full paper - ratcheting uplift of buried pipelines in sand (p. chitas)

Ratcheting Uplift of Buried Pipelines in Sand / August 2015 / 1

Ratcheting Uplift of Buried Pipelines in Sand

P. Chitas BSc, MSc

Often regarded as the "arteries" of the oil and

gas sector, offshore pipelines form arguably one

of the key contributing factors to the UK energy

industry, and in a larger scale, to the country's

gross domestic product (GDP). Accordingly, any

potential structural failure of a pipeline system

could have serious detrimental consequences,

both in financial and environmental level,

resulting in damage that could be amounted to

millions for the companies, while any

remediation attempt would be proved extremely

uneconomical. Offshore pipe burial, in turn, is

critical to the safety of the system during its

operation lifecycle, as it serves to provide both

protection from seabed hazards and appropriate

temperature levels for its function. One of the

main causes of failure for a subsea pipeline is

termed "ratcheting", a phenomenon described

by progressive loss of cover depth and may occur

after the system exposure to a certain number of

cycles of loading and unloading, as a result of the

high thermal stresses induced by the extremely

hot carried material (i.e. oil and gas). Based on

findings derived from a series of small-scale

laboratory tests at the University of Dundee, the

aims and objectives of the presented paper,

intend to shed light on dominant parameters that

take place during the investigated phenomenon,

validate current prediction methods and estimate

the adequate soil backfill required for sands of

various densities ranging from loose to medium

dense.

LIST OF NOTATION

D pipeline diameter

D10, 60 particle size

Dr relative density of soil

e in-situ void ratio of soil

emax, emin void ratio of soil in its loosest and

densest state

f empirical uplift resistance factor

Gs particle specific gravity

Gsec secant shear modulus

H depth of cover

𝐻 depth from seabed to pipe centre

H/D embedment ratio

IR relative dilatancy index

k soil permeability

K earth pressure coefficient

Ncycles number of cycles

L length of pipe segment

r roughness parameter

Rcyc cycle load amplitude

Rpeak peak uplift resistance

S soil shear resistance component

VSM vertical slip model

W’ submerged weight of soil block

γ dry unit weight of soil

δmob peak mobilisation distance

σc crushing strength of soil grains

(20 MPa for Quartz sand)

σ’h effective horizontal stress

σ’ν effective normal stress

φcrit soil critical state friction angle

φpeak soil apparent peak friction angle

ψ dilation angle of sand


1. INTRODUCTION

One of the primary causes of failure for offshore

buried pipelines is that of upheaval buckling, which

usually leads to a progressive unburial behaviour,

and forms a result of the vertical bending action

induced by the high internal thermal pressures, in

their attempt to find their way out of the section. The

first officially recorded case of buckling failure

occurred in 1986 in the North Sea, when a section

part of a 17 km long embedded pipe, was found to

stand arch-shaped out of the seabed (Nielsen et al.,

1990). As a result, this unburial behaviour can easily

lead to subsequent damage of the protruding pipe

section which then becomes prone to seabed

geohazards and trawling activities. Ever since the

first incident of buckling failure was recorded, a

considerable research amount of literature has been

concentrated on the establishment of a reliable

framework for the prediction of the required soil

uplift resistance which forms key parameter for the

assessment of the performance during the upheaval

buckling event. Nevertheless, the governing

standards for design against pipeline buckling

(DNV, 2007), are focused on the monotonic case,

without providing adequate information about the

cyclic mode of failure, which is commonly known

as “ratcheting”. In this regard, recently immerged

evidence, suggested that buckling failure of the pipe

can still occur after system’s long-term exposure to

a certain number of loading cycles, arising by forces

of lower load amplitudes than the peak design force.

With the exception of some limited literature

references on the subject (Wang, 2012; Bai and Bai,

2014; Yu et al., 2015), there are still no available

codes or criteria that could dictate reliable design

predictions to address the problem. In response to

the existing research gap on the phenomenon, a

series of small-scale pull-out tests of a pipe section

were conducted in dry sand, with the aid of a state-

of-the-art Instron apparatus (i.e. Figure 1.1), for

controlled material testing, so as to capture the

monotonic as well as the cyclic performance of the

soil, under various stress levels, relative densities

and load amplitudes.

Figure 1.1: Experimental setup

Hence, the aim here is the production of a physical

simulation of the “ratcheting” failure mechanism in

order to investigate the controlling parameters

during the phenomenon, validate existing design

approaches and predict the soil cover required to

provide sufficient resistance.

2. EXPERIMENTAL METHOD

In the presented paper, physical modelling was

adopted for the investigation of the phenomenon,

which was preferred over other methods as it

accounts for nonlinearity, while it may also allow

the capturing of certain aspects that govern soil

behaviour which cannot be modelled implicitly by

other approaches (e.g. finite elements). In this

regard, the problem was treated as a two-

dimensional plane-strain case, through the

conduction of repeated 1g pull-out tests on a short

circular pipe section positioned in a steel chamber

box and buried inside dry fine silica sand (HST95).


The performance was tested both under monotonic

and cyclic loading conditions, with the aid of

specialised Instron equipment provided by the

geotechnical lab of Dundee University, which

allowed controlled and precise measurements to take

place. Accordingly, the physical simulations both

for the “upheaval buckling” and “ratcheting” failure

mechanisms were carried out for loose and medium

density sand, and with the consideration of two

different shallow depths of cover corresponding to 1

and 2 pipe diameters respectively.

2.1 Experimental sequence

Initially, monotonic tests were conducted in order to

obtain representative reference values of the soil

peak uplift capacity and the mobilisation distance

for each case of sand and embedment ratio. In this

regard, the monotonic upheaval buckling

phenomenon was simulated through displacement-

controlled static pull-out tests, while maintaining a

constant pulling speed of 1 mm/min, sufficiently low

so as to ignore any rate effects and assure drained

conditions throughout the tests (Powrie, 2014). The

results from the monotonic case form a necessary

requirement in order to be able to proceed to the

cyclic phase, given that the ratcheting failure

mechanism occurs before the peak load phase is

reached.

The next stage, involving the ratcheting mechanism

simulation, was represented by the conduction of

load-controlled cyclic tests, with the application of

load amplitudes corresponding to 50%, 80% and

90% of the capacity, while the period of the sine

wave was kept constant at 30 seconds. The output

results were then imported into Excel for the

production of a hysteresis loop analysis that enabled

thorough examination of the stiffness degradation,

as well as the accumulated displacement history

after each cycle. All data readings were provided by

Bluehill 3, Instron's main software that enables real-

time, accurate and clear result presentation, during

any material testing through the device.

2.2 Soil sample preparation

The selected soil material for the tests was

represented by dry fine silica sand (HST95),

uniformly graded and rounded, with D10 = 0.10 mm

and D60 = 0.14 mm particle size features. The main

soil index properties comprised the critical state

friction angle, φcrit = 32 o, the particle specific

gravity, Gs = 2.63, and maximum and minimum

densities of 1792 and 1487 kg/m3 respectively. The

density states of the sand specimens were chosen so

as to get an approximate resemblance of the medium

to dense in-situ conditions encountered on the

seabed of the North Sea, as well as the loose soil

state after its placement as backfill at the post-

installation stage (Randolph and Gourvenec, 2011;

Liu et al., 2011; Shajarati et al., 2012).

During sample preparation, air pluviation methods

were considered with the aid of specialised pluviator

equipment allowing the required sample densities to

be achieved, through the simultaneous adjustment of

the drop height, the displacement rate, as well as the

apparatus slot opening. The relative density (i.e.

Equation 2.1) was measured and monitored through

the extraction of the in-situ soil density, while

weighing four small molds filled with sand and

estimating the net unit weight of the soil.

minmax

max

ee

eeDr

Equation 2.1

2.3 Testing arrangement

The main testing equipment used during the

conduction of the tests, consisted of an Instron 5900

series apparatus, a small steel-made rigid pipe model

and a 1g steel-made chamber box with a Perspex

side. Initially, the testing apparatus employed for the

conduction of the experiments, was the latest version

of Instron, a device especially designed for material


performance testing, and the ability to provide real-

time and precise load measurements.

The pipeline was represented by a steel-made rigid

pipe section of 90 mm diameter, 300 mm in length

and 3.9 kg unit weight. Extending 360 mm in height,

an iron frame attachment enabled the safe gripping

of the pipe model during the pull-out tests. For the

chamber to accommodate both the specimen and the

pipe model for the tests, a 1g steel-made box was

used. The box size was selected so as to enable

enough space during the tests and minimise the

boundary effects from wall friction to a sufficient

degree. A visual description of the representative

test arrangement is provided by Figure 2.1 below.

(a) (b)

Figure 2.1: Two-dimensional cross-sectional (a) and plan (b) views of the steel chamber

3. RESULTS PRESENTATION AND

ANALYSIS

In the context of the physical simulation, a total

number of 16 laboratory 1g tests were carried out,

including 4 monotonic displacement-controlled and

12 cyclic load-controlled tests. The first set of tests

was conducted in dry medium dense sand of 60%

relative density and the application of shallow

embedment depths equal to 1 and 2 pipe diameters.

The tests were then repeated for dry loose sand

specimen of 27% relative density and under the

same embedment ratios. Initially, the performance

was tested under monotonic loading conditions for

each case, providing an upheaval buckling

simulation, whereas the subsequent tests were

conducted under cyclic loading amplitudes of 50%,

80% and 90% of the peak load, in order to model the

ratcheting phenomenon. Comprehensive details

with regards to the main features of the experiments

carried out, are summarised in the Table 3.1 below.


Test No. Sand Description γ’

(kN/m3) Load Case Load (%) H/D

D

(mm)

01

Medium Dry

HST95 Sand

(Dr = 60%)

16.25

Monotonic 100

1 90 02 Cyclic 90

03 Cyclic 80

04 Cyclic 50

05 Monotonic 100

2 90 06 Cyclic 90

07 Cyclic 80

08 Cyclic 50

09

Loose Dry

HST95 Sand

(Dr = 27%)

15.25

Monotonic 100

1 90 10 Cyclic 90

11 Cyclic 80

12 Cyclic 50

13 Monotonic 100

2 90 14 Cyclic 90

15 Cyclic 80

16 Cyclic 50

Table 3.1: Summary of main test details

3.1 Monotonic test results

3.1.1 Peak uplift resistance

Soil peak uplift resistance, Rpeak, forms vital factor

when pipeline integrity is of interest, and thus, a

variety of contradictory methods and models have

been put forward by the researchers, most of which

seem to comply with the fact that the latter parameter

is mainly controlled by the depth of cover, H, the

unit weight of the soil, γ, the diameter of the pipe, D,

as well as the dimensionless uplift resistance

coefficient, f (White et al., 2001; DNV, 2007).

Generally, the main components that contribute to

the total resistance against upheaval buckling,

comprise the buoyant self-weight of the pipe, the

dead weight of the soil above the pipe, W’ and the

soil shear resistance, S. However, all the approaches

tend to differ with regards to the interpretation of the

parameter H and the uplift factor, f. A schematic

representation of the most widely accepted models

can be seen in Figure 3.1 below.

(a) Vertical Slip Surface Model (b) Inclined Slip Surface Model

Figure 3.1: Failure mechanisms: (a) Vertical Slip Surface Model; (b) Inclined Slip Surface Model;

(Thusyanthan et al., 2010)


The validation of the experimental results was

carried out through the comparison of the derived

outcomes against literature-based formulas,

including approaches of Trautmann, Schaminée,

White and the DNV guidelines, all of which are

based on the aforementioned design models. The

estimation of Rpeak, according to DNV standards

(2007), as well as based on Trautmann et al. (1985)

Vertical Slip Models (VSM), was implemented by

using the Equation 3.1, while predicting differently

the uplift resistance factor, f.

2

21

82

11'

H

D

D

Hf

H

DHDRpeak

Equation 3.1

With regards to the rest of the approaches, the peak

uplift resistance values were derived using relatively

similar formulas (i.e. Equations 3.2 (a) and (b)), with

the exception that Schaminée et al. (1990) takes H

as it is, while White et al. (2001) considers the

weight of the block from the waist of the pipe and

above or otherwise, H = H + D/2. The updated

formula of White et al. (2008) for the soil uplift

capacity, is given by Equation 3.3.

D

HfHDRpeak 1' Equation 3.2(a)

D

HfDHRpeak 1' Equation 3.2(b)

D

Hf

H

DDHRpeak

81'

Equation 3.3

The uplift resistance factor, f, for each case was

estimated in line with the formulas of Table 3.2

below. Where K0 and Kp, are the lateral earth

pressure coefficients for at-rest and passive

conditions respectively, φcrit, the critical state

friction angle of sand, φpeak, the friction angle at peak

stress, ψ, the dilatancy angle, and r is the roughness

parameter taken as -1.

Method Dimensionless uplift resistance estimation Selected f

DNV (2007)

22 )1tantan1(

tantan

rKf

critcrit

critcritP

(for medium to dense sandy backfill)

critcritcritKf tan)sin1(tan0

(for loose sandy backfill)

0.45 (Medium)

0.29 (Loose)

White et al.,

(2001) & (2008)

2

)2(cos)1(

2

1tantantan 00

KK

f peak

0.62 (Medium)

0.41 (Loose)

Schaminée et al.,

(1990) ].50 ,3.0[f 0.40 (Both)

Trautmann et al.,

(1985) critKf tan0 0.29 (Both)

Table 3.2: Dimensionless uplift resistance prediction methods


The values of peak and dilation angles, were

estimated following Bolton's (1986) concepts,

which are founded on correlations that equate

together, soil dilation angle with relative density, Dr,

and the in-situ mean effective stress, p'.

Rcritpeak mI 8.0 Equation 3.4

1'

ln

pDI c

rR

Equation 3.5

The Table 3.3 that follows, encompasses the

obtained Rpeak values normalised by the pipe unit

length, L, as those have been derived from the tests,

and according to the literature-based approaches,

while the percent error indicated with red colour, is

to enable direct proportional comparisons to be

made. Interpreting the following numerical data, it is

observed that all VSM-based methods tend to

produce the highest deviations from the actual

results, while the formula of the current DNV design

standards seems incapable of capturing the actual

performance owing to its empirical nature. In

contrary, the Inclined Surface Model of White et al.

(2008), represents the actual soil behaviour with the

highest levels of precision (i.e. the maximum

percent error is 3%), as it considers the uplift

coefficient as a function of relative density and the

ambient stress levels, accounting, thus, for dilatancy

effects in the estimations.

Medium sand (Dr = 60%) Loose sand (Dr = 27%)

H/D = 1 H/D = 2 H/D = 1 H/D = 2

Method Rpeak/L

(N/m)

Error

(%)

Rpeak/L

(N/m)

Error

(%)

Rpeak/L

(N/m)

Error

(%)

Rpeak/L

(N/m)

Error

(%)

Measured 322.04 - 793.20 - 247.37 - 552.77 -

DNV (2007) 277.88 -16% 646.11 -23% 217.52 -14% 486.18 -14%

White (2008) 329.58 2% 788.01 -1% 251.53 2% 570.75 3%

White (2001) 381.27 16% 839.69 6% 300.03 18% 619.26 11%

Schaminée (1990) 184.28 -75% 473.85 -67% 172.94 -43% 444.69 -24%

Trautmann (1985) 231.78 -39% 518.06 -53% 159.81 -55% 392.19 -41%

Table 3.3: Experimental and theoretical peak uplift resistance comparisons

Figure 3.2 on the left, aims to express

the non-dimensionalised form of the

results from the measured and the

theoretical peak uplift resistance (i.e.

the Rpeak normalized by the parameters

γ, H, D and L), as a function of the

embedment ratio, H/D.

Figure 3.2: Normalised peak uplift resistance vs H/D ratio


3.1.2 Mobilisation distance

The displacement required to mobilise the peak load,

δmob, is critical to the design against upheaval

buckling, as it forms a stiffness dependent parameter

and any underestimation of it could lead to unsafe

predictions (Randolph and Gourvenec, 2011;

Ivanovic and Oliphant, 2014). In this stage,

experimental results of the mobilisation distance are

normalized against the primary controlling

parameters that are the cover depth, H (i.e. Table

3.4), and the pipe diameter, D (i.e. Figure 3.3), so as

to enable direct comparisons with the literature data,

the DNV guidelines, as well as the recommended

framework derived from the Technip UK Ltd

database. The results from the tests, coincide well

with literature findings, whereas the DNV

guidelines provide high underestimations of the

parameter especially in loose sand. This is likely due

to the fact that the DNV method is strictly relying on

H, while ignoring the effects of other profound

factors, such as pipe diameter and soil state

parameters, which can easily lead to

misinterpretations.

Loose sand Medium sand

H/D = 1 H/D = 2 H/D = 1 H/D = 2

Method δmob/H

Measured [1.5%, 2%] [0.8%, 1.7%] [0.9%, 1.1%] [1%, 1.6%]

DNV (2007) [0.5%, 0.8%] [0.5%, 0.8%]

Trautmann (1985) 2.4% - 1.35% -

Thusyanthan (2010) 3.3% 2.7% - -

Wang (2012) 1.7% 2.3% - -

Table 3.4: Normalised mobilisation distance comparisons

Figure 3.3: Normalised mobilisation distance vs H/D for loose sand

R² = 0.4804

0%

1%

2%

3%

4%

5%

0.5 1 1.5 2 2.5 3 3.5

δm

ob/D

H/D

Measured

DNV (2007)

Thusyanthan et al.(2010)

Wang (2012)

Trautmann et al. (1985)

Ivanovic and Oliphant(2014)


3.2 Cyclic test results

3.2.1 Stiffness degradation

The results from the cyclic tests were analysed in

terms of the stiffness degradation and the

accumulated displacement at the end of each

hysteresis loop. Accordingly, the stiffness

degradation per loading cycle was investigated for

each case, through the derivation of the secant

stiffness parameter, Gsec, which forms product of the

peak shear stress amplitude divided by the

corresponding maximum shear strain for each loop.

The secant modulus graphs below (i.e. Figure 3.4)

indicate that most cases of the shallowest cover

depth can be affected by ratcheting, whereas in the

deepest cases the pipe cyclic movement leads to

densification and progressive stiffening of the

backfill. The difference between the responses for

the two embedment cases, can be explained also by

the common argument that soil stiffness scales up

with the square root of effective stress, which, in

turn, increases with depth.

(a) Loose sand

(b) Medium sand

Figure 3.4: Stiffness degradation as a function of cycle number:

(a) Loose sand; (b) Medium sand


3.2.2 Accumulated displacement

Reviewing the diagrammatical results for the

accumulated displacement normalized by the cover

depth from Figure 3.5, it may be seen that ratcheting

can affect the shallowest embedment cases

especially once δmob is exceeded. Also based on pipe

features, the phenomenon may still occur even if the

mobilisation due to ratcheting is as low as 1.2%D.

However, for the deepest embedment ratio the loss

of cover depth due to the residual cyclic

displacements is counterbalanced by the compaction

and stiffening of the soil.

(a) Loose sand

(b) Medium sand

Figure 3.5: Normalised accumulated displacement as a function of cycle number:



3.2.3 Post cyclic capacity

Worth examining at this point is also the post cyclic

available capacity after the application of 100

cycles, as the impact of this process was not adverse

in all cases (i.e. Figure 3.6). In particular, this is

evident in the case of the highest embedment ratio

for loose sand where the cyclic loading enhanced the

overall soil resistance by 157%, factor that indicates

the significant contribution of the effective stress

level to the soil response during the event.

(a) Loose sand

(b) Medium sand

Figure 3.6: Cyclic load-displacement curves and post cyclic capacity:


4. CONCLUSIONS

Based on conclusions drawn from the outcomes of

this study, the following recommendations are

proposed, which could find future application in the

industry:

a) In terms of the monotonic uplift resistance,

Rpeak, the revised formula proposed by

White et al. (2008), is experimentally

proven to be the most reliable

approximation for the prediction of the

parameter in both cases of sands and

embedment ratios. Additionally, the uplift

resistance coefficient, f, should be regarded

as a function of the soil relative density, Dr,

and the mean stress level at depth of

interest, p', and thus, it should be calculated

in line with Bolton's concepts (1986).


b) Regarding the mobilisation distance, δmob,

new suggested ranges are proposed by the

presented study, which fit well against the

most grounded literature findings as well as

the recommended zone defined by the

database of Technip Ltd, lab tests and real-

case studies.

Characterisation Dr (%) H/D = 1 H/D = 2

Loose sand 25-35 δmob[1.5%, 2%]·H δmob[0.8%, 1.7%]·H

Medium dense sand 40-60 δmob[0.9%, 1.1%]·H δmob[1%, 1.6%]·H

Table 4.1: Recommended ranges for mobilisation distance

c) In terms of the ratcheting phenomenon, in

both sands the H/D ratio of 1 is proven to

be the worst case scenario and under high

load amplitudes, the failure mechanism can

occur either when δmob is exceeded or when

the mobilisation is higher than 1.2%D. On

the other hand, for the H/D ratio of 2, the

compaction action of the cyclic pipe

movement has the effect of stiffening the

overlying soil and increasing its post cyclic

capacity, which can be proven favorable

for the long-term performance of the

pipeline.

d) The significant difference in the response

mechanism, which varied with respect to

the embedment ratio, and the fact that the

loss of cover depth was not proven to affect

all cases, as this was counterbalanced by

soil densification, leads to the assumption

that ratcheting is a stiffness controlled

process.

e) Factors observed to have pronounced

influence on the soil response under

ratcheting:

in situ soil relative density, Dr,

cover depth, H,

soil unit weight, γ,

mean effective stress at depth of

interest, p’.

REFERENCES

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Publishing.

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Design due to High Temperature/High Pressure. DNV, Høvik, Norway.

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full paper - ratcheting uplift of buried pipelines in sand (p. chitas)

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