The Impact of Submarine Slides on Pipelines: Outcomes from the COFS-MERIWA JIP

Abstract

This paper presents key outcomes of a 3-year Joint Industry Project funded by 6 Operators on the impact of submarine slides on pipelines. This JIP developed new techniques to simulate slide runout, and assess the resulting loading and deformation of seabed pipelines. The work was distilled into guidance for practical application, which has found adoption on projects.
The JIP spanned (i) characterization of soils at the solid-fluid transition, (ii) computational modelling of slide runout – via depth-averaged and continuum finite element methods, (iii) physical and numerical modelling of slide runout and pipeline impact, and (iv) analytical studies of pipeline response during slide loading. These elements combine to provide an improved practical basis for quantifying the risk associated with slide-pipeline interaction.
To characterize very soft seabed soils, a new geotechnically-based framework was devised based on extensive measurements of different soils. This framework spans the solid-fluid boundary that is crossed as slides evolve into a debris flow and turbidity current. It is shown that the geotechnical link between water content and shear strength extends continuously – with no phase transformation – far into the fluid domain, allowing a single rheology to be applied throughout.
Computational modelling of slide runout used a hierarchy of methods, from large deformation finite element analysis (LDFE) (with rate effects and softening at soil element level), through depth-averaged runout, to energy-based analytical solutions. In some regimes of behavior the simpler methods suffice, allowing efficient use of Monte Carlo methods to tackle uncertainty. More complex runout modes can be replicated by newly-developed LDFE techniques.
From a runout analysis results, pipeline impact loads can be assessed using new solutions for the bearing capacity and drag forces on pipelines developed from numerical and physical modelling, which again unify concepts from fluid dynamics and geotechnics. Finally, simple analytical methods for assessing the structural response of a pipeline to a known slide loading are provided. These solutions allow rapid assessment of the response of a pipeline to a specified slide loading.
These advances improve the methods available for quantitative assessment of slide runout and slide-pipeline interaction, allowing better determination of the resulting geohazard risk.

Full text

OTC-27034-MS
The Impact of Submarine Slides on Pipelines:
Outcomes from the COFS-MERIWA JIP
White D.J. (University of Western Australia, UWA), Randolph M.F. (UWA), Gaudin C. (UWA), Boylan N.P.
(Norwegian Geotechnical Institute, formerly with UWA), Wang D. (UWA), Boukpeti N. (UWA), Zhu H. (Fugro
Advanced Geomechanics, formerly with UWA) & Sahdi F. (Universiti Malaysia Sarawak, formerly with UWA)
Copyright 2016, Offshore Technology Conference
This paper was prepared for presentation at the Offshore Technology Conference held in Houston, Texas, USA, 2–5 May 2016.
This paper was selected for presentation by an OTC program committee following review of information contained in an abstract submitted by the author(s). Contents
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Abstract
This paper presents key outcomes of a 3-year Joint Industry Project funded by 6 Operators on the impact
of submarine slides on pipelines. This JIP developed new techniques to simulate slide runout, and assess
the resulting loading and deformation of seabed pipelines. The work was distilled into guidance for
practical application, which has found adoption on projects.
The JIP spanned (i) characterization of soils at the solid-fluid transition, (ii) computational modelling of
slide runout – via depth-averaged and continuum finite element methods, (iii) physical and numerical
modelling of slide runout and pipeline impact, and (iv) analytical studies of pipeline response during
slide loading. These elements combine to provide an improved practical basis for quantifying the risk
associated with slide-pipeline interaction.
To characterize very soft seabed soils, a new geotechnically-based framework was devised based on
extensive measurements of different soils. This framework spans the solid-fluid boundary that is crossed
as slides evolve into a debris flow and turbidity current. It is shown that the geotechnical link between
water content and shear strength extends continuously – with no phase transformation – far into the fluid
domain, allowing a single rheology to be applied throughout.
Computational modelling of slide runout used a hierarchy of methods, from large deformation finite
element analysis (LDFE) (with rate effects and softening at soil element level), through depth-averaged
runout, to energy-based analytical solutions. In some regimes of behavior the simpler methods suffice,
allowing efficient use of Monte Carlo methods to tackle uncertainty. More complex runout modes can
be replicated by newly-developed LDFE techniques.
From a runout analysis results, pipeline impact loads can be assessed using new solutions for the bearing
capacity and drag forces on pipelines developed from numerical and physical modelling, which again
unify concepts from fluid dynamics and geotechnics. Finally, simple analytical methods for assessing
the structural response of a pipeline to a known slide loading are provided. These solutions allow rapid
assessment of the response of a pipeline to a specified slide loading.
These advances improve the methods available for quantitative assessment of slide runout and slide-
pipeline interaction, allowing better determination of the resulting geohazard risk.

2 OTC-27034-MS
Introduction
The Australian oil and gas industry is moving into deeper waters, further from shore. Around 2007,
when the present Joint Industry Project was conceived, planned gas developments off the continental
shelf of North West Australia faced significant geohazards that had not been encountered in this region
previously. These geohazards included interaction of submarine slides and pipelines, motivating
research to address the uncertainties in engineering assessments of slide-pipeline risk, particularly in the
context of the carbonate soils found offshore Australia, which have unusual geotechnical properties.
There has been extensive research into certain submarine slides systems that lie in oil and gas
development regions, such as the Storegga slide off the coast of Norway (Bugge et al. 1998) and the
margins of the Mississippi delta where mudslides are triggered by major hurricanes (Gilbert et al. 2007).
The submarine slide hazards of Australia’s North West Shelf have received less attention until recently
(Hengesh et al. 2011, 2013, Zhang et al. 2015).
In regions where slides may cross pipeline routes, designers must estimate the likelihood and severity of
future slides, and the consequence of impact by the slide material. Recent reviews of submarine slide
behaviour are presented by Locat & Lee (2002) and Masson et al. (2006). Parker et al. (2009) and Evans
(2010) review the best practices and challenges of engineering practice related to slide-infrastructure
impact behaviour, and highlight the difficulties associated with the quantification of slide strength and
the resulting impact forces.
The process of slide failure, run-out and fluidisation is accompanied by a change in strength (or
mobilised shear stress) of more than three orders of magnitude. Also, the strain rate within the
deforming soil is far higher than the rates in usual geotechnical problems. Figure 1 shows an idealised
cross-section through a seabed slope. A submarine slide is shown with indicative values of the mobilised
shear strength, velocity and density, based broadly on reported values and typical conditions for fine-
grained sediments (Schwab et al. 1996, Marr et al. 2002, Locat & Lee 2002, Masson et al. 2006).
Evaluation of the risk associated with slide – pipeline impact requires assessment of:
 the likelihood of a submarine slide occurring;
 whether the slide (or distribution of probable slides) will impact the pipeline;
 the consequential (distribution of) loading and deformation/strain in the pipeline.
The JIP addressed the latter two steps, focusing on the geotechnical aspects of geohazard assessment.
Figure 1. Schematic of submarine slide behavior showing typical material parameter ranges

OTC-27034-MS 3
Project aims
The overall aim of the project was to address particular technical challenges of assessing the effect of
submarine slides on pipelines. These included evaluating how to:
1. Select a geotechnical model to capture the behavior of a submarine slide, from initiation
through the run-out process, as the material properties change from solid towards fluid.
2. Apply that geotechnical model in a simulation of slide runout, using various computational
techniques (continuum FE, depth-averaged and analytical energy-balance), to yield estimates
of slide geometry and properties on impact with a pipeline.
3. Assess the impact forces on a pipeline caused by a slide with given geometry and material
properties.
4. Assess the resulting strain and deformation in a pipeline caused by these impact forces, also
considering the soil restraint on the pipeline outside of the slide impact.
These four challenges represent key steps in assessments of the impact of a given slide on a pipeline.
The project tackled these aims through physical model tests, using the geotechnical centrifuge facilities
at UWA, and large deformation finite element analysis. In addition, a characterisation study was
performed to investigate the mechanical behaviour of soil across the solid-fluid boundary, and a new
depth-averaged approach was used to allow comparisons with industry standard procedures, and for this
modelling approach to be refined and extended.
The objectives and the associated activities are summarised in Figure 2. The following sections describe
the key activities and outcomes:
1. Soft soil characterization – models across the solid-fluid boundary
2. Modelling of slide runout, including (a) physical modelling, (b) large deformation finite
element analysis, (c) depth-averaged numerical analysis and (d) energy balance methods;
3. Modelling of slide-pipeline interaction, including (a) physical modelling and (b) finite
element analysis, leading to (c) practical calculation methods for design;
4. Modelling of pipeline structural response – simple design expressions.
Potential slide
Volume, weight
Effect on pipeline / infrastructure
Slide loading: magnitude, duration:
contributions from soil strength and inertia
Seabed erosion profile
Predicted slide
runout
Height, velocity,
internal strength
Site investigation
Seabed strength
characteristics
Project focuses on measuring
and quantifying these links, within a
geotechnical framework
Initial volume, strength
slide velocity, height, reduced strength
V
v
H
su1
su2
Centrifuge and numerical modelling,
distilled into simple methods for design
su1
su2
v
F
Slide properties
pipe force, movement, seabed erosion
Slide behaviour
Slide-pipeline
interaction
Figure 2. Project scheme showing targeted areas of uncertainty

4 OTC-27034-MS
Soft soil characterization – models across the solid-fluid boundary
Background
The rheology of soils across a range of water content spanning the conditions relevant to intact soil and
to a weak fluidised submarine slide was studied. A key aim was to devise a basis for characterising the
strength behaviour of both a fluid-like material (relevant to the final run-out stages of a slide) and a solid
material (representative of the intact material) using the same type of material model.
In current engineering practice geotechnical parameters such as shear strength, sensitivity and water
content are used to characterise the constitutive properties of the intact seabed from which submarine
slides form. These parameters form the basis of any assessment of the stability of a slope, and therefore
the tendency for a slide to be triggered. However, in current practice, these parameters are usually set
aside after failure, and the sliding material is then characterised by ‘fluid’ properties – such as yield
stress and viscosity, within a Bingham-type viscoplastic material model. These properties are usually
assessed based on a back-analysis of historic slide features, and are not connected to the geotechnical
parameters derived for the site.
The accuracy of hindcasting and forecasting methods would be improved if a connection could be
created between the geotechnical properties of the intact soil and the material properties that govern the
run-out behaviour as the slide transforms from the intact condition, through a debris flow, into a
turbidity current. To create this connection it is necessary to bridge the domains that are traditionally
termed ‘solid’ and ‘fluid’.
Material models for thick fluids and soft solids
Models for the strength of two-phase materials that are considered to be either thick fluids or soft solids
take into account two principal effects – the rate at which the material is being strained, and the relative
proportions of the solid and fluid fractions.
It is widely recognised that soils sheared at different strain rates show different shear strengths (Dayal &
Allen 1975, Biscontin & Pestana 2001). The equivalent relationships for a fluid have evolved from
Newton’s law of viscosity and in thick fluids, the shear stress, , must reach a minimum (yield) value,
y, before flow commences. This behaviour can be described by the general relation
n
y  (1)
where μ represents the viscous property of the fluid. With n equal to 1, Equation (1) represents the two-
parameter Bingham model, which assumes a linear variation of the shear stress with shear strain rate,
once the yield stress is exceeded. In its general three-parameter form, Equation (1) is referred to as the
Herschel-Bulkley model. The Herschel-Bulkley model, with y = 0 is exactly equivalent to the soil
mechanics power law model (e.g. Biscontin & Pestana 2001). The modelling of strain rate effects
therefore has obvious similarities within the solid and fluid domains.
The treatment of moisture content, w, or the relative proportions of the solid and fluid fractions, is not so
directly comparable, although similarities exist. For example, critical state soil mechanics theory leads to
exponential relationships between moisture content and shear strength (e.g. Wroth & Wood 1978) that
are well established for remoulded clays via index tests and the liquidity index parameter (LI). Within
the fluid mechanics literature, the equivalent behaviour is often described by empirical relationships that

OTC-27034-MS 5
link the yield stress and viscosity to the volumetric concentration of solids, Csv (O’Brien & Julien 1988,
Major & Pierson 1992). Csv is the quantity commonly used within the fluid dynamics literature instead
of moisture content. These relationships are often in the form of exponential functions such as
sv1sv2 C
1
C
2y e,e   (2)
with the parameters 1, 2, 1, 2 being determined by laboratory test measurements. Although the
functional form is different from the soil mechanics approaches, the general trends are similar. In order
to apply these relationships to submarine slides, in which material transitions from a solid to a fluid, the
key issue is whether these relationships are continuous across the solid-fluid boundary, or whether a
distinct phase transformation exists, marking a step change in strength properties as the moisture content
of a given solid-fluid mixture changes.
The JIP undertook a range of laboratory tests to characterise the strength behaviour of different soft soils
across the solid-fluid boundary. These tests included T-bar and ball penetrometer, vane shear,
viscometer and fall cone tests. The three soils were Burswood clay (a natural clay from nearby Perth),
kaolin clay, and a natural carbonate silt from offshore Australia. These soils were tested at 6-8 levels of
moisture content, with the tests being performed at up to 8 different rates (spanning approximately 2
orders of magnitude of strain rate). Theoretical solutions were used to link the rate at which the tests
were performed to a representative strain rate for the deforming soil. This allowed the effects of strain
rate to be identified across a wider range of strain rates, due to the different ranges achievable in each
type of test.
The principal observation from this study, which is described in more detail by Boukpeti et al. (2012a,
2015), was that there is a continuous relationship between moisture content – via whichever measure is
chosen – and shear strength across the solid-fluid boundary (Figure 3). The tests spanned a range of
liquidity index from 0.7 to 6 and a range of shear strength from 0.01 kPa to 10 kPa. For all three
soils, the variation in strength, su, with moisture content (or the related parameters, LI and Csv) was well
captured by any of the following three forms of equation:
1
b
u1
saw

 (3)

2
b
u2
saLI

 (4)
3sv
b
C
u3
sae (5)
where different values of the parameters a1-a3 and b1-b3 were found for each soil.
A notable difference between the soils is that at high moisture contents the carbonate silt is significantly
weaker than the other soils tested. This may be due to the lack of electrochemical effects present. Soil
particles with a clay mineralogy have attractive forces arising from the ionic layer present on the surface
of clay particles. In contrast, carbonate particles do not feature this effect. This implies that for a given
level of water entrainment – during a submarine slide, for example, but also during disturbance during
pipeline laying – a carbonate soil will be more susceptible to a loss of shear strength due to water
entrainment.
In all soils tested, the transition in the governing behaviour is gradual and continuous, and there is no
evidence of a distinct phase transformation at a boundary between regions that might be termed solid or
fluid behaviour.

6 OTC-27034-MS
The shear strength measurements obtained for different strain rates confirmed that the strength of the
material was higher at higher rates of shearing (Figure 4). This non-linearity of the strength – strain rate
relationship favours the use of the Herschel-Bulkley model (Eq. (1)), with normalized strain rate
parameters, over the simpler Bingham approach; or alternatively, a logarithmic relationship as shown in
Figure 4. These rate-dependency parameters did not vary across the range of moisture contents tested.
As a result of this study, a model has emerged that combines the influences of strain rate, remoulding
and changes in moisture content across gross changes in the operative soil strength, su-op, – spanning the
solid-fluid boundary. The model takes the general form


95
1ξ3ξ-
remrem
ref
1op-u eδ1δ
γ
γ
1s 

















n
b
wa 


(6)
where rem is the inverse of the soil sensitivity and , 95 are the cumulative shear strain and that required
to achieve 95 % softening. The final bracket links the accumulated plastic strain with the degradation
from the intact to the fully remoulded strength, following the approach used widely in numerical
analysis of large deformation problems (Wang et al. 2015). Parameters for a specific soil can be derived
from laboratory or in situ tests using penetrometers and viscometers, as well as conventional index tests.
0123
Water content, w
0.001
0.01
0.1
1
10
Shea
r
st
r
ength, s
u
(kPa)
kaolin
Burswood
carbonate soil
0123456
Li
q
u
i
d
i
t
y
i
n
d
ex, LI
0.001
0.01
0.1
1
10
100
Shea
r
st
r
ength, s
u
(
k
Pa)
kaolin
Burswood
Critical State Line
carbonate soil
s
u
= 170 exp(

4.6
.
LI)
Figure 3. Effect of water content on soil strength across the solid-fluid boundary – comparison of three soils
Figure 4. A unified relationship between strain rate and soil strength – comparing data obtained with kaolin tested across 5
orders of magnitude of strain rate and 3 orders of magnitude of strength

OTC-27034-MS 7
Equation (6), with a single set of the material parameters, can characterise the soil response across three
orders of magnitude in strength, and five orders of magnitude variation in strain rate. These relations are
amenable to inclusion in numerical modelling of large deformation processes. In these analyses, a
simple Tresca or von Mises failure criterion is commonly adopted for the material strength, updating the
strength according to the current strain rate or accumulated strain (e.g. Zhou & Randolph 2007). This
study highlighted how changes in moisture content could also be incorporated in the same numerical
modelling strategy, so the transition from solid to fluid can be captured without the need to consider
multiple material phases. The benefit of this unified framework is that it allows the full process of slide
triggering, run-out and resedimentation to be simulated without needing to distinguish between different
material phases, and without needing to discard the geotechnical framework that underpins conventional
engineering techniques for site characterisation and the design of seabed infrastructure.
Modelling of slide runout
Physical modelling of slide runout
An ambitious aim of this project was to simulate submarine slides in the geotechnical drum centrifuge at
UWA, to provide slide runout observations for well-characterised material in controlled conditions, to
compare with numerical analysis. Centrifuge modelling has not been widely used to model submarine
slides, despite being a common technique for simulation of other aspects of offshore geotechnics.
The majority of previous experimental studies of submarine slides begin with a fully fluidised slide
mass, with a strength of < 100 Pa (Mohrig et al., 1999; Mohrig and Marr, 2003; Mohrig et al., 1998 and
Ilstad et al. (2004a; 2004b; 2004c)). There had previously been no experiments that simulated the
transition from the intact (but post-failure) state to the fluidised condition, preventing a connection being
made between the in situ geotechnical properties and the fluidised behaviour in the subsequent run-out.
The experimental arrangement allowed slides to be triggered from an intact block within the drum
centrifuge channel, with the block then running out over an erodible seabed (Figure 5). A block of clay
was consolidated in a release box from a slurry. After consolidation, the box was rotated to an inclined
position, so that the block lost potential energy when the slide was initiated. The slide was triggered by a
door in the box being rapidly pulled upwards whilst concurrently a heavy mass was released to push
against the rear of the slide.
Figure 5. Arrangements for slide run-out modelling in drum centrifuge (Boylan et al. 2010)

8 OTC-27034-MS
The observed run-out distances, from the centre of the intact block to the toe of the run-out debris,
ranged from 30.7 m to 119 m (in prototype dimensions). The tests were rather complex, involving high
speed events with measurements of pressure, forces and displacements over a very short period
(typically lower than 200 ms), as well as concurrent high speed image capture. A novel high-speed
wireless data acquisition system was developed for this project (Gaudin et al. 2009). To differentiate the
seabed from the run-out material and to evaluate potential seabed erosion, a dyeing technique was
developed to colour the seabed (Sahdi et al. 2010).
Example profiles of post-test slide debris geometry and moisture content are shown in Figure 6. The
slides are thinner at the toe and show increasing moisture content with increasing runout distance –
indicative of water entrainment. The post-slide consolidation of the debris exceeded the wetting of the
material during the slide event (Figure 6b). The relict slide debris is denser, i.e. at lower moisture
content, than prior to failure because the undrained failure during run-out is accompanied by the
generation of positive excess pore pressure (because the slide material is initially on the wet side of the
critical state). Subsequent reconsolidation of the debris material leads to a reduction in moisture content
as the excess pore pressure dissipates. The final state is therefore denser than the initial state, and also
denser than the normally consolidated state at the same stress level (indicated by point A on Figure 7),
even if water is entrained during the run-out. This observation, and the theoretical background in Figure
7, should be considered when interpreting the present-day strength measurements of relict slides. A
similar effect has been hypothesised by Palmer (1997) as a tool for identifying relict iceberg scours
beneath the seabed.
Visual examination of the run-out material revealed morphological differences between the run-out
slides. At very low initial strength (< 1 kPa), the depositied material appeared homogenous and there
was little evidence of distinct features on the surface. At higher initial strength (> 1 kPa), a zone of
distinct compression ridges was evident at the terminal lobe, perpendicular to the run-out direction
(Figure 8a). The compression zones were similar in geometry to those identified in the field, such as the
examples shown in the southern flank of the Storegga slide, offshore Norway (Gafeira et al. 2007). At
very high initial strength (> 10 kPa), retrogressive wedge failures were observed (Figure 8b), similar to
retrogressive wedges identified in the upper part of the Storegga submarine slide (Bryn et al. 2004).
Figure 6. Example post-slide surface profiles and moisture content distribution

OTC-27034-MS 9
The mobility and scale of the centrifuge slides is characterised by the run-out ratio, H/L, and the slide
volume, V. To illustrate the relevance of the model slides to field conditions, these characteristics can be
compared with relict slides observed in the field (Elverhøi et al. 2005, Figure 9). These centrifuge results
fall within the range of field observations of submarine slides, albeit at the lower end of the volume
scale due to the limited volume of sliding material. Given that the overall behaviour appears broadly
representative of field conditions, the local measurements of changing strength and water content
provide valuable evidence to support the development of models for simulating these phenomena that
will ultimately be applicable to field scale events.
Figure 7 Simplified stress-specific volume path during slide run-out and debris reconsolidation
Figure 8 Morphological features in centrifuge models (a) compression ridges at toe (b) retrogressive wedge failures at scarp

10 OTC-27034-MS
Figure 9. Relationship between runout geometry and slide size: centrifuge results vs. field data (H is the maximum vertical
height of the debris flow source above the deposit, and L is the horizontal distance from source to deposit)
Numerical modelling of slide runout – large deformation finite element analysis
The two main objectives of the numerical modelling of slide runout in this project were to
 Investigate the feasibility of using modern large deformation finite element (LDFE) analysis to
model slide run-out, allowing conventional geotechnical soil models to be used for the flowing
material and the seabed.
 Develop a depth-integrated numerical approach (similar to BING and its variants, Niederoda et
al. 2003) with extended flexibility in terms of rheological models for the slide material and the
geometry of the seabed and the initial slide block.
The second objective is discussed in the following section.
The large deformation finite element (LDFE) approach developed during this project was a variation of
the so-called ‘remeshing and interpolation technique with small strain’ (RITSS) originally developed by
Hu and Randolph (1998). The overall scheme of the RITSS is to divide the overall displacements of the
continuum element into a series of incremental steps. The displacements of the element in each step
have to be small enough to avoid gross distortion of the soil element, in order that each small strain step
is accurate. A Lagrangian calculation is thus performed in each increment, followed by remeshing the
deformed geometry and Eulerian ‘convection’ of the stresses and material properties from the old mesh
to the new mesh.
The approach was implemented with the commercial finite element package, ABAQUS. A master
program, coded in Fortran, calls ABAQUS and the subroutine that performs interpolation of the field
variables. Several Python (the script language built into ABAQUS) files were written to extract the field
variables from the result files and to control the mesh regeneration.
For analysis of submarine slides, the technique was extended from static to dynamic analysis, using the
techniques described by Wang et al. (2013). In addition, a technique termed ‘element addition’ was

OTC-27034-MS 11
developed to improve the computational efficiency of both static and dynamic LDFE analyses that
involve moving boundaries. The RITSS approach is based on frequent mesh generation to avoid element
distortion. In dynamic RITSS, the field variables mapped from the old to the new mesh involve not only
the stresses and material properties, but also the nodal velocities and acceleration. Using the element
addition technique, new soil elements are attached to the domain boundaries periodically when the soil
near the boundaries becomes affected by large displacements of the sliding material.
The robustness of the techniques was validated and assessed through example problems including
vibration of an elastic plate, large-amplitude lateral displacement of a seabed pipeline, and movement of
a non-deforming landslide down a slope (Wang et al. 2013). Rate-dependent and softening undrained
shear strength was implemented in the LDFE analyses.
A parametric study was then undertaken considering slide mass running out over a firm (non-deforming)
seabed (Figure 10). The varied parameters included (i) initial strength (su), (ii) sensitivity (St), (iii)
brittleness (95), (iv) slide size (H, L) and (v) seabed slope (β). In addition, LDFE analyses that
represented each of the centrifuge model slides were performed using the relevant geometry and soil
properties.

Figure 10. Initial geometry explored in parametric study of slide run-out behaviour
The resulting mechanisms of runout fell into three distinctive patterns comprising (i) elongation (a single
stable slide mass), (ii) breakaway (where the head of the slide mass separated from the following
material) and (iii) block sliding (where the slide mass underwent little internal deformation but
accelerated down the slope in an unstable – accelerating – manner). The overall length of run-out varied
from as little as 20 m to over 2 km (at the point where the analysis had to be terminated – due to the
slide not decelerating even after several days of computation).
Example results highlighting these different mechanisms are shown in Figure 11 and Figure 12. Velocity
contours for an elongation case are shown for different stages of run-out in Figure 11. The LDFE shows
successive overtopping of the toe material by faster moving material behind the toe (see material at -
60 m at t = 10 s, and at -252 m at t = 50 s). Contrasting results for a breakaway case are shown in Figure
12 where separation of the faster moving front end of the slide occurs at times of 25.2 s (at -55 m) and
29.2 s (at -240 m).
Space constraints prevents a full description of all observations, but additional information is provided
in White et al. (2011). The parametric study showed that run-out length is a function of several
combined factors. It increases with decreasing interface strength, increasing slope inclination and
increasing height of the slide mass. These are generally intuitive. More surprising is the influence of the
internal shear strength. A higher internal shear strength of the soil mass (e.g. due to low sensitivity) can
limit the internal shearing (and thus elongation) resulting in a higher ratio of driving force to basal
resistance and a consequently longer run-out.
Elongation tends to occur when the basal resistance along the seabed is relatively large compared with
the driving force. For example, ductile elongation is more likely to occur as the interface shear strength

12 OTC-27034-MS
increases or the slope inclination is reduced. Decreasing the mobilised soil strength in the mass has a
similar effect to increasing the interface shear strength, since that leads to greater elongation and thus
increased sliding resistance.
The LDFE study demonstrated the capability of modern continuum-based solid-body finite element
techniques to mimic submarine slide run-out behaviour. This has been proven using a technology that is
already widely used for geotechnical boundary value problems involving gross deformations and
changes in geometry. This opens up the possibility of modelling submarine slide run-out using soil
constitutive models and input parameters that are consistent with geotechnical site characterisation
studies, albeit through considerable computational effort.
Figure 11. Velocity distributions (in m/s) for LDFE runout case with elongation and progressive failure developing in the
middle slide zone
Run-ou
t
(
m
)
Slide thickness (m)
-300 -250 -200 -150 -100 -50 050
0
2
4
6
Vel
9.0
8.0
7.0
6.0
5.0
4.0
3.0
2.0
1.0
Case 2, t = 10 s
Run-out
(
m
)
Slide thickness (m)
-65 -60 -55 -50 -45
0
2
4
6
Scaled size
Run-ou
t
(
m
)
Slide thickness (m)
-300 -250 -200 -150 -100 -50 050
0
2
4
6
Vel
6.0
5.0
4.0
3.0
2.0
1.0
Case 2, t = 60 s
Run-out (m)
Slide thickness (m)
-310 -305 -300 -295 -290
0
2
4
6
Scaled size
R
u
n
-ou
t
(
m
)
Slide thickness (m)
-300 -250 -200 -150 -100 -50 050
0
2
4
6
Vel
7.0
6.0
5.0
4.0
3.0
2.0
1.0
Case 2, t = 50 s
Run-ou
t
(
m
)
Slide thickness (m)
-258 -256 -254 -252 -250 -248
0
1
2
3
Scaled size

OTC-27034-MS 13
A remaining key limitation of the LDFE approach is that the influence of water entrainment and
changing moisture content, as identified in the centrifuge model tests, is not yet explicitly accounted for.
It can be incorporated implicitly by adopting a higher sensitivity parameter than is measured in
undrained (constant moisture content) conditions but this form of correction is somewhat artificial. The
effects of water entrainment will presumably be more significant near the top and base of the slide –
where water can enter the main body of material – and will be more significant in material that is more
susceptible to cracking and water ingress.
Run-ou
t
(
m
)
Slide thickness (m)
-250 -200 -150 -100 -50 050
0
2
4
6
Vel
6.0
5.5
4.9
4.4
3.8
3.3
2.7
2.2
1.6
1.1
0.5
Case 3, t = 10 s
Run-ou
t
(
m
)
Slide thickness (m)
-250 -200 -150 -100 -50 050
0
2
4
6
Vel
6.0
5.4
4.8
4.2
3.6
3.0
2.4
1.8
1.2
0.6
0.0
Case 3, t = 25.2 s
Run-ou
t
(
m
)
Slide thickness (m)
-250 -200 -150 -100 -50 050
0
2
4
6
Vel
5.0
4.5
4.0
3.5
3.0
2.5
2.0
1.5
1.0
0.5
0.0
Case 3, t = 29.2 s
Figure 12. Velocity distributions (in m/s) for LDFE runout case with repeated breakaway of outrunner blocks

14 OTC-27034-MS
Depth-averaged numerical modelling of slide runout
Numerical modelling of debris flow run-out is commonly modelled using finite difference flow models
based on depth-integrated equations of mass and momentum conservation, solved within a Lagrangian
framework. Examples of this approach include the program BING, developed by Imran et al. (2001).
Except for changes in the shear strength due to viscous strain rate effects, the shear strength remains
constant during the runout process. Although the BING software is freely available, it is not possible to
alter the rheological model or the other elements of the program.
For this project, coding was developed in MATLAB® entitled UWA-SM3 (Submarine Mass Movement
Modeller), which is described in Boylan and White (2016). This code allowed for the implementation of
additional features that are not included in BING, and permitted this project to explore in more detail the
underlying behaviour of the debris flow (e.g. different constitutive models, fluid drag on the slide front
and strain softening behaviour), with full control of the input and output.
The key building blocks of the depth-averaged numerical techniques for simulating slide run-out, such
as the BING program, are the vertical profiles of velocity and mobilised shear strength at any ‘slice’
within the slide. New solutions for these building blocks have been generated for additional forms of
rheological model not previously explored (Boukpeti et al. 2012b). These solutions provide a simple
method to explore the loading on a pipeline that is oriented across the path of the slide, for a given slide
height and speed, and given rheological properties. The solutions show the relative contributions of the
soil strength and the inertial drag, as well as the influence of the vertical position of the pipeline within
the slide (Figure 13).
04080120
Shear strain rate, (s
-1
)
0.0
0.4
0.8
1.2
1.6
2.0
Shear stress,

(kPa)
Bingham B1, B2, B3
Herschel-Bulkley
power
log
kaolin measurements
B2 B1
B3

b
= 1.28
12 13 14 15 16 17
q
(
k
Pa)
0.0
0.5
1.0
1.5
2.0
y (m)
B2
HB
power
log
y = 0.125
(a) (b)
Figure 13. Analytical modelling of steady slide flow: (a) rheological models fitted to test data, (b) consequent loading on a
pipeline as a function of distance above the rigid (non-eroding) seabed level (Boukpeti et al. 2012b)
In UWA-SM3, the debris is modelled using either a Bingham rheology, where the mobilised shear
strength varies linearly with the shear strain rate, or the Herschel-Bulkley model where the relationship
is non-linear. Softening of the slide material can also be included using parameters that are obtained
directly from the results of cyclic full flow penetrometer tests, while hydrodynamic drag forces on the
runout material can also be applied.
Figure 14 shows an example of the run-out of a slide down a 5º slope, modelled using the UWA-SM3
software, illustrating the influence of softening effects. Cases are shown for material with a shear
strength that remains intact (or constant), su-Intact, during the run-out, and for a softening case (su-With

OTC-27034-MS 15
Softening) so that the shear strength degrades towards the remoulded value. In this case, inclusion of
softening increased the run-out distance by a factor of 2.4 (Figure 14a).
Figure 14b shows the evolving slide shape for the softening case. The initial slide block collapses very
early and the run-out stretches almost uniformly. As well as providing information on the run-out
distance and the geometry of the debris, the analysis also provides the velocity of the slide mass. Figure
14c shows the time-velocity relationship for the front of the slide and at two fixed points relative to the
toe of initial slide block – 40 m and 150 m ahead of the initial toe position. These results highlight how
the time-velocity profile seen by an obstacle differs depending on the location within the slide run-out
path. For assessment of the impact loading from slides on infrastructure, the time-velocity profile at the
position of the asset should be considered rather than solely the toe velocity.
50 100 150 200 250 300
Position along run-out (m)
2
4
6
8
10
Height (m)
(b)
2
4
6
Velocity (m/s)
Front
40m from toe
150m from toe
10 20 30 40 50 60 70 80
Time (sec)
(c)
Evolution
of runout
Initial slide block
Final
geometry
50 100 150 200 250 300
Position along run-out (m)
2
4
6
8
10
Height (m)
Initial slide block
Run-out - s
u-Intact
Run-out - s
u-With Softening
(a)
Figure 14. Example analyses of submarine slide run-out, using UWA-SM3 depth-averaged software (a) Comparison of
runouts (b) Snapshots of softening case during runout (c) Time – Velocity profiles (White et al. 2013)

16 OTC-27034-MS
A second example is shown in Figure 15 using a cross-section profile obtained from a seismic survey of
a location from the North-West Shelf, offshore Australia. The water depth at the left hand extremity is
approximately 700 m, and the scarp has a maximum slope of 5, at the location of the initial slide.
The initial slide geometry was defined by examining the continuity of layers within the profile to
identify the level at which failures appear to be taking place. Different run-out cases were considered,
with (i) non-softening intact rate-dependent soil during run-out, (ii) soil softening, and (iii) inclusion of
soil softening and hydrodynamic drag effects. The results showed that inclusion of softening resulted in
greater run-out than for cases where the soil remained intact, while inclusion of both softening and drag
effects resulted in a shorter run-out than with softening alone. Whilst both of these results are intuitive,
the benefit of the analysis is that the relative influence of these effects can be quantified, and further runs
can be performed to explore the sensitivity of the response to the assumptions made when devising the
input parameters. These results illustrate how the additional features included in UWA-SM3 can
influence the run-out distance and their relative effects.
Figure 15. Example of debris flow run-out on cross-section of North West shelf, Australia using UWA-SM3
In practice, the process of predicting the run-out of a potential debris flow carries with it various levels
of uncertainty (i.e. volumes of debris mobilized, range of rheological properties, evolution of debris
properties during run-out). This uncertainty and the subtle interplay between these parameters
complicates the process of assessing whether a potential debris flow meets a particular characteristic (i.e.
passes a particular location, or applies an unacceptable impact to infrastructure).
To overcome this challenge, UWA-SM3 has the feature to conduct Monte Carlo type run-out
assessments where probability distributions describing the uncertainity attached to each parameter are
used as input. Many thousand run-out assessments are carried out, each using a random selection of
input parameters to evaluate all the different possibilities. By quantifying the uncertainty attached to
each input parameter in terms of probability distribution, the outputs of these simulations can be used to
assess the probability of specific aspects of these events. Figure 16(a) shows an example of a Monte
Carlo assessment, showing the variation in the slide toe velocity versus slide toe position for a large
number of simulations. These simulations used a natural seabed topography, which led to the irregular

OTC-27034-MS 17
initial acceleration of the slides. An example of the type of output produced by this analysis is on Figure
16(b), which shows the probability density function (PDF) related to the final slide toe position.
Alternative outputs include the slide thickness and velocity when passing a particular location where a
pipeline will be crossed. This information can in turn be converted into a distribution of expected
pipeline stress, using the methods described later in this paper.
Figure 16. Example of Monte-Carlo analysis using UWA-SM3 (a) Velocity at slide toe as a function of postion, (b)
Probability distribution of final position of slide toe
UWA-SM3 has also been used to examine the response of a flexible pipeline to the impact of a debris
flow normal to the pipeline. Using analytical expressions described later in this paper, the deflection of
the pipeline is tracked dynamically using an equation of motion that accounts for the portion of the slide
loading that is resisted by the membrane tension and internal bending moments in the pipeline. Figure
17 shows an example of the deflection calculated using the UWA-SM3 routine and the conventional
static approach, in which the instantaneous maximum slide load is used to assess the potential pipeline
deflection. Using the dynamic analysis, the deflection of the pipeline is 40% of that calculated using the
conventional approach.
The difference between the two results is due to the temporary nature of the maximum slide loading.
The pipeline does not instantaneously assume the equilibrium geometry that this load corresponds to,
but instead accelerates progressively, at a rate attenuated by the inertia of the moving pipe, ultimately
reaching a smaller displacement within the time-scale of the slide loading. This example analysis
highlights the conservatism of the conventional design approach, which is to perform a static analysis
based on the transient maximum load.
Using the UWA-SM3 software, a parametric study was undertaken to examine the effect on run-out
length of slope angle, initial slide volume, strength properties and the rate-dependency of the debris
strength. Back-analysis of the run-outs has shown that a simple energy balance approach compares
reasonably with predictions of the UWA-SM3 results. In the energy balance approach, the run-out
distance of the slide toe is estimated by assuming that the slide fails by uniform elongation, with the rear
of the slide fixed at the initial position (which is consistent with many of the LDFE analyses). Plastic
work is done through basal sliding and internal shearing, whilst potential energy is lost as the material

18 OTC-27034-MS
falls. This simple energy balance is a reasonable idealisation of slides that fall into the ‘elongation’ class.
It is not appropriate for slides that fail by block sliding (in which the rear is not anchored) or through
‘breakaway’ of successive outrunner blocks.
Geometric scaling relationships from the literature, linking slide size and run-out lengths based on field
observations gathered in the area of the Storegga slide off Norway, also show good agreement with the
trends shown by the UWA-SM3 results and the energy-based back-analysis. These tools – the depth-
averaged software and the simpler energy-based approach – therefore provide a basis for extending these
relationships from purely geometric considerations, to incorporate the influence of soil properties.
An extended version of the depth-averaged approach was developed following the JIP (named UWA-
SM4, White et al. 2013), which solved the governing equations using a smoothed particle
hydrodynamics approach and included lateral spreading of material over a three dimensional domain.
Example results for a slide spreading over a realistic seabed topography are shown in Figure 18.
0 0.5 1 1.5 2
Time (mins)
0.5
1
1.5
2
2.5
3
No
r
malised pipe displacement, y
pipe
/B (-)
Pipeline
Deflection from
static calculation,
y
max
/B
Figure 17. Comparison of pipeline deflection calculated using UWA-SM3 and conventional static calculation
Figure 18. Example of submarine slide runout calculated using UWA-SM4 (White et al. 2013)

OTC-27034-MS 19
Modelling of slide-pipeline interaction
Introduction
The prior sections have explored methods for assessing whether a slide will reach a given loacatoin, and
what the thickness, velocity and properties of the slide material will be in that event. To convert these
parameters into distributed loads on a pipeline requires solutions to link the strength, density and
velocity of the slide material to the net pressure (or force per unit length) on an impacted pipeline.
Activities on this topic included the following components:
1. Physical modelling of slide-pipeline interaction (for flow over a fully engulfed pipeline)
2. Numerical modelling of slide-pipeline interaction (to explore the influence of rate effects)
3. Collation of published data to develop an approach for non-perpendicular attack
They led to an approach for the drag on engulfed pipelines, taking proper account of the influences of
both the strength and self-eight of the flowing material, and also strain rate effects on strength.
Physical modelling of slide-pipeline interaction
A new experimental apparatus was developed to allow a buried model pipeline to be translated at
various velocities through a soil sample contained within the drum centrifuge channel, simulating a pipe
engulfed within a submarine slide flowing perpendicular to the pipe axis. By using a soil sample that
was initially unconsolidated, the tests were performed at different consolidation levels leading to varying
sample properties (density ρ and undrained strength su).
A total of 37 model pipe translation tests spanning a wide range of velocities and soil strengths. The
tests were designed so that the relative influence of ‘fluid’ drag and ‘solid’ bearing resistance varied.
This was achieved by covering 4 orders of magnitude of pipe velocity (from 0.0036 – 4.2 m/s) and two
orders of magnitude of soil strength (from 0.08 to 1.7 kPa), measured via extensive cyclic T-bar tests at
different degrees of consolidation (Sahdi et al. 2014a; Sahdi et al. 2014b). In all tests the pipe was
translated horizontally, at an embedment with the crown of the pipe located two diameters below the soil
surface.
The measured net lateral pressure on the pipe, qH, was affected by both the velocity and the soil strength,
but with no unique trend evident for either parameter (Figure 19). This indicates the need to separate the
resistance into two components to capture both the inertia-dominated and strength-dominated domains.
0.1
1
10
100
0.001 0.01 0.1 1 10
Horizontal pressure on model pipe, q
H
(kPa)
Horizontal velocity, v (m/s)
1-PS1 to 1-PS4 3-PS1 to 3-PS3
4-PS1 to 4-PS3 5-PS1
6-PS1 8-PS1 to 8-PS5
10-PS1 to 10-PS6 11-PS1 to 11-PS4
12-PS1
0.1
1
10
100
0.1 1 10
Horizontal pressure on model pipe, q
H
(kPa)
s
u-op
(kPa)
1-PS1 to 1-PS4 3-PS1 to 3-PS3 4-PS1 to 4-PS3
5-PS1 6-PS1 8-PS1 to 8-PS5
10-PS1 to 10-PS6 11-PS1 to 11-PS4 12-PS1
N
H
= 7.35
Figure 19. Drag forces during lateral pipe-slide interaction, for a range of strengths that cross the solid-fluid boundary

20 OTC-27034-MS
A better interpretation of the results links qH to both the operative soil shear strength su-op (adjusted for
strain rate effects) and also the soil density ρ and relative pipe-soil velocity v (which can be combined
into the non-Newtonian Reynolds number, v2/su-op).
Using this approach, all results follow a unique trend line (Figure 20) that supports a general expression
for active slide pipeline loading:
opuHDH sNvC 
 2
2
1
q

(7)
The full data set is predicted well using a unique combination of a bearing capacity factor NH of 7.35
(which is approapriate for this level of pipe embedment, Martin & White 2012) and a drag factor CDrag
of 1.06. Additional data is also shown on Figure 20 from flume tests and centrifuge tests reported by
Zakeri et al. (2008) and Zakeri et al. (2011), and fits the same trends with appropriately modified
parameters (see Sahdi et al. 2014c for further discussion).
The establishment of Equation (7) as a general approach for assessing active slide-pipeline loading
rationalizes the previous inconsistency between fluid-based (with only the CD term) and geotechnically-
based (with only the NH term) approaches. The components of loading from inertial drag and from the
material strength may be superposed additively, with the material strength being adjusted to account for
the imposed strain rate. This approach is robust in the limits of weightless or inviscid flow, and fits the
wide database of experimental results across the solid-fluid boundary.
1
10
100
1000
0.00001 0.0001 0.001 0.01 0.1 1 10 100 1000
Horizontal bearing capacity factor, NH
Renon-Newtonian = ρv2/su-op
Zakeri et al. (2008; 2011) fit - Eq.7 1-PS1 to 1-PS4
3-PS1 to 3-PS3 4-PS1 to 4-PS3
5-PS1 6-PS1
8-PS1 to 8-PS5 10-PS1 to 10-PS6
11-PS1 to 11-PS4 12-PS1
Zakeri et al. (2011) Zakeri et al. (2008)
Eq.7 (current data fit) Martin & White (2012) - lower bound
Martin & White (2012) - upper bound
0
20
40
60
80
100
120
10 30 50 70 90 110 130 150 170 190
Horizontal bearing capacity factor, N
H
Re
non-Newtonian
= ρv
2
/s
u-op
3-PS1 & 3-PS2
5-PS1
6-PS1
10-PS1 & 10-PS2
11-PS1
12-PS1
Eq.7 (current data fit)
Figure 20. Drag forces during lateral pipe-slide interaction, across the solid-fluid boundary (Sahdi et al. 2014)
Numerical modelling of slide-pipeline interaction
Two studies used finite element analysis to explore in more detail the interaction between a slide and
pipeline during perpendicular attack.
In the first study, large deformation finite element (LDFE) analysis was used to simulate static
deformation of a landslide moving past a partially-embedded pipeline, to quantify the loads and

OTC-27034-MS 21
displacements imposed on pipe. These analyses used an elastic perfectly plastic soil model and plane
strain conditions. The pipeline was restrained by a set of springs so that the active load built up to a
stable value, representing the limiting load at which the debris flowed over the pipeline. A parametric
study was undertaken by varying the pipeline embedment and the relative strengths of the debris and
seabed, which is reported by Zhu & Randolph (2010).
The analyses showed that the various combinations of soil strength and embedment depth lead to
different debris-pipeline movement patterns and consequently lead to rather different magnitudes of the
loads imposed on pipelines. An example of different pipeline and seabed responses is shown in Figure
21. For debris that is much stronger than the seabed, erosion can lead to the pipeline becoming entirely
engulfed by the debris (Figure 21c), whereas for a relatively stronger seabed (or weaker debris) after
some erosion the debris flows over the top of the pipeline (Figure 21d). The pipeline is subjected to the
largest load (an equivalent pressure of 11.5 times the debris strength) from the landslide when it rests on
the weakest seabed. The pressure is proportional to the debris material strength but varies inversely with
the seabed strength for partially embedded pipelines. For all strength combinations, there is a critical
embedment depth beyond which the force on the pipeline reduces to a very small magnitude due to the
protective effect of the initial embedment. The parametric study showed that the seabed mudline
strength and initial pipeline embedment are the two main parameters that affect the debris-pipeline
movement patterns and consequently the magnitude of loads imposed on pipelines.
(a) Initial state
(c) Pipeline engulfed by debris
(b) Debris eroding the seabed
(d) Debris flowing over top of pipeline
Figure 21. Different debris flow patterns for an initial pipeline embedment zi/D = 1.1
In the second study, the influence of strain rate effects on active slide-pipeline loading was examined.
The purpose of this work was to establish the link between the rate-dependency of the soil strength and
the resulting rate-dependency of the active slide-pipeline loading (Zhu & Randolph 2011). These

22 OTC-27034-MS
analyses adopted a fully engulfed (i.e. buried) pipe and the normalised form of the Herschel-Bulkley
model, expressed as:
uy
ref
us1s 




















 (8)
where  represents a viscosity parameter and the parameter β is often referred to as the shear-thinning
index. The relative magnitudes of ‘static’ and ‘viscous’ resistance for a given rate of soil flow is
expressed by the Oldroyd number (Od) given by:








D/v
1
Od ref

(9)
where v/D represents a nominal strain rate in the soil around the pipe, and the other parameters are rate-
dependency properties of the soil.
The LDFE analyses investigated the undrained limiting loads on a cylinder moving steadily through
imponderable (i.e. weightless) soft rate-dependent material. The flow mechanism and the effects of the
shear-thinning index and Oldroyd number on the shear zones were explored. The average rate of strain
experienced by the soil flowing past the cylinder was estimated for a given flow velocity and an
expression in the form of a conventional bearing capacity equation, but with shear strength linked
directly to the normalised flow velocity, was established to predict the magnitude of the viscous force
exerted by the debris flow.
Figure 22 shows the effect of velocity, and thus strain rate, on the failure mechanism. At very low
velocity (Od = 104) the mobilized viscous component of soil strength is negligible, and so the classical
plasticity solution of a concentrated velocity discontinuity (Martin & Randolph 2006) is recovered. At
very high velocity (Od = 0.1), where the viscous enhancement is significant, the shear strains become
much more diffuse.
In spite of the changing shear strain patterns, it was found that the force exerted on the pipe could be
expressed in terms of a constant bearing factor, NH, and a nominal shear strength based on a strain rate
of D/fv
, where f is an adjustment factor lying between 0.8 and unity. Thus, for the Herschel-
Bulkley rheological model, the net pressure on the cylinder can be expressed as
op,uH
ref
uyHH sN
d/vf
1sNq 




















 (10)
Figure 23 shows that the nominal bearing factor, NH, (taking f = 1) becomes increasingly independent of
the velocity (expressed in terms of the Oldroyd number) as the value of the strain rate power
coefficient reduces. For values of  relevant to real soils (for which  < 0.3, generally), it is therefore
sufficient to evaluate the impact force by adopting a constant, velocity independent, bearing factor based
on plasticity solutions but using an operative shear strength based on a strain rate of v/D. This finding
underpins the general expression, for weighty soil, given by Equation (7).

OTC-27034-MS 23
Figure 22. Effect of velocity on shear strain pattern around a smooth pipe (

= 0.26)
4
6
8
10
12
14
16
0.1 1 10 100 1000 10000
Oldroyd number, Od
N
nom
Ful ly bonded
Fully smooth
 = 1
 = 0.26
 = 0.5  = 0.75
Figure 23. Variation in nominal resistance factor with velocity for different values of

Non-perpendicular slide-pipeline interaction: Inclined attack
To extend the solution for perpendicular slide-pipeline loading given by Equation (7) to other angles of
attack, data from computational fluid dynamic (CFD) numerical analyses of slide-pipeline interaction
forces, published by Zakeri (2009b), were reviewed. This allowed a re-interpretation of the data within
the new framework proposed in the project, in which the geotechnical resistance and inertial fluid drag
components of pipeline loading are treated separately (Randolph & White 2012). Dr Zakeri assisted by
providing additional information beyond his original publication.
NH

24 OTC-27034-MS
Zakeri (2009b) used the CFD software ANSYS CFX to analyse the pipeline response, extracting net
forces normal and parallel to a pipe segment placed at different angles to the flow direction. His data
have been replotted in Figure 24, considering separately the resulting drag factors for flow normal to the
pipe (angle of 90º) and parallel to the pipe (angle of 0º).
The results for normal flow at low values of non-Newtonian Reynolds number, i.e. low v2/su-op, are
consistent with the conventional geotechnical treatment of flow around a cylindrical object, with a
bearing resistance factor, Np, (the ratio of net pressure to material strength – at strain rate of v/D –
equivalent to the gradient of the lines shown) of 12 being identified. At higher rates of flow, a
consistent drag factor of 0.4 was found, with a geotechnical resistance component that remained
consistent with the resistance factor of 12 and a rate-enhanced soil strength (Figure 24). The
geotechnical resistance component is dominant in most cases.
For axial flow the deduced frictional drag can be expressed as Fa = fasu,opD, where su,op is the rate-
enhanced shear strength for a strain rate of v/D. The theoretical value of fa for the power law rheological
model adopted is about 1.4 (Einav & Randolph 2006), which is approximately 10 % below the gradient
from Zakeri’s numerical results. This discrepancy may be attributed to the high spatial gradients of
velocity very close to the pipe, which are only approximated within the numerical analysis.
To allow the interaction forces to be determined for any angle of attack, the variation of bearing and
friction factors with the angle of the flow relative to the pipeline axis were interpreted in terms of a
failure envelope, as shown in Figure 25. The dashed curve is a fit to the numerical data, expressed as
1
N
N
f
f1
90,p
p
3
0,a
a
















 (11)
where, consistent with the data, fa,0 was taken as 1.54 and Np,90 was taken as 12.8.
Although the envelope provides a reasonable fit to the data points for 45º and 60º, the data point
for 30º lies well within the envelope, and indeed implies some concavity in the envelope. The
reason for this is unclear, and requires further detailed numerical simulation for flow angles
lying between 0 and 45 º to resolve.
y = 12.82 9x
R² = 0.9984
y = 11.721x
R² = 0.9892
y = 10.602x
R² = 0.9957
y = 7.4724x
R² = 0.9766
0
2
4
6
8
10
12
14
16
18
00.511.52
Lateral drag coefficient, C
d-90
Factored inverse of Reynolds number: 2/R
e
90 degrees 60 degrees
45 degrees 30 degrees
Normal component
y = 1.5359x, R
2
= 0.9996
y = 0.9616x, R
2
= 0.996
y = 0.8797x
R
2
= 0.9967
y = 0.7045x, R
2
= 0.992
0
0.5
1
1.5
2
2.5
00.511.5
Axial drag coefficient, C
d-0
Factored inverse of Reynolds number: 2/R
e
0 degrees 30 degrees
45 degrees 60 degrees
Theory Theory
Axial component
(a) Flow components normal to pipe (b) Flow components parallel to pipe
Figure 24. CFD Results: Replotted drag coefficient data for general flow conditions

OTC-27034-MS 25
Figure 25. Failure envelope derived from CFD results for varying flow angle relative to pipe axis
A design envelope is then proposed that is constrained to the theoretical limiting bearing factors of
fa,0 = 1.4 and Np,90 = 11.9 (see the solid curve in Figure 25). In addition to the failure envelope, a
relationship is required to determine the position on this envelope that corresponds to a particular angle
of flow, and hence provide the resulting bearing factors Np (and fa). The resultant force on the pipe,
which is the vector sum of Fn and Fa, is not parallel to the flow direction.
Instead, the bearing factor, Np, may be expressed as
 
7.07.0
90,pp sin9.11sinNN  (12)
with the corresponding value of fa then being found from Equation (11). Equation (12) is a purely
empirical relationship, which appears to give reasonable predictions of the numerical data. The proposed
failure envelope, together with spot values of the factors for the same angles as examined by Zakeri
(2009b), are shown in Figure 25. The calculated factors all lie within 10 % of those deduced from the
numerical data.
This methodology provides, for the first time, a robust approach for deriving the drag forces from debris
loading of a pipeline under an arbitrary angle of attack. For the realistic range of rheological parameters
considered, the total drag load is dominated by the component derived from the rate-enhanced material
strength (as opposed to the inertia). This supports the use of a geotechnically-based interpretation, using
bearing and ‘skin friction’ factors akin to the loading on piles. As identified in the LDFE study
illustrated in Figure 22 and Figure 23, it is appropriate to use a rate-enhanced material strength
corresponding to a strain rate of approximately v/D.
Modelling of pipeline structural response
The previous sections described new techniques developed within this project to assess the local loading
on an element of pipeline that is impacted by a slide. In order to assess the severity of this geohazard, the
next task is to assess the structural response of the pipeline, and the resulting bending and tensile stresses
as it is deformed. A new simplified technique for performing this assessment was developed within the
project, as summarised here. Additional information is provided in Randolph et al. (2010).

26 OTC-27034-MS
The structural response of a pipeline impacted by a slide was investigated analytically, with additional
validation using the finite element software, ABAQUS. The slide was modelled as ‘active’ loading over
a finite length of the pipeline, balanced by passive resistance and friction provided by the seabed on each
side of the slide, extending the work of Sweeney et al. (2004). The bending and membrane tension
induced in the pipeline were evaluated by means of analytical non-dimensional solutions. A schematic
of the problem is shown in Figure 26.
As a pipeline impacted by a submarine slide deforms, it will be restrained by lateral and axial resistance
in adjacent passive zones. Ultimately the pipeline may come to a stable deformed shape where continued
active loading from the slide is equilibrated by membrane tension in the pipeline in addition to the
passive resistance. Parametric solutions of the problem in terms of general values of the pipeline
properties and pipe-soil load transfer parameters were developed during this project (Randolph et al.
2010). The aim was to provide generic design charts giving the resulting pipeline response – the stresses
resulting from bending moments and tensile forces, and lateral displacement – as a function of the input
parameters, but with all quantities expressed in terms of appropriate non-dimensional groups. In spite of
some idealisations of the problem, the charts provide an initial design basis and a framework for more
detailed numerical analysis as necessary. They are simple and rapid to apply, and therefore amenable to
Monte Carlo simulations within a quantitative geohazard assessment – where full FE structural
modelling would not be practical.
The pipeline is assigned an elastic axial stiffness, EA, bending stiffness, EI, and diameter, D (although,
for a given wall thickness ratio, D/t, the last quantity may be derived from the ratio of I/A). Focusing on
slide loading normal to the pipeline, the main output quantities of interest are: the induced tension, T,
and bending moment, M; the maximum lateral motion, ymax; and the length of pipeline, s, that feeds into
the deformed shape local to the slide.
Flow direction
q
y
q
x
p
f
p
f
Active region
loaded by slide
Passive region
resisting movement
Passive region
resisting movement
Flow direction
q
y
q
x
p
f
p
f
Active region
loaded by slide
Passive region
resisting movement
Passive region
resisting movement
(a) Schematic of general slide loading on pipeline
y
x
q
p
B
x
a
x
m
y
x
q
p
B
x
a
x
m
(b) Nomenclature for analysis of normal loading acting on pipeline
Figure 26. Idealisation of pipeline loading by slide and resulting deformation

OTC-27034-MS 27
The outputs may be non-dimensionalised and expressed in terms of non-dimensionalised inputs:




































EI
EA
B,
p
f
,
q
p
,
EA
Bq
,
EA
qB
g
B
s
;
EI
EA
B,
p
f
,
q
p
,
EA
Bq
,
EA
qB
g
B
y
EI
EA
B,
p
f
,
q
p
,
EA
Bq
,
EA
qB
g
EI
MB
;
EI
EA
B,
p
f
,
q
p
,
EA
Bq
,
EA
qB
g
EA
T
f
4
f
3
max
f
2
f
1
(13)
Analytical solutions, validated from numerical analysis, were developed to evaluate these quantities,
which were then expressed as functions of the input quantities using multiple regression analysis. A
typical set of results is shown graphically in Figure 27, where the vertical axes shows the magnitude of
tensile and bending strain, /E, induced in the pipe by slides of different magnitude, qB/EA, and
different widths, B/D.
Regression analysis allowed the tensile and bending strains to be expressed as functions of the non-
dimensionalised input quantities according to:
2
k0.22
0.075 0.75
t
10.2
2
BqBpf
0.28
EDEAqp
BqB
where k 0.075 2.0 DEA


  

  
   
 

 
 
(14)
2.006.0
27.098.0
b
p
f
q
p
EA
qB
D
B
E































 (15)
Corresponding expressions for other output quantities are provided in Randolph et al. (2010).
These solutions provide an effective design tool for quantitative assessments of submarine slide hazards.
Based on an active loading over a defined length and specified values of passive lateral and axial soil
resistance outside the slide zone, the analysis allows the maximum deflection and the maximum stresses
within the pipeline to be estimated directly – for example, using a simple spreadsheet. In an extension to
previously published solutions, the bending resistance of the pipeline has been considered in addition to
the tensile response.
The inclusion of bending effects showed that for narrow slides and stiff pipelines, the bending stresses
dominate the structural load, whereas for wider slides and more flexible pipelines the maximum stress is
predominantly from tensile loading due to stretching of the pipeline. The effect of the passive soil
restraint on the maximum pipeline stress also depends on the width of the slide. For narrow slides, high
soil restraint leads to reduced pipeline stresses whereas for a wide slide the maximum pipeline stress is
reduced by lower soil restraint. Randolph et al. (2010) provide methods to extend the analysis to account
for a finite length of anchorage and a non-zero initial tension in the pipeline.

28 OTC-27034-MS
0.00001
0.0001
0.001
0.01
10 100 1000 10000
Normalized debris flow width, B/D
Computed strain,
/E
Bending
Tension
Combined
qB/EA =
0.001
0.0005
0.0002
0.0001
0.00005
0.00002
p/q = f/p = 0.5
Figure 27. Effect of slide loading and width on maximum pipeline strains
Conclusions and outcomes
The assessment of geohazards, such as submarine slides and their potential impact on pipelines, is a
highly multi-disciplinary activity, requiring contributions from a variety of geo-specialists including
geologists and geophysicists, as well as geotechnical engineers. Naturally, the assessment of slide-
pipeline interaction also involves pipeline engineering.
This Joint Industry Project was focused on geotechnical aspects of this problem, and has made research
contributions that have already been adopted in design practice for projects located in Australia and
elsewhere globally.
The project has established a framework for the material response of seabed sediments that is applicable
from their intact condition through to a fully mobilised fluid state of debris flow. By characterising the
seabed material in this framework, the material response during relict slides can be constrained,
enhancing the back-analysis of existing slides. Once these refined back-analyses are complete, new
techniques of slide simulation can be applied, including enhanced depth-averaged methods, and also
full-blown continuum LDFE. This project has demonstrated that continuum LDFE can recreate a range
of realistic flow mechanisms.
Once the potential slide hazard has been quantified in terms of the expected size, speed and strength of
the flows, a much more sophisticated assessment of the resulting slide-pipeline interaction can be
performed using techniques developed in this project. New simple analytical solutions have been
derived to assess the pipeline loading for given slide characteristics and the relevant pipeline properties
and orientation. These solutions can also be applied to other infrastructure in the path of a slide. In
addition to the local loading, solutions are also available for the consequent structural response –
capturing the influence of both tensile and bending stresses on the pipeline response.
These outcomes have been applied to projects in various ways. They have led to better-constrained
characterisation of relict slides, and more detailed assessment of the impact of potential future slides. By

OTC-27034-MS 29
eliminating some of the assumptions within current practice, and using more refined calculation models,
improved design efficiency has been achieved. The simple analytical solutions for slide loading and
pipeline response have unlocked the possibility of Monte Carlo-based quantitative assessments of
potential pipeline damage – allowing true risk-based decision making. As a result, designers have a basis
to demonstrate an acceptable probability of survival of a slide hazard, rather than needing to route a
pipeline entirely around the hazard.
Bibliography
Additional information on the activities performed during this Joint Industry Project can be sourced
from the cited references or the JIP Final report, published by MERIWA:
White D.J., Randolph M.F., Gaudin C., Boylan N., Wang D., Boukpeti N. & Sahdi F. 2011. MERIWA-administered Joint
Industry Project: The impact of submarine slides on pipelines: Final report to MERIWA and the project participants.
UWA GEO report 10538 1025 pp., MERIWA report 290.
Acknowledgements
This JIP was administered by the Minerals and Energy Research Institute of Western Australia
(MERIWA), which also contributed to the project funding. Six oil and gas Operators were project
sponsors: BP, BHP Petroleum, Chevron, Petrobras, Shell and Woodside. Additional support was
received from the CSIRO Flagship Collaboration on Subsea Pipelines, hosted at UWA, and the ARC
Federation and Future Fellowship Schemes (supporting Mark Randolph and David White, respectively).
The opinions expressed in this paper and in the JIP reports and publications are those of the authors and
may not reflect the views of the sponsors or MERIWA.
The authors acknowledge the advice and support received from the project steering committee, which
included Pam Smith (MERIWA), Paul Dimmock (BP), Trevor Evans (BP), Philippe Jeanjean (BP), Joe
Straub (BHP), Steve Paulson (Chevron), Jen-Hwa Chen (Chevron), Rob Little (Chevron), Leopoldo
Paganelli (Petrobras), Jason Newlin (Shell), TG Tan (Shell), Andrew Pearce (Woodside) and Marc
Senders (Woodside).
This paper forms part of the activities of the Centre for Offshore Foundation Systems (COFS), currently
supported as a node of the Australian Research Council Centre of Excellence for Geotechnical Science
and Engineering (grant CE110001009) and through the Fugro Chair in Geotechnics (held by the second
author), the Lloyd’s Register Foundation Chair and Centre of Excellence in Offshore Foundations and
the Shell EMI Chair in Offshore Engineering (held by the first author).
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