RECOMMENDED PRACTICE
DNVGL-RP-F114
Edition May 2017
Pipe-soil interaction for submarine
pipelines
The electronic pdf version of this document, available free of charge
from http://www.dnvgl.com, is the officially binding version.
DNV GL AS
FOREWORD
DNV GL recommended practices contain sound engineering practice and guidance.
©
DNV GL AS May 2017
Any comments may be sent by e-mail to rules@dnvgl.com
This service document has been prepared based on available knowledge, technology and/or information at the time of issuance of this
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for loss or damages resulting from any use of this document.
Changes - current
CHANGES – CURRENT
This is a new document.
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Changes – current.................................................................................................. 3
Section 1 Introduction............................................................................................ 6
1.1 General............................................................................................. 6
1.2 Objective...........................................................................................6
1.3 Scope................................................................................................ 6
1.4 Application........................................................................................ 6
1.5 Contributions from joint industry projects........................................7
1.6 Structure of this recommended practice...........................................8
1.7 Referenced standards and recommended practices.......................... 8
1.8 Definitions.........................................................................................9
Section 2 Modelling pipe-soil interaction.............................................................. 16
2.1 General........................................................................................... 16
2.2 Pipe-soil interaction within the design process...............................16
2.3 Establishing a geotechnical model.................................................. 19
2.4 Finite element modelling to assess pipe-soil interaction................. 20
Section 3 Material properties required for design and assessment....................... 21
3.1 General........................................................................................... 21
3.2 Geotechnical field and laboratory testing....................................... 21
3.3 Geotechnical properties.................................................................. 22
3.4 Pipe properties............................................................................... 23
Section 4 Exposed pipelines..................................................................................24
4.1 General........................................................................................... 24
4.2 Pipe embedment............................................................................. 25
4.3 Axial pipe-soil interaction............................................................... 35
4.4 Lateral pipe-soil interaction............................................................ 42
4.5 Soil stiffness................................................................................... 57
4.6 Soil damping................................................................................... 61
Section 5 Buried and covered pipelines................................................................ 64
5.1 General........................................................................................... 64
5.2 Effect of trenching method............................................................. 64
5.3 Axial pipe-soil interaction............................................................... 65
5.4 Lateral pipe-soil interaction within rock berms...............................68
5.5 Uplift resistance..............................................................................69
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Contents
CONTENTS
5.7 Downward resistance and stiffness................................................ 75
Section 6 Treatment of uncertainties.................................................................... 76
6.1 General........................................................................................... 76
6.2 Considerations for pipeline design and assessment........................ 76
Section 7 Special considerations........................................................................... 80
7.1 On-bottom stability.........................................................................80
7.2 Free spanning pipelines.................................................................. 81
7.3 Design philosophy of support fills.................................................. 84
7.4 Penetration of falling objects......................................................... 86
7.5 Carbonate soils............................................................................... 88
Section 8 Bibliography.......................................................................................... 90
8.1 Bibliography.................................................................................... 90
Appendix A Soil models from DNV-RP-F109 (October 2010).................................94
A.1 General........................................................................................... 94
A.2 Initial penetration.......................................................................... 94
A.3 Passive soil resistance....................................................................94
A.4 Nomenclature................................................................................. 95
Changes - historic.................................................................................................96
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Contents
5.6 Rock fill over backfilled clay........................................................... 75
SECTION 1 INTRODUCTION
1.1 General
Pipe-soil interaction is an important aspect of a pipeline system as it may have a large influence on both the
structural behaviour and integrity of the pipeline during installation and operation.
Knowledge about the soil conditions along the pipeline route is essential to evaluate the pipe-soil interaction,
and the planning of soil investigations should be tailor-made for the conditions encountered during the
lifetime of the pipeline. Soil variability is inevitable over large distances and is especially the case in the
surficial soils. The variation in soil parameters seen in a pipeline development project is thus larger compared
to traditional foundation design.
During installation of an exposed pipeline, the soil around the pipe will be disturbed, affecting both the
strength and stiffness properties as well as the seabed configuration close to the pipe. These installation
effects are difficult to predict, as they are highly governed by the pipe motions during laying. For buried
pipelines, the state of the backfilled material is challenging to predict.
The complexity and uncertainty in pipe-soil interaction are significant, and require simplifications and
assumptions in the engineering models. The effort spent on pipe-soil interaction should however reflect the
sensitivity to the pipeline design.
This recommended practice involves guidance related to pipe-soil interaction for submarine pipelines and
supersedes the pipe-soil information given in the following recommended practices:
—
—
—
—
—
—
DNVGL-RP-F105
DNVGL-RP-F107
DNVGL-RP-F109
DNVGL-RP-F110
DNVGL-RP-F111
DNVGL-RP-F113
Free spanning pipelines
Risk assessment of pipeline protection
On-bottom stability design of submarine pipelines
Global buckling of submarine pipelines
Interference between trawl gear and pipelines
Pipeline subsea repair
Hence, the pipe-soil interaction parts in the above recommended practices will be removed in next respective
revision and a reference to this recommended practice will be included.
The pipe-soil interaction guidance for exposed pipelines included in this recommended practice has to a far
extent been developed by the SAFEBUCK joint industry project (JIP). Contributions from other JIPs to the
recommended practice are included from the PIPESTAB JIP, HOTPIPE JIP and GUDESP JIP.
This recommended practice follows the same principles as outlined in ISO 19901-4 and API RP 2GEO, but
with more specific guidance.
1.2 Objective
The objective of this recommended practice is to provide guidance related to pipe-soil interaction relevant for
the various conditions experienced during the lifetime of a pipeline system according to the requirements set
out in DNVGL-ST-F101.
1.3 Scope
This recommended practice gives recommendations on how to evaluate pipe-soil interaction for various
design situations or assessments relevant for exposed and buried submarine pipelines.
1.4 Application
This recommended practice is written primarily for qualified geotechnical engineers. Hence, basic
geotechnical terms are not always explained. It is however important that the geotechnical engineer
cooperates with the pipeline engineer to understand the design situations.
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As far as possible, engineering models based on geotechnical principles should be used. In some cases, it
is difficult to establish theoretical models. In such cases, empirical models are necessary. Empirical models
should be used with care and the geotechnical engineer should evaluate the validity of the model for the
problem at hand, understanding its limitations.
Due to the complexity in the pipe-soil interaction assessments, this recommended practice cannot be as
prescriptive as other recommended practices. The calculation models presented herein can therefore be
considered as examples. In general, more than one model should be evaluated, see Sec.6 for discussion
about uncertainties related to pipe-soil interaction assessments.
For scenarios involving pipe-soil interaction which are not captured in this recommended practice, specific
evaluations by the geotechnical engineer is required. In general, when well-established methods are not
available, the consequences of this uncertainty should be evaluated. Alternative design solutions may then be
considered. If the consequence of an unfavourable incident do not jeopardize the pipeline integrity, a survey
plan in combination with mitigation measures may also be a viable solution.
Guidance for exposed pipelines obtained through the SAFEBUCK JIP has been included. Also, calculation
models based on geotechnical principles have been included, and where possible, they have been compared
with recognized empirical formulations used in the industry.
Alternative calculation methods may also be used than those provided in this recommended practice.
However, it is recommended that they have a sound theoretical basis, or that they capture better some
relevant data or conditions, such as project-specific model testing of PSI resistance, or relevant field
observations of existing pipelines from the same region. By use of new calculation methods, its applicability
should be documented in a transparent manner which allows for independent verification.
The full interaction between the pipe and the soil accounting for the stiffness of the pipe and the loads
acting upon the pipe, is not covered in this recommended practice other than as brief discussions where
relevant, giving reference to other relevant recommended practices. The recommendations provided in this
recommended practice are primarily related to the interaction between the pipeline and the soil per unit
length of the pipe.
Likewise, the integrated interaction between a pipeline or flowline including spools and connected
structures is covered in respective standards or recommended practices for pipelines and structures. The
recommendations given herein may be used as input for analysis of such integrated interaction.
This recommended practice does not consider the following items that may be included in future editions:
—
—
—
—
geohazards
earthquake design and assessment of pipelines
riser-pipe interaction
pipe-soil-structure interaction.
1.5 Contributions from joint industry projects
1.5.1 SAFEBUCK JIP
The SAFEBUCK JIP was a joint industry project, which had the aim of developing design methodologies
related to high pressure, high temperature (HPHT) pipelines susceptible to lateral buckling. Extensive
research with respect to pipe-soil interaction was carried out as a part of the JIP. New calculation models
were developed based on small and large scale tests. The JIP mainly focused on exposed pipelines placed on
deep water clays, hence the soil conditions that is covered by the SAFEBUCK database is limited accordingly.
The database consists of tests primarily carried out on soft West African clays with high plasticity.
The vertical embedment model proposed in SAFEBUCK is based on theoretical considerations, but the
approach to use remoulded shear strength to account for cyclic effects during laying is highly empirical.
The JIP recommended to perform specialized interface tests to measure the axial interface strength directly.
The proposed lateral soil resistance models are empirically calibrated towards a limited database of tests. and
is therefore expected to show some bias to the underlying database conditions.
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1.5.2 PIPESTAB JIP
The PIPESTAB JIP was a joint industry project which was part of the development of DNVGL-RP-F109.
The soil models developed from this work have been reviewed and adopted as found appropriate in this
recommended practice.
1.5.3 HOTPIPE JIP
The HOTPIPE JIP was a joint industry project which was part of the development of DNVGL-RP-F110. Pipesoil interaction guidance on uplift resistance of buried pipelines has been reviewed and adopted as found
appropriate in this recommended practice.
1.5.4 GUDESP JIP
The GUDESP JIP was a joint industry project which was part of the development of DNVGL-RP-F105.
Guidance on simplified soil damping is included in this recommended practice.
1.6 Structure of this recommended practice
The recommended practice is structured as follows:
1)
Introduction (this section)
2)
Presents the overall objective, scope and applicability of the recommended practice, as well as relevant
abbreviations and symbols (giving all the symbols in the equations). Referenced standards are listed
in this section and referred to by its acronyms while bibliographies and reports are listed in Sec.8 and
referenced by reference numbers.
Modelling pipe-soil interaction
3)
Presents an introduction to pipe-soil interaction and how soil resistance curves may be included in the
pipeline analysis
Material properties required for design and assessment
4)
Provides guidance with respect to soil investigations relevant for pipe-soil interaction and the pipe
properties required for pipe-soil assessment
Exposed pipelines
5)
Provides guidance for evaluating pipe-soil interaction for exposed pipelines
Buried pipelines
6)
Provides guidance for evaluating pipe-soil interaction for buried pipelines
Treatment of uncertainties
7)
Discusses in general terms the different sources of uncertainties in geotechnical design, and highlight
special considerations related to pipe-soil interaction
Special considerations
8)
Presents other issues related to pipe-soil interaction which may not naturally be placed in Sec.1 to Sec.6.
Bibliography
1.7 Referenced standards and recommended practices
1.7.1 General
In case of conflict between this recommended practice and referenced DNV GL standards, the standard or
recommended practice with the latest edition date shall prevail.
The latest valid edition of each of the DNV GL reference documents applies.
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Referenced relevant standards and recommended practice documents are given in [1.7.2] and [1.7.3] while
bibliography is given in Sec.8.
1.7.2 DNV GL standards and recommended practices
Document code
Title
DNVGL-OS-C101
Design of offshore steel structures, general – LRFD method
DNVGL-ST-F101
Submarine pipeline systems
DNVGL-RP-C207
DNVGL-RP-C212
Statistical representation of soil data
*)
Offshore soil mechanics and geotechnical engineering
DNVGL-RP-F105
Free spanning pipelines
DNVGL-RP-F107
Risk assessment of pipeline protection
DNVGL-RP-F109
On-bottom stability design of submarine pipelines
DNVGL-RP-F110
Global buckling of submarine pipelines
DNVGL-RP-F111
Interference between trawl gear and pipelines
DNVGL-RP-F113
Pipeline subsea repair
*)
DNVGL-RP-C212 will soon replace DNV Classification Notes 30.4
1.7.3 Other standards and recommended practices
Document code
Title
ANSI/ASME B46.1
Surface Texture (Surface Roughness, Waviness and Lay)
API RP 2GEO
Geotechnical and Foundation Design Considerations
ISO 19901-4
Petroleum and natural gas industries – Specific requirements for offshore structures – Part 4:
Geotechnical and foundation design considerations
1.8 Definitions
1.8.1 Abbreviations
Abbreviation
Description
BE
best estimate
CPT
cone penetration test
FE
finite element
FEED
front end engineering design
HE
high estimate
HPHT
high pressure, high temperature
JIP
joint industry project
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Abbreviation
Description
LE
low estimate
NC
normally consolidated
OCR
overconsolidation ratio
PCPT
cone penetration test with pore pressure measurements
PIP
pipe-in-pipe
PLEM
pipeline end manifold
PLET
pipeline end termination
PSI
pipe-soil interaction
SHANSEP
stress history and normalized soil engineering properties
TUM
terrain unit mapping
1.8.2 Symbols – Greek characters
Symbol
Description
α
pipe-soil adhesion or roughness factor
γ'
submerged unit weight
γc
cyclic shear strain
γ'fill
submerged unit weight of backfilled material
γpre
consolidation preloading effect
γrate
rate factor
γ'seabed
submerged unit weight of seabed soil
δf
failure mobilization distance
δpeak
peak interface friction angle
δres
residual interface friction angle
εres
residual reduction factor
ζ
wedging factor
ζsoil
soil damping ratio
κa
active earth pressure coefficient
κp
passive earth pressure coefficient
μA,brk,d
axial breakout friction factor in drained conditions
μA,brk,u
axial breakout friction factor in undrained conditions
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Symbol
Description
μA,res,d
axial residual friction factor in drained conditions
μA,res,u
axial residual friction factor in undrained conditions
μfill
friction coefficient between pipe and backfilled material
μseabed
friction coefficient between pipe and seabed soil
ν
Poisson’s ratio
ρ
gradient of undrained shear strength profile with depth
ρs/ρ
specific mass ratio between the pipe mass and the displaced water
σa
atmospheric pressure (100kPa)
σ'h
horizontal effective stress
σ's
mean effective stress in soil
σ'v
vertical effective stress
τ
shear stress
φ
drained friction angle
φfill
friction angle of backfilled material
φpeak
peak friction angle of the soil
φres
residual friction angle of the soil
ω
angular frequency
1.8.3 Symbols – Latin characters
Symbol
Description
a
horizontal oscillation amplitude prior to lateral breakout
Aberm
displaced soil area creating a berm adjacent to the pipe
Abm
penetrated cross sectional area of the pipe
Ap
cross-sectional area of the pipe
Apipe
plugged area of falling pipe
B
pipe-soil contact width
c
viscous damping coefficient
CL
lateral dynamic stiffness factor in simplified evaluation for free spanning pipelines
CV
vertical dynamic stiffness factor in simplified evaluation for free spanning pipelines
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Symbol
Description
D
pipe outer diameter (including coating)
dca
depth correction factor for clay
dq
depth correction factor for sand
Dref
reference diameter
EDissipated
dissipated energy within one hysteretic loop
EElastic
elastic energy within one hysteretic loop
Ep
energy absorbed in gravel
EI
pipe bending stiffness
F
bearing capacity factor clay (accounting for pipe roughness and soil strength gradient)
f
uplift resistance factor
FA
axial resistance
FA,brk
axial breakout resistance
FA,brk,d
axial breakout resistance factor in drained conditions
FA,brk,u
axial breakout resistance in undrained conditions
FA,deep,d
axial resistance for buried pipelines in drained conditions (deep failure mode)
FA,shallow,u
axial resistance for buried pipelines in undrained conditions (shallow failure mode)
FL,brk,d
lateral breakout resistance in drained conditions
FL,brk,d,fric
frictional part of the lateral breakout resistance in drained conditions
FL,brk,d,passive
part of the lateral breakout resistance in drained conditions which involves passive soil resistance
FL,brk,u,fric
frictional part of the lateral breakout resistance in undrained conditions
FL,brk,u,remain
part of the lateral breakout resistance in undrained conditions which involves active and passive
soil resistance and soil weight
FL,brk,u
lateral breakout resistance in undrained conditions
FL,res,d
lateral residual resistance in drained soil condition
FL,res,u
lateral residual resistance after breakout in undrained soil conditions
Fp
drained passive resistance as proposed in /22/
Fuplift,d
uplift resistance in drained conditions
Fuplift,global,u
uplift resistance in undrained conditions (global failure mode)
Fuplift,local,u
uplift resistance in undrained conditions (local failure mode)
G
shear modulus
Gmax
shear modulus at small strains
H
cover height
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Symbol
Description
Ip
plasticity index
K
lateral earth pressure coefficient (general)
k
linearized spring stiffness
K0,fill
lateral earth pressure coefficient of backfilled material
K0,seabed
lateral earth pressure coefficient of seabed soil
KL,d
lateral dynamic stiffness
KL,s
lateral static stiffness
klay
touchdown lay factor
klay,1, klay,2
touchdown lay factors (used to determine the initial embedment)
Kp
passive earth pressure coefficient
Kv,d
vertical dynamic stiffness
KV,s
vertical static stiffness
L
span length
Lsh
span support length on one shoulder (for transfer of one-half the weight of the free span)
m
factor that accounts for long term effect of overconsolidation
Nγ
bearing capacity factor for sand
Nc
Theoretical bearing capacity factor for clay (for constant undrained shear strength)
Nq
bearing capacity factor for sand
p
pipe-soil contact arc length
Qv
vertical penetration resistance (including depth effects)
Qv0
vertical penetration resistance (not including depth effects)
r
roughness parameter according to /21/
Ra
pipe coating roughness
rpipe-soil
pipe-soil roughness factor
St
sensitivity
Su
undrained shear strength
su,0
undrained shear strength at reference level for depth effects
su,1
average undrained shear strength above the foundation level
su,2
average undrained shear strength below the foundation level
u,active
average undrained shear strength within the active failure zone
u,backfill,reconsolidated
average value of reconsolidated shear strength along the vertical failure plane in the backfilled
material
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Symbol
u,bottom,reconsolidated
su,intact
u,passive
su,r
u,reconsolidated
Description
average value of reconsolidated shear strength at the bottom half of the pipe
intact undrained shear strength
average undrained shear strength within the passive failure zone
remoulded undrained shear strength
average value of reconsolidated shear strength
su,z=0
undrained shear strength at seabed (z=0)
T0
horizontal effective lay tension in the pipe during installation at touchdown point
V
vertical pipe-soil force
Wf
submerged pipe weight during hydrotest (flooded/waterfilled condition)
Wi
submerged pipe weight during installation
Wop
submerged pipe weight during operation
xbrk
mobilization displacement required to mobilize the axial breakout resistance
xfailure
lateral extent of passive failure surface
xmob
axial mobilization displacement
xres
mobilization displacement required to mobilize the axial residual resistance
ybrk
mobilization displacement required to mobilize the lateral breakout resistance
yres
mobilization displacement required to mobilize the lateral residual resistance
z
pipe invert embedment (general)
z0
reference level for depth effects in sand
zf
pipe invert embedment after flooding
zfailure
vertical extent of passive failure surface
zini
initial pipe invert embedment after laying
zmod
modified height taking into account presence of a berm when calculating lateral breakout
zop
pipe invert embedment during operation
zsu,0
reference level for depth effects in clay
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1.8.4 Definitions of verbal forms
Term
Definition
shall
verbal form used to indicate requirements strictly to be followed in order to conform to the document
should
verbal form used to indicate that among several possibilities one is recommended as particularly
suitable, without mentioning or excluding others, or that a certain course of action is preferred but not
necessarily required
may
verbal form used to indicate a course of action permissible within the limits of the document
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SECTION 2 MODELLING PIPE-SOIL INTERACTION
2.1 General
Key pipe-soil interaction (PSI) parameters are the soil resistances during vertical, axial and lateral pipe
movement. In this recommended practice and in most pipeline engineering practice, the PSI resistances
related to exposed pipelines are described in terms of equivalent friction coefficients, defined as the available
resistance to axial or lateral movement divided by the current submerged pipe weight.
The vertical pipe-soil interaction is particularly important during installation, when the pipe penetrates
the seabed. After installation, the vertical pipe-soil interaction is usually less important, but the resulting
embedment from the installation phase is important for the subsequent lateral and axial resistances.
In reality, the lateral and axial resistances are not solely dependent on pipe weight, but are influenced by
pipe embedment, soil type, drainage condition, interface condition and the previous history of loading and
pipe movement. It is therefore important to recognize that a pipe-soil equivalent friction factor is not a soil
property, but depends on the soil properties, the pipeline properties and the mode and history of loading.
PSI may be included in the analysis of a pipeline in various ways, which are listed below in order of increasing
complexity:
a)
b)
c)
d)
As a single limiting value of axial or lateral resistance (or friction factor).
As force-displacement responses in the axial and/or lateral directions, within a finite element analysis of
the pipeline (similar to the t-z and p-y load transfer methods of analysing pile response). Simple forcedisplacement (or friction-displacement) responses are bi-linear (elastic perfectly-plastic), tri-linear (with
an initial peak) or piecewise linear. Additional rules may define the cyclic behaviour.
As a general vertical-lateral response model, based on plasticity theory, implemented within a finite
element analysis of the pipeline via a force-resultant macroelement.
Through explicit modelling of the soil continuum, pipe and pipe-soil interface by using a finite element
analysis software. This is computationally expensive given the need to model typically several kilometres
of pipeline in a single model. This approach is rarely used except for research.
For buried pipelines, the uplift resistance is often modelled as a bi-linear or a tri-linear curve.
The adopted modelling approach should reflect the current project requirements, recognizing the project
stage, risks and opportunities for optimization.
In any design case, the geotechnical engineer needs to consider whether the soil behaviour is drained,
undrained or partially drained and select appropriate calculation models. Note that different classification
systems exist in different countries with respect to soil characterization, and that the same soil can behave
differently for different rate of loading. If there are uncertainties in the soil behaviour, the geotechnical
engineer needs to take this uncertainty into account in the further assessments.
2.2 Pipe-soil interaction within the design process
During the different stages of a pipeline project, PSI should be addressed with an increasing degree of detail,
sufficient to optimize the design through reduced uncertainty. The flowchart in Figure 2-1 illustrates a project
PSI workflow for on-bottom pipeline design. Three sets of PSI recommendations are passed through to the
pipeline engineering workflow at the different project stages – desk study, preliminary and detailed design.
As indicated in Figure 2-1, readily available data may be used to obtain preliminary values for the PSI
parameters at early stages of the design process. Increasingly detailed estimates may then be obtained
through more complex testing and analysis, as described in [3.3.2].
The level of PSI analysis performed in a project should be chosen to suit the project requirements, both to
minimize risk and to maximize added value. During a project, as the geotechnical input and pipeline design
conditions are refined, more complex PSI analysis becomes possible and the level of uncertainty in that
modelling reduces. This potentially leads to more optimized designs. However, the cost benefit of overall
optimization should be assessed along with the cost of engineering to try to achieve those optimizations.
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For re-assessment or modification during operation, field observations from installation and operation may be
used to perform back-calculations to update PSI parameters for use in pipeline operational assessments and
as feedback to future projects.
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Figure 2-1 Pipe-soil interaction workflow during a project
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2.3 Establishing a geotechnical model
When assessing any geotechnical problem, the first step is usually to establish a geotechnical model where
all parameters and boundary conditions affecting the result are defined. This geotechnical model should
contain all assumptions in the geotechnical analysis. This should ideally also be the first step in pipe-soil
assessments.
In traditional foundation design, usually the intact soil strength and boundary conditions are relatively well
defined. However, for a pipeline, given the spacing of soil data along the route and the unknown effects of
laying it is challenging, if not impossible to identify all possible scenarios the pipe could experience. As such,
when using analytical or empirical methods, it is important to evaluate the assumptions and background for
the development of the methods. When a model is based on test results, the model uncertainties are usually
related to differences in how the geotechnical model is defined and how the tests are carried out. The main
effects to consider when establishing a geotechnical model for pipe-soil interaction for an exposed pipeline
are as follows with some of the effects illustrated in Figure 2-2:
— drainage conditions and loading rate (drained, partially drained or undrained soil behaviour)
— geometrical boundary conditions (e.g. ideal soil contact or pipe placed inside a trench)
— shear strength underneath the pipe (may be affected by remoulding during installation and subsequent
thixotropy effects and consolidation from pipe weight)
— shear strength on the outside of the pipe
— pipe-soil interface roughness
— durations of and time between pipelay, flooding, hydrotest, dewatering and start-up.
Figure 2-2 Geotechnical model for an exposed pipeline
For buried pipelines, many of the same items as listed above for exposed pipelines are also important. The
soil resistance is extremely dependent on the type and state of the backfill material. The soil properties of
natural backfill are uncertain as they could be altered significantly from the in-situ properties.
Hypothetically, assuming all aspects of the geotechnical model is well known, which may be the case for
other types of foundations, the best way to calculate soil resistances would be to employ finite element
analyses. However, as this is never the case for pipe-soil interaction assessment, finite element analyses
should not be the only design tool. Finite element analyses may however help to understand the physics
involved and is considered to be the best way to perform sensitivity studies, allowing the user to investigate
how different scenarios (changes in the geotechnical model) would affect the result. Evaluations using
finite element analyses may be relevant for calculating breakout resistances, but not necessarily residual
resistances as large displacement analyses are encumbered with more uncertainty. In those cases, empirical
based methods are needed.
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In general, due to the difficulties in establishing the geotechnical model, more than one model should be
used in PSI assessments, see also Sec.6.
2.4 Finite element modelling to assess pipe-soil interaction
If pipe-soil interaction is evaluated using finite element (FE) analysis one should thoroughly evaluate possible
sources of error and their effect on the results. The following issues are of particular concern in this context:
— the constitutive soil model should represent the soil behaviour needed for the problem at hand
— the iteration procedure should not result in an overshoot of failure loads
— the mesh should be sufficiently fine with proper width/length/height ratios of the elements to ensure a
proper load distribution throughout the soil.
When establishing an FE model, several assumptions need to be made. The influence of the model
assumptions should be investigated and evaluated. The model assumptions include the representation of the
pipe, loading conditions, soil behaviour and soil parameters. The model should be able to capture vertical
and horizontal strength variations. For pipe-soil assessments, the geotechnical model is never known, and
FE analyses will not necessarily give the correct result, but is better suited to evaluate different effects and
changes in the boundary conditions for the pipe, compared to analytical or empirical models with prescribed
failure modes.
When non-linearities associated with large displacements are to be studied by means of an FE model, the
non-linearities may be represented explicitly in the FE model or the underlying large displacements may
be simulated in a wished-in-place analysis. The modelling limitations for the particular analysis should be
assessed. Non-linear large displacement effects are encountered in several situations, including:
— changes in boundary conditions, such as changes in contact area due to large displacements
— displacement-dependent loads, such as displacement-dependent changes in load direction
— large strains.
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SECTION 3 MATERIAL PROPERTIES REQUIRED FOR DESIGN AND
ASSESSMENT
3.1 General
This section describes the required material parameters to carry out pipe-soil interaction assessments, both
related to geotechnical properties and pipe properties.
3.2 Geotechnical field and laboratory testing
For evaluating pipe-soil interaction, the near surface soil properties are critical and require specific pipelinefocused activities in a site investigation and laboratory test program. Uncertain soil conditions will result in
a large range between low estimate and high estimate pipe-soil interaction parameters, which will increase
the uncertainty and potential mitigation costs. In some cases, it may be difficult or costly to demonstrate a
robust design solution.
Planning of soil investigations for a pipeline should be performed with focus on the design scenarios of
importance for the pipeline. These scenarios include, but are not necessarily limited to:
—
—
—
—
—
—
—
—
axial expansion and walking
lateral and upheaval buckling
on-bottom stability and route curve pull-out
free spans
pipeline supports
trenching and back-filling
potential external impacts like trawling equipment and anchors
potential geohazards like landslides from surrounding areas that could hit the pipeline.
The soil influenced by the pipe−soil interaction for exposed pipelines is normally within the upper one
metre and for many pipelines within a few tens of centimetres. Thus, the sampling and testing should have
particular focus on the shallowest soils. Note however that the pipe embedment within a lateral buckle may
achieve depths down to two metres due to cyclic movements.
Where possible and in particular in soft clay, box coring should be performed obtaining blocks of up to half a
meter side dimensions, from which samples can be taken for laboratory testing, or within which small scale
in-situ testing may be performed. Some deeper coring should be performed in addition to the box corings.
Subsea in-situ testing should be performed in addition to the corings. This could primarily consist of PCPT
testing and in clays also T-bar testing, which near the surface can provide more reliable interpretation of
undrained shear strength than PCPT testing will allow for. Alternatively, ball penetrometer testing may be
performed. Ball penetrometer or T-bar penetrometer cyclic tests provide the fully remoulded soil strength,
which is of relevance for the assessment of pipeline embedment and cyclic lateral pipeline response.
For analysis of stability of pipeline supports as well as for evaluation of trenching capabilities, soil information
to somewhat larger depths, i.e several metres, would be required to capture the soil strength within the
predicted failure zone. A geotechnical engineer should in cooperation with the pipeline engineer be involved
in the planning of the soil investigations to make sure that the information required in the subsequent design
analyses is obtained.
For pipelines that are buried by ploughing or jetting, the largest uncertainty is related to how the trenching
method has affected the in-situ strength and stiffness parameters of the backfilled soil. The soil investigation
program needs to consider both the intact soil conditions and the soil conditions following a trenching/jetting
operation. The latter may require the construction of a certain length of dummy trench as part of the soil
investigation program.
Due to general soil variability, it will be practically impossible to obtain very accurate soil data for each
location of interest where the scenarios listed above may be relevant. Thus, a proper strategy for planning
the soil investigations would be to identify from geophysical surveys, possibly combined with relevant other
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information, the various soil units at or very close to the surface along the pipeline route, and to perform soil
sampling to identify the range of characteristics for each unit along the route.
For a survey particularly aiming to provide a basis for pipeline routing and design, sub-bottom profiling
should be included unless available knowledge is such that inhomogeneous conditions of the top soils can
be precluded. Note that it is very challenging to characterize the upper soil very close to the seabed and the
profiler frequency should be targeted specifically for this purpose.
Recognized standards shall be used to carry out laboratory testing. Particular attention should be given to the
planning and execution of tests required to determine very low soil and interface strengths corresponding to
the very low contact stresses between the pipe and the soil.
3.3 Geotechnical properties
3.3.1 General
Geotechnical characteristics necessary for evaluations of all relevant loading conditions shall be determined
for the soils along the pipeline route, including possible unstable soils in the vicinity of the pipeline.
Geotechnical properties may be obtained from generally available geological information, results from
geophysical surveys, including seabed topographical surveys and sub-bottom profiling, and from geotechnical
in-situ tests and laboratory tests on sampled soil. Supplementary information may be obtained from visual
surveys.
Soil parameters of main importance for the pipeline response are:
— shear strength parameters (intact and remoulded undrained shear strength for clay, and angle of friction
for sands)
— deformation characteristics (stress-strain relationships)
— drainage characteristics (permeability and coefficient of consolidation).
These parameters should preferably be determined from adequate laboratory tests or from interpretation of
in-situ tests. In addition, classification and index tests should be considered, such as:
—
—
—
—
—
—
unit weight
water content
liquid and plastic limit
grain size distribution
carbonate content
other relevant tests.
Such tests are important to assess spatial variability of the soil conditions along the pipeline route and to
validate in-situ or laboratory test results from empirical correlations or using such correlations to supplement
the laboratory tests.
Laboratory tests with specific modifications to suit low stress testing are the preferred method to determine
the undrained strength and friction angle of pipe-soil interfaces. Interface materials that are representative
of the planned pipeline coating should be used as interface roughness has a strong influence on interface
friction. Guidelines regarding specialized interface tests are given in [3.3.2].
3.3.2 Specialized pipe-soil interaction testing
The following specialist laboratory equipment for low-stress shear testing may be used to derive the soil-pipe
interface parameters:
— tilt table device for drained interface strength /1/
— low-stress shear box, for drained and undrained interface strength /3/.
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For interface tests, the roughness characteristics of the interface should be representative of the planned
pipe coating or covering a range of potential coating roughness. Roughness characteristics should be
documented.
Large and small-scale pipe models may also be used as background for evaluating the vertical, axial and
lateral pipe-soil resistance in case the existing engineering models are not considered appropriate or if a
certain aspect is particularly important for the project. Four types of tests allow such measurement:
—
—
—
—
small scale centrifuge tests (e.g. /4/)
large scale ex-situ pipe-soil tests (e.g. /5/)
small scale ex-situ pipe-soil tests
large scale in-situ pipe-soil tests (e.g. /6/).
This allows for project specific refinement (calibration) of the engineering models and generally leads to a
reduction in the model uncertainty and the range of PSI parameters.
Tests should cover the range of normal effective stress that can be expected between the pipe and the soil.
Testing at normal effective stresses far from the expected value should be avoided whenever possible to
avoid extrapolation errors. As far as possible, the tests should be carried out in a way replicating the actual
loading history of the soil underneath the pipe, e.g. accounting for penetration arising from motions during
pipeline laying and for the pre-consolidation effect from the water filled/flooded condition during the pipeline
hydrotest.
Laboratories executing such tests should have the correct knowledge about the equipment and procedures to
execute and measure the interface strength at low stress levels.
3.4 Pipe properties
Typical pipeline properties required for pipe-soil interaction assessments are:
—
—
—
—
—
submerged pipe weights, Wi, Wf, Wop
pipe outer diameter including coating, D
pipe coating roughness, Ra
pipe bending stiffness, EI
horizontal effective lay tension in the pipe during installation at touchdown point, T0.
A distinction should be made between the following pipe weights:
— The submerged pipe weight at installation, Wi, which is usually the empty weight.
— The water flooded submerged pipe weight, Wf.
— The operating submerged pipe weight, Wop. A range may be required, considering the range of product
density during the operating life and also potential separation on shutdown, e.g. liquid hold up.
The roughness (Ra) of the planned pipe coating should be determined or estimated when planning interface
testing. To avoid confusion, the pipe coating roughness should be defined as the average deviation from the
mean height, in accordance with ANSI/ASME B46.1. Profilometer or laser interferometer devices provide
rapid quantification of interface roughness, and may be used to gather data from existing pipe coating
samples, and to assess the roughness of interfaces available for laboratory testing.
The pipe bending stiffness, EI, should consider the composite behaviour including coatings (if significant),
inner and outer pipes for a pipe-in-pipe (PIP) line, etc.
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SECTION 4 EXPOSED PIPELINES
4.1 General
Pipe-soil interaction is a key element in the assessment of exposed pipelines. Typical scenarios involving
pipe-soil interaction are lateral buckling, end expansion, pipeline walking, route-curve pullout, flow line
anchoring, on-bottom stability, trawl impact and development of free spans. The consideration of low and
high estimates of pipe-soil interaction parameters is generally required during design and assessment
to satisfy all limit states. If not accounting for a defined range of low and high estimates of resistance,
this should be a conscious choice where possible consequences are evaluated and it is possible to rectify
unfortunate incidences based on survey and contingency plans.
The content of this section is summarized in Table 4-1. In general, the soil behaviour may be drained,
partially drained or undrained depending on the loading rate and drainage conditions and should for each
scenario be considered when selecting calculation models.
Table 4-1 Pipe-soil responses for exposed pipelines
Response
Description
Embedment
(see [4.2])
The initial embedment is controlled by the soil conditions and the loads during and following
installation. It has a significant influence on the subsequent axial and lateral response.
Axial friction
(see [4.3])
Axial breakout response
An initial peak in resistance that is mainly relevant to the first
load response
Axial residual resistance
The large displacement response as the pipe expands or contracts
Cyclic axial response
The long term cyclic response under repeated expansion and
contraction
Lateral breakout response
An initial peak in resistance as the pipe first displaces from the
as-installed position
Lateral residual resistance
The large displacement resistance
Cyclic lateral response
The long term cyclic response, when the pipe becomes embedded
in a trench within a buckled pipe section and soil berms grow
causing a rise in lateral resistance
Vertical stiffness
Static and dynamic stiffness
Lateral stiffness
Static and dynamic stiffness
Lateral resistance
(see [4.4])
Soil stiffness
(see [4.5])
Soil damping
(see [4.6])
Soil damping may be introduced in dynamic analyses.
Specific guidance regarding cyclic resistances (axial and lateral) are not provided in this recommended
practice, as this is an area of ongoing research without clear conclusions. The process is complex, as the
development of a trench at a lateral buckle will reduce the vertical reaction in the midspan, transferring
vertical reaction forces towards the inflection points. When investigated, it should be followed up during
operation through surveys.
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4.2 Pipe embedment
4.2.1 General
The pipe embedment is an important factor influencing the pipe-soil interaction as it will determine the
boundary conditions around the pipe for the subsequent loading scenarios. The initial pipe embedment is
governed by the contact stresses imposed on the soil during laying. The embedment will be influenced by
the installation method (e.g. J-lay, S-lay or reeling), the weight and stiffness properties of the pipe and the
sea state/vessel motions during laying and lay rate. As such, the sea state is considered the biggest source
of uncertainty and a large range in predicted embedment depths is needed to cover possible scenarios. The
embedment will be governed by the bearing capacity of the seabed soils, and different scenarios may occur:
— Pure vertical penetration (laying in calm waters, spools lowered by a crane etc.), see Figure 4-1.
— Combination of vertical and horizontal motions, reducing the vertical bearing capacity, leading to higher
penetrations and also creating more complex boundary conditions, see Figure 4-2.
Figure 4-1 Vertical penetration process
Figure 4-2 Combined vertical and horizontal penetration process
After the pipe is laid on the seabed, usually a pressure test with water-filled pipe is performed. In some
cases, the increased weight of the pipe will lead to further pipe penetration.
The boundary conditions (pipe penetration, trench development etc.) may vary during the operational life,
and regular surveys are recommended to ensure that the boundary conditions around the pipe are covered
by the original design assumptions.
It should be realized that any embedment model will be a simplification as it is not possible to model the true
soil behaviour during an unknown installation scenario including pipe motions in the touchdown zone. In this
section, models for calculating the pure vertical penetration resistance only is given. Discussion of the static
and cyclic effects of laying is included in [4.2.5].
When regional or local pipe embedment measurements are available from existing pipelines, e.g. post-lay or
operational survey data, they can be used to narrow the uncertainty range in the calculation methodology
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when applied to new pipelines nearby. The selection of key parameters such as undrained shear strength, su
or su,r, unit weight, γ’, and touchdown lay factor, klay may be refined using these data, before the calculation
model is applied to the new pipelines under consideration. Observed embedment data should not be applied
directly to new pipelines unless all conditions including pipe characteristics and above mentioned key
parameters are closely the same at the two locations. Instead, the calculation method should be calibrated
and then reapplied, to scale correctly for the differences in the pipeline and soil characteristics and laying
conditions. However, uncertainties related to sea state, type of vessel and the corresponding motions of the
pipeline cannot be fully known and back-calculations should be used with care. Changes in seabed conditions,
e.g. due to seabed mobility, during the operational phase may also be a source of uncertainty when backcalculating embedment from survey data.
In general, due to the uncertainties in the calculation models, more than one model should be evaluated, see
Sec.6. The proposed methods in this section can therefore be considered as examples. Other methods may
also be relevant, see [1.4].
For re-assessments of existing pipelines, embedment measurements may be used directly in PSI
assessments. Embedment measurement error and scatter should then be considered.
Back-analysis of embedment can provide a critical review of parameters initially considered in design and
help in refining the assessed PSI parameters and the subsequent buckling and walking behaviour. This type
of back-analysis is an important aspect of a pipeline integrity management system, especially if the design is
sensitive the PSI parameters.
4.2.2 Definition of pipe embedment
Nominal pipe embedment is defined as the depth of penetration of the invert (bottom of pipe) relative to the
undisturbed seabed (sometimes termed the far embedment in surveys), see Figure 4-3. Pipeline embedment
influences the pipe-soil contact area, which affects the axial and lateral resistance.
Heave of soil during penetration increases the local embedment of the pipe (sometimes termed the near
embedment in surveys). Data indicates that in cohesionless soil this heave may reduce with time and may
then not be reliable in providing additional axial or lateral resistance. The nominal embedment is therefore
the conventional embedment definition in design and assessment to define the pipe-soil contact arc length,
p.
Figure 4-3 Terminology for pipeline embedment
The pipe embedment may vary during the lifetime of the pipeline, changing the axial and lateral pipe-soil
resistance. In this document, the embedment at various stages of the pipeline life cycle is defined as follows:
- Zini
initial pipe invert embedment after laying
- Zf
pipe invert embedment after flooding
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- Zop
pipe invert embedment during first operation (start-up)
-Z
pipe embedment (general). Can be replaced in the calculation methods by Zini, Zf, or Zop depending
on the case considered.
4.2.3 Embedment assessment in undrained conditions
4.2.3.1 General
There are different calculation models available for calculating vertical embedment in undrained conditions.
For a pipe pushed vertically into the soil, the embedment depth will be the depth where the pipe contact
force is in equilibrium with the bearing capacity of the seabed soil. In this section, two approaches are given.
Both models give comparable results for normal conditions, but could deviate for special conditions. It is
recommended to evaluate various models to assess model uncertainties, see Sec.6. Finite element analyses
may also be used as stated in [2.4].
4.2.3.2 Model 1
The penetration resistance may be estimated using the following approach, applying bearing capacity
principles from /8/. The vertical force, Qv, required to penetrate the pipe to the embedment, z, assuming
linear increase in shear strength with depth, may be calculated as:
Qv = Qv0 ∙ (1+dca) + γ’ ∙ Abm
(4.1)
where:
Qv0
γ’
dca
Abm
is the bearing capacity (not including depth effects or soil buoyancy)
is the soil submerged unit weight
is a depth correction factor
is the penetrated cross-sectional area of the pipe, see Equation (4.7)
Qv0 = F ∙ (Nc ∙ su,0 +
ρ ∙ B/4) ∙ B
(4.2)
where:
F
is a function of pipe roughness and of ρB⁄su,0 and can be taken from Figure 4-4. It should be noted
that the roughness in this respect is related to the degree of mobilized shear stress at the pipe-soil
interface. The remoulding process of the soil underneath the pipe during installation will likely be
close to a smooth foundation
Nc
is a bearing capacity factor for clay. For pipes considered as smooth, the bearing capacity factor,
Nc, may be taken as 5.14 for small penetrations, but could reduce to 4 when the pipe embedment
is equal to z=D/2 due to the circular arc shaped foundation base. More detailed discussions can be
found in /7/
su,0
ρ
B
is the undrained shear strength at the reference z-level for depth effects, see Figure 4-5
is the gradient of the undrained shear strength with depth
is the pipe-soil contact width.
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The relation between contact width, B, and the embedment, z, is:
for z < D/2
(4.3)
for z ≥ D/2
B=D
where:
D
z
is the pipe outside diameter including coating
is the pipe embedment.
Figure 4-4 Correction factor according to /8/
It is assumed that there is no depth effect provided that the penetrated pipe is inside the active Rankine
zone. The reference z-level for depth effects, zsu,0, is taken as the seabed for shallow penetrations. For
deeper penetrations, zsu,0 is taken as the depth where a tangent to the pipe at 45° intersects the vertical
line through the edge of the soil/pipe contact, see Figure 4-5. The reference z-level for depth effects may be
expressed as follows:
Zsu,0 = 0
(4.4)
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The shear strength su,0 at the reference z-level for depth effects is taken as:
su,0 = su,z=0 +
ρ ∙ zsu,0
(4.5)
Figure 4-5 Reference level for depth effects in undrained conditions
The depth correction factor, dca, is taken in accordance with DNVGL-RP-C212 as
dca = 0.3 ∙ su,1 ⁄ su,2 ∙ arctan(zsu,0 ⁄B)
(4.6)
where su,1 = (su,z=0 + su,0)⁄2 is the average shear strength above the reference foundation level and su,2 =
Qv0 ⁄ (B ∙ Nc) is the average shear strength below the reference foundation level.
The penetrated cross-sectional area of the pipe, Abm, is taken as:
2
Abm = arcsin(B⁄D) ∙ D ⁄ 4 – B ∙ D⁄4 ∙ cos(arcsin(B⁄D))
Abm =
π∙D
2
for z < D/2
for z ≥ D/2
⁄8 + D ∙ (z – D⁄2)
(4.7)
4.2.3.3 Model 2
An alternative model is described in /9/, /10/ and /11/.
The vertical force required to penetrate the pipe to the embedment, z, is:
(4.8)
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where:
su
D
z
γ’
Abm
is the soil undrained shear strength at pipe invert (and therefore a function of z)
is the pipe outside diameter including coating
is the pipe embedment
is the soil submerged unit weight
is the pipe submerged cross-sectional area (function of z, Equation (4.7)).
The first term of Equation (4.8) represents the soil resistance to pipe penetration. The second term accounts
for the soil buoyancy which is enhanced by soil heave by a factor of 1.5 (based on a best fit value to
numerical analysis, /10/). More information on this parameter can be found in /11/.
At very high embedment ratios (z⁄D>0.5) Equation (4.8) may underestimate the penetration resistance and
the penetration estimate should be used with caution. Alternative bearing capacity factors may be found
in /31/.
4.2.4 Embedment assessment in drained conditions
There are different calculation models available for calculating vertical embedment in drained conditions. The
embedment depth will be the depth where the pipe contact force is in equilibrium with the bearing capacity of
the seabed soil. The static penetration resistance may be estimated using the following approach:
The relation between the contact width, B, and the embedment, z, can be taken from Equation (4.3).
The vertical force, Qv, required to penetrate the pipe to the embedment, z, is calculated as:
Qv = 0.5 ∙ γ' ∙ Nγ ∙ B + z0 ∙ γ' ∙ Nq ∙ dq ∙ B
2
(4.9)
where:
γ’
is the submerged unit weight of soil
Nq
is a bearing capacity factor, see Figure 4-6
Nγ
is a bearing capacity factor /12/, /13/, see Figure 4-6
Nγ = 1.5 ∙ (Nq–1) ∙ tanφ
Nγ = 2 ∙ (Nq + 1) ∙ tanφ
φ
B
dq
z
z0
is the friction angle of the soil
is the pipe-soil contact width
is a factor accounting for depth effects
is the embedment at pipe invert
is the reference z-level for depth effects.
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Figure 4-6 Bearing capacity factors for drained conditions
It is assumed that there is no embedment effect as long as the penetrated pipe is inside the active Rankine
zone, see Figure 4-7. This reference depth, z0, will be dependent on the friction angle of the soil, and can be
found by:
z0 = 0
for z < D⁄2 ∙ [1 – cos (π ⁄4 +
φ⁄2)]
for z > D⁄2 ∙ [1 – cos (π ⁄4 +
φ⁄2)]
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Figure 4-7 Reference level for depth effects in drained conditions
The depth factor can be taken as:
(4.11)
It should be noted that the depth effect is usually not important in sands, because of the large penetration
needed before it is relevant.
The expression for Qv is based on bearing capacity formulae for ideal 2D foundations. Note that if this model
is used to predict the expected penetration z for a given contact force, Qv, it may lead to underestimation of
the true penetration due to effects of the pipe laying process, see [4.2.5].
4.2.5 Effect of laying process on embedment
4.2.5.1 General
Observations show that the as-laid pipeline embedment is typically much greater than would be expected
from the static weight alone, due to motions of the pipeline during laying and the interaction between the
pipe and the soil in the touch down zone /14/. Vertical and horizontal motions of the pipeline within the
touchdown zone may have significant effects on the penetration of the pipe. The penetration of the pipeline
is a result of a complex process, the outcome of which depends on intact and remoulded soil properties, the
weight and stiffness characteristics of the pipeline, the pipe motions and the mechanisms of the gradual cycle
by cycle additional penetration.
The dynamic motions of the pipeline owing to vessel motions are dependent on the sea state during laying,
the water depth and the lay tension and may thus vary significantly.
4.2.5.2 Static touchdown factor
A reference static touchdown lay factor, klay, can be used to account for the increased vertical pipe-soil force.
In the absence of project-specific pipe lay analysis klay may be estimated as /9/:
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(4.12)
where:
klay
EI
Wi
zini
T0
is the touchdown lay factor (= Qv/Wi)
is the pipe bending stiffness
is the submerged pipe weight during installation
is the initial pipe embedment after laying
is the horizontal component of the effective lay tension in the pipe at touch-down point during
installation.
Typical values of klay lie between 1 and 3 depending on the seabed stiffness, lay tension, departure angle and
pipeline bending stiffness. The touchdown factor increases in stiffer soils where the touchdown reaction is
concentrated over a shorter length of pipe. In softer soils, where the reaction is spread over a longer length
of pipe, the touchdown factor converges towards unity (but can never be less than one).
The parameters 0.6 and 0.4 used in Equation (4.12) have been derived from curve fitting to numerical
analyses of the catenary response, with a linear variation of seabed resistance with pipe embedment depth.
0.5
2⁄3
Equation (4.12) applies only for T0>[3∙(EI) ∙Wi]
/9/.
During conceptual design and in the absence of a pipe lay analysis, the horizontal component of lay tension,
T0, can be uncertain. It is then recommended to consider a range of possible lay tension.
The embedment is found by first establishing klay,1=Qv/Wi with penetration depth from [4.2.3.2] or [4.2.3.3].
klay,2 is found by inserting klay,1 in the right hand side of Equation (4.12):
klay,1=Qv/Wi
(4.13)
(4.14)
The compatible embedment and touchdown factor is found when klay,1= klay,2, as illustrated in Figure 4-8.
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Figure 4-8 Example of determining touchdown factor and inital embedment
4.2.5.3 Dynamic effects
In undrained conditions, /14/ recommends using a vertical penetration model for assessing the initial
embedment combined with the remoulded strength to account for dynamic laying effects. In lack of other
data this may be a reasonable approach. However, it should be emphasized that this is a simplification,
as the pipe will not be pushed vertically into fully remoulded soil, but gradually digs itself down due to a
combination of vertical and horizontal pipe movements, see Figure 4-2. In this process the soil is gradually
remoulded but engaging new intact soil as the pipe penetrates. The simplification has been proved to
provide a reasonable fit for a limited number of installed pipelines, /14/ however there is a need to extend
the database with more examples. As-laid surveys will therefore be very valuable in order to increase
the confidence in this simplification. For pipe-laying in calm sea states, the use of remoulded strength
could overestimate the embedment. As such, when a range in embedment is established based on the
above approach, the low estimate embedment should be compared with the static penetration using
intact undrained shear strength to represent laying in calm conditions. Note that the approach using the
remoulded shear strength is only valid for embedment prediction during installation. When evaluating
additional penetration during the hydrotest, a regain of soil strength needs to be considered, and could lead
to penetration into intact soil.
In drained conditions, the observed embedment of pipelines is often higher than predicted using the
submerged pipe weight in a vertical penetration model. The combination of lateral and vertical pipe motions
occurring in the touchdown zone during laying may explain the differences. The embedment is therefore
strongly related to the sea state during laying. The embedment predictions in drained conditions should be
evaluated carefully.
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4.3 Axial pipe-soil interaction
4.3.1 Description of axial response
A typical axial pipe-soil response is illustrated schematically in Figure 4-9. The response is described by an
equivalent friction factor, FA/V.
The soil behaviour during axial pipe movements may be drained or undrained, depending on the rate and
duration of pipe movement, the drainage characteristics of the soil and the pipe surface coating.
During an axial pipe movement in undrained conditions, particularly in the first cycle, an initial peak is often
observed, followed by decay to a steady residual value. During axial pipe movement in drained conditions,
the response is generally ductile with no peak. In analyses of global buckling, a breakout peak generally has
little influence on long-term cyclic walking. However, it may affect the axial force profile along the pipeline,
end expansions during start-up and the possibility of rogue buckle formation, depending on the brittleness of
the response and the magnitude of the pipeline displacements involved.
The difference between the drained and undrained resistance is due to the generation of excess pore
pressure around the pipe in undrained conditions, and therefore is dependent on the soil state, the tendency
for contraction or dilation, and the rate of drainage relative to movement. The soil state may change over
many cycles of movement, as the soil surrounding the pipe is repeatedly failed and consolidated. This causes
a change in the undrained resistance towards the drained value. These mechanisms are discussed in detail
in /15/ and /16/.
The axial PSI is usually idealized in structural modelling with an elastic-plastic model that consists of two
parameters: the limiting (or residual) axial resistance, FA (or equivalent friction, FA/V), and a mobilization
distance, xmob. An initial breakout peak can be incorporated using a piecewise linear axial PSI response.
Figure 4-9 Illustration of axial pipe-soil interaction response
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4.3.2 Framework for axial pipe-soil interaction
4.3.2.1 General
The model for axial pipe-soil interaction (PSI) is based on the following key concepts, which are summarized
in Figure 4-10. Further background information and research is described in /3/ and /17/. This research has
mainly focussed on the residual resistance. However, the same factors are believed to affect the breakout
resistance and will therefore fit into the same framework.
On fine-grained silty or clayey soils the axial response may be undrained, drained or in the intermediate
transitional zone. In this case, design and assessment should as a minimum be based on separate
assessments of drained and undrained axial resistance, with a range bounding both cases being adopted.
On coarse-grained soils without silt or clay, it is generally only necessary to consider the drained resistance.
Figure 4-10 Conceptual model for axial residual resistance
4.3.2.2 Assessment of drained and undrained resistance
Assessments of drained and undrained axial resistance should consider the following:
— The undrained resistance consists of a peak (breakout) and a residual value. A peak may be important in
buckling design, and the effect of both including and neglecting a peak should be checked.
— The drained resistance is usually not significantly affected by large soil displacements/strains.
— The axial resistance is affected by the pipe-soil interface roughness and the effective stress level (the
undrained strength is also affected by any overconsolidation of the soil).
— The axial resistance may be enhanced by a wedging effect, which causes the integrated normal contact
force between the pipe and the soil along the contact area to exceed the pipe weight by a wedging factor
denoted by ζ.
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4.3.2.3 Assessment of transitional drainage and consolidation behaviour
Characteristic values of axial resistance may be refined by considering the level of drainage and consolidation
at the pipe-soil interface. Drainage and consolidation cause the undrained resistance to converge towards the
drained resistance. Specific guidance on how to quantify axial friction in the drained-undrained transition are
not provided in this recommended practice, as this is an area of ongoing research without clear conclusions.
When investigated, the undrained-drained transitional axial friction should be bounded by the undrained
and drained axial resistance and should be based on project specific pipe-soil interface testing program and
followed up during operation through surveys.
When it is not possible to define whether the pipe will respond fully drained or fully undrained, the undraineddrained transitional axial friction may be implemented in the design, alternatively both drained and undrained
response should be considered.
4.3.3 Axial breakout resistance
4.3.3.1 Undrained resistance
The undrained shear strength underneath the pipe is dependent on the load history the soil has been
subjected to. During pipe penetration, the soil underneath the pipe will be at failure until the vertical bearing
capacity is high enough to support the contact force during installation. As such, the soil strength at the
pipe-soil interface will be degraded towards the remoulded value. Subsequently, the soil will reconsolidate
to a higher strength dependent on the soil-pipe contact stresses. It can be assumed that the soil will be
reconsolidated in accordance with the normally consolidated strength ratio (su⁄σ’v)NC.
The pipe is usually pressure tested with water prior to operation which can lead to further penetration. If the
water-filled period is long enough, the soil at the pipe-soil interface will be consolidated to a higher stress
level. This can be treated as an overconsolidation compared to the operational case and the effect on the
shear strength can be taken in accordance with the SHANSEP methodology /18/. The lay-induced pipe-soil
normal stress should not be considered in the assessment as it is only applied for a short period, preventing
full consolidation.
In absence of specialized interface testing, see [3.3.2], a method to estimate the breakout axial resistance is
given in Equation (4.15), written as an equivalent friction factor. Equation (4.15) contains the factors that are
considered to affect the pipe-soil axial resistance in undrained conditions.
(4.15)
where:
α
is the pipe-soil adhesion or roughness factor, representing the reduction in soil-interface
strength compared to soil-soil strength
(Su/σ’v)NC
is a ratio for the normally consolidated shear resistance versus the consolidation vertical
stress. Note that the ratio is stress dependent /19/, as illustrated in Figure 4-11, and at low
stress levels the factor is significantly higher than usually reported in literature for traditional
geotechnical design situations. Typical range is from 0.25 to 0.5 for non-carbonate soils.
V
γpre
is the static vertical pipe-soil force for the condition considered, e.g. operation
is the consolidation preloading effect taken as the ratio between the preloading (e.g. water
filled condition) and the static pipe-soil force, V, for the condition considered (e.g. operation).
This represents the overconsolidation ratio, OCR, of the soil underneath the pipe.
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m
ζ
is a preloading factor that accounts for the long term effect of overconsolidation, being a
number less than 1.0. Typical range is from 0.65 to 0.9.
, where
is the wedging factor, taken as
β is the angle defined in Figure 4-12. For
z>D/2, the wedging factor is constant.
γrate
is a rate factor to account for the speed of loading to undrained failure (to be taken as 1.0 for
a reference speed of 2 hours to failure, may be increased by 10-15% per log cycle of the rate
of loading).
Figure 4-11 Stress dependency on shear strength ratio, typical range
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Figure 4-12 Definition of angle used to determine the wedging factor
Note that recommendations for the normalized shear resistance including preloading effects, effect of pipe
roughness and the dependency of the effective stress level are not generally available in the literature.
Thus, specific soil tests, e.g. special interface tests described in [3.3.2], are recommended. By doing
such tests, the stress dependency and preloading effect can be captured. If such tests are performed it is
recommended to specify testing procedures representative for the actual conditions the pipe will experience
(e.g. consolidate for the stress conditions under the flooded weight and unloading to the operational stress
condition prior to shearing). In lack of such tests, each parameter should be assessed based on engineering
judgement to establish a conservative range. Such judgement could favourably make use of available tests
for similar conditions related to type of clay, preloading effects and roughness of pipe surface. There is a
need in the industry to establish a database containing such test results which can be used to assess the
above-mentioned factors in routine design.
In soft soils, the soil is disturbed and remoulded by the laying process and then reconsolidated by the pipe
weight. Both processes alter the soil strength from the in-situ condition and it is not recommended to relate
the long-term axial resistance to the intact undrained shear strength simply by using an adhesion factor,
α∙su,intact.
In stronger soils, experience shows that the pipe does not become fully bonded to the seabed and therefore
the in-situ soil strength (even after adjustment for interface roughness) cannot be mobilized at the pipesoil interface. The apparent lack of bonding may also be related to the uncertainties in defining the shear
strength in the upper soil, and that the real strength underneath the pipe is not captured in the soil
investigation.
During laying, the soil underneath the pipe will be highly disturbed and any consolidation effects cannot be
relied upon. The axial resistance will thus be lower than the subsequent consolidated axial resistance due to
excess pore pressure in the soil.
4.3.3.2 Drained resistance
The drained breakout resistance is governed by the submerged pipe weight, the drained friction angle of the
soil and the interface properties, and can be expressed by a friction factor:
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(4.16)
where:
δpeak
φpeak
rpipe-soil
ζ
is the peak interface friction angle
is the peak friction angle of the soil
is the pipe-soil roughness factor
is the wedging factor, as given in [4.3.3.1].
Specialized testing (as described in [3.3.2]) for directly assessing the interface friction as a function of
stress-level is recommended. For a given soil, the axial friction for a smooth pipe coating may be as low as
30% of the value for a rough coating /20/. For this reason, it is important to use a representative interface
material when performing laboratory testing. In case such tests are not available, conservative assumptions
should be made for each of the parameters in Equation (4.16).
4.3.4 Axial residual resistance
4.3.4.1 Undrained resistance
The undrained residual friction is believed to follow the same trends as the breakout resistance with respect
to stress-dependency. Specialized testing, as described in [3.3.2], for directly assessing the interface friction
as a function of stress-level is recommended. The sensitivity of the clay is the measure of how much the
intact shear strength degrades when remoulded. As such, it is believed that the soil sensitivity, St, is a
parameter that affects the residual friction, however not necessarily being a direct correlation.
The residual friction can then be expressed as:
μA,res,u = εres ∙ μA,brk,u = εres ∙ α ∙ (su⁄σ'v)NC ∙ γprem ∙ ζ ∙ γrate
(4.17)
where:
μA,brk,u
εres
is the equivalent peak(breakout) axial friction (from Equation (4.15))
is the residual reduction factor, believed to be in the range 1/St<εres<1.
In order to optimize the outcome of the soil investigations/laboratory tests it is recommended to involve
relevant stakeholders in an early phase of a project. This will help defining a scope that accounts for
project-specific data, such as type of coating (including surface roughness) and any foreseen level of preconsolidation from the water-filled condition.
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Guidance note:
As part of the SAFEBUCK JIP, several small scale tests were performed within or made available to the JIP focusing on the residual
axial friction. Based on these tests, the preloading factor m was found to range between 0.35 and 0.6, which is significantly lower
than expected. Experience from SHANSEP tests shows a range from 0.65 to 0.9. The low m values calibrated from the tests are
most likely due to an overestimation of the actual OCR. The reported OCR was related to the consolidation pressure applied for
preparation of the clay in the test bin before the pipe was pushed in-place. A comment to the obtained low m factors was that they
could be explained by the pipe penetration process, i.e. the pipe is erasing some of the in-situ OCR when disturbing the soil during
penetration. The SAFEBUCK tests show the importance of trying to replicate the real stress history of the pipe when carrying out
model tests/interface tests. The OCR underneath the pipe usually originates from the pipe weight during the water filled period and
the tests should replicate this effect as accurate as possible.
---e-n-d---o-f---g-u-i-d-a-n-c-e---n-o-t-e---
4.3.4.2 Drained resistance
The drained residual resistance is usually similar to the breakout resistance, but could potentially be lower
due to a lower residual interface friction angle. Equation (4.16) can be used, replacing the peak friction angle
with the residual friction angle, φres:
μA,res,d = tanδres ∙ ζ = rpipe-soil ∙ tanφres∙ζ
(4.18)
Specialized testing, as described in [3.3.2], for directly assessing the interface friction as a function of stresslevel is recommended.
Published studies (/1/, /2/, /3/ /17/ and /20/and experience indicate that the drained residual interface
friction lies in the range of 0.3<μres<1.0 (or 15°<δres<45°) for non-carbonate soils. Results for carbonate
clays and silts extend to a higher upper limit of
combinations may lie outside this range.
μres~1.4 (δres~55°), see also [7.5]. Other pipeline and soil
4.3.5 Axial mobilization displacements
In structural modelling of a pipeline, the axial response is usually modelled using a bi-linear elasticperfectly plastic model, which requires specification of the mobilization displacement, xmob. The mobilization
displacement defines the initial stiffness and the unload-reload stiffness. If a breakout peak is modelled, xmob
is replaced by two mobilization distances, xbrk and xres. The actual response is generally non-linear, with a
reduction in tangent stiffness as the axial resistance is mobilized.
In assessments of pipe walking, a low value of xmob creates a higher rate of axial walking. To be
conservative, a bi-linear fit to the non-linear response should be a tangent fit to the initial part of the
axial force-displacement response, which represents the elastic recoverable part, see Figure 4-9. In other
situations, such as lateral buckling and feed-in of pipe into the buckle, a higher value of xmob can be more
onerous, and the bi-linear fit should be a secant fit to the displacement when FA,brk or FA,res is fully mobilized.
In the absence of project-specific assessments, the mobilization displacements can be selected using the
advice in Table 4-2. The recommendations may be inaccurate for pipeline properties and soil conditions that
are outside of the model test database, see [1.5.1]. For drained conditions, the tabulated high estimates are
probably too high. In absence of detailed investigations all values of axial breakout and residual mobilization
distance are possible and should be considered in design.
Table 4-2 Axial mobilization displacement
Axial model
Uncertainty case
Parameter
1
Bi-linear
(i.e. with no breakout
peak)
Low estimate, LE
Best estimate, BE
Typical values
Minimum of 1.25 mm and 0.0025∙D
1
xmob
2
High estimate, HE
Minimum of 5 mm and 0.01∙D
Maximum of 250 mm and 0.5∙D
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Axial model
Uncertainty case
Parameter
1
Low estimate, LE
Best estimate, BE
Tri-linear
(i.e. with breakout
peak)
Minimum of 1.25 mm and 0.0025∙D
1
xbrk
2
High estimate, HE
Minimum of 5 mm and 0.01∙D
Maximum of 50 mm and 0.1∙D
1
Low estimate, LE
Best estimate, BE
Typical values
Minimum of 7.5 mm and 0.015∙D
1
xres
2
High estimate, HE
Minimum of 30 mm and 0.06∙D
Maximum of 250 mm and 0.5∙D
Notes:
1)
Represents a tangent fit to the initial part of the axial force-displacement response
2)
Represents a secant fit to the displacement when FA,brk or FA,res is fully mobilized
4.4 Lateral pipe-soil interaction
4.4.1 Description of lateral response
There are two stages of lateral pipe-soil response behaviour:
— First load displacement (monotonic), characterized by a breakout resistance and a steady residual
resistance.
— Cyclic displacement characterized by the growth of soil berms at the limits of the pipe displacement
range, with the pipe descending into a shallow trench. This load scenario may be valid for repeated
heating and cooling corresponding to breaks in production at the location of the pipe where lateral buckle
displacements occur.
The stages above are integrated in the schematic model presented in Figure 4-13. The red line represents
the first load response and the blue line the cyclic response.
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Figure 4-13 Illustration of lateral pipe-soil response
This model is appropriate to light pipes that tend to rise after breaking out from the as-laid position. Heavier
pipelines (heavy pipes) can move downward in the soil after the initial breakout resistance is mobilized. This
downward movement, coupled with the growth of the soil berm ahead of the pipe, leads to a continuous
increase in the lateral resistance, rather than a steady value. The model presented in this recommended
practice is not applicable to heavy pipes beyond mobilization of the breakout resistance. Limited data is
available on heavy pipe behaviour and no general calculation method to assess their behaviour after breakout
is included in this recommended practice. If heavy pipe behaviour is expected, alternative solutions should be
considered, such as initiation of buckles on gravel carpets or sleepers.
If the pipeline tends to penetrate deeper within a lateral buckle during the first buckle initiation or in
subsequent cycles of cooling and heating, such penetration will be counteracted by transfer of vertical
reaction forces towards the shoulders of the buckle. Such 3D effects are not accounted for in typical 2D
testing conditions. Further collection of relevant survey data would improve the background for developing
recommended practice. A particular concern of a high berm resistance relates to a situation where the
production pressure or temperature is increased at a late stage of the operation of the pipeline. Regular
surveys should focus on monitoring berm developments.
In clayey soil, heavy pipe behaviour is observed when the pipe-soil vertical force is more than approximately
half of the ultimate vertical bearing capacity, as calculated at a penetration at half a diameter (V>2∙su∙D).
This is however not a strict limit but special attention should be given when approaching this limit. Heavy
pipe behaviour is not observed in sandy soil.
In some cases, the vertical pipe-soil force may differ from the pipeline weight (e.g. the touchdown area on
either side of free spans, distributed buoyancy sections and natural uneven seabed). In these cases, the
recommendations provided in this recommended practice can be followed by adjusting the vertical pipe-soil
force V.
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This section discusses the parameters:
— Breakout resistance: calculation models to assess the breakout resistance are given in [4.4.2].
— Residual resistance after breakout: calculation models to assess the residual resistance are given in
[4.4.3].
— Mobilization displacements: mobilization displacements are discussed in [4.4.4].
In general, due to the uncertainties in the calculation models, more than one model should be evaluated, see
Sec.6. The proposed methods in this section can therefore be considered as examples. Other methods may
also be relevant, see [1.4].
4.4.2 Lateral breakout resistance
4.4.2.1 General
The boundary conditions around the pipe are very important when evaluating the lateral breakout resistance.
Depending on the pipe motions during laying, different boundary conditions may apply as illustrated in Figure
4-14.
Different models are available for calculation of the breakout resistance. It is appropriate to use a theoretical
approach or a semi-theoretical approach with empirical adjustment factors, if the boundary conditions of the
model are known, e.g. shear strength profile, penetration, contact area, trench geometry, pipe to soil surface
roughness. It is recommended to use a model that is capable of varying the above mentioned effects and
assess how they influence on the break-out resistance. As the geometrical boundary conditions (e.g. trench
geometry) are not known prior to the pipe installation, this represents a challenge in the design phase so the
range in the breakout resistance should include different scenarios. Two models are proposed in this section.
It is recommended to evaluate various models to assess model uncertainties, see Sec.6. The use of finite
element analysis where all parameters affecting the resistance can be defined is recommended for sensitivity
studies, see [2.4]. The challenge however remains to identify the range of assumptions needed to capture
the likely conditions the pipeline could experience.
The boundary conditions (e.g. pipe penetration, trench development etc.) may vary during the operational
life, and regular surveys are necessary to ensure that the boundary conditions around the pipe are covered
by the original design assumptions.
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Pipe in perfect contact with the
seabed – installed in calm sea state
Pipe inside a trench – installed in rough sea state
Figure 4-14 Effect of boundary conditions and laying effects on the lateral resistance
4.4.2.2 Undrained resistance
Model 1
The undrained lateral breakout resistance, FL,brk,u, can be taken as the sum of:
— a shear resistance on a horizontal surface underneath the pipe, FL,brk,u,fric, similar to the axial friction
(from Equation (4.15)) only without the wedging factor
— a remaining resistance, FL,brk,u,remain, due to mobilizing the soil in front of the pipe and at the rear of the
pipe (suction). Suction may be considered if the pipe is installed by vertical penetration, assuring good
contact with the soil at both sides of the pipe, or when a high lateral resistance is unfavourable.
FL,brk,u = FL,brk,u,fric + FL,brk,u,remain
FL,brk,u,fric =
(4.19)
α ∙ (su⁄σ'v)NC ∙ γprem ∙ γrate
(4.20)
(allowing for suction at the rear of the pipe)
(4.21)
(not allowing for suction at the rear of the pipe)
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where:
κa, κp
are soil pressure resistance coefficients, in the range from 2 to 2.5
is the average undrained shear strength within the active failure zone, typically at a depth equal to
z/2
is the average undrained shear strength within the passive failure zone, typically at a depth equal to
z/2.
The possibility to use anisotropic shear strength being different in the passive and active zone is included.
However, it may often be more realistic to use an equivalent average isotropic strength.
If there is no suction at the rear, the resistance due to the weight of the soil is added to the passive
resistance. When the active suction is included the soil weight terms from the two zones cancel each other
out.
A simplification in the model is that the passive resistance is taken in accordance with classical earth
pressure theory against a vertical surface in front of the pipe. As such, it does not fully capture the real
failure surface being affected by the shape of the pipe.
Model 2
An alternative undrained lateral breakout resistance model is as follows:
(4.22)
where:
FL,brk,u
su
D
z
V
γ’
is the lateral breakout resistance in undrained soil conditions
is the soil undrained shear strength at the pipe invert depth
is the pipe outer diameter including coating
is the pipe embedment
is the static vertical pipe-soil force for the condition considered, e.g. operation
is the soil submerged unit weight at the pipe invert depth.
In this equation, the first term reflects the passive resistance related to the shear strength of the soil berm
pushed in front of the pipe. The second term is a frictional component and the third term captures the
passive self-weight resistance from the soil ahead of the pipe.
Equation (4.22) was fitted to a database which comprised of 67 pipe tests on mainly very soft West African
clays (su between 0.4 kPa and 9 kPa at mudline). The pipes were generally embedded by monotonic
penetration and then broken out laterally at a relatively rapid rate. The nominal pipe bearing pressure, V/D,
was generally in the range between 1 kPa and 7 kPa. Results from 5 tests showing heavy pipe behaviour was
included in the calibration of this equation. The model uncertainty factor associated with this model is found
to be 1.5, meaning that the LE and HE resistance respectively is found by dividing and multiplying Equation
(4.22) by 1.5. Note however, that this model uncertainty is related to scatter in test results. Whether these
tests have captured all likely scenarios the pipeline will experience in the field is still uncertain. Uncertainties
in soil parameters along the pipeline route should also be considered.
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Discussion
A disadvantage of model 2 is that the stress-dependency of the frictional part is fixed and does not account
for pre-consolidation effects. Model 1 encourages the engineer to use a consistent frictional term both in the
axial and lateral direction. Also, the passive resistance relating to the shear strength of the pipe is for model
2 using the shear strength at the invert of the pipe thus not reflecting the shear strength variation between
the seabed and the pipe invert, which sometimes may be of significance. The two models are compared in
Figure 4-15 for one specific case, however leaving out the frictional term for simplicity. In model 1 the effect
of allowing for suction at the rear of the pipe can be investigated and this effect is included in the figure.
Allowing for suction to develop at the rear of the pipe could be relevant when a high lateral resistance is
unfavourable. The comparison is included to illustrate the effect of suction at the rear of the pipe and may to
some extent explain the associated model uncertainty with model 2.
Figure 4-15 Comparison of model 1 and model 2 for a typical normally consolidated soil
4.4.2.3 Drained resistance
Different models exist for evaluating lateral pipe resistance in siliceous sand. Most methods are empirical
formulations fitted to a limited number of tests. However, the testing conditions and the validity range for
a given empirical based model are often not well described which makes it difficult to judge whether all
relevant scenarios are captured by the test program.
The empirical equations are based on the breakout resistance, and have therefore taken the effect of the
berm into account. Hydrodynamic effects such as sediment transport and scour could easily with time
wash away or build up a berm, leading to changed boundary conditions for the pipe. In order to take
these changes into account, the methods have to be based on geotechnical theories rather than empirical
equations.
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Model 1
The drained lateral breakout resistance, FL,brk,d, can be taken as the sum of:
— friction between pipe and soil underneath the pipe
— passive earth pressure when pushing a wedge of soil laterally
— the effect of the additional weight of the berm (due to soil heave during penetration).
FL,brk,d = FL,brk,d,fric + FL,brk,d,passive
(4.23)
The passive earth pressure can be taken as:
(4.24)
The passive earth pressure coefficient, Kp, is dependent on the roughness defined as the shear mobilization
on the vertical plane in front of the passive wedge, and can be taken according to classical earth pressure
theory, see Figure 4-16. Note that this roughness parameter, r= τv/(σh' ∙ tanφ) is different from the pipesoil interface roughness or adhesion factor. It is not a material parameter, but rather a parameter that
reflects the direction of the force transferred to the soil in front of the pipe (or in general terms in front of
the foundation) and is as such related to the direction of the load acting on the pipe, see Figure 4-17. The
roughness is an uncertain value and requires special attention. For more details regarding how the roughness
parameter is defined and how it influences the passive earth pressure coefficient, see /21/.
A simplification in the model is that the passive resistance is taken in accordance with classical earth
pressure theory against a vertical surface in front of the pipe. As such, it does not fully capture the real
failure surface being affected by the shape of the pipe.
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Figure 4-16 Passive earth pressure coefficient according to classical earth pressure theory
Figure 4-17 Definition of roughness parameter
When mobilizing the passive resistance, shear stresses in the upward direction acting from the soil in front of
the pipe will develop and reduce the effective normal force acting on the sliding plane underneath the pipe.
The roughness on the vertical plane in front of the passive wedge is then of importance when determining
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the effective normal force underneath the pipe. The upward force acting on this plane cannot be greater than
the pipe-soil contact force. Hence, both the roughness and the passive resistance are in principle limited
by the pipe weight, and will reduce the uncertainty related to the roughness parameter. For light pipes, the
roughness will be close to zero.
The frictional term can be taken as:
FL,brk,d,fric = rpipe-soil ∙ tanφ ∙ (V – r ∙ tanφ ∙ FL,brk,d,passive)
(4.25)
where the term (V – r ∙ tanφ ∙ FL,brk,d,passive) cannot be negative. If that is the case, the assumed roughness
is too high and shall be reduced. Note that rpipe-soil is a roughness factor between the pipe and soil and is
different from r.
Figure 4-18 Illustration of lateral failure mechanism in drained conditions, including effect of
berm next to the pipe
In addition, as illustrated in Figure 4-18, the pipe will get additional resistance from any berm that is built up
next to the pipe. A berm will form from soil heave during penetration, and possibly also due to oscillations
during laying. To account for the additional berm resistance, the displaced area can be idealized as:
(4.26)
where Abm is the penetrated cross-sectional area of the pipe, see Equation (4.7), and the second term is the
displaced soil due to a horizontal amplitude a. The berm area can then for simplicity be evenly distributed
over the extent of the failure surface, xfailure, leading to a modified height, zmod, which is replacing z in
Equation (4.24). Note that the failure surface in accordance with classical earth pressure theory is also
related to the roughness parameter, as indicated in Figure 4-19. How xfailure varies with friction angle and
roughness is given in Figure 4-20.
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Figure 4-19 Failure surfaces for different roughness parameters (example for φ=40°)
Figure 4-20 Normalized failure surface, (xfailure/z), as function of friction angle and roughness
Model 2
A relationship to assess drained lateral breakout resistance based on calibration to full scale model tests
performed on siliceous sands was proposed in /22/. The breakout resistance is divided into a frictional and a
passive component that provides increasing resistance with penetration.
FL,brk,d = 0.6 ∙ V + Fp
(4.27)
where the passive resistance, Fp, is defined as:
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(4.28)
and
FL,brk,d
V
γ’
D
z
is the lateral force at breakout in drained soil conditions
is the vertical pipe-soil force
is the soil submerged unit weight
is the pipe outer diameter including coating
is the pipe embedment.
It should be noted that the tests were performed by first oscillating the pipe laterally until a distance of half
the diameter was obtained, and by that creating a berm in front of the pipe. Then it was further pushed
through the berm and the breakout resistance was recorded. The published method only relates to achieved
penetrations without any evaluation of the effect of how the penetration is achieved.
It should also be noted that the tests were carried out using steel pipes. Hence, for other types of coating
material, the specified friction factor of 0.6 may not be relevant.
Discussion
In order to illustrate the effect of the oscillations prior to breakout in the tests reported in /22/, the two
models are compared in Figure 4-21. The input parameters are given in Table 4-3.
Table 4-3 Input parameters used for comparison of models
Pipe outer diameter, D
0.5 m
Submerged pipe weight, V
1 kN/m
Submerged unit weight of soil, γ’
9 kN/m
Friction angle of sand,
3
φ
40°
Pipe-soil interface roughness factor, rpipe-soil
0.6
As the penetration of the pipe gets deeper, the passive resistance becomes more important. In general, the
mobilized roughness on the vertical passive surface has large impact on passive earth pressure, as seen from
Figure 4-16. Similarly, the roughness is very important for large penetrations for the geotechnical model.
However, when limiting the possible vertical shear onto the vertical passive surface to the submerged pipe
weight, possible roughness values are limited and accordingly the range in passive resistance reduces as
shown in the Figure 4-21 below. No effect of a berm in front of the pipe has been accounted for in model 1.
It is seen that the model 2 predicts significantly higher resistance. However, by including the effect of a berm
due to oscillations prior to breakout (a/D = 0.5), a good fit can be obtained, as shown in Figure 4-22. The
lateral motions imposed to the pipe prior to breakout in /22/ has a large impact.
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Figure 4-21 Comparison of model 1 and model 2, using compatible roughness
Figure 4-22 Comparison of model 1 and model 2, using compatible roughness and accounting for
test conditions used in /22/
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It is important to be aware of the uncertainty related to the seabed configuration around the pipe in sandy
sediments and how this affects the resistance. Where the pipe after laying lays in a bowl shaped trench
rather than with perfect soil contact, see [4.4.2.1], the pipe may have to displace and rise within the bowl
before breaking through, in the same way as experienced for the cyclic tests with large amplitude oscillations
in /22/. For such conditions the embedment as well as the effect of the berm would have to be adjusted
when calculating the breakout resistance.
Hydrodynamic effects like scour could change the seabed geometry around pipelines, and these effects will
vary both along the pipeline and in time. If the berms created during laying are removed due to scour, the
lateral breakout resistance would be significantly reduced. In contrast, in areas of limited water depth, where
there is a significant hydrodynamic effect with varying erosion and deposition, the resulting effect with time
could be a significant self-burial. In this case the breakout lateral resistance could be significantly higher. For
a deterministic design and assessment situation this uncertainty has to be captured by analysing different
seabed configurations. Model 1 is as such considered more appropriate to use when evaluating different
seabed configurations, but it is based on certain assumptions and simplifications.
4.4.3 Lateral residual resistance
4.4.3.1 Undrained resistance
The undrained lateral residual resistance may be calculated as:
(4.29)
where:
FL,res,u
V
D
z
is the lateral residual force after breakout in undrained soil conditions
is the vertical pipe-soil force
is the pipe outer diameter including coating
is the pipe embedment prior to lateral movement.
The numerical parameters in this equation were obtained by fitting the formulation to small scale model test
data, with some constraints to reflect theoretical considerations. The equation may therefore show some bias
to the underlying database conditions. The model uncertainty factor associated with this model is found to
be 1.5, meaning that the LE and HE resistance respectively is found by dividing and multiplying Equation
(4.29) by 1.5. Note however, that this model uncertainty is related to the scatter in the test results, and
whether these tests have captured all likely scenarios the pipeline will experience in the field is still uncertain.
Uncertainties in soil parameters along the pipeline route needs also to be considered.
Equation (4.29) was fitted to a database which comprised of 67 pipe tests on mainly very soft West African
clays (su between 0.4 kPa and 9 kPa at mudline). The pipes were generally embedded by monotonic
penetration and then broken out laterally at a relatively rapid rate. The nominal pipe bearing pressure, V/D,
was generally in the range between 1 kPa and 7 kPa. During lateral movement, a light pipe rises from the aslaid position. However, the model was calibrated based on the initial embedment at the start of the lateral
movement. As the model does not include any physical parameters as soil strength or weight of the pipeline,
it is difficult to assess its applicability outside the conditions of the test database. To increase confidence
in the method, there is a need to test the model further against a larger database, possibly also modifying
the model to account for other parameters. The method is applicable for pipes where V < 2 ∙ su ∙ D, thus
excluding heavy pipes on soft soil. Alternative models to calculate the lateral residual resistance may be
found in /52/ and /53/.
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It should be noted that the tests were performed by moving the pipe sideways but allowing the pipe to move
vertically. For global buckling, buckles are not necessarily initiated by the pipe first moving laterally. For
some buckles, the pipes may rather lift upwards and then move out of the original trench before they transit
into a horizontal buckle. In such cases, initial penetration may not be relevant for determining the residual
resistance and Equation (4.29) should be used with care. This would also be the case for the resistance at
very large displacements.
4.4.3.2 Drained resistance
Model 1
In case of light pipe behaviour, the residual resistance may be considered using the same formulation as
for the breakout resistance, see [4.4.2.3], but using a shallower penetration depth. For a conservative low
estimate, only the frictional component can be used assuming zero embedment.
Model 2
The drained lateral residual resistance may be calculated as:
(4.30)
where:
FL,red,d
V
γ’
Ap
is the lateral residual force in drained soil conditions
D
Dref
is the pipe outer diameter including coating
is the vertical pipe-soil force
is the soil submerged unit weight
is the cross sectional area of the pipe (π ∙ D /4)
2
is reference diameter taken as 20 inches (508 mm).
The formula was obtained by statistical analysis of available laboratory test data for sand of various densities
and pipes from 25 mm to 225 mm diameter, and back-calculated resistances from fitting lateral buckles in
existing pipelines. Note that Equation (4.30) gives the mean or best estimate value.
Low and high estimates can be taken as:
(4.31)
(4.32)
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4.4.4 Lateral mobilization displacements
4.4.4.1 Breakout mobilization
The mobilization displacement to reach the breakout resistance, ybrk, is difficult to predict. Experience shows
that ybrk primarily depends on embedment, z, apart from for very low values of z/D. Low, best and high
estimates of ybrk are given in Table 4-4. These estimates are selected to fit a database for clays, see [1.5.1].
In absence of detailed investigations all values of breakout and residual mobilization distance are possible
and should be considered in design.
The embedment process (static or dynamic) appears to affect ybrk. Model tests in which the pipe is
monotonically pushed to the initial embedment show much smaller ybrk values (and consequently a stiffer
lateral response) than model tests in which the dynamic embedment process is simulated.
For drained conditions, there are less data available regarding mobilization distances. However, the large
ranges given in Table 4-4 are believed also to be representative for drained conditions.
Table 4-4 Lateral mobilization distance to breakout resistance
Parameter
3
Uncertainty case
Typical values
1
Low estimate, LE
ybrk
Best estimate, BE
2
2
High estimate, HE
Notes:
1)
The low estimate is a minimum value which considers the model test results from statically embedded pipe data
2)
The best and high estimates consider statically and dynamically embedded model test data
3)
All values represent a secant fit to the displacement when FL,brk is fully mobilized. No distinction is made between
drained and undrained behaviour.
4.4.4.2 Residual mobilization
The displacement to mobilize the residual resistance, yres, can be estimated using the values in Table
4-5, which are derived from results in /23/ and the database described in [1.5.1]. In absence of detailed
investigations all values of breakout and residual mobilization distance are possible and should be considered
in design.
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Table 4-5 Lateral mobilization displacement to residual resistance
Parameter
Uncertainty case
Typical values
Low estimate, LE
yres
Best estimate, BE
High estimate, HE
4.5 Soil stiffness
4.5.1 General
The soil stiffness should be evaluated differently for static and dynamic analyses. Unless the non-linear soil
response can be explicitly accounted for in the pipeline analysis and elastic springs are used to represent
the soil response, the non-linearity of the soil has to be accounted for. The stiffness should in that case be
a secant stiffness representing the expected load level in the pipeline analysis. The static soil response will
relate to a static loading situation, e.g. maximum load. Dynamic stiffness will be characterized mainly by the
unloading/re-loading situation.
4.5.2 Static soil stiffness
4.5.2.1 Vertical soil stiffness
The static vertical stiffness is a secant stiffness representative for penetration conditions such as during
installation and during development of free spans due to erosion. The static vertical stiffness is defined as:
KV,s = Qv/z
(4.33)
where:
Qv
z
is the static vertical soil reaction per unit length of pipe
is the vertical penetration of the pipe required to mobilize this reaction.
The penetration curve can be established according to [4.2]. Examples of an equivalent secant stiffness at
different load levels are illustrated in Figure 4-23.
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Figure 4-23 Examples of vertical static secant stiffness for different load levels
The penetration curves given in [4.2] are based on a wished-in-place philosophy and an ideally plastic failure
situation, which means not considering the additional distance required to mobilize the failure mechanism at
any given penetration depth. As such, the stiffness may be adjusted by adding a mobilization distance:
KV,s = Qv/(z+
δf)
(4.34)
where:
δf
is the failure mobilization distance at any given penetration depth, which may typically be taken as
10% of the pipe-soil contact width, B.
4.5.2.2 Lateral soil stiffness
The models proposed in [4.4] should be used to establish the lateral resistance curve. The static lateral
stiffness, KL,s, should be estimated from these models. Examples of an equivalent static secant stiffness at
different load levels relating to a bi-linear lateral resistance model are illustrated in Figure 4-24.
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Figure 4-24 Examples of lateral static secant stiffness for different load levels
4.5.3 Dynamic soil stiffness
4.5.3.1 General
The soil stiffness may be evaluated from an equivalent shear modulus of the soil. The shear modulus G,
defined as a secant modulus, is a decreasing function of the shear strain amplitude, γc, in the soil. The shear
modulus, Gmax, for sands at small strains may be calculated from the following expression /24/:
(4.35)
where:
σa
σs
es
is the atmospheric pressure, 100 kPa
is the mean effective stress in soil
is the void ratio
For clays, the small-strain shear modulus Gmax may alternatively be calculated from the undrained shear
strength, su, in the following manner, as a best estimate approximation to laboratory test data /19/:
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(4.36)
where Ip denotes the plasticity index (in absolute numbers), and OCR is the overconsolidation ratio of the
clay.
The relation between the secant shear modulus G and the cyclic shear strain amplitude,
γc, is typically
expressed as a curve of G/Gmax versus γc, with a typical range given in Figure 4-25. More details may be
found in e.g. /25/. Such relations should be used with care in particular at large strains, where it is important
to assure that shear stresses exceeding the shear strength are not obtained.
Figure 4-25 G/Gmax as function of cyclic shear strain, typical range
Note that other approaches may also be relevant, and the models presented above can be considered as
examples. The soil behaviour during dynamic loading is complex and require specific assessments. Further
guidance may be found in DNVGL-RP-C212.
For free-span specific scenarios, see [7.2].
4.5.3.2 Vertical soil stiffness
For determination of the dynamic vertical stiffness, KV,d, the following expression may be applied:
(4.37)
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which is based on elastic half space theory for a rectangular foundation under assumption of a pipe length
that equals 10 times the contact width between pipe and soil. The Poisson’s ratio for soil, v, can be taken
from 0.3 to 0.35 for sand and from 0.45 to 0.5 for clay. More details about elastic soil stiffness can be found
in /26/.
The main challenge is to estimate an equivalent shear modulus representative for the elastic half space.
4.5.3.3 Lateral soil stiffness
For determination of the dynamic lateral stiffness, KL,d, the following expression may be applied:
KL,d = 0.76 ∙ G ∙ (1 + ν)
(4.38)
which is based on elastic half space theory for a rectangular foundation under assumption of a pipe length
that equals 10 times the contact width between pipe and soil. The Poisson’s ratio of soil, v, can be taken
from 0.3 to 0.35 for sand and from 0.45 to 0.5 for clay. More details about elastic soil stiffness can be found
in /26/.
The main challenge is to estimate an equivalent shear modulus representative for the elastic half space.
4.6 Soil damping
The soil damping is generally dependent on the dynamic loads acting on the soil. It can be distinguished
between two different types of soil damping mechanisms:
— radiation damping associated with propagation of elastic waves through the yield zone
— material damping associated with hysteresis effects taking place close to the yield zone in contact with the
pipe.
The radiation damping may be evaluated from available solutions for elastic soils using relevant soil modulus
reflecting the soil stress (or strain) levels. The radiation damping depends highly on the frequency of the
oscillations, and is more important for high frequency oscillations. Soil damping for free spanning pipelines is
normally governed by soil material damping.
The case-specific modal soil damping ratio,
ζsoil, due to pipe-soil interaction may be determined by:
(4.39)
where the soil damping per unit length, c(s), may be defined on the basis of an energy balance between the
maximum elastic energy stored by the soil during an oscillation cycle and the energy dissipated by a viscous
damper in the same cycle.
The equation may be solved from a finite element analysis of the pipe modelled with discrete soil supports.
The viscous damping coefficient ci of support no. i can be calculated from:
(4.40)
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where:
ki
ζsoil,i
ω
is the linearized spring stiffness at support no. i
is the damping ratio representing support no. i
is the angular frequency of the mode considered.
Knowing the non-linear hysteretic reaction of a support length the damping ratio representing the support
can be calculated as:
(4.41)
where:
EDissipated
EElastic
is the energy dissipation at support no. i, as illustrated on Figure 4-26
is the equivalent elastic energy at support no. i, as illustrated on Figure 4-26.
Figure 4-26 Energy dissipation at soil support, shown in the load-displacement space
Because of the soil non-linearity, the equivalent spring stiffness and the damping ratio are dependent on the
displacements at the support. For a case-specific determination of the modal soil damping ratio this needs to
be taken into account. An iterative solution will be required to assure compatibility between:
— the dependency of the
— the dependency of the
— the dependency of the
support displacements
— the dependency of the
mode-shape on the equivalent support springs
oscillation amplitude on the modal damping
equivalent springs and the damping ratio of the discrete soil supports on the cyclic
modal damping ratio on mode-shape and on support springs and damping ratio.
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As basis for the iterations non-linear relationships for spring stiffness, ki, and damping ratio, ζi, as function
of pipe penetration and cyclic displacements for the relevant soil and pipe diameter are required. Such
relationships are qualitatively shown in Figure 4-27, and may be determined either experimentally or
analytically.
For free-span specific scenarios, see [7.2].
Figure 4-27 Non-linear characteristics of soil stiffness and damping
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SECTION 5 BURIED AND COVERED PIPELINES
5.1 General
This section gives recommendations for design conditions relevant for buried pipelines. Unlike exposed
pipelines, buried pipelines are designed to stay in place. Thus, the focus in the pipe-soil interaction
assessment should be on the behaviour prior to breakout.
The condition of the soil surrounding the pipe is the most important aspect in any assessment of buried
pipelines. This will to some extent be influenced by the trenching method. Further, the boundary conditions
around the pipe (e.g. depth to intact soil conditions below the pipe in the trench) will add to the uncertainties
related to selection of geotechnical parameters.
Unexpected and poorly defined ground conditions are generally the most commonly re-occurring causes
of delays in project schedule and budget challenges. In many cases marine pipeline projects, including
buried pipelines, will be particularly exposed to such risks due to their reliance on satisfactory seafloor and
subsurface soil properties for trouble free installation and operation.
For pipelines that are buried by ploughing or jetting, the largest uncertainty is related to how the trenching
method has affected the in-situ strength and stiffness parameters. The soil investigation program needs
to consider both the intact soil conditions and the soil conditions following a trenching/jetting operation.
The latter may require the construction of a certain length of dummy trench as part of the soil investigation
program.
In the calculation models presented in this section, it is assumed infinite width of the backfill material.
In case of a rock berm with limited width, the contact stresses around the pipe can be different from the
assumption. Finite element analyses are then recommended to assess the normal stresses acting on the pipe
surface.
In general, due to the uncertainties in the calculation models, more than one model should be evaluated, see
Sec.6. The proposed methods in this section can therefore be considered as examples. Other methods may
also be relevant, see [1.4].
In case pipelines are covered with rock, the effect of soil settlements should be considered.
5.2 Effect of trenching method
5.2.1 Jetting
When jetting in soft clay a water-clay suspension is expected to prevail in the trench immediately after
trenching. The pipe will be completely surrounded by this material when it is installed in the trench, and
the water/clay suspension will gradually settle. Minor penetration of the pipe into the intact material at the
bottom of the trench may also be expected. The shear strength of the clay surrounding the pipe will gradually
increase from practically zero to that of a normally consolidated clay, depending on the coefficient of
consolidation and the thickness of the clay layer. When estimating the reconsolidated strength, uncertainties
in both the normally consolidated shear strength ratio, su⁄σ'v, the unit weight and the degree of backfill
should be considered. Following the consolidation process the height of the clay backfill will decrease.
With time, the clay in the trench will regain shear strength. The regained shear strength is eventually
expected to reach a constant level.
The upward displacements required to reach the maximum uplift resistance are also likely to be affected by
the method of trenching. Jetting may introduce water filled voids in the soil in the trench, but generally the
soil in the trench will form a homogeneous material.
It is not recommended to rely on any upheaval resistance in a jetted cohesive soil shortly after installation.
The upheaval resistance at time of operation needs to rely on the degree of consolidation of the backfilled
material.
When trenching by jetting in sandy soil, the sand is fluidized before settling which is similar to the process
of preparing soil for determination of minimum density in the laboratory /54/. Thus initially after trenching
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the sand will have close to zero relative density. Any kind of subsequent loading may densify the sand.
One potential source is varying hydrodynamic pressure on the seabed due to waves. This will cause some
vertical gradients of the hydrodynamic pore pressures, but more important also impose shear stresses to
the soil /55/, which may build up pore pressures in the sand. Such pore pressure build up with subsequent
drainage may lead to a densification of the sand, which will be favourable if occurring prior to the operation
of the pipeline. However, increased pore pressures will reduce the upheaval resistance compared to
that calculated for a fully drained condition. This should be carefully considered if water depth and wave
conditions are such that pore pressures may build up during a storm. The shear stress level in the soil may
be evaluated from /55/.
5.2.2 Ploughing
When the pipe is placed in a soft clay trench formed by a ploughing device with subsequent backfill of the
trenched material, the water content of the clay will not increase relative to that of the intact material. Thus
the remoulded resistance of the clay as established through sensitivity measurements is likely to represent
an expected minimum strength.
The regained shear strength with time is eventually expected to reach a strength which is proportional to the
effective stresses in accordance with theory for normally consolidated clays, or equal to the remoulded shear
strength, whichever is the greater.
Ploughing is expected to change the macro structure of the clay by introducing cracks and water-filled
voids. When ploughing in a stiff clay, the clay is likely to break up forming lumps of clay, and the upheaval
resistance would be related to the interface shear resistance between lumps of clay rather than the shear
strength in the clay material.
Ploughing in sandy conditions will affect the relative density, however less severe compared to jetting.
Subsequent densification due to wave loading may increase the relative density with time, but such positive
effects are difficult to quantify.
5.3 Axial pipe-soil interaction
5.3.1 General
The axial resistance of a buried pipeline can be determined by investigating two different failure modes; a
deep and a shallow mode (as illustrated in Figure 5-1 where a two-layer model is shown). The difference
between the modes is whether the soil above the pipe slides axially together with the pipe, or if the shear
resistance of the backfilled material is high enough to prevent soil movement of the soil above the pipe.
The shallow mode will be governing for low cover heights.
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Figure 5-1 Illustration of axial failure modes for buried pipelines
For each scenario the engineer needs to assess the material surrounding the pipe, drainage conditions and
loading rate to evaluate whether the soil surrounding the pipe will behave drained or undrained.
5.3.2 Drained resistance
A two-layer model, assuming ideal conditions, as given in Figure 5-1, can be used to estimate the axial
friction in drained conditions. The equation for the deep mode has been found by integration of contact
stresses around the pipe circumference assuming two different layers. The equation for the shallow mode is
found by replacing the formula for the upper half of pipe with the vertical shear resistance through the upper
layer. The ultimate resistance will be the lowest resistance obtained by the two failure modes.
Deep failure mode:
(5.1)
Shallow failure mode:
(5.2)
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where:
μseabed
μfill
γ’seabed
γ’fill
K0,seabed
K0,fill
H
D
V
φfill
is the friction coefficient between the pipe and the seabed soil
is the friction coefficient between the pipe and the backfilled material
is the submerged unit weight of the seabed soil
is the submerged unit weight of the backfilled material
is the lateral pressure coefficient of the seabed soil
is the lateral pressure coefficient of the backfilled material
is the cover height (above pipe)
is the outer pipe diameter including coating
is the submerged weight of pipe
is the friction angle of the backfilled material.
K0 may be taken as 1 – sinφ of the respective material. Note that Equation (5.1) and Equation (5.2) assume
that the pipe is penetrated D/2 into the seabed, which is most likely not the case. Lower penetrations are
expected for sandy soils. For pipelines covered with rock and a low friction is unfavourable, it is conservative
to use K0,fill also for the lower half of the pipe, as it results in lower normal stresses acting on the pipe.
For guidance on axial friction coefficients in rock cover, /27/ presents results from relevant full-scale tests.
5.3.3 Undrained resistance
An undrained resistance may be relevant to consider in case of short term loading (e.g. trawl impact for
partly exposed pipes lying on clay seabed, see Figure 5-2). For pipes covered with rock Equation (5.1) and
Equation (5.2) for drained conditions can be used for the upper half of pipe.
The undrained shear resistance around the pipe will be governed by the effective stresses in the
reconsolidated soil underneath the pipe, controlled by the same factors as for axial friction for exposed pipes,
see [4.3.3.1]. Buried pipes are, however, exposed to a higher stress level, and the weight of the rock fill and
the pipe weight should be used to assess the reconsolidated strength in the underlying clay.
In case the backfill material consists of clay, the shear strength acting on the upper half of the pipe will be
determined by the weight of the backfilled material, but as mentioned in [5.2], it is challenging to estimate
the shear strength of backfilled clay.
The same failure modes as for the drained resistance model should be checked, and the ultimate resistance
is the lowest resistance obtained from the two modes.
Deep failure mode:
(5.3)
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Shallow failure mode:
(5.4)
where:
α
u,reconsolidated
u,bottom,reconsolidated
u,backfill,reconsolidated
H
D
is the pipe-soil roughness or adhesion factor
is the average value of reconsolidated shear strength around the pipe
is the average value of reconsolidated shear strength at the bottom half of the pipe
is the average value of reconsolidated shear strength along the vertical failure plane
in the backfilled material
is the cover height (above the pipe)
is the outer pipe diameter including coating.
5.4 Lateral pipe-soil interaction within rock berms
The lateral resistance for a buried pipe is usually not a problem when the pipe is protected from anchor and
trawl loads by a sufficient burial depth. For buckling analyses, the uplift resistance is usually lower and thus
more important, unless the lateral out-of-straightness is high.
Exposed pipes that could be exposed to trawl loads are sometimes covered with rock berms at a certain
spacing to withstand these loads, see Figure 5-2. Spot dumping is also sometimes used to ensure on-bottom
stability.
Figure 5-2 Example of trawl loading on partly exposed pipeline, top view
The lateral resistance within a rock berm is generally governed by its passive resistance. However, two
possible failure modes need to be checked, as illustrated in Figure 5-3:
— Failure through the rock berm.
— Failure in the underlying material. Close to the toe of the rock berm, the failure will go through the rock
due to a lower shear strength.
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Figure 5-3 Possible lateral failure modes for a pipe inside a rock berm with low cover heights
The ultimate resistance will be the lowest resistance obtained by the two modes. The failure through the rock
berm will be governed by the friction angle of the rock berm material and the effective stresses along the
failure line. The inclination of the failure plane that gives the lowest resistance should be sought. The failure
in the underlying material will be governed by a lower friction angle and effective stresses (if sand) or the
consolidated undrained shear strength (if clay). The resistance can be obtained through a limit-equilibrium
analysis or modelled explicitly in a finite element program. Note that the plane failure surfaces indicated
in Figure 5-3 inside the rock berm are not valid for large cover heights. For large cover heights, the failure
planes within the rock berm will be curved and require the use of finite element programs.
5.5 Uplift resistance
5.5.1 Drained resistance
In global buckling analyses of buried pipelines, the uplift resistance is the critical aspect. There are several
calculation models available to estimate the uplift resistance, and most include the resistance from the
weight and shear resistance through the overlying soil. In /28/, a vertical slip surface through the overlying
soil is assumed, and is a simplification, as research has indicated that the slip surfaces are inclined. In /29/,
an uplift resistance including the angle of dilatancy is proposed as a more theoretical model.
Regardless of the method used to calculate the uplift resistance, it should be kept in mind that in accordance
with DNVGL-RP-F110, the low estimate soil resistance should be selected as two standard deviations below
the mean or with a fractile in the order of 2% to 5%. Thus best fit relations should not be used without
evaluating the uncertainties involved. The main uncertainty for a trenched and backfilled pipe will be the
characteristics of the backfilled soil and how this affects the calculated resistance or compares to any test
database.
Assuming vertical slip surfaces, the total resistance, in terms of weight and integrated shear resistance as
illustrated in Figure 5-4 can be written as:
(5.5)
where:
H
γ’
is the cover height (above the pipe)
is the submerged weight of soil
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K
φ
D
is the lateral earth pressure coefficient also accounting for increase in vertical stress during uplift
is the angle of internal friction of the backfilled material
is the outer pipe diameter including coating.
Figure 5-4 Vertical slip model
Equation (5.5) can be rewritten as:
(5.6)
where:
f
is the uplift resistance factor.
The key issue is to assess an adequate uplift resistance factor, f. The uplift resistance factor may be
calculated from a drained (peak) friction angle and a lateral earth pressure coefficient or estimated from
model test results.
If the friction angle is used as basis for the assessment, it should be emphasized that the friction angle
should be established at the relevant stress level. Relative density may be used for correlation to drained
friction angle when considering the following aspects:
— The drained friction angle of a backfill will be dependent on the relative density of the material. This may
be found by carrying out laboratory tests, but the in-situ density for the trenched and backfilled material
has to be estimated.
— The relative density may in-situ be established through cone penetration test (CPT) measurements,
considering though that established relations between cone resistance and density generally have less
accuracy in the low stress region. Backfilled material will have different density than the in-situ soil, so insitu CPT tests would be non-conservative to use.
— Relative density is a poor measure for gravel and rock.
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— As a consequence, any interaction between the uplift resistance factor and other relevant soil parameters
shall be evaluated carefully.
Equation (5.5) was fitted to uplift tests, and it was found that different values of the lateral earth pressure
were needed for different types of sand. The assumptions needed to obtain the best fit with tests are given
in Table 5-1. It should be noted that a good fit for medium dense and dense sand was found by adopting
a passive earth pressure coefficient using a roughness close to -1. This relates to having a full mobilization
of shear resistance along the assumed vertical failure surface when mobilizing the passive wedge on the
outside.
Table 5-1 Results for best estimates of uplift resistance factor
φ(°)
K model
Roughness r
Best estimate f
Loose
30
K0
N/A
0.29
Medium
35
Kp
-1.00
0.47
Dense
40
Kp
-0.97
0.62
Sand type
A slightly simpler, empirical model for the low estimate uplift resistance factor, fLE, is found to be:
(5.7)
where the friction angle,
φ, is given in degrees.
Best estimate and high estimates can be found by adding 0.19 and 0.38 respectively to the low estimate
values. Figure 5-5 presents the variation in the uplift resistance factor with peak friction angle.
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Figure 5-5 Uplift resistance factor vs peak friction angle
For gravel and rock backfill, the uplift resistance factor is expected to be similar to dense sand, i.e. between
0.6 and 1.0.
The displacement required to mobilize the uplift resistance, Fuplift, is uncertain. In /56/ it is indicated that the
mobilization distance can be as high as 0.1 times the height H for loose sands. Note that in such loose sands,
the soil above the pipe may be compressed before the failure surfaces are developed, requiring additional
displacements to fully mobilize the uplift resistance compared to denser material. For pipes deeply embedded
in very loose sand, failure may be triggered by a flow around mechanism before changing to a failure mode
extending to the surface. Additional discussion regarding uplift resistance and mobilization distances may be
found in /57/.
For gravel and rock backfill, a mobilization distance equal to 0.01 times the height H is considered
appropriate.
A bi-linear or tri-linear curve is typically used. The initial stiffness in a tri-linear model should be carefully
assessed.
The validity range of Equation (5.7) with respect to soil cover ratios is as follows:
Loose sand:
Medium/dense sand/rock fill:
When backcover or rock dumping to above the virgin seabed is performed, such cover material should be
scour resistant. Tolerances for rock dumping should be accounted for when specifying cover heights.
Note, as discussed in [5.2.1], that for pipes being buried by jetting in sand the backfilled sand will be
very loose and may not be covered by the above experimental tests. Such very loose sand also has a high
potential for liquefaction, whereby no uplift resistance will result. Buried pipelines with low specific weight
would under such a condition float up even when not exposed to high temperature or pressure.
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5.5.2 Undrained resistance
This subsection applies to pipelines buried by jetting or by ploughing in soft clay where the backfilling
operation form a uniform material with no major voids. Ploughing in stiff clay with subsequent backfilling will
result in a backfilling consisting of lumps of clay, where the upheaval resistance will be governed by the shear
strength at the interface between the lumps.
There are mainly two different failure modes that govern the development of uplift resistance in undrained
conditions:
— a local soil failure mode where the soil above the pipe is displaced around and beneath the pipe
— a global soil failure mode where a wedge extending to the soil surface is lifted together with the pipe.
The local soil failure mode will be a simple function of the shear strength at the depth of the pipe, whereas
the global soil failure mode implies a combination of weight and shear resistance in the overlying soil. The
two failure modes are illustrated in Figure 5-6. The ultimate resistance will be the lowest resistance obtained
by the two modes.
For a sustained uplift force, a drained model may be more appropriate also in clay. Regarding the applicability
of a drained or undrained model in clay, it will be up to the geotechnical engineer in cooperation with the
pipeline engineer to assess the loading and drainage conditions in any specific case. Effect of undrained creep
should also be considered when selecting an appropriate undrained shear strength profile.
It should be noted that there are significant uncertainties related to the properties of backfilled clay, and
the alternative solution to install the pipeline on the clay seabed and cover it with rock is the recommended
approach to provide more reliable uplift resistance.
Regardless of the method used to calculate the uplift resistance, it should be kept in mind that in accordance
with DNVGL-RP-F110, the low estimate soil resistance should be selected as two standard deviations below
the mean or with a fractile in the order of 2% to 5%.
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Figure 5-6 Vertical failure modes for uplift resistance in undrained conditions
Local failure mode:
The local failure mode will be a function of the reconsolidated shear strength within the failure mode around
the pipe:
(5.8)
where the bearing capacity factor, Nc, corresponding to a deep failure mode is in the range between 9
and 12, /30/. The case considered in /30/ is related to the deep failure mode of laterally loaded piles with
constant undrained shear strength. The last term in Equation (5.8) represent a soil buoyancy effect.
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Global failure mode:
The global failure mode assumes that the soil above the pipe is lifted together with the pipe, and two vertical
failure planes are formed. The total resistance is similar to the drained resistance model. However, the shear
resistance through the overlying soil is governed by the average undrained shear strength along the assumed
failure planes:
(5.9)
For both the global and local failure mode, the average shear strength within the failure surface should be
used. The user should have the failure mode in mind when selecting an appropriate average shear strength,
. The average shear strength to be used in the two equations is likely to be different.
The assumption of vertical slip planes is a simplification. Other formulae or approaches may be more relevant
for pipelines, e.g. finite element analyses to determine uplift resistance for various shear strength profiles
that account explicitly for linearly increasing undrained shear strength with depth /31/. Regardless of the
chosen methodology to calculate the uplift resistance, the main challenge is to define the shear strength of
backfilled clay, see [5.2].
5.6 Rock fill over backfilled clay
In case backfilled clay is covering the pipe and additional rock is used to increase the uplift resistance, the
full effect of the additional weight and shear resistance from the rock cannot be mobilized until the pipe has
moved through the clay and obtained contact with the rock particles. The initial failure mode will in that case
be a local failure within the backfilled clay, and the only benefit of adding rock on top will be that the added
weight will consolidate the soil, leading to a higher shear strength of the clay. When taking account for such
increase of the shear strength the time for consolidation should be considered.
Also when adding rock on top of a very soft clay within a narrow trench, full effect of the additional rock may
not be achieved because of the much higher compressibility of the soft clay inside the trench as compared to
the intact clay at both sides of the trench. This may cause part of the rock fill weight above the trench to be
transferred to the intact clay on both sides. A 2D consolidation analysis should be performed to evaluate such
an effect. The possibility of an inclined failure should also be evaluated, depending on soil conditions, trench
geometry, out-of-straightness, lateral extent of rock cover etc.
5.7 Downward resistance and stiffness
Pipelines laid on soft clay, may buckle downwards instead of upwards. Pipelines on soft clay may also
experience downward displacements that influence the initial out-of-straightness of the pipeline. Also for
upheaval buckling, the downward stiffness at the shoulders is important to estimate the pipe curvature
and bending stresses. The load-displacement curve is non-linear and is dependent on the undrained shear
strength underneath the pipe. The soil strength underneath the pipe is further determined by its load history.
The undrained shear strength will deviate from the in-situ shear strength because of effects of pipe laying,
trenching, hydrotest, weight of backfill/rock fill etc. The downward resistance may be calculated similar to the
models for penetration resistance for exposed pipes in [4.2], but accounting for the overburden pressure and
possibly a deep failure mode.
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SECTION 6 TREATMENT OF UNCERTAINTIES
6.1 General
Soil data to be used as a basis for geotechnical analysis are usually encumbered with uncertainty, both in
terms of natural variability and in terms of limited amounts of data. The selection of characteristic values
of soil properties for use in geotechnical design and assessment is often based on subjective judgment and
accumulated experience, and the uncertainties which are involved with the soil properties are only to a
limited extent brought into the picture when the characteristic values are chosen.
Uncertainties associated with geotechnical data have many sources and may be divided into the following two
main types of uncertainty:
— aleatory uncertainty, i.e. physical uncertainty
— epistemic uncertainty, i.e. uncertainty related to imperfect knowledge.
Aleatory uncertainty is also known as inherent uncertainty and intrinsic uncertainty and is a natural
randomness of a quantity such as the variability in the soil strength from point to point within a soil volume.
Such physical uncertainty or natural variability is a type of uncertainty which cannot be reduced.
Epistemic uncertainty consists of statistical uncertainty, model uncertainty and measurement uncertainty,
which are all classified as a type of uncertainty associated with limited, insufficient or imprecise knowledge.
Epistemic uncertainty can in principle be reduced by collection of more data, by improving engineering
models and by employing more accurate methods of measurement.
Statistical uncertainty is uncertainty due to limited information such as a limited number of observations of
a quantity, e.g. a limited number of soil strength values from a limited number of soil samples tested in the
laboratory.
Model uncertainty is uncertainty due to imperfections and idealizations made in:
— Applied engineering models for representation and prediction of quantities, such as soil resistances.
— Choices of probability distribution types for representation of uncertain quantities. Model uncertainty
involves two elements: a bias if the model systematically leads to overprediction or underprediction
of a quantity in question and a randomness associated with the variability in the predictions from one
prediction of that quantity to another.
Advice on the statistical treatment of uncertainty in soil data is given in DNVGL-RP-C207.
6.2 Considerations for pipeline design and assessment
6.2.1 General
Shear strength variations along a long pipeline route may be large. Typically, the obtained data from soil
investigations will not cover the potential local variability as shown in the example in Figure 6-1, where soil
investigations are performed every second kilometer. By looking at all the data from the soil investigations,
variations in shear strength will be large and could be represented by a wide probability distribution function.
If however, the design situation is governed by the soil conditions locally compared to the spacing of the soil
investigations, the range established for the entire route would be too wide. For example, in the formation
of a lateral buckle, the soil conditions within 100 m to 200 m would be governing for the response of the
pipeline. If soil investigations had been made to capture the soil conditions within such a short distance, the
range would be narrower. Ideally, such detailed soil information would be preferable, but is not possible to
obtain along the entire route. For each design situation, the designer needs to evaluate whether a high or low
value of shear strength is unfavourable and estimate a range to be used based on the available information.
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Guidance note:
High and low estimates of characteristic values are cautious estimates, i.e. they are – in statistical terms – estimates with
confidence. For estimation of characteristic values with confidence, see DNVGL-RP-C207 and DNVGL-RP-C212. Note that the terms
lower and upper bound are also frequently used in the industry. However, these terms may give the impression that we have
perfect knowledge and that these bounds are absolute limits for the quantity in question. This is never the case, and the terms low
and high estimates are considered more appropriate and used consistently throughout this recommended practice.
---e-n-d---o-f---g-u-i-d-a-n-c-e---n-o-t-e---
Figure 6-1 Example of how soil conditions may vary along the pipeline route with respect to the
available soil data
The range in soil resistance shall in addition to soil variability also reflect other uncertainties, such as
assumptions of laying effects and uncertainties related to selection of calculation method. Note that the
specific pipeline recommended practices may have different approaches on how to combine low, best and
high estimates for different pipe-soil interaction parameters.
Because of the uncertainties in governing soil parameters, load effects, idealization of calculation models etc.,
it is difficult to define universally valid methods for simulation of pipe-soil interaction effects. The limitations
of the methods used, whether theoretically or empirically based, shall be thoroughly considered in relation to
the problem at hand. Extrapolation beyond documented validity of a method shall be performed with care, as
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shall simplifications from the problem at hand to the calculation model used. In general, the use of more than
one calculation approach is recommended.
It is recommended to enhance the confidence in the chosen engineering models by carrying out field
monitoring. It can then be confirmed if the operating behaviour of the pipeline is consistent with the design
assumptions. General requirements related to the execution of surveys can be found in DNVGL-ST-F101.
A contingency plan for intervention, appropriate for the uncertainty in the design assumptions is also
recommended. Note that it may be difficult to back-calculate specific pipe-soil interaction parameters as
there are many sources of uncertainties related to installation effects and local variations.
6.2.2 Model uncertainties
When evaluating the uncertainties or a probability distribution related to a certain type of soil resistance it
is important to take into account that there may be model uncertainties that need to be accounted for in
addition to the uncertainties related to the parameters that enter into the model. Such uncertainties relate
to:
— The scatter around the fit between measured and calculated resistance when empirical methods are
applied. The shape of the probability distribution associated with this model uncertainty can be obtained
from a relevant test database.
— Different boundary conditions” between the test database and the real conditions for the pipeline,
including when model parameters for the design condition is outside the range found in the test database
and when the tests do not include load steps which real pipeline will experience. Such differences between
the testing conditions and the real conditions may also lead to systematic errors. Therefore, it is important
that the background for any method is clearly defined. Boundary conditions in this context include shear
strength variations around the pipe accounting for effects of remoulding and penetration during laying,
subsequent reconsolidation including consolidation for a waterfilled condition during pressure testing,
depth of penetration, contact area, trench geometry, pipe to soil surface roughness etc. Such boundary
conditions may be a larger source of uncertainty than the calculation algorithm for defined boundary
conditions or the uncertainty related to physical, e.g. geotechnical, parameters entering the algorithm.
It is important that the engineer understands the limitations of a given model. Model uncertainties may
be captured by using alternative models. The shape of the probability distribution of these effects is
often unknown and is most probably correlated (e.g. a large laying effect will give higher penetration
and probably also change the lateral geometrical boundary conditions). These effects will not necessarily
create a wider range around the best estimate value, but rather shift the range in one direction as
compared to any single calculation model. As many of the sources to the uncertainty have unknown
probability distributions, a deterministic approach is favourable compared to a probabilistic approach. A
suitable design range will in that case include a certain degree of engineering judgement. If employing
probabilistic analyses a model uncertainty probability distribution should be rather wide and would not
necessarily be centred about the applied method if only one method is considered.
As there are many different calculation models available and most are empirical, it is important that for each
specific model, a sensitivity check is performed with reasonable ranges of input parameters entering into the
model to identify the uncertainty of the total resistance. It is recommended to evaluate different calculation
models in order to better capture the modelling uncertainty of a given design situation, see also [1.4].
For the models based on geotechnical principles provided in this recommended practice, the modelling
uncertainty is unknown. However, the idea is that if the models are capable of representing different
scenarios and include the most important input parameters, a suitable range in the quantity (e.g. soil
resistance) can be found by defining ranges for relevant input soil parameters and parameters that defines
the model conditions. An example of different model conditions can be seen in Equation (4.21) where
the user can either account for suction at the rear of the pipe or not in the calculation of lateral breakout
resistance. Finite element analysis where all parameters and conditions affecting the soil resistance can be
defined is considered to be the preferable approach to carry out sensitivity studies, see [2.4].
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6.2.3 Combinations of uncertainties
It should be kept in mind that it could be overly conservative to combine the extremes of each parameter
range entering into the calculation models as the extremes will usually not occur in the same position.
For example, there is negative correlation between shear strength and embedment depth, such that an
extremely large embedment is not likely to occur in a position with an extremely high shear strength. In
general, for a specific calculation model, the input parameters which have the largest effect on the end result
should be given more weight than less important parameters.
There may also be correlations between different pipe soil interaction characteristics (e.g. lateral breakout
resistance and lateral residual resistance). In some cases, the uncertainties can be reduced by carrying out
specific studies, see /32/.
It should be noted that it is not possible to identify all uncertainties and corresponding probability
distributions related to parameters affecting pipe-soil behaviour. Using a full statistical treatment to obtain
ranges to be used in design is not considered applicable as engineering judgement will still be required
to assign reasonable probability distribution functions for all relevant input parameters including model
uncertainty. It may however still be used by as a tool in PSI assessments by qualified geotechnical engineers.
In any case, when evaluating range of resistances, it is important to use engineering judgement to
qualitatively consider possible correlations between parameters. Considering the complexity in pipe-soil
interaction, it is obvious that the ranges used in design should be larger than what is required for traditional
foundation design.
It is recommended to perform sensitivity studies of the pipeline performance at an early phase of the project
to identify the pipe-soil interaction parameters that are most important for the behaviour and integrity of
the pipeline, to understand which parameters require the most attention during the subsequent detail design
phase.
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SECTION 7 SPECIAL CONSIDERATIONS
7.1 On-bottom stability
7.1.1 Comment to the generalized approach in DNVGL-RP-F109
On-bottom design according to DNVGL-RP-F109 are based on empirical soil models. When using empirical
models, it is in general very important to understand their background and the limitations with respect to
validity ranges of the models. Tables presented in DNVGL-RP-F109 for the generalized method are based
on dynamic analyses following strictly the methodology in /33/ for clay and /22/ for sand, allowing for an
increased penetration for small oscillations prior to breakout. The research which formed the basis for DNVRP-F109 was tailor-made for this purpose, and the models in /33/ and /22/ are best fits to large-scale
pipe test results. Consequently, the generalized approach in DNVGL-RP-F109 does not consider the model
uncertainty of the soil model. However, despite its empirical basis, it has been used since the early 1980s.
It should be noted that in the development of DNVGL-RP-F109 the empirical models in /33/ and /22/ were
slightly modified. Also, slightly different formulations exist in various published sources. To avoid confusion,
the full set of equations from /58/ is included in App.A
The generalized method in DNVGL-RP-F109 does not relate to absolute stability, but is a prescribed
equivalent stability approach allowing for large displacements which is calibrated against dynamic time
series analyses of several sea states where extended models of those presented in /33/ and /22/ were used.
Although these models not being state-of-the-art methods, experience gained shows that the overall design
approach using the design tables in DNVGL-RP-F109 are conservative for most cases. This may not only be
due to the geotechnical model but also due to conservatisms in the hydrodynamic model. The generalized
method in DNVGL-RP-F109 using the tabulated values may therefore still be used.
The soil model for clay is developed based on tests primarily in soft clay (su < 8 kPa), but also on some tests
on stiff clay (su = 70 kPa). The tests cover a large range, but only extremes within the expected range of soil
strengths. It should also be noted that there is limited information available about other soil characteristics of
the tested soils. The proposed friction factor is in /33/ stated to be 0.2, which indicates that pre-consolidation
effects are not included. Consolidation due to the pipe weight and pre-consolidation due to an increased
weight in the water-filled period during the hydrotest would give a more beneficial initial condition for the
pipeline. Note however, that for assessment of stability in the short period after laying or flooding, a friction
factor less than 0.2 is possible due to the excess pore pressures remaining in the soil.
The soil model for sand in /22/ is based on tests on sand with a large range in initial relative density varying
from 5% to 100%. It is important to note that the model for estimating breakout resistance have been based
on tests where soil berms are built up by oscillating the pipe prior to breakout. As shown in [4.4.2.3], this
has a large impact and the model should not be used to assess absolute stability, i.e. when assuming no
pipe displacements prior to breakout. The way the tests were performed probably also changed the relative
density from its initial value.
For on-bottom stability assessments, the soil models given in [4.4] may be used to document absolute
stability.
On-bottom stability design according to DNVGL-RP-F109 assumes an unstable pipe on a stable seabed.
For sandy soils susceptible to scour and sediment transport, this is not the case as the seabed particles
are more unstable than the pipe. The recent research activities within on-bottom stability of pipelines have
focused on fluid-structure-seabed interaction. In sandy soils, the development of scour underneath the pipe
may positively lead to self-burial of the pipe, protecting it from hydrodynamic loads and also increase the
soil resistance. Scour on either side of the pipe may negatively reduce the effective embedment and soil
resistance. This is a complex process, and design methodologies taking these effects into account have
not yet been developed. For more information, see /34/. The physical phenomenon would be that scour
underneath parts of the pipeline would result in free spans. As the scour develops further it will create a
longer free span, and eventually the pipe will sag into the free span. When a free span develops, the pipesoil contact forces at the shoulders will increase as it also needs to carry the weight of the free-spanning pipe
section. As sand is a frictional material, it will increase the soil resistance at the shoulders.
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7.1.2 Soil liquefaction
Soil liquefaction denotes the phenomenon that the soil loses a significant part of or all its shear strength.
Residual liquefaction may occur due to cyclic shear stresses, imposed by waves or earthquakes, that
generate excessive pore pressure until the soil loses a significant part of its shear strength. Instantaneous
liquefaction may occur if a steep wave travels over a loose soil inducing an upward-directed pressure gradient
under the wave trough /55/.
Obviously, soil liquefaction will affect both vertical stability, i.e. sinking and floatation, and lateral stability.
Depending on the specific gravity of the pipe, soil liquefaction may make a heavy pipe, laid on the seabed,
sink into the soil and bury itself, or make a light, buried pipe float up through the soil.
Models do exist to predict possible liquefaction. However, some rather sophisticated in-situ and laboratory soil
tests are required to quantify input parameters to these models, see /35/ and /36/.
7.2 Free spanning pipelines
7.2.1 General
Pipe-soil interaction is important when evaluating the static equilibrium configuration and dynamic response
of a free spanning pipeline. The following functional requirements apply for modelling of soil resistance:
— The seabed topography shall be well represented by a vertical profile along the pipeline route. The spacing
of data points characterising the profile should be sufficiently small to describe the actual roughness of the
seabed.
— The modelling of soil resistance shall account for non-linear contact forces vertical to the pipeline, e.g. lift
off.
— The modelling of soil resistance shall account for sliding in the axial direction. For force models this also
applies in the lateral direction.
— Appropriate short- and long-term characteristics for stiffness and damping shall be applied, i.e. for static
and dynamic conditions.
7.2.2 Simplified soil damping
If no detailed assessment according to [4.6] is carried out, the soil contribution to modal damping,
be taken from Table 7-1 or Table 7-2. Interpolation is allowed.
ζsoil, may
Note that the simplified soil damping given in Table 7-1 and Table 7-2 are valid only for single spans and to
the first mode of oscillation. Whereas they are considered to be conservative for a traditional single span
situation, the tabulated values could be non-conservative to use for a multi-span situation or for investigating
higher modes of a long single span.
Table 7-1 Soil contribution to modal damping [%] for sand
φ
Horizontal (in-line) direction
L/D
[°]
<40
100
>160
<40
100
>160
Loose
28-30
3.0
2.0
1.0
2.0
1.4
0.8
Medium
30-36
1.5
1.5
1.5
1.2
1.0
0.8
Dense
36-41
1.5
1.5
1.5
1.2
1.0
0.8
Friction angle,
Sand type
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Vertical (cross-flow) direction
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Table 7-2 Soil contribution to modal damping [%] for clay
Undrained
shear strength,
su
Clay type
Firm-stiff
Very stiff – hard
Vertical (cross-flow) direction
L/D
<40
100
>160
<40
100
>160
<25
4.0
2.0
1.0
3.0
2.0
1.0
25-100
2.0
1.4
0.8
1.2
1.0
0.8
>100
1.4
1.0
0.6
0.7
0.6
0.5
2
[kN/m ]
Very soft – soft
Horizontal (in-line) direction
L/D
For pipes supported by rock, values for the soil contribution to modal damping may be taken as for dense
sand.
The damping recommendations are based on work described in /37/. Two types of tests were performed:
— A string-model test where instrumented pipes with a cantilevered free end into a pre-made ditch were
excited both vertically and laterally to establish the global damping response of the pipe.
— A sectional test where a pipe was penetrated vertically and loaded cyclically in vertical and lateral
direction while measuring the force displacement relationship. Here the sectional response (per unit
length) of the pipeline was established for different levels of cyclic force or displacement. The response
measured was the integrated response of the surrounding soil.
Tests were performed on two types of soil: a soft clay and a medium-to-dense sand. The resulting global
behaviour of the string model was checked against the sectional model using a finite element beam model
with discrete modelling of soil stiffness and damping.
The results related to soil stiffness and damping from the tests were then used to analyse real free span
conditions of single spans with varying length-to-diameter ratios in a finite element beam model with discrete
modelling of soil stiffness and damping. In this way, modal damping ratios were calculated for soil conditions
similar to those tested. Assumptions were made to cover other soil conditions.
7.2.3 Soil stiffness
When evaluating Gmax from Equation (4.35) the mean effective stress, σs, in the soil at the span support may
be calculated from the stress conditions at a representative depth below the pipe. The representative depth
may be assumed equal to the contact width B in Equation (4.3). The following formula may then be applied:
(7.1)
where:
K0
γ’
V
B
L
Lsh
is the coefficient of earth pressure at rest
is the submerged unit weight of soil
is the submerged weight of pipe per unit length
is the pipe-soil contact width
is the span length
is the span support length on one shoulder (for transfer of one-half the weight of the free span).
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Note that for pipes on clay, when estimating Gmax from Equation (4.36), the clay might not be consolidated
for the weight of the pipe in the temporary phase immediately after pipe laying.
The span support length, Lsh, which is the contact length between pipe and soil on one shoulder, depends on
span length, soil stiffness of the shoulders, soil type, shoulder geometry, and submerged weight and stiffness
of the pipe.
7.2.4 Simplified soil stiffness
The vertical and lateral dynamic stiffness, KV,d and KL,d, may be calculated in a simplified manner as follows
when the topographical conditions are not complex, when the soils are non-stratified and homogenous, and
when no detailed analysis is carried out for determination of the soil stiffness according to [4.5.3]:
(7.2)
(7.3)
in which D, is the outer pipe diameter including coating [m] and the coefficients CV and CL are taken
according to Table 7-3 and Table 7-4. ρs/ρ is the specific mass ratio between the pipe mass (not including
added mass) and the displaced water. It shall be assessed whether the soil behaviour is drained or
undrained. Typical values of Poisson’s ratio are given in [4.5]. The expressions are valid for 1.2 < ρs/ρ ≤
2.0. The tabulated values in Table 7-3 and Table 7-4 are based on the formulae in [4.5.3], using typical
parameters for each soil type under the assumption of homogenous conditions.
Table 7-3 Simplified dynamic stiffness factor and static stiffness for pipe-soil interaction in sand
Sand type
Friction angle,
φ
Cv
5/2
[kN/m
[°]
CL
]
KV,s
5/2
[kN/m
]
[kN/m/m]
Loose
28-30
10500
9000
250
Medium
30-36
14500
12500
530
Dense
36-41
21000
18000
1350
Table 7-4 Simplified dynamic stiffness factor and static stiffness for pipe-soil interaction in clay
with OCR=1
Clay type
Undrained shear strength, su
Cv
KV,s
[kN/m
<12.5
600
500
50-100
Soft
12.5-25
1400
1200
160-260
Firm
25-50
3000
2600
500-800
Stiff
50-100
4500
3900
1000-1600
[kN/m ]
Very soft
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5/2
CL
2
]
5/2
[kN/m
]
[kN/m/m]
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Clay type
Very stiff
Undrained shear strength, su
2
[kN/m ]
Cv
5/2
[kN/m
CL
]
KV,s
5/2
[kN/m
]
[kN/m/m]
100-200
11000
9500
2000-3000
>200
12000
10500
2600-4200
Hard
For free spans supported by sand, the lateral dynamic stiffness, KL,d, should be calculated under the
assumption of loose sand properties to properly account for effects of complex soil mobility, including erosion
and self-burial.
The axial dynamic stiffness is usually not important. However, when long free spans are considered, it is
important to include an axial soil-support model with friction and stiffness. If no information is available
about the axial dynamic soil stiffness, it may be taken as equal to the lateral dynamic soil stiffness, KL,d.
7.2.5 Axial and lateral soil resistance
The available capacity for transfer of axial loads between pipe and soil should be considered. Axial frictional
stresses may exist between pipe and soil, e.g. formed by residual stresses from the pipelay. When such axial
stresses are present, they may prevail over long distances far back from the point of separation between pipe
and soil on the shoulder of the free span.
The axial friction will be limited by the interface shear resistance, see [4.3]. In areas with lateral movements
of the pipe, typically on the shoulders near the free span, the available surface friction needs to be shared
between axial and lateral stresses. This interaction will probably reduce both the axial and lateral resistance
at the shoulders and should be considered.
7.3 Design philosophy of support fills
7.3.1 General
The purpose of this section is to discuss the design of and the safety aspects related to stability of rock fills in
order to assure a safe situation for the pipeline while avoiding overly conservative and cost-driving design of
gravel fills with regard to stability. In such an evaluation it is necessary to consider the consequence of a rock
fill failure as well as the probability of failure. The consequences and probabilities of failure may vary in time
and a risk assessment may offer possibilities to optimize the design.
In general, the requirements to a low probability of failure should be stricter the higher the consequences.
Therefore, the purpose of the fills related to the pipeline design is essential. A pipe support that is required
to fulfil the ultimate limit states (ULS) operational safety requirements of the pipeline should have a low
probability of failure, whereas e.g. supports required in order to increase fatigue life, could be allowed to
have a higher probability of failure. This would especially be so for those fills where the pipe has relatively
high fatigue life of say more than one or two years in unsupported condition, considering the possibility for
inspection and re-establishing any failed supports, and the fact that for a fill where all static loads have been
applied, the major cause for failure in operational condition would be earthquake loading.
Carrying out a risk assessment study in the early phases of the project taking the factors mentioned in this
section into account could lead to significant cost savings in areas where rock fills are required and there is
possibility of foundation failure for such fills. An optimization of safety and cost should involve:
— Identifying fills that are essential with respect to pipeline safety, for which one should assure that
foundation failure does not occur. Detailed stability analyses are required based on reliable soil data.
— Evaluating costs related to detailed soil investigations and stability analyses against possible future costs
related to rectification for foundation failures taking place either during construction of the fills or later,
e.g. due to earthquake.
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7.3.2 Design calculations
The design calculations should be carried out with appropriate methods capturing the critical failure
mechanisms in the rock berms and underlying soil. Recognized limit-equilibrium analyses may be used for
this purpose. When soil layering affects the geometry of the failure mode, it is recommended to use finite
element analyses.
Different parts of the gravel fill will either be driving or resisting failure, in particular when counter fills
are required to achieve stability. However, all parts of the fill are recommended to be modelled by their
characteristic load, i.e. with a load factor γf = 1.0. For any loads on top of the fill a load factor larger than
γf = 1.3 is recommended for foundation design in DNVGL-OS-C101. A material
γm should be applied to the shear strength of the supporting soil and the rock fill. For rock fills with
high consequence of failure a material factor not lower than γm = 1.4 is recommended, but this needs to be
1.0 should be applied, and
factor
evaluated case by case, and the choice of material factor is particularly dependent on how the characteristic
shear strength is selected in relation to the available soil data, see [7.3.3].
In shallow waters, hydrodynamic forces from wave and current may give a net driving force on the rock fill
that should be taken into account. The design wave should be taken in accordance with DNVGL-ST-F101.
When applying hydrodynamic pressure, this should be applied to the entire fill and to the seabed outside the
fill, in order also to account for the stabilising part of the loading. The most critical phase of the wave should
be sought.
Any rock fill for which failure during operation has a large consequence should be designed to resist
earthquake. When failure of the rock fill has minor consequences to the integrity of the pipeline and when the
support can be re-established as required after the earthquake, design for earthquake may be omitted. The
response of the rock fill may be limited displacements or complete failure as a result of degradation of the
soil strength. Failure may occur during the earthquake or shortly after. The latter may for example occur in
clay where the post-earthquake static strength will reduce with time shortly after the earthquake because of
creep effects.
For verification of stability against earthquake quasi-static analyses applying accelerations from a free-field
response spectrum may be performed. This may be a solution where the rock fill has a high safety against
failure for static loading. Alternatively, a more explicit earthquake analysis should be performed as a 2D or
3D site response analysis modelling the geometry of the fill and where the earthquake motions are applied
as input motions at the bedrock or at a sufficiently deep stiff soil reference layer. Cyclic degradation of the
strength may be difficult to account for explicitly in the analyses, but could be accounted for iteratively based
on the cyclic and static strains in the soil obtained from the analyses.
7.3.3 Uncertainties in soil conditions
The soil conditions of the upper soil layers vary along the pipeline route. Traditional soil investigations
performed for pipeline design do not allow for an optimal design of each individual support because soil
data is not site specific for every respective fill. Therefore, the general approach is to check the foundation
stability of the rock fills by using the established lower bound shear strength along parts of the route having
similar soil deposits. By doing this, one will most often use a conservative shear strength for all of the rock
fills, and for the majority of fills, this will actually be significantly conservative.
For an optimal design of a specific rock fill where uncertainties with respect to soil shear strength is reduced
to a minimum, several soil borings and/or in-situ tests, e.g. CPT, have to be performed within the area of a
potential soil failure surface. This should be considered for rock fills where the consequence of failure is large,
either by imposing a pipeline ULS failure, or where there are large economic impacts of having to rectify a
stability failure.
7.3.4 Likelihood of soil failure versus time
Soil failure may take place during, and most likely towards the end of the rock-dumping process. Since the
largest uncertainty relates to the soil conditions, this is also the most likely time for any failure to occur. If
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the soil is almost at failure at the immediate time of completing the dumping, undrained creep of the soil
(clay) may cause failure within the first few days after dumping.
After the first week, consolidation of the soil for the new stress situation will tend to increase the shear
strength. For failure to occur in this situation, an additional load will have to be applied.
The only additional static loads applied are loads from the pipe and from the special structures (PLETs,
PLEMs, etc.) placed on top of fills. The rock fills will normally be constructed a certain time period before
pipelines or structures are installed and additional loading are applied to the rock fills. The effect of
consolidation during this time period will vary, depending on the magnitude of effective stresses as compared
to pre-consolidation stresses, but will in general improve the strength of the soil and thereby also make the
fills more stable with time. When placing structures with significant weight onto a fill this would still be the
critical condition for which the fill should be designed.
Early failures will be detected during dumping or by the post-dump survey, and the consequences would
be to re-establish the failed fills with more gentle slopes or using additional counter fills. If the detailed
geometry survey is performed too shortly after the dumping in relation to possible late undrained creep
failure, a rough survey to detect such failures is suggested a week or more after dumping.
7.4 Penetration of falling objects
Rock or gravel cover is the most common protection method for pipelines. Based on full-scale tests the
energy absorbed in rock material, Ep, when a falling pipe penetrates, can be described as:
(7.4)
where:
γ’
D
Apipe
z
Ny, Nq
is the effective unit weight of the soil material
is the outer diameter including coating of the faling pipe
is the plugged area of the falling pipe
is the penetration depth
are bearing capacity factors.
The bearing capacity coefficients can be chosen as Nq = 99 and Nγ = 137 for falling objects onto rock fills.
Τhe effective unit weight for a typical rock fill may be taken as 10 kN/m3.
Guidance note:
The use of the plugged area of the pipe in Equation (7.4) should be seen in context with the size of the stones in the rock fill
material. If the pipe diameter is small compared to the diameter of the stones, the full cross section of the pipe can be used. If the
pipe diameter is large, however, an equivalent area where the pipe circumference is multiplied by the stone diameter may be used.
---e-n-d---o-f---g-u-i-d-a-n-c-e---n-o-t-e---
For non-tubular objects, like containers, the energy absorption in rock can become higher. The following two
equations are proposed for penetration of side edges and corners of containers:
for side edges of containers
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for corners of containers.
(7.6)
Where sγ is a shape factor equal to 0.6, and L is the length of the impacting side.
The energy absorption of different objects is given in Figure 7-1.
Figure 7-1 Absorbed energy in gravel
Energy absorption in natural back-filled sand is considerably lower than for rock. Natural back-filled sand is
very loose, and pipes may not plug in sand. Bearing capacity factors for sand can be taken from Figure 4-6.
Effective protection against dragged commercial ship anchors can be obtained by burying the pipeline. The
required depth will depend on the size of the anchors of the passing ships and the local soil conditions, i.e.
how deep the anchors will penetrate.
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7.5 Carbonate soils
7.5.1 Definition
A soil is classified as carbonate soil if it contains more than 50% of calcium carbonate /38/. Most carbonate
soils are composed of large accumulations of the skeletal remains of marine organisms, such as coralline
algae, coccoliths, foraminifera and echinoderms, although they also exist as non-skeletal material in the form
of oolites, pellets and grape-stone. Carbonate deposits are abundant in the warm, shallow tropical waters,
such as offshore Africa, Australia, Brazil and the Middle East.
7.5.2 Characteristic features
Carbonate soils have important characteristic features that distinguish them from terrigenous silica seabed
materials. Particles of skeletal carbonate material can be highly angular with rough surfaces and intra-particle
voids, leading to a soil fabric that is very open and compressible. Non-skeletal carbonate particles may be
rounded and solid, but are still susceptible to crushing or fracturing due to the low hardness that calcium
carbonate has as compared to quartz.
A typical granular carbonate soil will be characterized by a high void ratio (and low density), high
compressibility due to the high initial void ratio and the crushable nature of the individual particles, and a
high friction angle due to angularity, roughness and interlocking of the particles.
Carbonate sediments are susceptible to transformation by biological and chemical processes over time. Older
sedimentary deposits that may have been sub-aerially exposed at some stage in geological life are prone to
cementation, which completely alters the mechanical properties of the sediment. The cementation may occur
over a flat horizon, forming a caprock layer near the seabed, or may occur in irregular discontinuous lenses.
The engineering properties of uncemented carbonate soils differ from silica rich soils in the following ways:
— Calcium carbonate has a low hardness value compared to quartz. This leads to relatively high crushability
of carbonate soils at relatively low stress levels.
— Carbonate soils often have large porosity, resulting in high void ratio and low density, and are therefore
more compressible than soils from silica deposits. However, they generally exhibit higher friction angles
than equivalent silica soils.
— The undrained cyclic strength of carbonate soils is generally lower than for most silica soils and
permeability also tend to be lower. Consequently, carbonate soils are generally more susceptible to
liquefaction under the action of cyclic loading.
— More detailed description of these features and experimental data can be found in /38/ to /44/.
7.5.3 Pipeline response on carbonate soils
A pipeline may be viewed as a special type of strip footing that has a circular cross-section. Therefore, its
behaviour in carbonate sediments is in many ways similar to that of shallow foundations in the same soil.
Experimental evidence indicates that pipeline response in carbonate soils is characterized by the following
features, see /45/ to /48/:
— Approximately linear vertical load-displacement response.
— Relatively large lateral displacement to achieve the ultimate resistance. A pipeline on low density
calcareous sand may typically move laterally two or more diameters before developing ultimate soil
resistance, compared with one half to one diameter for a pipeline on silica sands.
— Cyclic loading induces larger embedment than in silica sand.
— Load-displacement response typically exhibits a ductile strain hardening response (except where the pipe
is embedded below its equilibrium depth), unlike in silica sands that generally exhibit post-peak strain
softening behaviour.
— Significant excess pore pressure may accumulate under cyclic environmental loads acting on the pipe, and
wave pressure loading imposed to the seabed, compared to silica soils. This is because such sediments
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are more prone to degradation and compaction under cyclic loading and tend to have larger coefficients
of consolidation than typical silica sands, and consequently, carbonate soils have a higher propensity for
liquefaction than silica soils.
7.5.4 Pipe-soil interaction model
Theoretical models assuming drained soil behaviour, have been developed for pipe-soil interaction in
carbonate sands within the framework of soil plasticity. These models link the displacement increments of a
pipeline with the load increments through a non-linear stiffness or flexibility matrix. This approach enables
straightforward incorporation of the pipe-soil interaction model into the full structural analysis of a pipeline.
Simplified expressions for calculating the ultimate lateral soil resistance of pipelines have also been derived
for basic pipeline stability assessments. Detailed descriptions of these models are given in /47/ and in /49/
to /51/.
In most real situations, pipeline-seabed interaction in carbonate soils is much more complex than assumed
in a model assuming drained conditions. Local pipe-soil interaction and large scale wave-soil interaction shall
both be considered, and this is further complicated by partial undrained behaviour of the seabed soil in many
cases. Therefore, a comprehensive model is required to describe pipe-soil interaction for submarine pipelines
under realistic wave and soil conditions. An appropriate model should be able to predict:
— Accumulation of excess pore pressure in the seabed soil due to wave actions acting directly on the seabed
and due to cyclic loading on the pipeline.
— Resulting strength/stiffness degradation of the seabed soil in the vicinity of the pipeline.
— Pipeline embedment, i.e. self-burial of the pipeline.
— Increase in lateral resistance due to self-burial.
Such a model has not been presented within the public domain. As a general guide, local properties of the
carbonate soils should be determined, which allows development of a model for local use.
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SECTION 8 BIBLIOGRAPHY
8.1 Bibliography
/1/
Najjar SS, Gilbert RB, Liedtke E, McCarron B, Young AG. Residual shear strength for interface
between pipeline and clays at low effective normal stresses. ASCE Journal of Geotechnical and
Geoenvironmental Engineering 133, June 2007, p. 695-706.
/2/
Ganesan S, Kuo M, Bolton M. Influences on Pipeline Interface Friction Measured in Direct Shear
Tests. ASTM Geotechnical Testing Journal Volume 37, Issue 1, January 2014, p. 1-13.
/3/
White DJ, Campbell ME, Boylan NP, Bransby MF. A new framework for axial pipe-soil interaction
illustrated by shear box tests on carbonate soils. Proc. Int. Conf. on Offshore Site Investigation
and Geotechnics, SUT, London, 2012, p. 379-387.
/4/
White DJ, Gaudin C. Simulation of seabed pipe-soil interaction using geotechnical centrifuge
modelling. Proc. 1st Asia-Pacific Deep Offshore Technology Conference, Perth, 2008, session 7, p.
1-26.
/5/
Langford, TE, Dyvik R, Cleave R. Offshore pipeline and riser geotechnical model testing: practice
and interpretation. Proc. of the Conference on Offshore, Marine and Arctic Engineering, OMAE,
2007, San Diego, p. 443-452.
/6/
Ballard JC, De Brier C, Stassen K, Jewell RA. Observations of pipe-soil response from in-situ
measurements. Proc. Offshore Technology Conference, OTC, Houston, 2013, Paper no. 24154.
/7/
Gao F, Wang N, Zhao B. Ultimate bearing capacity of a pipeline on clayey soils: Slip-line field
solution and FEM simulation. Ocean Engineering, Volume 73, 2013, p. 159-167.
/8/
Davis EH, Booker JR. The Effect of Increasing Strength with Depth on the Bearing Capacity of
Clays. Géotechnique, Volume 23, Issue 4, 1973, p. 551-563.
/9/
Randolph MF, White DJ. Pipeline Embedment in Deep Water: Process and Quantitative
Assessment. Proc. Offshore Technology Conference, OTC, Houston, 2008, Paper no. 19128.
/10/
Merifield RS, White DJ, Randolph MF. The effect of pipe-soil interface conditions on undrained
breakout resistance of partially-embedded pipelines. Int. Conf. on Advances in Computer Meth. &
Analysis in Geomechanics, Goa, 2008, p. 4249-4256.
/11/
Chatterjee S, Randolph MF, White DJ. The effects of penetration rate and strain softening on
the vertical penetration resistance of seabed pipe, Géotechnique, Volume 62, Issue 7, 2012, p.
573-582.
/12/
Brinch Hansen, J. A Revised and Extended Formula for Bearing Capacity. Geoteknisk Institut,
Bulletin No. 28, Copenhagen, 1970, p. 5-11.
/13/
Caquot A, Kerisel J. Sur la Terme de Surface dans le Calcul des Fondations en Milieu Pulvérult,
rd
Proc. Of the 3 International Conference on Soil Mechanics and Foundation Engineering, Vol. 1,
Zürich, 1953, p. 336-337.
/14/
Westgate ZJ, White DJ, Randolph MF, Brunning P. Pipeline Laying and Embedment in Soft
Fine-grained Soils: Field Observations and Numerical Simulations. Proc. Offshore Technology
Conference, OTC, Houston, 2010, Paper no. 20407.
/15/
White DJ, Cathie DN. Geotechnics for subsea pipelines. Proc. 2nd Int. Symp. on Frontiers in
Offshore Geotechnics, ISFOG, Perth, 2011, p. 87-123.
/16/
White DJ, Westgate ZJ, Ballard J-C, De Brier C, Bransby MF. Best Practice Geotechnical
Characterisation and Pipe-soil Interaction Analysis for HPHT Pipeline Design. Proc. Offshore
Technology Conference, OTC, Houston, 2015, Paper no. 26026.
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/17/
Hill A, White DJ, Bruton DAS, Langford T., Meyer V., Jewell R. & Ballard J-C. A new framework
for axial pipe-soil interaction illustrated by a range of marine clay datasets. Proc. Int. Conf. on
Offshore Site Investigation and Geotechnics, SUT, London, 2012, p. 367-377.
/18/
Ladd CC, Foott R. New Design Procedure for Stability of Soft Clays. Journal of the Geotechnical
Engineering Division, ASCE, Volume 100, No. GT7, 1974, p. 763-786.
/19/
Andersen KH. Cyclic soil parameters for offshore foundation design. The 3rd McClelland Lecture.
Proceedings of the Frontiers in Offshore Geotechnics III, ISFOG, 2015, p. 5-82.
/20/
Oliphant J, Maconochie A. The axial resistance of buried and unburied pipelines. Proc. Int. Conf.
on Offshore Site Investigation and Geotechnics, SUT, London, 2007, p. 125-132.
/21/
Janbu N, Bjerrum L, Kjærnsli B. Veiledning ved løsning av fundamenteringsoppgaver, Soil
Mechnics Applied to some Engineering Problems. NGI Publication No. 16, Norwegian Geotechnical
Institute, 1956.
/22/
Verley RLP, Sotberg T. A Soil Resistance Model for Pipelines Placed on Sandy Soils. Proc. of the
Conference on Offshore, Marine and Arctic Engineering, OMAE, Volume 5, 1992, Copenhagen, p.
123-131.
/23/
Chatterjee S, White DJ, Randolph MF. Numerical simulation of pipe-soil interaction during large
lateral movements on clay. Géotechnique, Volume 62, Issue 8, 2012, p. 693-705.
/24/
Hardin BO. The nature of stress-strain behavior for soils. Proceedings ASCE Geotech. Engrg. Div.
Speciality Conf. on Earthquake Engineering and Soil Dynamics, Volume 1, 1978, p. 3-90.
/25/
Vucetic M, Dobry R. Effect of Soil Plasticity on Cyclic Response. Journal of Geotechnical
Engineering, Volume 117, No.1, 1991, p. 89-107.
/26/
Guha I, Randolph MF, White DJ. Evaluation of Elastic Stiffness Parameters for Pipeline-Soil
Interaction. ASCE Journal of Geotechnical and Geoenvironmental Engineering, Volume 142, Issue
6, 2016.
/27/
Eiksund G, Langø H, Hove F. Full-Scale Tests of Axial Friction on Pipelines in Rock Berms.
International Journal of Offshore and Polar Engineering. Volume 23, No.1, 2013, p. 63-68.
/28/
Trautman CH, O’Rourke TD, Kulhawy FH. Uplift Force-Displacement Response of Buried Pipe.
Journal of Geotechnical Engineering, Volume 111, No. 9, 1985, p. 1061-1075.
/29/
White DJ, Barefoot AJ, Bolton MD. Centrifuge modelling of upheaval buckling in sand.
International Journal of Physical Modelling in Geotechnics, Volume 2, 2001, p. 19-28.
/30/
Randolph MF, Houlsby GT. The limiting pressure on a circular pile loaded laterally in cohesive soil.
Géotechnique, Volume 34, Issue 4, 1984, p. 613-623.
/31/
Martin CM, White DJ. Limit analysis of the undrained bearing capacity of offshore pipelines.
Géotechnique, Volume 62, Issue 9, 2012, p. 847-863.
/32/
White DJ, Westgate ZJ, Tian Y. Pipeline Lateral Buckling: Realistic Modelling of Geotechnical
Variability and Uncertainty. Proc. Offshore Technology Conference, OTC, Houston, 2014, Paper no.
25286.
/33/
Verley RLP, Lund, KM. A Soil Resistance Model for Pipelines Placed on Clay Soils. Proc. of the
Conference on Offshore, Marine and Arctic Engineering, OMAE, Volume 5, 1992, Copenhagen, p.
225-232.
/34/
Luo C. On-bottom stability of submarine pipeline on mobile seabed. PhD thesis, University of
Western Australia, 2013.
/35/
ASCE Journal of Waterway, Port, Coastal and Ocean Engineering. Special Issue: Liquefaction
Around Marine Structures. Processes and Benchmark Cases. Editor: D. M. Sumer, Volume 132,
Issue 4, 2006, p. 225-335.
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/36/
Damsgaard JS, Palmer AC. Pipeline Stability on a Mobile and Liquefied Seabed: A Discussion of
Magnitudes and Engineering Implications. Proc. of the Conference on Offshore, Marine and Arctic
Engineering, OMAE, 2001, Rio de Janeiro, p. 195-204.
/37/
Tura F, Dumitrescu A, Bryndum MB, Smeed PF. Guidelines for Free Spanning Pipelines: The
GUDESP Project. Proc. of the Conference on Offshore, Marine and Arctic Engineering, OMAE,
Volume 5, 1994, Houston, p. 247-256.
/38/
Clark AR, Walker BF. A Proposed Scheme for the Classification and Nomenclature for Use in the
Engineering Description of Middle Eastern Sedimentary Rocks. Géotechnique, Volume 27, Issue 1,
1977, p. 93-99.
/39/
Jewell RJ, Andrews DC, Khorshid MS. Engineering for Calcareous Sediments. 1st International
Conference on Engineering for Calcareous Sediments, Perth, Vols. 1 and 2, 1988, Balkema,
Rotterdam.
/40/
Datta M, Gulhati SK, Rao GV. Crushing of Calcareous Sand During Shear. Proc. Offshore
Technology Conference, OTC, Houston, 1979, Paper no. 3525.
/41/
Finnie IMS, Randolph MF. Bearing Response of Shallow Foundations in Uncemented Calcareous
Soil. Proc. International Conference Centrifuge ’94, Volume 1, Singapore, 1994, p. 535-540.
/42/
Poulos HG. The Mechanics of Calcareous Sediments. Australian Geomechanics, 1988, p. 8-41.
/43/
Joer HA, Bolton MD, Randolph MF. Compression and Crushing Behaviour of Calcareous Soils.
International Workshop on Soil Crushability, IWSC’99, Ube, 2000, p. 10-1 – 10-16.
/44/
International Conference on Engineering for Calcareous Sediments. Engineering for calcareous
sediments: Proceedings of the Second International Conference on Engineering for Calcareous
Sediments, Editor: Al-Shafei KA, Bahrain, 21-24 February 1999.
/45/
Wallace LTI. Pipeline Performance in Calcareous Soil. Honours Thesis, the University of Western
Australia, 1995.
/46/
Browne-Cooper E. The Vertical and Horizontal Stability of a Pipeline in Calcareous Sand. Honours
Thesis, the University of Western Australia, 1997.
/47/
Zhang, J. Geotechnical Stability of Offshore Pipelines in Calcareous Sand. PhD Thesis, the
University of Western Australia, 2001.
/48/
Zhang J, Stewart DP, Randolph MF. Centrifuge Modelling of Drained Behaviour for Offshore
Pipelines Shallowly Embedded in Calcareous Sand. International Journal of Physical Modelling in
Geotechnics, Volume 1 No. 1, 2001, p. 25-39.
/49/
Zhang J, Stewart DP, Randolph MF. Vertical Load-Displacement Response of Untrenched Offshore
Pipelines on Calcareous Sand. International Journal of Offshore and Polar Engineering, Vol. 12 No
1, 2002, p. 74-80.
/50/
Zhang J, Stewart DP, Randolph MF. Modelling of Shallowly Embedded Offshore Pipelines in
Calcareous Sand. ASCE Journal of Geotechnical and Geoenvironmental Engineering, Volume 128,
No. 5, 2002, p. 363-371.
/51/
Zhang J, Stewart DP, Randolph MF. A Kinematic Hardening model for Pipeline-Soil Interaction
under Various Loading Conditions. International Journal of Geomechanics, Vol. 2, No. 4, 2002, p.
419-446.
/52/
Bruton D, White D, Cheuk C, Bolton M, Carr M. Pipe/Soil Interaction Behavior During Lateral
Buckling, Including Large-Amplitude Cyclic Displacement Tests by the Safebuck JIP. Proc. Offshore
Technology Conference, OTC, Houston, 2006, Paper no. 17944.
/53/
Cardoso CO, Silveira RMS. Pipe-Soil Interaction Behavior for Pipelines Under Large Displacements
on Clay Soils – A Model for Lateral Residual Friction Factor. Proc. Offshore Technology Conference,
OTC, Houston, 2010, Paper no. 20767.
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/54/
Cathie DN, Jaeck C, Ballard J-C, Wintgens J-F. Pipeline geotechnics – state-of-the-art. Proc.
Frontiers in Offshore Geotechnics, ISFOG, London, 2005, p. 95-114.
/55/
Verruijt A, de Bruijn VMPJ, Muralidhar B. Liquefaction Around Buried Pipe Lines. Behaviour of
Offshore Structures – BOSS ’97, Volume 1, p. 99-108.
/56/
Thusyanthan NI, Mesmar S, Wang J, Haigh SK. Uplift Resistance of Buried Pipelines and DNV-RPF110 Guidelines. Proc. Offshore Pipeline Tech. Conf., Amsterdam, 2010.
/57/
Stuyts B, Cathie B, Powell T. Model uncertainty in uplift resistance calculations for sandy backfills.
Canadian Geotechnical Journal 53(11), 2016, p. 1831-1840.
/58/
DNV, On-Bottom Stability Design of Submarine Pipelines, October 2010, DNV-RP-F109.
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APPENDIX A SOIL MODELS FROM DNV-RP-F109 (OCTOBER 2010)
A.1 General
The development of DNVGL-RP-F109 was based on the work presented in /22/ and /33/. However, the
equations used for passive resistance for both sand and clay were slightly modified compared to those
presented in /22/ and /33/.
As there exists some confusion with regards to the formulations of the empirical methods, the full set of
equations from /58/ are given below. The nomenclature in [A.4] is valid for Equation (A.1) to Equation (A.4).
A.2 Initial penetration
Sand:
(A.1)
Clay:
(A.2)
A.3 Passive soil resistance
Sand:
for
κs ≤ 26.7
(A.3)
for
κs > 26.7
Clay:
(A.4)
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A.4 Nomenclature
D
Fc
FR
Fz
pipe outer diameter (including coating)
= wS – FZ
passive soil resistance
vertical hydrodynamic (lift) load
GC
Su
zpi
zp
WS
γs
γ’s
undrained clay shear strength
initial pipe penetration
pipe penetration (general)
submerged pipe weight
dry unit soil weight
submerged unit soil weight
κC
κs
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Changes - historic
CHANGES - HISTORIC
There are currently no historical changes for this document.
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