that in very loose sand fast-rate tests maycause liquefaction, and drastically reduce the uplift resis-tance. Given that it is an extremely complex task to quantify

the pore pressure effects across varying densities, and esti-mate the rate of uplift that might occur during buckling, itmakes little sense to attempt to predict this partially drained response. Instead, it follows that the design should stipulatea minimum backfill density for a given embedment ratio,which will ensure that dilation occurs upon shearing of thesoil. This, in turn, means that the flow-around failuremechanism at peak resistance will probably be avoided, and the conservative drained resistance can then be predicted using the aVSM and the DNV guidelines.

CONCLUSIONSThe results from this scale study of pipe uplift resistance

at various embedment ratios, uplift rates and relative densi-ties illustrate the importance of the backfill density incontrolling the failure mechanism and corresponding peak resistance. Given these findings, it is practical to ensure thata minimum relative density of the backfill is reached toavoid the negative effects of contraction of very loose soil.This study identifies approximate transition densities for onesoil (and pipe diameter); further work is required to relatethis transition between mechanisms to measurable param-eters, and thus extend the framework to a variety of soils.Additional research is needed to determine field installation backfill densities for pipes installed by jet-trenching and ploughing, as well as more details of the time-dependentdensification processes after installation. Finally, it is ac-

knowledged that the application of this work in designwould require consideration of scaling issues, as the experi-ments were conducted on smaller-diameter pipes than mighttypically be used in the field.

ACKNOWLEDGEMENTSThe first author gratefully acknowledges the generous sup-

port from the Clarendon Fund at Oxford University. Fig. 3 isreproduced with the permission of the copyright owner, OTC,and further reproduction is prohibited without permission.

NOTATION

D diameter of piped pipe displacement

d f mobilisation displacementemax maximum void ratioemin minimum void ratio F ult peak uplift force

f empirical uplift factor f r empirical post-peak uplift factor

G s specific gravity H initial embedment depth to pipe centre

H c initial embedment depth to pipe crown K lateral at-rest earth pressure coefficient L segmental length of pipe

N ult soil breakout factor (normalised peak uplift force) R uplift force

Rmax DNV peak uplift force

RD relative densityue excess pore water pressure z embedment depth

Æ, shape parameters for DNV tri-linear uplift curveÆf , d fr post-peak shape parameters for DNV uplift curve

ª9 effective unit weight of soilªUR DNV uplift resistance safety factor 9 effective friction angle

peak peak friction anglecrit critical-state friction angleł dilation angle

REFERENCES

Bolton, M. D. (1986). The strength and dilatancy of sands. Gé o-technique 36, No. 1, 65–78, http://dx.doi.org/10.1680/geot.1986.36.1.65.

Bransby, M. F. & Ireland, J. (2009). Rate effects during pipelineupheaval buckling in sand. Proc. Instn Civ. Engrs – Geotech.

Engng 162, No. 5, 247–256.Bransby, M. F., Newson, T. A., Davies, M. C. R. & Brunning, P.

(2002). Physical modelling of the upheaval resistance of buried offshore pipelines. Proc. 1st Int. Comf. on Physical Modelling inGeotechnics, St Johns, 899–904.

Byrne, B. W., Schupp, J., Martin, C. M., Oliphant, J., Maconochie,A. & Cathie, D. (2008). Experimental modeling of the unburial

behaviour of pipelines. Proc. Offshore Technol. Conf., Houston,TX , paper OTC-2008-19573.

Byrne, B. W, Schupp, J., Martin, C. M., Oliphant, J., Maconochie,A. & Cathie, D. (2013). Uplift behaviour of shallowly buried

pipe sections in saturated very loose sand. Gé otechnique 63 , No.5, 382–390, http://dx.doi.org/10.1680/geot.11.P.016A.

Cheuk, C. Y., White, D. J. & Bolton, M. D. (2008). Upliftmechanisms of pipes buried in sand. J. Geotech. Geoenviron.

Engng 134, No. 2, 154–163.DNV (2007). DNV-RP-F110, Global buckling of submarine pipe-

lines: Structural design due to high temperature/high pressure.Baerum, Norway: Det Norske Veritas.

Matyas, E. L. & Davis, J. B. (1983). Experimental study of earthloads on rigid pipes. J. Geotech. Engng 109, No. 2, 202–209.

Schaminée, P. E. L., Zorn, N. F. & Schotman, G. J. M. (1990). Soilresponse for pipeline upheaval buckling analyses: full-scalelaboratory tests and modelling. Proc. Offshore Technol. Conf.,

Houston, TX , paper OTC-1990-6486.Schupp, J. (2009). Upheaval buckling and flotation of buried off-

shore pipelines. DPhil thesis, Univeristy of Oxford, UK.

Thusyanthan, N. I., Mesmar, S., Wang, J. & Haigh, S. K. (2010).Uplift resistance of buried pipelines and DNV-RP-F110 guide-lines. Proc. 33rd Ann. Offshore Pipeline Technol. Conf., Amster-dam, 1–20.


8/19/2019 geot63-09461

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Trautmann, A. M., O’Rourke, T. D. & Kulhawy, F. H. (1985). Upliftforce–displacement response of buried pipe. J. Geotech. Engng 111, No. 9, 1061–1076.

Villalobos, F. A., Byrne, B. W. & Houlsby, G. T. (2005). Momentloading of caissons installed in saturated sand. Proc. Int. Symp.on Frontiers in Offshore Geotechnics (ISFOG 2005), Perth,411–416.

Wang, J., Ahmed, R., Haigh, S. K., Thusyanthan, N. I. & Mesmar,S. (2010). Uplift resistance of buried pipelines at low cover-

diameter ratios. Proc. Offshore Technol. Conf., Houston, TX , paper OTC-2010-20912.

Wang, J., Haigh, S. K., Forrest, G. & Thusyanthan, N. I. (2012).Mobilization distance for upheaval buckling of shallowly bur-ied pipelines. J. Pipeline Syst. Engng Practice 3, No. 4, 106– 114.

White, D. J., Barefoot, A. J. & Bolton, M. D. (2001). Centrifugemodelling of upheaval buckling in sand. Int. J. Phys. Modelling Geotech. 1, No. 2, 19–28.

White, D. J., Cheuk, C. Y. & Bolton, M. D. (2008). The upliftresistance of pipes and plate anchors buried in sand. Gé otechni-

que 58, No. 10, 771– 779, http://dx.doi.org/10.1680/geot.2008.3692.


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