GEOTECHNICAL PERFORMANCE OF SPUDCAN ......Guidance for site assessment of jackup operability is...

Guidance Notes on Geo techn ica l Per fo rmance o f Spudcan Foundat i ons

GUIDANCE NOTES ON

GEOTECHNICAL PERFORMANCE OF SPUDCAN FOUNDATIONS

JANUARY 2017 (Updated March 2018 – see next page)

American Bureau of Shipping Incorporated by Act of Legislature of the State of New York 1862

2016 American Bureau of Shipping. All rights reserved. ABS Plaza 16855 Northchase Drive Houston, TX 77060 USA

Updates

March 2018 consolidation includes: • January 2017 version plus Corrigenda/Editorials

F o r e w o r d

Foreword For the last decade, ABS has been involved in numerous joint industry projects on jackup spudcan foundations. ABS has outlined this knowledge herein to provide guidance covering the design and installation of jackup spudcan foundations addressing the following topics:

• Site assessment

• Spudcan penetration prediction in a single layer of soil, such as clay or sandy soil

• Punch-through prediction and mitigation

• Foundation stability assessment

• Foundation fixity

• Spudcan-jacket pile interaction

• Spudcan footprint interaction

These Guidance Notes become effective on the first day of the month of publication.

Users are advised to check periodically on the ABS website www.eagle.org to verify that this version of these Guidance Notes is the most current.

We welcome your feedback. Comments or suggestions can be sent electronically by email to [email protected].

Terms of Use

The information presented herein is intended solely to assist the reader in the methodologies and/or techniques discussed. These Guidance Notes do not and cannot replace the analysis and/or advice of a qualified professional. It is the responsibility of the reader to perform their own assessment and obtain professional advice. Information contained herein is considered to be pertinent at the time of publication, but may be invalidated as a result of subsequent legislations, regulations, standards, methods, and/or more updated information and the reader assumes full responsibility for compliance. This publication may not be copied or redistributed in part or in whole without prior written consent from ABS.

ABS GUIDANCE NOTES ON GEOTECHNICAL PERFORMANCE OF SPUDCAN FOUNDATIONS . 2017 iii

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T a b l e o f C o n t e n t s

GUIDANCE NOTES ON

GEOTECHNICAL PERFORMANCE OF SPUDCAN FOUNDATIONS

CONTENTS SECTION 1 Introduction ............................................................................................ 1

1 General Comments ............................................................................. 1 3 Scope and Application ........................................................................ 1 5 Terms and Definitions ......................................................................... 1 7 Symbols and Abbreviations ................................................................ 2

7.1 Symbols ........................................................................................... 2 7.3 Abbreviations ................................................................................... 5

SECTION 2 Site Assessment ..................................................................................... 6

1 General ............................................................................................... 6 3 Geotechnical Site Investigation .......................................................... 6 5 Sampling and Field Testing ................................................................ 7 7 Laboratory Testing .............................................................................. 8 9 Soil Engineering Parameter Selection ................................................ 9

9.1 Identification of Soil Type and Layers .............................................. 9 9.3 Clay ............................................................................................... 10 9.5 Sand .............................................................................................. 10 9.7 Derivation of Soil Strength Profile .................................................. 12

TABLE 1 Relative Reliability of Tests Measuring the Strength of Clay

Soils .......................................................................................... 8 TABLE 2 Index Properties for Jackup Foundation Site Specific

Assessment ............................................................................... 9 TABLE 3 Values of φcv and Qcrushing Derived from Triaxial Compression

Tests ....................................................................................... 12 SECTION 3 Spudcan Penetration in Single Layer Soil .......................................... 14

1 Background ....................................................................................... 14 3 Numerical Simulations ...................................................................... 14 5 Simplified Prediction Methods ........................................................... 15

5.1 General .......................................................................................... 15 5.3 Spudcan Penetration in Clay ......................................................... 15 5.5 Spudcan Penetration in Sand ........................................................ 18 5.7 Lattice Leg Effect on Cavity Depth and Spudcan Penetration ....... 20

iv ABS GUIDANCE NOTES ON GEOTECHNICAL PERFORMANCE OF SPUDCAN FOUNDATIONS . 2017

7 Alternative Approach: Direct Correlation with Penetrometer Tests ................................................................................................. 22

9 Consequences for Observation ........................................................ 23

TABLE 1 Bearing Capacity Factors Nc for Conical Shaped Footings .... 18 TABLE 2 Bearing Capacity factors Nγ .................................................... 20 FIGURE 1 Embedded Spudcan with Open Soil Cavity ............................ 16 FIGURE 2 Definition of Equivalent Cone .................................................. 17 FIGURE 3 Backflow in Sand..................................................................... 19 FIGURE 4 Lattice Leg View ...................................................................... 21 FIGURE 5 Top Mounted Skirt on Spudcan .............................................. 22

SECTION 4 Punch-through in Strong Soil Overlying Soft Soil ............................ 24

1 General ............................................................................................. 24 3 Prediction Methods ........................................................................... 24

3.1 Sand over Clay .............................................................................. 24 3.3 Strong Clay over Soft Clay ............................................................ 27

5 Mitigation Method – Perforation Drilling ............................................ 28 7 Remedial Action during Punch-through ............................................ 29 FIGURE 1 Nomenclature for Spudcan Penetration in Sand over Clay .... 24 FIGURE 2 Values for Spudcan Bearing Capacity Calculation when

h ≥ hlayer [Ref.18] ...................................................................... 26 FIGURE 3 Nomenclature of Spudcan Penetration in Sand over Clay

when h ≥ hlayer .......................................................................... 27 FIGURE 4 Nomenclature of Spudcan Penetration in Strong Clay over

Soft Clay .................................................................................. 28 SECTION 5 Foundation Stability Assessment ....................................................... 30

1 Approach ........................................................................................... 30 3 Acceptance Criteria ........................................................................... 30 FIGURE 1 Foundation Stability Checks ................................................... 31

SECTION 6 Foundation Fixity ................................................................................. 32

1 Introduction ....................................................................................... 32 3 Foundation Fixity ............................................................................... 32

3.1 Initial Stiffness ............................................................................... 32 3.3 Yield Surface under Combined Loadings ...................................... 35 3.5 Calculation Procedures ................................................................. 36

5 Consolidation Effect .......................................................................... 37 7 Lattice Leg Effect on Foundation Fixity ............................................. 37

ABS GUIDANCE NOTES ON GEOTECHNICAL PERFORMANCE OF SPUDCAN FOUNDATIONS . 2017 v

FIGURE 1 Spudcan Soil Rotational Stiffness ........................................... 32 FIGURE 2 Depth Factors kd1 for Vertical Spring Stiffness ........................ 34 FIGURE 3 Depth Factors kd2 for Horizontal Spring Stiffness .................... 34 FIGURE 4 Depth Factors kd1 for Rotational Spring Stiffness .................... 34 FIGURE 5 Yield Envelopes for Conical Footings ..................................... 36

SECTION 7 Spudcan-Pile Interaction ..................................................................... 38

1 Introduction ....................................................................................... 38 3 Soil Flow Mechanism for Spudcan in Soft Clay ................................ 40

3.1 Spudcan Penetration ..................................................................... 40 3.3 Spudcan Operation and Extraction ................................................ 40

5 Dimensionless Charts ....................................................................... 41

FIGURE 1 Potential Soil Loading Effects on Jacket Platform

(after [Ref.29]) ......................................................................... 39 FIGURE 2 Spudcan-Pile Interaction ......................................................... 39 FIGURE 3 Incremental Soil Movement Trajectories at Different

Distances from Spudcan ......................................................... 40 FIGURE 4 Nomenclature for a Socketed Pile .......................................... 42

SECTION 8 Spudcan-Footprint Interaction ........................................................... 43

1 Introduction ....................................................................................... 43 3 Footprint Characteristics and Its Influence on Spudcan-Footprint

Interaction ......................................................................................... 44 5 Mitigation Methods ............................................................................ 46 FIGURE 1 Bathymetry of an Established Site .......................................... 43 FIGURE 2 Generalized Soil Condition of a Footprint

(for τ1 < 0.002; τ2 < 0.2) ........................................................... 45 FIGURE 3 Simplified Diagram of Probable Soil Failure Mechanisms

During Penetration at 0.5D from Footprint Center .................. 46 FIGURE 4 Stomping Process ................................................................... 47

APPENDIX 1 References ............................................................................................ 48

vi ABS GUIDANCE NOTES ON GEOTECHNICAL PERFORMANCE OF SPUDCAN FOUNDATIONS . 2017

S e c t i o n 1 : I n t r o d u c t i o n

S E C T I O N 1 Introduction

1 General Comments For the last decade, ABS has been involved in numerous joint industry and government sponsored projects on jackup spudcan foundations. The current state of the art in spudcan design and installation represents worldwide contributions by a large number of investigators from government agencies, marine warranty survey groups, classification societies, contractors, shipyards, universities and petroleum companies.

ABS has gathered this body of knowledge to provide a guideline on geotechnical performance of spudcan foundations, including site assessment, spudcan penetration prediction, punch-through problem, foundation stability assessment, foundation fixity, spudcan-jacket pile interaction and spudcan footprint interaction.

3 Scope and Application The Guidance Notes apply to the geotechnical design, installation and operation of jack up spudcan foundations. These recommendations are to be used in conjunction with Part 3 of the ABS Rules for Building and Classing Mobile Offshore Drilling Units (MODU Rules). As MODU Classification is not site-specific in nature, the Owner specifies conditions for which the unit is to be installed.

Guidance for site assessment of jackup operability is given in Section 2, followed by spudcan penetration prediction in single layer soil, such as clay and sand in Section 3. In strong over soft soil where punch-through failure is a major concern on design of jackup spudcan foundation, the recommended prediction methods to calculate spudcan bearing capacity in layered soils are introduced in Section 4, as well as the mitigation method for reducing punch-through failure. Three steps are suggested in the order of increasing complexity and reducing conservatism for the overall foundation stability assessment in Section 5. The procedure for calculating foundation fixity is presented in Section 6. If the proximity of the spudcan to the pile is less than one spudcan diameter, detailed analysis is recommended to estimate the severity of spudcan-pile interaction problems as given in Section 7. The footprint characteristics and its influence on spudcan-footprint interaction are presented in Section 8 as well as possible effective mitigation methods in various field conditions.

These Guidance Notes are applicable to self-elevating units which undertake drilling, construction, support, wind turbine installation or other offshore activities.

5 Terms and Definitions Backfill: Soil material used to refill the cavity above the spudcan due to spudcan penetration.

Backflow: Soil flows from bottom of spudcan back into the cavity on the exposed top of the spudcan.

Breakout Force: The maximum uplift force during spudcan extraction.

Cavity Depth: The depth of an open cavity forms above the spudcan after penetration.

Consolidation: The process in which soil reduction in volume takes place by expulsion of water under long term static loads.

Footprint: After a jackup unit is removed from a site, depressions are left in the seabed where the spudcans were located. The soil in way of the footprints has modified physical profiles of the seabed and soil properties.

Foundation Fixity: Also called spudcan-soil rotational stiffness, is the rotational restraint offered by the soil supporting the foundation.

ABS GUIDANCE NOTES ON GEOTECHNICAL PERFORMANCE OF SPUDCAN FOUNDATIONS . 2017 1

Section 1 Introduction

Mobile Jackup Platform: A self-elevating unit with movable legs capable of raising its hull above the surface of the sea.

Perforation Drilling: A technique performed to perforate the stiff layer with drill holes and remove the soil as one of the mitigation methods to reduce the punch-through failure.

Prescriptive Preloading: During preloading, the hull is kept at, in or close to water level, with each individual leg preloaded by sequential filling and discharge of selected preload ballast tanks.

Punch-Through Failure: Unexpected sudden and rapid penetration of the spudcan through soil occurs when strong soil is over soft soil and is a major risk for the stability and equilibrium of the jackup structure.

Rack Phase Difference (RPD): The difference in elevations between the chords of any one leg.

Reaming: A technique that involves forcing a leg into position by incremental vertical reciprocation to penetrate the spudcan into the soil at the required position. Reaming is one of the mitigation methods for reducing spudcan footprint interaction.

Remoulded Undrained Shear Strength: The magnitude of the shear stress that a disturbed soil can sustain in an undrained condition.

Simultaneous Preloading: During preloading the hull is held with minimal draft or air gap and the preload is incrementally increased on all the legs simultaneously.

Soil Sampling: A suitably stored, small amount of soil for visual inspection and laboratory testing for the determination of the soil unit geological provenance, characteristics and geotechnical engineering design parameters.

Spudcan: A large inverted cone that is roughly circular in-plan with a shallow conical underside and a sharp protruding spigot. It is mounted at the base of a jackup’s leg, and is primarily considered to provide sliding and bearing resistance to the jackup rig when deployed into the sea bed.

Stomping: A process where a spudcan is initially emplaced away from the center of an old footprint and then the spudcan is used to displace soil towards the old footprint at desired positions to widen the disturbed region. In soft to stiff clay, stomping is very effective in mitigating spudcan footprint interaction.

Triaxial Test: A common laboratory testing method widely used for obtaining shear strength parameters for a variety of soil types under drained or undrained condition.

7 Symbols and Abbreviations

7.1 Symbols A = spudcan maximum cross sectional area

Aa = area ratio

Ae = spudcan effective bearing area based on cross-section taken at uppermost part of bearing area in contact with soil

As = spudcan laterally projected embedded area

cul = undrained cohesive shear strength at spudcan tip

cu0 = undrained cohesive shear strength at lowest maximum bearing area

cv = coefficient of consolidation of clay

D = maximum spudcan diameter below mudline

Deff = equivalent spudcan diameter at mudline

Df = diameter of the spudcan which forms the footprint

e = opening ratio

2 ABS GUIDANCE NOTES ON GEOTECHNICAL PERFORMANCE OF SPUDCAN FOUNDATIONS . 2017


Eu = Young’s modulus of clay in undrained condition

FH = leg horizontal reaction associated with FV

FM = leg moment reaction associated with FV

Fmob = mobilization factor to reduce the calculated bearing capacity of spudcan foundation

fr = reduction factor

FV = leg vertical reaction

G = soil shear modulus

hc = soil back flow depth

hlayer = sand layer thickness

HL0 = horizontal foundation capacity

hplug = height of sand plug

ht = height of spudcan widest diameter

I = second moment of area

ID = relative density

IRD = relative dilatancy

k0 = coefficient of earth pressure at-rest

K1 = vertical stiffness

K2 = horizontal stiffness

K3 = rotational stiffness

Ks = punching shear coefficient

m = 3 for failure under triaxial or general loading conditions and 5 under plane-strain conditions

ML0 = moment capacity of foundation

Nc = bearing capacity factor at shallow spudcan penetration prior to any backflow

Nkt = cone factor

Nγ, Nq = bearing capacity factors

p' = mean effective stress

pr = 100 kPa (0.010197 kgf/mm2, 14.5 lbf/in2)

qc = cone resistance

qnet = net cone resistance, which is the measured cone resistance being corrected for the pore pressure effects and the total in-situ vertical stress

Qcrushing = particle crushing strength on a natural log scale

Qv = ultimate vertical bearing capacity of the spudcan foundation

R = radius of the circular footing

Rd = radial distance from spudcan center to footprint center

r = shear strength ratio



rf = failure ratio

Sua = undrained shear strength averaged from (h – ht) to (h + hplug)

Subs = shear strength intercept at the clay layer’s surface in strong overlying soft soil

Su_footprint = footprint undrained shear strength

SuH = shear strength at the back flow depth hc

Sum = shear strength intercept at the layer’s surface

Su,plugbase = undrained shear strength corresponding to the level of the sand plug base

Sutop = undrained shear strength of the top strong clay

Su_undisturbed = undisturbed undrained shear strength

Su0 = undrained shear strength at the lowest depth of the maximum plan area of spudcan

uS = mean value of undrained shear strength

t = consolidation time

V = volume of the conical spudcan below mudline when z = ym

Vc = volume of the conical spudcan below mudline

VL0 = maximum vertical foundation load during preloading

Vsoil = volume of the backfill soil that rests on the spudcan

z = soil depth

z = mean values of soil depth

γ' = submerged soil unit weight

ν = Poisson’s ratio

ξsq, ξhγ = empirical shape factors

ξhq = empirical depth factor

ρ = shear strength gradient with soil depth

σ = standard deviation

vσ ′ = effective stress at the level of the sand plug base

0vσ ′ = effective overburden stress

τ1 = adjusted time factor for soil consolidation during the operational period

τ2 = adjusted time factor for soil consolidation during the elapsed time after a footprint is formed

φ' = friction angle

φcv = critical state friction angle



7.3 Abbreviations CAUC Anisotropically consolidated undrained triaxial compression test

CAUE Anisotropically consolidated undrained triaxial extension test

CID Isotropically consolidated drained test

CoV Coefficient of variation

DS Direct shear test

DSS Direct simple shear test

MinV Miniature vane

MODU Mobile offshore drilling unit

MV Motor vane

PP Pocket penetrometer

RPD Rack phase difference

SI Site investigation

TV Torvane

UU Unconsolidated undrained triaxial tests


S e c t i o n 2 : S i t e A s s e s s m e n t

S E C T I O N 2 Site Assessment

1 General A clear understanding of the seabed and sub-seabed conditions is critical for the site specific assessment of a jackup’s suitability during installation, elevated operations and leg extraction. The jackup foundation site assessment procedure usually comprises:

• Acquisition of regional and local geological and geotechnical data (to include previous jackup foundation performance if available)

• Geological data review to develop the ground model

• Geophysical site survey to refine the ground model

• Geotechnical Site Investigation (SI) to further refine the ground model

• Generation of geotechnical design profiles with engineering parameters

• Performance of jackup foundation site-specific assessment using the design profiles.

The Geophysical Site Survey report should describe the interpreted ground model; include an explanation of the geological setting, depositional environment and history together with descriptions of any potential geohazards which could influence jackup operations.

Intrusive geotechnical SIs are conducted in order to ground-truth the geophysical data and to obtain the required geotechnical index and strength measurements. The geotechnical SI allows confirmation or further refinement of the interpreted ground model. Adequate data are required to facilitate detailed engineering characterization of each soil layer and to provide understanding of the spatial variation of these parameters.

Preferably the geophysical and geotechnical components are planned together as integrated parts of the same investigation. The soil design profiles with associated engineering strength parameters are then developed for use in the predictive bearing capacity analyses. The actual scope of work developed will depend upon the vertical and lateral variability of the soil as well as the presence of any geohazards.

The main requirements of site investigation are to be in accordance with 3-2-5/3 of the ABS Rules for Building and Classing Offshore Installations.

3 Geotechnical Site Investigation Ideally, the geotechnical SI should be conducted well in advance of the jackup deployment to the field in order to allow time for adequate data interpretation and site specific assessment. The scope of in-situ tests, soil sampling, soil index, strength and advanced laboratory test data required for analytical purposes will depend on the ground conditions. For a continuous layer of homogeneous soil, limited data acquisition will be necessary. However for highly variable and complex ground conditions where advanced foundation performance modeling is to be conducted, a considerably greater amount of information may be required. The SI scope of work should be planned according to the expected circumstances with the provision for amendment should the actual ground conditions differ from that expected during the investigation.

The geotechnical report should include borehole logs and soil laboratory test procedures and results. If piezocone or other in-situ penetrometer tests are performed, the test records should be documented together with descriptions of test procedures, tool calibration certificates, field offsets and processing information, as applicable. All reports should contain water depth data and geographic coordinates of the location with the measurement datum clearly stated.


Section 2 Site Assessment

5 Sampling and Field Testing Sampling and in-situ testing tools can be deployed through a drill string supported by a drilling derrick mounted on a floating vessel, and these are referred to as “down hole operations”. For marine operations, the drilling system should be heave compensated to minimize sample disturbance resulting from vessel movement. Alternatively, in-situ testing and sampling operations can be conducted using remotely operated systems deployed at the mudline in the “seabed mode” which do not require heave compensation. These systems may provide improved depth control. However, due to their limited thrust, the borehole may not be able to advance to the required depth.

Details of the field investigation and subsequent testing program depend on soil variability, environmental design conditions, structure type and geometry, presence of geo-hazards, etc. At least one borehole with sampling and laboratory testing, or in-situ testing is provided at each platform location. If the foundation area shows significant variation, it may be necessary to perform more than one boring/in-situ testing to verify the irregular soil conditions.

Soil sampling provides material for visual inspection and laboratory testing for the determination of the soil’s geological provenance, characteristics and geotechnical engineering design parameters. Additionally, such parameters can be evaluated from in-situ field testing. Continuous in-situ test profiling may also be used to more accurately determine layer boundaries and material variation.

As the quality of the SI significantly affects the quality of the spudcan penetration prediction, sampling and field testing should be planned accordingly. Jackup SI may be conducted from the jackup itself where there is reliable knowledge of the ground conditions and local jackup operating experience, otherwise relying only on such cursory data should only be considered where it is certain that this data will be adequate for the intended purpose.

During soil sampling material disturbance should be minimized since the greater the soil disturbance the less representative the laboratory test results will be of the in-situ condition. Jackup spudcan penetration performance can be significantly influenced by minor variations in the soil properties and the acquisition of continuous vertical soil profiles is recommended. When continuous profiling is not possible, the aim should be to minimize the gap between data points. The target data gaps should not exceed 0.2 m, although it is noted that for some geotechnical systems this may be unachievable, in which case the maximum allowable gap may have to be increased to 0.5 m. Due consideration should be given to the consequence of data gaps in terms of their location and size, and how this may influence interpretation of the ground model and the necessity to acquire additional data.

Cone penetration tests provide measurements of tip penetration resistance, sleeve friction and pore pressure and are conducted either continuously in the seabed mode or in up to 4.5 m strokes in the down hole mode. Unlike soil sampling and testing data the continuous cone penetration test plots allow for the identification of stratigraphic boundaries and soil strength trends, and the data are usually compared with the laboratory strength test data. The cone penetration tests can be further extended to incorporate dissipation tests to estimate the consolidation coefficient of soils.

Where site specific borehole data are required for a jackup’s operating site assessment, the borehole or piezocone penetration target depth is the greater of either 30 m or 1.5 × spudcan diameter beneath the calculated spudcan tip penetration depth at the maximum preload, with all relevant soil properties being sufficiently investigated. Where appropriate, the target depth may be extended deeper to account for future jackup installation operations where a different rig subjected to a higher preload level may be employed.

A full-flow penetrometer; such as the cylindrical T-bar and spherical ball penetrometer, offer some advantages over the conventional cone in accurate strength interpretation, measurement of remoulded strength and hence soil sensitivity in soft soils. They are not commonly employed in geotechnical SI for jackup assessment and the understanding of the penetration behaviors of these penetrometers is restricted to clay soil conditions.

In-situ vane tests provide only discrete measurements of the intact and remoulded undrained shear strength of clays, and thereby soil sensitivity. The vane testing procedure is important and the strength measured may be sensitive to drainage effects, material heterogeneity and in-situ stress anisotropy. Note that the in-situ vane tests may not be suitable for soil conditions which are heterogeneous due to the test nature of providing only discrete measurement. In comparison, the various penetrometer systems are more versatile than the vane testing systems, and as such, their use is favored.



7 Laboratory Testing Soil samples are extruded from the sample tubes on the SI vessel, (in the same direction which they were taken to minimize disturbance by stress reversal), where they are carefully separated from drill cuttings and any heavily disturbed material. The samples are described, photographed and catalogued in accordance with recommended industry practice. Selected undisturbed samples are stored in sealed containers, which are usually wax-filled, to preserve the moisture content and limit further disturbance.

A range of different soil tests are available from simple laboratory index tests to advanced stress path tests conducted in specialist onshore soil testing laboratories. Typically offshore tests include moisture content and density determination, carbonate content, particle size distribution together with simple strength tests such as pocket penetrometer, torvane, laboratory vane, fall cone, unconsolidated undrained triaxial tests (UU). Index tests are conducted to assist with the soil type classification and also to provide preliminary parameters for initial jackup foundation assessment.

While these undrained soil strength tests are comparatively quick and relatively easily conducted in the offshore laboratory they measure shear strength by different failure mechanisms, and may provide significant strength data scatter and may not provide an accurate indication of sample quality. Some offshore soil laboratory tests, such as the fall cone and the motor vane, can be used to provide an indication of the soil sensitivity.

Advanced onshore soil laboratory testing is encouraged as these strength tests [e.g., anisotropically consolidated undrained triaxial compression test (CAUC), anisotropically consolidated undrained triaxial extension test (CAUE), direct simple shear (DSS) test] provide data for calibration of the in-situ penetrometer results. Oedometer tests provide information about the stiffness and drainage behavior of the soil. Simple offshore laboratory strength tests may provide adequate information for relatively simple soil conditions which lend themselves to modeling by simple bearing capacity formulations. However, for more complex soil conditions, simple soil test results with the application of simple models become less reliable and advanced soil testing is advisable. The relative reliability of various tests in measuring the undrained strength of clay is presented by [Ref.1] in Section 2, Table 1.

TABLE 1 Relative Reliability of Tests Measuring the Strength of Clay Soils

Test Type Soil Profiling Intact Su(1)

Remoulded Su(1) <40 kPa 41-80 kPa >80 kPa

Piezocone 1 2 2 2 4-5 T-bar & ball penetrometers

1 (with pore pressure

measurement)

1-2 1-2 1-2 1-2

In-situ vane --- 1-2 1-2 1-2 3 UU(2) --- 4-5 3-5 2-5 2-3 Motor vane(2) --- 3-5 3-5 4-5 2-3 Torvane(2) --- 3-5 3-5 4-5 --- Pocket penetrometers(2) --- 4-5 4-5 4-5 --- CIU/CAU/DSS --- 1-2 1-2 1-2 2

Notes: 1 Rating ~ 1 High reliability; 2 High to moderate reliability; 3 Moderate reliability; 4 Moderate to low reliability;

5 Low reliability.

2 The test result reliability is dependent on the sample quality (or degree of sample disturbance) and soil homogeneity.



9 Soil Engineering Parameter Selection Section 2, Table 2 lists the basic soil properties which should be obtained for jackup foundation assessment purposes for homogeneous clay or sand. For silty material, the properties defined for clay may be applicable but with consideration of partial drainage characteristics. For a highly variable and complex ground where advanced foundation performance modeling is to be conducted more soil properties may be obtained.

TABLE 2 Index Properties for Jackup Foundation Site Specific Assessment

Soil Type Strength Properties Index Properties and Additional Parameters

Clay Undrained shear strength Su Remoulded shear strength and soil sensitivity St

Water content Plastic limit Liquid limit Submerged unit weight of clay Coefficient of consolidation Over consolidated ratio

Sand Critical state angle of friction Crushing strength

Particle size distribution curve Relative density Submerged unit weight of sand Over consolidated ratio

9.1 Identification of Soil Type and Layers In order to determine the soil stratification, the desk study, geophysical site survey report and geotechnical testing results should be viewed in a holistic and integrated manner. As part of this process, variation of the subsoil conditions across the planned jackup spudcan locations should be assessed. Soil classification charts are empirical correlations which should be verified against the site geotechnical data (i.e., borehole logs and laboratory test data). For offshore conditions, pore pressure data may be generally considered more reliable than friction sleeve data for soil classification.

Continuous penetrometer profiling is suggested in order to provide a full profile for site characterization. Where intermittent piezocone, penetrometer and soil sampling is used in an alternating manner, data gaps are introduced and precise identification of layer boundaries may not be possible. Furthermore, where intermittent penetrometer profiling is used, portions of test data may be of reduced quality and reliability due to soil disturbance from the drilling and soil sampling operations. Discrepancies in the penetrometer resistance profiles between two sequential penetrometer test strokes may be difficult to resolve, and this may introduce uncertainty in the soil layer interpretation. If intermittent profiling is unavoidable, the interval between strokes should be kept as small as practicable and drilling disturbance should be minimized.

It is also suggested that penetrometer tests be conducted adjacent to continuous sampling borehole(s) with a separation on the order of 5 m to avoid interference. Also obtaining pore pressure can aid in the identification of soil type and layering.

The following paragraphs consider strength measurement in:

• Clay, where undrained conditions are assumed for both testing and spudcan penetration;

• Sand, where drained conditions are assumed for both testing and spudcan penetration.



9.3 Clay The undrained shear strength, Su, can be established from a number of commonly conducted strength tests as follows.

• In-situ testing: field penetrometers (cone), vane shear

• High quality laboratory testing: CAUC, CAUE, DSS

• Simple laboratory testing: unconsolidated undrained triaxial test (UU), miniature vane (MinV), motor vane (MV), torvane (TV), pocket penetrometer (PP)

The reliability of the interpreted strength parameters relies heavily on the quality of test and sample conditions. The simple laboratory tests typically should be conducted immediately after the samples are extruded in the laboratory. However, the test results may tend to show significant variability both within a given type of test and between different tests, and are ‘operator dependent’. Tests of this type may be sufficient where reasonable knowledge of the ground conditions and knowledge of successful jackup operations in the locality are available; the data may also be used to extend or interpolate between results from previous high quality laboratory tests. As samples are not reconsolidated prior to testing, the reliability of these tests suffers from unquantifiable sample quality.

The knowledge of remoulded shear strength and sensitivity is necessary for consideration of strain softening effects, particularly for highly sensitive clays. Strain softening reduces the average operational strength during spudcan penetration, thus increasing the penetration. Partial remoulding of the soil will also affect the magnitude of strength recovery or enhancement with time and this is relevant for spudcan breakout force assessment and bearing capacity evaluation for jackup revisits.

In-situ cyclic full-flow penetrometer tests (using T-bar or Ball) with 10 cycles of penetration and extraction, will generally show a well-defined remoulded penetration resistance. The ratio by which the penetration resistance decreases between initial and post-cyclic penetration resistance will be less than the actual sensitivity at the elemental test level, largely due to the partial remoulding that occurs during initial penetration. Such tests do, however, provide an appropriate measure of remoulded shear strength that is directly applicable to spudcan performance.

Where the sediments may be susceptible to significant strength loss during cyclic loading under environmental or seismic loading conditions, cyclic shearing tests should be undertaken in the laboratory in addition to monotonic tests. In the depth range of expected spudcan penetration, sufficient cyclic tests should be undertaken to establish a “cyclic fatigue” curve, showing how the normalized shear stress to cause a given magnitude of strain varies with the number of cycles. This information may then be used to assess an appropriate cyclic shear strength for use in quantifying the jackup’s performance during installation and under ultimate cyclic loading conditions.

9.5 Sand The most common laboratory tests for determining effective strength parameters in sand are CID and DS tests, which represents triaxial and plane strain condition, respectively. Sample disturbance is inevitable when sampling cohesionless material from the seabed. The samples are reconstituted to their approximate in-situ state, with the relative density generally estimated from the cone resistance. Appropriate effective stresses are then applied before the shearing stage. During shearing, it is important to verify that the shearing rate applied is slow enough to prevent the development of excess pore pressure. Theoretically, under a given stress level, it is expected that the friction angle φ' from DS tests will be greater than that from CID tests.

To better account for the stress level effect on φ', the design value may be estimated from the value of the relative density, ID, and the critical state friction angle, φcv, using an appropriate strength-dilatancy relationship that takes account of the mean effective stress p' during bearing failure.

Since the value of critical state friction angle φcv lies within a small range, at least for silica sand, it is possible to estimate the in-situ directly from the cone resistance, qc. The following expression has been applied widely to sandy sites in the North Sea by [Ref.2].



ID = a1lnr

c

pq – a2ln

r

v

pk

3)21( 00 +′σ – a3

where

a1 = 0.336

a2 = 0.154

a3 = 1.91

0vσ ′ = effective overburden stress, in kPa (kgf/mm2, lbf/in2)

qc = cone resistance, in kPa (kgf/mm2, lbf/in2)

pr = 100 kPa (0.010197 kgf/mm2, 14.5 lbf/in2)

k0 = coefficient of earth pressure at-rest, normally taken as 0.5 and 1.0 in order to generate upper and lower bounds of ID, respectively

For a given ID, the design value of φ' can be determined using the general strength-dilatancy framework established by [Ref.3] which makes allowance for different sand types and loading conditions as follows.

φ' = φcv + mIRD

IRD = ID[Qcrushing – ln(p')] – 1 0 ≤ IRD ≤ 4

where

m = constant

= 3 for failure under triaxial or general loading conditions

= 5 under plane-strain conditions

IRD = relative dilatancy

Qcrushing = particle crushing strength on a natural log scale, in kPa (kgf/mm2, lbf/in2)

p' = mean effective stress which in turn depends on the value of φ', in kPa (kgf/mm2, lbf/in2).

As an approximation, the recommendation is to treat p' as the maximum preload pressure.

The value of φcv may be obtained from direct shear tests on disturbed sand, from the “steady state” friction angle in the later stages of the test. Some of the reported values for φcv and Qcrushing by [Ref.4] are given in Section 2, Table 3. The above procedures provide an estimate of the peak angle of friction, including the effect of the average stress level in the soil. It is important to note, however, that as a spudcan continuously penetrates the soil, the peak strength is not mobilized simultaneously throughout the deforming soil. As a result, calculations of spudcan resistance based solely on peak strength of a rigid-plastic soil can result in overestimates of resistance. [Ref.5] addresses this issue by employing reduced friction angles. An alternative approach is suggested in Section 3 where a mobilization factor is applied to the calculated resistance.



TABLE 3 Values of φcv and Qcrushing Derived from Triaxial Compression Tests

Sand Mineralogy Qcrushing φcv (°) Ticino Siliceous (containing comparable amounts of

quartz and feldspar grains) 10.8 33.5

Toyoura Quartz 9.8 32 Hokksund Siliceous 9.2 34 Mol Quartz 10 31.6 Ottawa Quartz (with varying fine content from 0 to 20%) 9.8 to 10.9 30 to 33.5 Antwerpian Quartz & Glauconite 7.8 to 8.5 31.5 Kenya Calcareous 8.5 40.2 Quiou Calcareous 7.5 41.7

9.7 Derivation of Soil Strength Profile Statistical methods can be used to more objectively derive a mean undrained shear strength profile as a function of depth in an extent of homogeneous soil. Assuming a linear variation with depth, the profile can be described by:

Su = Sum + ρz

where

Sum = strength intercept at the layer’s surface, in Pa (kgf/mm2, lbf/in2)

ρ = shear strength gradient with soil depth z, in Pa/m (kgf/mm3, lbf/in3)

z = soil depth, in m (mm, in.)

For a set of n data points in a homogeneous layer, the parameters Sum and ρ can be evaluated as:

ρ = ( )( )

( )∑∑

=

=

−

−−n

i i

n

i uiui

zz

SSzz

12

1 ,

Sum = uS – ρ z

where the mean values of depth , z , and undrained shear strength, uS , are calculated as:

z = ∑=

n

iiz

n 1

1

uS = ∑=

n

iiuS

n 1,

1

These formulations simply fit a straight line through n data points with all the raw data points assumed to be correlated and each data point in the layer assumed to be of equal importance.

Detailed knowledge of the local soil characteristics can be incorporated into the soil strength derivation process if the data can be presented in a quantitative form such as by test reliability (see, for example, Section 2, Table 1). This can then be incorporated into the statistical method used to determine the strength profile and be achieved with the use of weighting factors to modify discrete data values. However, manipulation of raw data in this fashion has to be justifiable and conducted with caution.



Further information provided by the discrete set of strength points includes an estimate of the variation of the undrained shear strength measurements. This is commonly expressed in terms of standard deviation σ and coefficient of variation (CoV) as follows.

σ = ( )[ ]∑=

+−n

iiumiu zSS

n 1

2,

1 ρ

CoV = uSσ

The approach of adopting a conservative estimate of the shear strength profile, as is common practice in other assessments of bearing capacity, is not appropriate for the jackup installation calculation, where an accurate estimate of the actual penetration is required. However, the use of confidence bands on soil strength profiles, making reasonable allowance for uncertainties in soil strength measurement, is beneficial for assessment purposes.

Confidence bands placed on the strength profile should reflect the consequence of the final bearing capacity analysis. For instance, in situations of soft clay where final penetration depth is critical so that the jackup does not run out of leg length, it is essential that the strength profile is lower than the best estimate. However, worst case scenarios, for cases with punch-through potential, may be to consider upper bound strengths in one layer and lower bound strengths in the next. It is indeed possible that many different profiles may require analysis, each reflecting bounds on the problem and consequence.

Consideration of uncertainty in parameters other than strength may also be required. For example, the depth of a layer interface may be a critical factor for punch-through calculations. If there is uncertainty in its depth (or in the thickness of a layer), this should be accounted for in the final bearing capacity analysis. Uncertainties may arise due to spatial variability or uncertainty in interpretation of the test data.

For relatively homogeneous soil layers confidence bands can be determined in a meaningful way by a proportion of the standard deviation of the measurements above and below the mean profile, where the chosen proportion should reflect the uncertainty and severity of consequences. Calculating at ±1 standard deviation from the mean undrained shear strength is often employed in industry practice.


S e c t i o n 3 : S p u d c a n P e n e t r a t i o n i n S i n g l e L a y e r S o i l

S E C T I O N 3 Spudcan Penetration in Single Layer Soil

1 Background Before a jackup is installed at a site, a prediction of the spudcans’ penetrations into the seabed as a function of the imposed loads should be made. The actual load-penetration response is recommended to be monitored on site, and this response should be compared with the predictions.

The purposes of this load-penetration prediction are to:

• Evaluate whether the rig may be able to operate at the site (e.g., whether the leg length available is sufficient).

• Identify any potentially hazardous conditions (e.g., the possibility of punch-through), so that plans can be made to mitigate risks.

• Provide a benchmark against which the actual load-penetration performance can be compared. Deviations from the predictions may indicate an inadequate understanding of ground conditions. In this latter case the consequences depend critically on the nature of the deviations, and whether there are possible implications of hazards.

To make prediction the following information is required:

• Geometry of the spudcans

• A ground model and design soil profile

• Expected light-ship load and maximum preload on each spudcan

Recommended best practice is that a complete load penetration curve should be developed, normally extending to a depth at least, the greater of:

• The predicted penetration at 1.5 times the maximum preload value

• 0.5 times the spudcan diameter below the predicted penetration at the preload value

Analysis and design tools to determine the spudcan penetration and resistance curve can be classified as two general methods (e.g., advanced numerical analysis and simplified prediction methods).

3 Numerical Simulations Lagrangian finite element analysis of large deformation problems such as spudcan penetration often encounters numerical problems arising from excessive element distortion causing degradation of accuracy and convergence difficulty. The Eulerian formulation does not face this problem as it permits the mesh to deform independently of the material. It is widely used in fluid mechanics applications, and has also been applied to other problems like deep penetration analysis involving geological media. A numerical simulation method is convenient for simulating single layer or double layer soil with complete soil parameters. The simulation should be assessed and verified by centrifuge test result or field data.

The spudcan is normally simulated as a rigid body using Lagrangian element, which means that the deformation of the spudcan during penetration is neglected. A void mesh layer allows the soil to heave and flow into the initially empty Eulerian elements. The soil domain is simulated as Eulerian elements because of severe soil distortion. Typically, commercially available software employs explicit Eulerian analysis, which uses a total stress approach and provides no information on excess pore water pressure.


Section 3 Spudcan Penetration in Single Layer Soil

The undrained, effective stress Eulerian finite element analysis of spudcan penetration in clay was developed, which allows excess pore pressure to be computed [Ref.6]. Three elastic–plastic soil models, namely the modified Cam-Clay, Mohr-Coulomb and Drucker-Prager models, are implemented in ABAQUS and applied to the spudcan penetration problem. Comparisons with centrifuge model data indicate that provided the undrained shear strength profile of the ground is well-replicated, so will the penetration resistance and pore pressure response. Both measurement and computation show significant levels of excess pore pressure being generated during preload. It is, therefore, reasonable to expect that the subsequent dissipation of this excess pore pressure will lead to significant changes in the strength of the soil around the embedded spudcan. Hence, computing pore pressure build-up is an essential step in assessing the long-term performance of spudcan foundations.

5 Simplified Prediction Methods

5.1 General The following Subsection deals with the calculation of the load-penetration curve for a deposit that can be treated as a single layer of clay or sand. Multi-layers of soil will be discussed separately in the next Section.

Penetrations in carbonate sands are highly unpredictable, and they may be small in strongly cemented materials, or large, in un-cemented materials. Extreme care should be exercised when operating in these materials. For silts, it is recommended that upper and lower bound analyses for drained and undrained conditions are performed to determine the range of expected penetrations. The upper bound solution is modeled as loose sand and the lower bound solution as soft clay.

5.3 Spudcan Penetration in Clay Load-penetration predictions for spudcans can be based on the application, at each depth of penetration of a bearing capacity calculation. The penetration of a spudcan into the seabed is conventionally defined as the distance of the tip (lowest point on the spudcan) beneath the mudline; see Section 3, Figure 1. In bearing capacity analysis it is often more convenient to use the lowest depth of the maximum plan area instead. Careful distinction between the two variables z and h is essential. The soil undrained shear strength is assumed to increase linearly as follows. The case that strength decreases with depth and hence negative ρ needs special treatment.

Su = Sum + ρz

where

Sum = undrained shear strength at the soil’s surface, in N/m2 (kgf/mm2, lbf/in2)

ρ = shear strength gradient with soil depth z, in Pa/m (kgf/mm3, lbf/in3)

Most spudcans are polygonal in plan. For the purposes of bearing capacity calculations each plan section should be converted to an equivalent circle of diameter, D, enclosing the same area as the polygon.

An important difference, by comparison with more conventional foundations, is that a spudcan is continuously pushed into the seabed, displacing soil as it goes, resulting in differences in the displacement mechanisms in the soil. An important feature is that after a certain penetration depth, the soil flows back into the hole above the spudcan. A relationship of normalized cavity depth hc/D, with SuH/γ'D was proposed [Ref.7], where SuH is the shear strength at the back flow depth hc, see Section 3, Figure 1.

Dhc =

55.0

′D

SuH

γ –

DSuH

γ′41

This does not give hc/D directly in terms of Sum, and some iteration is needed, making use of:

SuH = Sum + ρhc



FIGURE 1 Embedded Spudcan with Open Soil Cavity

hc

hz

ym

mudline

SuSu0SuHSum

Su avg

z

ρ

D

The ultimate vertical bearing capacity of the spudcan foundation Qv in clay can be expressed as follows.

• If z ≤ ym, then there is partial penetration and no backfill can occur:

Qv = Su0NcAeff + γ'Vc

where

Su0 = undrained shear strength at the lowest depth of the maximum plan area of spudcan, see Section 3, Figure 1

Nc = bearing capacity factor at shallow spudcan penetration prior to any backflow

γ' = submerged soil unit weight, in N/m3 (kgf/mm3, lbf/in3)

Deff = equivalent spudcan diameter at mudline, in m (mm, in.), see Section 3, Figure 2

Aeff = 4

2effDπ

, in m2 (mm2, in2)

Vc = volume of the conical spudcan below mudline, in m3 (mm3, in3)

• If z ≥ ym and h ≤ hc, then:

Qv = Su0NcA + γ'(V + Ah)

where

A = spudcan maximum cross sectional area, in m2 (mm2, in2)

V = volume of the conical spudcan below mudline when z = ym, in m3 (mm3, in3)



• If z ≥ ym and h > hc, then:

Qv = Su0NcA + γ'(V + Ahc)

FIGURE 2 Definition of Equivalent Cone

z ym

Deff

Mudline

Volume Vc

The above calculations make the approximations that, until the critical depth, there is no backflow; but after the critical depth, there is full incremental backflow around the spudcan. This is clearly an idealization, as in reality a more gradual transition will occur.

Nc is estimated using the tabulated values suggested [Ref.8], see Section 3, Table 1. The bearing capacity factor is a function of the cone angle β, the dimensionless embedment depth h/D, the roughness factor α, (0 ≤ α ≤ 1, a value of 0.5 is recommended in the absence of evidence that supports any other value), and the dimensionless measure of the rate of increase of strength with depth ρD/Sum. Alternative published values of bearing capacity factors may be used if they can be justified.



TABLE 1 Bearing Capacity Factors Nc for Conical Shaped Footings

ρD/ Sum h/D

β = 30 β = 60 β = 90 β = 120 β = 150 α = 1 α = 0 α = 1 α = 0 α = 1 α = 0 α = 1 α = 0 α = 1 α = 0

0.0 0.0 8.79 4.61 6.68 4.45 6.17 4.64 6.05 4.95 6.06 5.32 0.1 8.95 4.80 6.89 4.68 6.40 4.90 6.29 5.22 6.31 5.59 0.25 9.18 5.05 7.17 4.98 6.71 5.22 6.61 5.57 6.61 5.94 0.5 9.50 5.41 7.56 5.41 7.13 5.68 7.04 6.03 7.05 6.41 1.0 10.03 5.98 8.18 6.07 7.78 6.37 7.71 6.73 7.72 7.12 2.5 11.10 7.12 9.39 7.33 9.04 7.64 8.98 8.06 9.00 8.46

1.0 0.0 14.47 7.53 8.87 5.81 7.53 5.56 7.09 5.68 6.96 5.93 0.1 14.13 7.45 8.88 5.92 7.65 5.74 7.25 5.88 7.15 6.16 0.25 13.72 7.38 8.91 6.04 7.79 5.93 7.45 6.11 7.35 6.40 0.5 13.24 7.28 8.94 6.2 7.97 6.16 7.65 6.39 7.59 6.70 1.0 12.68 7.21 9.03 6.43 8.20 6.49 7.97 6.79 7.94 7.12 2.5 12.19 7.34 9.38 6.97 8.77 7.24 8.61 7.52 8.60 7.90

2.0 0.0 20.10 10.45 10.98 7.14 8.82 6.46 8.03 6.37 7.73 6.50 0.1 18.40 9.65 10.50 6.92 8.64 6.41 7.97 6.41 7.74 6.59 0.25 16.72 8.89 10.02 6.74 8.64 6.40 7.93 6.46 7.76 6.68 0.5 15.08 8.20 9.60 6.59 8.34 6.40 7.91 6.56 7.81 6.83 1.0 13.54 7.60 9.29 6.55 8.32 6.53 8.03 6.80 7.98 7.10 2.5 12.35 7.37 9.37 6.99 8.71 7.15 8.53 7.42 8.52 7.80

3.0 0.0 25.71 13.36 13.09 8.49 10.08 7.35 8.93 7.04 8.43 7.03 0.1 22.00 11.51 11.84 7.77 9.44 6.99 8.57 6.83 8.20 6.93 0.25 18.85 9.98 10.81 7.24 8.93 6.69 8.27 6.71 8.03 6.87 0.5 16.18 8.75 10.00 6.82 8.56 6.54 8.07 6.65 7.93 6.90 1.0 13.98 7.79 9.42 6.61 8.38 6.55 8.06 6.80 7.99 7.23 2.5 12.42 7.40 9.36 6.99 8.68 7.11 8.50 7.38 8.49 7.76

4.0 0.0 31.32 16.27 15.18 9.83 11.33 8.22 9.81 7.69 9.09 7.54 0.1 25.08 13.10 12.99 8.51 10.13 7.48 9.03 7.20 8.58 7.22 0.25 20.44 10.83 11.41 7.61 9.29 6.94 8.52 6.87 8.22 7.01 0.5 16.91 9.11 10.26 6.97 8.71 6.63 8.17 6.72 8.01 6.95 1.0 14.23 7.91 9.50 6.64 8.42 6.56 8.07 6.80 8.00 7.09 2.5 12.46 7.40 9.35 6.86 8.67 7.04 8.48 7.38 8.47 7.71

5.0 0.0 36.92 19.18 17.26 11.17 12.56 9.11 10.66 8.34 9.73 8.04 0.1 27.75 14.48 13.99 9.14 10.74 7.87 9.45 7.52 8.90 7.46 0.25 21.68 11.46 11.86 7.90 9.56 7.12 8.71 7.01 8.37 7.12 0.5 17.43 9.37 10.44 7.08 8.81 6.69 8.25 6.76 8.06 6.98 1.0 14.40 7.98 9.55 6.66 8.44 6.57 8.08 6.80 8.00 7.08 2.5 12.48 7.40 9.35 6.85 8.66 7.03 8.47 7.34 8.45 7.70

5.5 Spudcan Penetration in Sand Normally soil backflow starts immediately after the widest cross-section of the spudcan is below the ground surface, and any unsupported sand above the level of the spudcan flows freely into the hole above the spudcan. The backfill sand would usually be assumed to rest at the “angle of repose”, which may be taken as approximately the critical state friction angle φcv, see Section 3, Figure 3. However, predicted penetrations in sand will often be sufficiently small that only partial penetration occurs and backflow is not relevant.



• If z ≤ ym, then there is partial penetration and no backfill can occur and the following bearing capacity expression should be used:

Qv = 21 γ'DeffNγFmobAeff + γ'Vc

• If z > ym, then the following bearing capacity expression should be used:

Qv = Fmob

′+′ hqsqqh hNDN ξξξ γγ

21

γγ A + γ'(V – Vsoil)

where

Vsoil = volume of the backfill soil that rests on the spudcan, in m3 (mm3, in3), see Section 3, Figure 3

Nγ, Nq = bearing capacity factors

ξsq, ξhγ = empirical shape factors

ξhq = empirical depth factor

Fmob = mobilization factor

FIGURE 3 Backflow in Sand

Mudline

Volume Vsoil

φcv

The bearing capacity factor, Nγ, is obtained in Section 3, Table 2 [Ref.9]. These are calculated directly for conical shaped footings, and are presented in terms of the cone apex angle, β, the roughness factor, α, (0 ≤ α ≤1, a value of 0.5 is recommended in the absence of evidence that supports any other value), and the friction angle, φ'. The values [9] are only presented for surface footings (h = 0). For h > 0, the value ξhγ = 1.0 is recommended [Ref.10] (i.e., no adjustment should be made to this term to account for depth of embedment).

The bearing capacity factor, Nq, is only relevant once the lowest maximum bearing area is below the mudline (h > 0). The following equation was suggested for a plane strain surface footing [Ref.10]:

Nq = eπtanφ'tan2

4π +

φ′

2

The empirical shape factor ξsq = 1 + tan φ' is applied to convert the plane strain condition to a value appropriate for axial symmetry.



A depth factor ξhq = 1 + 2tanφ'(1 – sinφ')2tan-1(h/D) allows for an increase in the contribution of this factor with depth. However, no account is taken of the influence of cone angle or roughness on the Nq factor.

If a calculation is based on realistic values of the peak angle of friction, for a penetration process such as that of an approximately conical spudcan, the soil resistance will be significantly overestimated since the peak strength is not mobilized simultaneously through the deforming soil. In the SNAME 5-5A, this issue is addressed by artificially reducing the angle of friction used in the bearing capacity calculation. This procedure involves the use of unrealistically low angles of friction that may bear little resemblance to measured values. Here a mobilization factor, Fmob, is used to the calculated resistance. Based on back analysis of 10 case records [Ref.1], of which eight are from the North Sea Region, and the others from the Gulf of Mexico and offshore Australia, a value of the reduction factor of 0.25 to 0.5 is suggested. It would be expected that lower values of Fmob would be applicable to more compressible materials (e.g., carbonate sands) and higher values for stiffer materials, but little quantitative evidence exists. If alternative values of the mobilization factor are better established they should be used. It is arguable that, for consistency, a similar procedure should be used in clays, but experience shows that for clays a calculation that does not employ any such factors provides sufficient accuracy, so that the added complication is not justified.

Note that the above procedures are intended solely for the estimation of the load-penetration response of the spudcan during preloading. Calculation procedures are set out in SNAME 5-5A and ISO 19905-1 for determining fixity of spudcans under combined horizontal and moment loading, and these require selection of appropriate bearing capacity factors. However, calculation procedures in SNAME 5-5A and ISO 19905-1 are not consistent with the above procedure, and the different modeling methods should not be combined.

TABLE 2 Bearing Capacity factors Nγ

Friction Angle φ' (degrees)

Cone Apex Angle β = 30 β = 60 β = 90 β = 120 β = 150

Roughness Factor α = 1 α = 0 α = 1 α = 0 α = 1 α = 0 α = 1 α = 0 α = 1 α = 0

5 2.63 1.87 1.07 0.86 0.62 0.52 0.39 0.33 0.22 0.19 10 5.29 2.83 2 1.3 1.17 0.82 0.78 0.56 0.55 0.37 20 20.86 6.62 7.33 3.08 4.54 2.11 3.37 1.69 2.73 1.43 30 89.8 16.26 31.99 7.95 21.12 6.22 17.58 5.77 15.93 6.27 40 504.1 45.24 209.2 23.75 142.8 22.13 129.4 25.84 128.1 34.36 50 6504.3 145 2650 108.8 1923.3 115.2 1905.4 180.2 1981.3 347.2

5.7 Lattice Leg Effect on Cavity Depth and Spudcan Penetration In current design, the effects of the lattice legs are often ignored. The authors of [Ref.12] presented their findings on monotonic lateral load on a spudcan with lattice leg, and concluded that almost all the lateral load is transferred to the clay by the lattice leg, with minimal moment transfer at the spudcan. The effects of lattice legs on penetration resistance were ignored by previous depth-resistance calculation methods [Ref.13].

The lattice leg effect on spudcan penetration resistance was studied using centrifuge tests and numerical simulations [Ref.14]. The main parameters of the lattice legs studied are the opening ratio e of 0, 0.25, 0.75 and 1 and the area ratio of 0.61 and 1. Section 3, Figure 4 shows that the opening ratio, e, which is defined as the ratio of the total area of the openings on the lattice to the total surface area of the lattice. The area ratio, Aa, is defined as the area enclosed by the sleeve or truss relative to the spudcan bearing area.



In comparison to spudcan penetration without a lattice leg, backflow is found to be partially inhibited; and a lattice leg with a smaller opening ratio tends to induce a deeper cavity formation and yield larger transient vertical bearing capacity. The cavity induced by spudcan penetration is more likely to develop in stiff clay than in soft clay. However, cavity formation appears to account for only a part of the bearing capacity improvement. The other part of the increase is likely to be due to the change in the back flow pattern due to the presence of the leg itself. Moreover, the simulated leg displaying the same radius with the spudcan footing (Aa = 1) is shown to yield the largest vertical bearing capacity compared to one with smaller leg radius (Aa = 0.61). A transient vertical bearing capacity improvement of approximately 30% can be inferred compared to the spudcan without the leg effect considered.

Usually the opening ratio, e, of a lattice leg is about 0.8 to 0.9. To reduce the opening ratio one possible way is to employ a top-mounted skirt on a spudcan. In this arrangement, the spaces between chords of a conventional lattice leg are closed by installing a plate between each of two adjacent chords. The plates extend upward from the top of the spudcan to the soil surface, see Section 3, Figure 5. The area ratio is about 0.5. In comparison with conventional spudcans without skirts and spudcans with downward skirts extending to the cone tip, a top-mounted skirted spudcan has improved global bearing capacity, spudcan fixity, and resistance against punch-through, without compromising hydrodynamic performance [Ref.15]. If the top-mounted skirted spudcan also has downward skirts, the effect of the above three benefits will be amplified.

FIGURE 4 Lattice Leg View

83

83

83 60

R60

Opening

Lattice leg

Lattice leg

Spudcan

(a) Plan View in Relation with Spudcan (b) Elevation View



FIGURE 5 Top Mounted Skirt on Spudcan

(a) Plan View (b) Elevation View

7 Alternative Approach: Direct Correlation with Penetrometer Tests If high quality, almost continuous, penetration testing data are available, which are typically from piezocone penetration test, using a T-bar or ball penetrometer, then an alternative is available. It leads to a quick and simple means of estimating spudcan resistance in clay soils. This method should be used with great caution and it is not recommended in the layered soils.

The Nc value for a typical deeply penetrated spudcan is approximately 9.0, and the Nkt value recommended for cone interpretation is approximately 13.5 (if calibrated against average strength of CAUC, CAUE and DSS) or 18.6 (if converted to equivalent UU strength using average conversion factor of 1.38). Combination of the bearing capacity expressions for the spudcan and for the piezocone leads to:

Qv = kt

c

NN qnetA + γ'(hA + V – Vsoil)

where

qnet = net cone resistance, which is the measured cone resistance being corrected for the pore pressure effects and the total in-situ vertical stress, in N/m2 (kgf/mm2, lbf/in2)

The Nc/Nkt ratio may range between 0.48 and 0.67.

The self-weight and buoyancy terms in the spudcan analysis have been included as γ'(hA + V – Vsoil), in N (kgf, lbf).

In the cases where the soil is sufficiently coarse partial drainage will occur during the penetrometer test, even though the spudcan penetration will be undrained. The Nc/Nkt ratio needs to be reduced to account for this effect.



9 Consequences for Observation The penetration at lightship load is highly recommended to always be measured as well as the penetration at full preload; and ideally, continuous measurements should be taken throughout the installation process. The observed spudcan penetration response should be compared with the predictions particularly at still water and full preload conditions.

If the trend of the measurements follows that predicted then it may be assumed that the analytical model may reasonably represent the actual soil failure mechanisms. If the model and actual mechanism correspond then any discrepancy between prediction and measurement is best explained by a systematic offset, such as the soil strengths being consistently over-estimated. Depending on the magnitude of the difference, further investigation may be appropriate to explain the discrepancy.

If the observed spudcan load penetration path does not follow predictions, then this would provide an immediate warning to use extreme caution. Even if the final penetration is within the expected range, if the trend is not predicted correctly then the model may not be capturing the appropriate mechanisms, and the agreement may be no more than coincidental. If the measurements are far from the prediction and the trend is incorrect, further investigation should be pursued.


S e c t i o n 4 : P u n c h - t h r o u g h i n S t r o n g S o i l O v e r l y i n g S o f t S o i l

S E C T I O N 4 Punch-through in Strong Soil Overlying Soft Soil

1 General Punch-through potentially occurs in ground exhibiting post-peak reduction in bearing resistance, in which exceeding the peak bearing resistance results in excessive and uncontrollable spudcan settlement and loss of hull trim. Punch-through failure is possibly the most hazardous of all the geohazards for jackup foundations, and is responsible for more than half the events accounting for loss of life, injury, structural damage, etc. Prediction methods, as indicated below, may be used to evaluate such risk. Other methods can be used if they are justified.

3 Prediction Methods

3.1 Sand over Clay For a thin top sand layer of hlayer/D ≤ 0.5, punch-through failure would not occur for a spudcan on loose sand overlying normally consolidated clay (Section 4, Figure 1) [Ref.16]. With a thicker top sand layer of hlayer/D > 0.5, a significant peak value can be experienced in the penetration resistance response of a spudcan.

In the calculation the sand is treated as drained with strength represented by friction angle φ' or relative density ID, and critical state friction angle, φcv; and the clay is treated as undrained with strength Sub, see Section 4, Figure 1. The resistance in the two-layer system is bounded by the resistance in the lower clay (as lower bound) and uniform sand (as upper bound) at all times. Backflow will occur for a foundation in sand. Formation and subsequent progression of the sand plug trapped below a penetrating spudcan into the underlying clay should be considered.

FIGURE 1 Nomenclature for Spudcan Penetration in Sand over Clay


Section 4 Punch-through in Strong Soil Overlying Soft Soil

3.1.1 Bearing Capacity Calculation when h < hlayer For loose to medium dense sand (ID < 0.65 or φ' < 36°), the punching shear method is proposed to calculate the bearing capacity Qv in spite of the fact that it suffers from the limitation of not being able to model the progressive change of failure mechanism or to model the plug progression [Ref.17]. Qv is the upper bound value in uniform sand.

Qv = NcSubsA + ( )22

2hhlayer

sand −γ′ πDKstanφ' + sandγ′ (hA + V – Vsoil)

where

Ks = punching shear coefficient. For cases with 25° < φ' < 35° and

0.05 < D

S

sand

u

γ′ < 0.5, Kstanφ' ≈ 2.5

6.0

γ′ D

S

sand

u

Su = undrained shear strength at h + D/4 below mudline, in N/m2 (kgf/mm2, lbf/in2)

Vsoil = volume of the backfill soil that rests on the spudcan, in m3 (mm3, in3), see Section 3, Figure 3

Other parameters are as defined in Section 4, Figure 1. For determination of bearing capacity factor, see Section 4, Figure 2 and for the selection of φ', see 2/9.5 [Ref.8].

For dense sand overlying soft clay, back-analysis of centrifuge test results indicates that the above method underestimates peak resistance. Calculation for such soil profiles can be improved by using methods in [Ref.18] and [Ref.19]. These methods provide estimates of bearing capacity when the spudcan is at different depths, taking into consideration the change in shear strength of sand, and accounting for the progressive change of failure mechanism.

3.1.2 Bearing Capacity Calculation when h ≥ hlayer Beyond the sand-clay interface, the resistance is assessed as a foundation in the clay, and the methods in 3/5.3 may be used, with the upper sand layer contributing to the overburden stress. However, sand from the upper layer may be trapped below the penetrating spudcan (see Section 4, Figure 3). Allowance for the trapped sand plug increases the predicted bearing capacity for h ≥ hlayer. The increase is mainly due to the side friction on the sand plug and the depth effect on the bearing capacity factor.

The following is a possible method for assessment of spudcan bearing capacity with inclusion of sand plug [Ref.18 & 20].

• For h = hlayer

Qv =

+′+

DhS

SN pluguavplugbaseuc

4, σ A – γ'Vsoil

• For h ≥ hlayer +ht

Qv =

++′+

DhhS

SN tpluguavplugbaseuc

)(4, σ A – γ'Vsoil

where

Nc = bearing capacity factor

Su,plugbase = undrained shear strength corresponding to the level of the sand plug base, in N/m2 (kgf/mm2, lbf/in2)



= effective stress at the level of the sand plug base, in N/m2 (kgf/mm2, lbf/in2)

Sua = undrained shear strength averaged from (h – ht) to (h + hplug), in N/m2 (kgf/mm2, lbf/in2)

hplug = height of the sand plug, in m (mm, in)

ht = height of the spudcan widest diameter, in m (mm, in)

A = spudcan widest area, in m2 (mm2, in2)

γ' = unit weight of backfill soils, in N/m3 (kgf/mm3, lbf/in3)

Vsoil = volume of backfill soils above spudcan, m3 (mm3, in3)

For dense sand overlying soft clay with hlayer/D < 1, hplug was measured to range between 0.6 to 1.0hlayer [Ref.18,19&20]. A smaller hplug is expected for loose to medium dense sand over clay.

The values back-analyzed from centrifuge data generally fall between the values in Section 4, Figure 2 [Ref. 7&8]. Applying the former set of Nc values in [Ref.8] in the calculation gives lower bound predictions of Qv, and hence these values should be preferred.

FIGURE 2 Values for Spudcan Bearing Capacity Calculation when h ≥ hlayer [Ref.18]



FIGURE 3 Nomenclature of Spudcan Penetration in Sand over Clay when h ≥ hlayer

z

SubSubs

1ρ

hlayer hQv

φcv

CLAY[Sub or (Subs, ρ), γ′clay]

SAND[γ′sand, φ′ or (ID, φcv)]

Trapped soil plug

3.3 Strong Clay over Soft Clay For strong clay overlying soft clay, the resistance in the two-layered system is bounded by resistance in the lower clay as lower bound and upper clay as upper bound. The possibility of soil backflow in the upper layer may be assessed using the procedures for a single layer as in 3/5.3. Soil backflow may occur in the lower layer when hc determined following procedure is greater than hlayer. In this case, hc should be determined using the following equation [Ref.21]:

5.0

,

1

γ4.1

+

×′

≈

ubs

layer

bclay

utop

c

SD

Dh

DS

Dh

ρ

where the parameters are as defined in Section 4, Figure 4.

Formation and progression of the clay plug trapped below a penetrating spudcan into the underlying clay should be considered. The method for calculating resistance in strong clay over soft clay includes three stages [Ref. 21&22]:

• Stage 1 – Calculation of initial to peak bearing capacity

• Stage 2 – Calculation of post-peak bearing capacity

• Stage 3 – Calculation of bearing capacity when h ≥ hlayer



FIGURE 4 Nomenclature of Spudcan Penetration in Strong Clay over Soft Clay

D

z

SubSubs

1ρ

hlayerh′layer

h

Qv

Soft CLAY[γ′clay,b, Sub or (Subs, ρ)]

Strong CLAY[γ′clay,top, Sutop]

5 Mitigation Method – Perforation Drilling Spudcan skirts conventionally project downward from the rim of the footing. Such skirts can increase the jackup’s penetration resistance against punch-through and spudcan-soil rotational stiffness. However, skirts should be sufficiently long to be effective, but a skirt extending appreciably beneath the tip of a spudcan cone can increase hydrodynamic drag, making towing of the unit difficult [Ref.16]. Thus, a skirt’s benefits need to be considered against problems arising from its length.

Perforation drilling has been performed to perforate the stiff layer with drill holes and remove the soil. It has had varying degrees of success for predominantly clay conditions rather than sand ground conditions. In clean sand, it is difficult that the drilled holes will remain open and that the sand is not mixed with the underlying clay. The ability to reduce the punch-through risk by perforation drilling depends on the percentage reduction in the ratio of the reduced peak to the minimum post-peak bearing resistances. A reduced ratio of unity or less means that the punch-through potential has been completely eliminated, although this may not be either achievable or necessarily required.

The perforation drill hole distribution commonly used is an equilateral triangular grid within the spudcan footprint. A greater reduction in the punch-through resistance ratio may be achievable with an increased density of perforations located within and outside the spudcan periphery. The actual hole pattern may not be as regular as planned due to drill bit positioning control, hole collapse, hydraulic fracturing and linkage between adjacent perforations due to excess drilling (water or air) pressure, etc. The expected reduction in the punch-through resistance ratio should be re-evaluated on completion of the drilling operation.

The drill hole “perforations” are typically constructed by open flush rotary drilling methods using a 26-inch bit sometimes coupled with a 36-inch hole opener, which forms a hole of the bit diameter (or hole opener diameter if it is used) or perhaps a hole slightly greater due to drill string flexibility and soil wash-out.

The drill hole spacing (center-to-center) determines the efficiency of the perforation drilling scheme. Experience suggests that the efficacy of perforation drilling reduces once the spacing falls below about 2.3 hole diameters. However, increasing the hole separation results in fewer perforations and less total soil removal. If an inadequate number of perforation holes are drilled then the process will not succeed in removing the punch-through risk.



To date, perforation drilling has been achieved using the rotary open hole water flush method. If insufficient flush rates are used then the disturbed soil may not be removed from the hole. On the other hand, if excess drilling pressures are applied then this may lead to hydraulic fracturing and linkage between adjacent perforations. Both scenarios significantly reduce the effectiveness of the operation.

By using reverse circulation (i.e., using an airlift to remove cuttings from the base of the hole up through the center of the drill string to surface) the effectiveness of the excavation, and hence overall process, may be increased as the material removal efficiency will improve. Since there is little experience with this method for perforation drilling, it is not possible to confirm its efficiency and reliability on vessels so far used for this operation. Experimental data suggest that factors such as perforation distribution pattern in particular, perforation depth, drilling methods, etc., may significantly affect the reduction of peak bearing resistances.

7 Remedial Action during Punch-through Leg runs and punch-through usually occur on a single leg at any one time, although occasionally more than one leg may simultaneously suffer such an event. On completion of the punch-through and leg settlement, the hull may be partially in the water, listing with the legs inclined. The legs may be resisting large bending moments at the hull interaction points (i.e., the guide wear plates and jacking pinions).

The implementation of recovery procedures, following a punch-through, should be carefully planned and executed with expert advice sought where damage may have occurred, or may occur during the recovery process. If inappropriate action is taken during the recovery operation then greater structural damage than caused by the punch-through may result. The recovery plan will depend on the post punch-through situation, the particular jackup design and possibly other factors.

Where a punch-through has occurred and not all the legs have been forced through the strong layer at the end of the preloading operations then unless the unequal spudcan load vs. penetration behavior can be adequately accounted for by the geological model and installation method, the safety of the installation should be reconsidered.


S e c t i o n 5 : F o u n d a t i o n S t a b i l i t y A s s e s s m e n t

S E C T I O N 5 Foundation Stability Assessment

1 Approach The overall foundation stability assessment may follow SNAME 5-5A and ISO 19905-1. There are three steps recommended in the order of increasing complexity and reducing conservatism (See Section 5, Figure 1):

i) Step 1

a) Preload check of leeward leg for pinned spudcan

b) Sliding check of the windward leg for pinned spudcan

ii) Step 2

a) Foundation capacity and sliding check for pinned spudcan

b) Foundation capacity and sliding check for spudcan with moment fixity and vertical and horizontal stiffness

c) Foundation capacity and sliding check for spudcan assuming non-linear foundation stiffness

iii) Step 3

Foundation displacement check

3 Acceptance Criteria The adequacy of a foundation to resist bearing and sliding loads, and when considered rotational stiffness, should be verified using the criteria specified in SNAME 5-5A, ISO 19905-1 or the ABS Rules for Building and Classing Offshore Installations.


Section 5 Foundation Stability Assessment

FIGURE 1 Foundation Stability Checks

Perform structural analysis

Step 1aPreload check

Step 1b Sliding check

Step 2aFoundation capacity and

sliding check (pin)

Structural analysis assuming fixity with elastic vertical and horizontal springs

Step 2bFoundation capacity and

sliding check

Structural analysis assuming full non-linear stiffness

Step 2cFoundation capacity and

sliding check

Step 3Displacement check

on all legs

Foundation acceptableFoundation NOT acceptable

OK

OK

OK

OK

Not OK

Not OK

Not OK

Not OK

Not OK



S e c t i o n 6 : F o u n d a t i o n F i x i t y

S E C T I O N 6 Foundation Fixity

1 Introduction Foundation fixity, so called spudcan soil rotational stiffness, is the rotational restraint offered by the soil supporting the foundation. Traditional self-elevating unit (SEU) design usually assumes a pinned condition to represent spudcan and soil interaction, this being equivalent to zero foundation fixity as shown in Section 6, Figure 1. Since 2003, the MODU Rules permit consideration of foundation fixity for cases involving dynamic response. ABS Rules allow consideration of a range of fixity values up to the “fully fixed” condition. An SEU Owner is to verify that the conditions for which the SEU has been approved are satisfied. Actually, the spudcan/soil rotational supports are partially fixed, see Section 6, Figure 1.

FIGURE 1 Spudcan Soil Rotational Stiffness

2 = 0

Pinned Fully Fixed Partially Fixed

1 3 < 1

At present, to assess the suitability of an SEU for installation at a specific site, and hence the adequacy of its spudcan foundation, the industry relies extensively on SNAME 5-5A. The value of SNAME defined rotational stiffness is approximately 20%-40% of the “fully fixed” value, while the measured dynamic fixity ranges from 30% to 80% of the “fully fixed” value [Ref.24]. The key parameters affecting spudcan fixity include soil permeability, footing embedment, lattice legs, spudcan skirt, time lag between installation and operation, etc.

3 Foundation Fixity

3.1 Initial Stiffness The guidelines provided in SNAME 5-5A and ISO 19905-1 are probably the most widely used in the offshore industry for spudcan analysis and design. The spudcan is typically analyzed as a structural entity which interacts with the soil through vertical, horizontal and rotational springs. The effect of fixity is accounted for by the stiffness of the rotational spring.

Section 6 Foundation Fixity


For a rigid circular footing resting on an elastic half-space, the vertical, horizontal and rotational spring constants K1, K2, and K3, are given by:

Vertical stiffness K1 = 11

2dk

GD

in N/m (kgf/mm, lbf/in)

Horizontal stiffness K2 = 287

)1(16dk

GD

in N/m (kgf/mm, lbf/in)

Rotational stiffness K3 = 3

3

)1(3 dkGD

in N-m/rad (kgf-mm/rad, lbf-in/rad)

where

G = soil shear modulus, in N/m2 (kgf/mm2, lbf/in2)

= Poisson’s ratio of the half space medium

D = footing diameter, in m (mm, in.)

kd1, kd2, kd3 = depth factors for the vertical, horizontal and rotational spring stiffness respectively, as shown in Section 6 Figures 2, 3 and 4. Case (a) in the Figures represent an open hole above the spudcan, case (b) in the Figures represent a back filled hole [Ref.25].

h = spudcan embedment depth from mudline to spudcan maximum lowest bearing area, in m (mm, in.) see Section 3, Figure 1

3.1.1 Selection of Shear Modulus G for Clay

The value of initial, small strain shear modulus G for clay may be calculated as follows.

G = Gmax = Su 25.0OCR

IrNC in N/m2 (kgf/mm2, lbf/in2) with G < SuIrNC

where

Su = undrained shear strength measured at the depth of spudcan lowest maximum bearing area plus 0.15 times spudcan diameter, in N/m2 (kgf/mm2, lbf/in2)

OCR = over consolidation ratio

IrNC = rigidity index for normally consolidated clay. For extreme loading conditions, and in absence of other data, IrNC should be conservatively limited to 400 based on over-consolidated clay sites with plasticity index of up to 40% [Ref. 26]. For regions with low OCR and plasticity index less than 40% like the Gulf of Mexico, IrNC could be up to 600.


FIGURE 2 Depth Factors kd1 for Vertical Spring Stiffness

(a) Open hole above spudcan (b) Back filled hole

FIGURE 3 Depth Factors kd2 for Horizontal Spring Stiffness


FIGURE 4 Depth Factors kd1 for Rotational Spring Stiffness




3.1.2 Selection of Shear Modulus G for Sand The value of initial, small strain shear modulus G for sand may be calculated from

apG

= j50.

a

swl

ApV

where

pa = atmospheric pressure, in N/m2 (kgf/mm2, lbf/in2)

j = dimensionless stiffness factor

= 230

+500

90 rD.

Dr = relative density of sand, in percent

A = spudcan effective bearing area, in m2 (mm2, in2)

Vswl = seabed vertical reaction under still water conditions, in N (kgf, lbf)

3.3 Yield Surface under Combined Loadings For shallow embedment in both sand and clay, the yield interaction is defined by the following expression:

−

−

+

00

2

0

2

014

L

V

L

V

L

M

L

H

VF

VF

MF

HF = 0 .............................................................................. (1)

where

FV = leg vertical reaction, in N (kgf, lbf)

FH = leg horizontal reaction associated with FV, in N (kgf, lbf)

FM = leg moment reaction associated with FV, in N*m (kgf*mm, lbf*in)

VL0 = maximum vertical foundation load during preloading, in N (kgf, lbf)

HL0 and ML0 are defined as follows:

• For sand:

HL0 = 0.12 VL0

ML0 = 0.075VL0 D

• For clay:

HL0 = cu0 A + (cu0 + cul)As

ML0 = 0.1VL0 D

where

As = spudcan laterally projected embedded area, in m2 (mm2, in2)

cul = undrained cohesive shear strength at spudcan tip, in N/m2 (kgf/mm2, lbf/in2)

cu0 = undrained cohesive shear strength at lowest maximum bearing area, in N/m2 (kgf/mm2, lbf/in2)

The yield envelopes for a spudcan in soil are illustrated in Section 6, Figure 5(a) & (b).



FIGURE 5 Yield Envelopes for Conical Footings

–1

0.51

1

0.51

00

1

0

FM/(FHB) = 0

FH/HL0

FV/VL0

2

–2

–2 20

0

FV/VL0 = 0.25

FV/VL0 = 0.5

FM/ML0

FH/HL0

(a) Elevation View (b) Cross Sections at Constant FVHM /VL0

3.5 Calculation Procedures K3 is an initial estimate of the rotational spring constant, see Section 6/3.1. SNAME 5-5A contains guidelines on how the value of K3 should be adjusted based on whether the initial load combination FV, FH, and FM fall within, on or outside the yield surface. A general description of the procedure is given below.

1. Determine yield surface based on preload using Section 6, Equation (1).

2. The vertical, horizontal and initial rotational stiffness are first evaluated and included in the analytical model, together with the applied gravity, metocean and inertial loadings. Based on these stiffness and the applied loadings, the yield interaction function value is calculated. For extreme wave analysis, it is likely that the force combination will fall outside the yield surface, in which case the initial rotational stiffness is arbitrarily reduced and the analysis is repeated.

3. Continue with Step 2 until the force combination acting on the spudcan lies on the yield surface.

4. If the load combination of (FV, FH, FM) lies inside the yield surface, the initial estimate of rotational stiffness should be reduced by a factor, fr. The reduction factor is equal to unity when the moment and horizontal forces are zero. It is given by the following expression:

fr = (1 – n)rf/ln

−

−

f

f

rnr

11

where

rf = failure ratio

=

−

+

00

5.02

0

2

0

14L

V

L

V

L

M

L

H

VF

VF

MF

HF

≤ 1.0

n = parameter to adjust spudcan moment-rotation curves with various degrees of curvature change, and -1 ≤ n ≤ 1



When site specific or regionally applicable information is sufficient to establish a value of n, then that value should be used. Finite element analysis for Gulf of Mexico clay indicates the range of n is between –0.25 to –1.0, with n = –0.5 providing the best overall representation [Ref.27]. When no such information is available, a default value of n = 0 could be used. In special cases where an extremely conservative stiffness reduction is desired a value of n = 1.0 could be used, which leads to:

fr = 1 – rf

5 Consolidation Effect The zone of clay being remoulded during spudcan penetration will reconsolidate after completion of the spudcan installation. The extent of consolidation depends on the period between spudcan installation and operation, soil consolidation characteristics, penetration depth, spudcan geometry etc. The effect of the consolidation on excess pore pressure response and settlement of spudcan was studied, as well as foundation fixity by means of centrifuge tests and numerical simulations in Kaolin clay [Ref.28].

In “no dissipation” tests spudcan rocking commenced immediately after installation, since there is no consolidation. In “full dissipation” tests the excess pore pressure was completely dissipated before commencement of rocking, which is full consolidation. Compared with the settlements in all the “no dissipation” tests that show a non-linear increase, the settlements in all the “full dissipation” tests show much smaller settlements.

The yield envelope, Section 6, Equation (1), is considered to be relatively conservative which cannot accommodate a long-term combined force locus in “full dissipation” test [Ref.28]. Actually, the penetration resistance at a specific depth is changing with the excess pore pressure due to pressure dissipation. A ratio of 1.7 between the long-term and short-term resistance is obtained from the test results. The yield envelop based on long-term penetration resistance is found to accommodate the entire force locus.

The initial foundation fixity for the “full dissipation” tests is about double that from the “no dissipation” test since consolidation of the surrounding soil leads to a significant increase of bending moment resistance in “full dissipation” tests. With an increase in the number of cyclic loadings, foundation fixity in “full dissipation” tests shows an approximately constant value, while the “no dissipation” tests show an increasing tendency. This increase is partially caused by the dissipation of excess pore pressure in the soil surrounding the foundation, and also arises because of the relatively large settlement deeper into the soil specimens. After one thousand loading cycles the magnitude of the foundation fixity in the “no dissipation” tests is gradually close to that in the “full dissipation” tests. Hence the post-installation, long-term consolidation can substantially enhance the loading capacity and rotational fixity of spudcan footing. If such beneficial effects can be reliably obtained and accounted for, they can benefit the design.

7 Lattice Leg Effect on Foundation Fixity For the spudcans with lattice, studies show that the lateral resistance at the top section of lattice is small, this being consistent with the softer soil near the ground surface [Ref.28]. However, large lateral resistance on the middle and bottom sections is implied by the measured moments from the strain gauges. Furthermore, the bending moments next to spudcan are not necessary the largest. This indicates that lattice can help to reduce the footing moment at the connection of spudcan and lattice, and the point of maximum moment is up-shifted. Lattice helps to reduce the settlement of spudcan. With a high opening ratio, the effects of lattice on settlements are not significant. However, with a lower opening ratio and higher area ratio, the lattice can largely and efficiently reduce the settlements. A significant increase in foundation fixity, of the order of 50%, is conferred by the presence of the lattice leg.


S e c t i o n 7 : S p u d c a n - P i l e I n t e r a c t i o n

S E C T I O N 7 Spudcan-Pile Interaction

1 Introduction A jackup rig is often used to do drilling or work-overs adjacent to existing pile supported platforms. Also the jackup may be sited adjacent to a pile supported structure to provide additional accommodations, power generation or fabrication space. Furthermore, some modular packages, such as those for drilling apparatus or construction cranes, are sometimes transferred from a jackup rig to a fixed platform for construction, drilling or reworking of an existing well. Prior to these works, the jackup should first be positioned adjacent to the fixed platform. Depending on the platform’s footprint and the location of the jackup rig, the spudcan foundation of the jackup rig may be close to the permanent pile foundations of the fixed platform, as shown in Section 7, Figure 1. The proximity of the spudcan to the existing piled platform would induce stresses, and may affect the performance of the pile foundations, and subsequently causes distress to the superstructure, see Section 7, Figure 2. If the piles have not been designed to withstand these stresses, the structural integrity of the piles may be threatened.

According to SNAME 5-5A, if the foundation materials comprise either a deep layer of homogeneous firm to stiff clay or sand, and if the proximity of the spudcan to the pile is greater than one spudcan diameter, no significant pile loading is expected. When the proximity is closer than one spudcan diameter, then analysis by the platform owner is recommended to determine the consequences of the induced pile loading.

The induced loading on the pile is the function of:

• Spudcan-pile clearance

• Spudcan and pile diameters

• Spudcan penetration depth

• Pile length for a floating pile

• Pile socket length for a socketed pile

• Upper clay thickness for a dual layered soil profile

• Soil rigidity index

• Pile bending stiffness


Section 7 Spudcan-Pile Interaction

FIGURE 1 Potential Soil Loading Effects on Jacket Platform (after [Ref.29])

Jack-up Spud Can

Remolded Zone

Seafloor

NOT TO SCALE

Pile

Conductors

Displaced Pile

Gap

Forcesin

Bracing

Deformation in Leg

Jacket Translation

Shear Forces and Deformations in Conductors

FIGURE 2 Spudcan-Pile Interaction

Spudcan Activity

Induced Soil Movement

Installation

Removal

Operation

Settlement

Vertical MovementLateral Movement

Axial Force

DeflectionBending Moment

Distress to Superstructure

Assessment Improvement

Pile Response



3 Soil Flow Mechanism for Spudcan in Soft Clay

3.1 Spudcan Penetration Two distinct forms of soil behavior are typically observed during spudcan penetration [Ref.30]. At shallow spudcan penetration, the soil movements are governed by the general bearing capacity failure mechanism. In contrast, localized deep plastic flow mechanism is dominant at deep penetration depths.

The incremental soil movement trajectories at various distances away from the spudcan edge (i.e., 0.25D, 0.5D, 0.75D and 1D, where D is maximum spudcan diameter) at selected spudcan penetration depths of 3 m and 12 m are shown in Section 7, Figure 3. In general, the magnitudes of both lateral and vertical soil movements decrease with increasing spudcan-pile clearance and the zone with significant soil backflow is confined to an area approximately 0.25 diameters away from the spudcan edge. The influence zone at shallow penetration depths is about 0.83D away from spudcan edge and nearly 1D below the spudcan base; while at deep penetration depths, the influence zone is found to be within about 0.42D away from spudcan edge and 0.75D below spudcan.

3.3 Spudcan Operation and Extraction During the jackup’s operation the soil moves laterally inwards and vertically downwards due to cavity collapse and soil reconsolidation. Both the maximum lateral and vertical soil movements occur at the soil surface and gradually decrease with soil depth. For a longer operating period, the magnitude of soil movement becomes larger and gradually stabilizes.

During initial extraction, the soil being dragged underneath the spudcan is primarily caused by suction. After the breakout, suction becomes less prominent and the soil flow is triggered by the cavity below the spudcan with the soil mainly from above the spudcan. The soil flows towards the spudcan base to fill the cavity.

Compared to the operating and extraction stages, the penetration stage is identified as the most critical in terms of adverse impact on adjacent piles.

FIGURE 3 Incremental Soil Movement Trajectories at Different Distances from Spudcan

0.25D 0.5D 0.75D 1D

0.25D 0.5D 0.75D 1D

(a) 3 m Penetration Depth (b) 12 m Penetration Depth



5 Dimensionless Charts Systematic, dimensionless charts may be used to perform preliminary estimates on the severity of spudcan-pile interaction problems in various field conditions [Ref.31]. However, a detailed three-dimensional finite element analysis should be performed in situation where higher accuracy is required.

The maximum bending moment is estimated by multiplying the bending moment for a case of a socketed pile having spudcan-pile clearance of 1D with correction factors to account for the effects of spudcan-pile clearance, pile socket length and thickness of upper clay layer for a socketed pile, clay rigidity index and pile bending stiffness, where appropriate, see Section 7, Figure 4. Each of the above parameter is represented by an adjustment factor.

The bending moments in the charts are for a fixed-headed pile. It is postulated that this would result in a larger induced pile head bending moment when compared to the actual situation where the pile head’s rotational degrees of freedom are considered but the translational degrees of freedom are partially fixed due to the fact that the pile head can deflect together with the platform leg. As such, the measured induced pile bending moments should generally be on the conservative side. The bending moment responses of piles were compared with free head and totally fixed head, which represents for the two extreme cases [Ref.32]. Different from the fixed headed pile, the elevation of maximum bending moment for free-headed pile is located around the center of the pile rather than at the pile head. This is because the free restraint at both ends causes the maximum moment to be focused at the center. However, the magnitude of the maximum bending moment for a free-headed pile is about one third that for a fixed-headed pile. Therefore, the dimensionless charts lead to a conservative estimation of pile bending moment due to spudcan installation. On the other hand, as the soil moves continuously upward, the fixed-headed pile is found to be in compression due to the restraint at the pile top; while the free-headed pile is in tension. However, the magnitude of axial force is comparable for the two kinds of piles.

The charts are derived from specific soil conditions consisting of normally consolidated kaolin clay with typical strength gradient of 1.6 kPa/m depth overlying a layer of sand which represents a hard soil. In cases where multiple soil layers or soil with irregular shear strength are present, additional care should be exercised to evaluate the validity of the procedure and to decide whether a detailed analysis should be conducted.



FIGURE 4 Nomenclature for a Socketed Pile

Sand

Pile Socket Length

ClaySpudcan

Penetration Depth

Spudcan-Pile Clearance

Pile Length

Sand Thickness

Clay Thickness


S e c t i o n 8 : S p u d c a n - F o o t p r i n t I n t e r a c t i o n

S E C T I O N 8 Spudcan-Footprint Interaction

1 Introduction After a jackup is removed from a site, spudcan footprints are left in the seabed, as shown in Section 8, Figure 1. The spudcan footprints are ground conditions that experience:

• Changes in physical profile of the seabed (existence of depression); and

• Changes in soil properties.

The soil beneath a depression may be highly non-uniform due to back flow of remoulded soils during and after spudcan penetration and extraction, and reconsolidation of the soil. If the subsequent positioning of another jackup is very close to or partially overlapping the footprints, the slope of the footprint and the varying soil strength inside and around the footprint results in an eccentric/inclined soil reaction on the spudcan. This can cause a spudcan to slide towards the footprint and hence leads to overloading the leg. Thus pre-loading of jackups near existing footprints can result in uncontrolled penetration, slewing of the rig and even excessive structural stresses in the legs, which might even lead to catastrophic failure. The footprint feature is dependent on various factors such as footing shape and size, soil types and strength, previous spudcan size and penetration, and the elapsed time of previous spudcan operation and after extraction.

FIGURE 1 Bathymetry of an Established Site

(a) Plan View (b) Section View 1-1

SNAME 5-5A suggests that a jackup identical in footing geometry is installed and located in exactly the same position as that of previously installed unit. It is important to verify that the footing geometry and position of the rig are exactly similar to the previous jackup. If it is not possible to locate exactly in the previous location, a minimum distance of one spudcan diameter from the edge of bearing area to the edge of the footprint is recommended.


Section 8 Spudcan-Footprint Interaction

3 Footprint Characteristics and Its Influence on Spudcan-Footprint Interaction The soils within a footprint can be heavily remoulded with highly variable shear strength and physical profile. The spudcan penetration depth and the subsequent spudcan extraction significantly affect the physical profile of the footprint. In softer clay with relatively deep spudcan penetration, a bowl-shaped spudcan imprint is more likely to be formed after spudcan extraction. In firmer clays with a relatively shallow spudcan penetration, a nearly vertically sided cylinder with a fairly firm base indentation is more likely to be formed. The undrained shear strength beneath the footprint immediately after its formation is lower than that of the undisturbed soil and the reduction in undrained shear strength decreases with increasing radial distance from the footprint center. The critical center to center spacing between the initial penetration and re-penetration is found to be 0.5 to 1.0 times spudcan diameter, inducing the peak lateral load and significant bending moment in the leg [Ref.33&34].

Section 8, Figure 2 shows the generalized short-term soil condition of a footprint formed in soft to firm clays for τ1 < 0.002 and τ2 < 0.2, where τ1 and τ2 are the adjusted time factors for soil consolidation during the operational period and during the elapsed time after a footprint is formed, respectively [Ref.33].

τ = cv 2Rt

where

cv = coefficient of consolidation of clay, in m2/s (mm2/s, in2/s)

t = consolidation time, in sec

R = radius of the circular footing, in m (mm, in.)

The footprint generally consists of a crater with the lowest point at a depth of about 0.2D, and a highly non-uniform soil beneath the crater surface, see Section 8, Figure 2. The degree of soil disturbance is classified by the shear strength ratio r, which is the ratio of the footprint undrained shear strength, Su_footprint, to the undisturbed undrained shear strength, Su_undisturbed, at the same elevation.

r = ndisturbeduu

footprintu

SS

_

_

A strength ratio of less than 0.5 is categorized as heavily remoulded, which is basically confined within the spudcan area Rd/Df ≤ 0.5 (Rd is the radial distance from spudcan center to footprint center and Df is the diameter of the spudcan which forms the footprint). The soil with strength ratio of 0.5 to 0.7 is classified as moderately remoulded, whereas a ratio range of 0.7 to 0.9 is less remoulded. The extent of radial soil disturbance varies with depth. At depths of up to 0.5Df, the radial disturbance is found to extend an Rd/Df ratio of 1.5. Below this depth, the major radial disturbance (r < 0.7) is found to be confined within Rd/Df of 0.75. In terms of vertical extent, it is found that the soil is heavily remoulded to a depth of up to 0.9de and moderately remoulded from 0.9 to 1.1de, where de is the penetration depth to spudcan base level. Below this depth, a minor soil strength reduction (0.7< r <0.9) extending up to 0.3Df is observed.

Based on the observations from centrifuge tests, a simplified soil failure mechanism at different penetration depths is postulated [Ref.33]. For the penetration depth above the crater depth, the horizontal forces and moments are caused by the non-uniform soil bearing resistance where the resultant soil reaction inclines at an angle of θ to vertical and an eccentricity of e from the spudcan center, as indicated in Section 8, Figure 3(a). This inclined eccentric soil reaction tends to push and rotate the spudcan from the stronger soil side towards the weaker side. On the other hand, when the spudcan penetrates further below the crater, the soil bearing capacity failure along the sliding surface is initiated that drives the spudcan towards the footprint center. At the same time, the resistance from the soil outside the sliding plane tends to push the spudcan towards the footprint center, as shown in Section 8, Figure 3(b). If the horizontal and rotational movements are restricted, the tendency of horizontal movement and rotation are then transmitted as horizontal forces and rotational moments acting at the spudcan. A combination of these two forces has the effect of magnifying the magnitude of induced horizontal force (as both forces tend to move the spudcan towards the footprint center) while diminishing the magnitude of induced rotational moment (caused by rotation in opposite direction).




FIGURE 2 Generalized Soil Condition of a Footprint

(for 1 < 0.002; 2 < 0.2)

r < 0.5(Heavily remoulded zone)

0.5 < r < 0.7(Moderately disturbed zone)

0.7 < r < 0.9(Less disturbed zone)

r =Su_footprint

Su_undisturbed

de = penetration depth to spudcan base levelDf = diameter of spudcan used to create footprint

Rd = radial distance

WhereLegends

0.4~0.5D

0.8~0.9de

1~1.1de

~0.2D

0.2~0.3D

1.501.251.000.750.500.250

Rd/Df

Crater

In normally consolidated clay the soil strength variation is the key factor influencing spudcan footprint interaction. The spudcan will experience relatively high horizontal and moment loadings below the previous rig’s penetration depth. Hence one may consider a rig with lower required preload pressure where the rig installation can be terminated at a sufficient distance above the previous penetration depth.

In over consolidated clay the physical profile of the depression can be the dominant factor in spudcan-footprint interaction. Hence the spudcan will likely experience high horizontal and moment loadings above the previous penetration depth. If these forces are found to be hazardous to the new rig installation, avoiding the footprints should be considered or use an identical rig to re-install at the same position of the previous rig as recommended in SNAME 5-5A. If the spudcan-footprint interaction is unavoidable due to site constraints, one may consider to carefully position the new rig in such a way that the critical center to center spacing between the initial penetration and re-penetration of 0.5 to 1.0 times spudcan diameter can be avoided. If that fails an effective mitigation measure to overcome the spudcan-footprint interaction should be pursued.


FIGURE 3 Simplified Diagram of Probable Soil Failure Mechanisms

During Penetration at 0.5D from Footprint Center

MbearingHbearing

θ

Rse

Model Ground

FootprintCL

(a) Shallow Penetration

Hslide

Rs1e1

Model GroundFootprintCL

MbearingHbearing

Mslide

Sliding Soil Mass Sliding Failure Plane

(b) Deep Penetration

5 Mitigation Methods The Rack Phase Difference (RPD) system is an effective way to monitor the sliding of a spudcan into an old footprint and hence prevent the leg damage [Ref.35]. RPD is the difference in elevation between the chords of any one leg, and it is a direct measure of the inclination of the leg with respect to the hull. RPD monitoring during the jacking process enables the operator to monitor that the RPD limits are not exceeded, and to take necessary steps otherwise by stopping the jacking operation. Thus monitoring gives real time control of the operation.

Infilling of old footprints formed in a coarse grained sea bed with the granular material has been carried out successfully, and it is well understood that infilling will not pose problems if the material characteristics are similar. The studies confirmed that the difference in the stiffness between the footprint and the infill material greatly affects the effectiveness of infilling and the infilling of footprints in layered cohesive materials, and therefore is not an advisable solution for the footprint interaction problem [Ref.36].

In firm to stiff clay, the physical profile of the depression whose profile is steep and deep is significant in influencing spudcan footprint interaction. To minimize the effect of a footprint, evening out the depression by excavation might be done. Leveling the seabed profile is unlikely to be effective in soft clay as the critical horizontal and moment loadings are due to the variation of soil strength well below the crater depth.



In soft to stiff clay, “stomping” is reportedly very effective in mitigating spudcan footprint interaction. It is a process where the footings are initially emplaced further away from the old footprint center and then used to displace soil towards the old footprint at desired positions to widen the disturbed regions by controlled additional disturbance (Section 8, Figure 4). Reaming, also known as leg working or leg reciprocation is another possible mitigation method, although its effectiveness is not as good as stomping. To be effective, reaming should be executed with small penetration-extraction increments [Ref.37]. The major limitations for these options are the significant rig time involved, and careful planning is also needed for the program. This may only be possible under mild environmental conditions, and the operation also depends on the clearance between the jackup and the adjacent fixed structure.

FIGURE 4 Stomping Process

After First Stomp Before Second Stomp


A p p e n d i x 1 : R e f e r e n c e s

A P P E N D I X 1 References

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2. Jamiolkowski, M.B., Lo Presti, D.C.F. and Manassero, M. (2003) Evaluation of relative density and shear strength of sands from cone penetration test (CPT) and flat dilatometer (DMT), Soil Behaviour and Soft Ground Construction, Eds. J.T. Germain, T.C. Sheahan and R.V. Whitman, ASCE, GSP 119, 201-238.

3. Bolton, M.D. (1986). The strength and dilatancy of sands. Géotechnique, 36 (1), 65-78.

4. Randolph, M.F., Jamiolkowski, M.B. and Zdravkovi, L. (2004). Load carrying capacity of foundations. Proc. Skempton Memorial Conf., London, Vol. 1, 207-240.

5. SNAME (Society of Naval Architects and Marine Engineers). (2008). Recommended practice for site specific assessment of mobile jack-up units, SNAME Technical and Research Bulletin 5-5A, Rev. 3, New Jersey.

6. Yi J.T., Lee F.H., Goh S.H., Zhang X.Y. and Wu J.F. (2011). Eulerian Finite Element Analysis of Excess Pore Pressure Generated by Spudcan Installation into Soft Clay. Computers and Geotechnics.

7. Hossain, M.S., Randolph, M.F., Hu, Y., White, D.J. (2006). Cavity Stability and Bearing Capacity of Spudcan Foundations on Clay. OTC 17770.

8. Houlsby, G.T and Martin, C.M. (2003). Undrained bearing capacity factors for conical footings on clay. Géotechnique, 53 (5), 513-520.

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15. Zhang X.Y., Li Y.P., Yi J.T., Lee F. H., Tan P.L., Wu J.F. & Wang S. Q., (2015). A Novel Spudcan Design to Enhance Foundation Performance. ISFOG 2015, Norway.

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Appendix 1 References

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22. Hossain, M.S. and Randolph, M.F. (2010). Deep-penetrating spudcan foundations on layered clays: numerical analysis. Géotechnique, 60(3), 171-184.

23. ABS Rules for Building and Classing Mobile Offshore Drilling Units (2014). Part 3 Hull Construction and Equipment.

24. Temperton I, Stonor RWP & Springett CN. (1999). Measured spudcan fixity: analysis of instrumentation data from three North Sea jack-up units and correlation to site assessment procedures, Marine Structure 12: 277-309.

25. Bell R.W. (1991). The Analysis of Offshore Foundations Subjected to Combined Loading. MSc. Thesis presented to the University of Oxford.

26. Noble Denton Europe and Oxford University (2005). The Calibration of SNAME Spudcan Footing Equations with Field Data. Report No L19073/NDE/mjrh, Rev.4, dated 21st Nov. 2005.

27. Templeton, J.S. (2009). Spudcan Fixity in Clay, Further Results from a Study for IADC. Proc. 12th International Conference, The Jack-Up Platform, City University, London. 2009.

28. Lee F.H., (2013). Final Report for Research Project: Spudcan Fixity for Realistic Design of jack-Up Units. Project No. 082 135 0042. National University of Singapore.

29. Mirza, U.A., Sweeney, M. and Dean A.R. (1988). Potential effects of jack-up spud can penetration on jacket piles. Offshore Technology Conference. OTC 5762, 147-157.

30. Leung C.F., Chow Y.K., Somsak, S., Tho K.K. & Xie Y. (2009). Spudcan-Pile Interaction Joint Industry Project. National University of Singapore.

31. Leung C.F., Tho K.K., Chow Y.K., Xie Y, Wong P.C. & Purwana O.A. (2012). Experimental and Numerical Studies of Spudcan-Pile Interaction. OTC 23053.Houston, Texas, USA.

32. Xie Y. (2009). Centrifuge Model Study on Spudcan-Pile Interaction. Ph.D Thesis. National University of Singapore.

33. Gan C.T. (2009). Centrifuge Model Study on Spudcan-Footprint Interaction. Ph.D Thesis. National University of Singapore.

34. Stewart, D.P. and Finnie, I.M.S. (2001). Spudcan-footprint interaction during jack-up workovers. International Society of Offshore and Polar Engineers (ISOPE), Cupertino, California. Vol. 1. pp. 61-65.

35. Foo, K.S., Quah, M.C.K., Wildberger, P. and Vazquez, J.H. (2003). Spudcan footprint interaction and rack phase difference. Proc. 9th Int. Conf. Jackup Platform Design, Construction and Operation, City University, London, UK

36. Jardine, R.J., Kovacevic, N., Hoyle, M.J.R., Sidhu, H.K. and Letty, A. (2001). A study of eccentric jack-up penetration into infilled footprint craters. Proc. 8th International Conference the Jack-up Platform, eds: C.D’Mello and L. Boswell, City University, London.

37. Hartono, C.F. Leung, K.K. Tho and Y.K. Chow. (2013). Reaming as a mitigation measure for jackup reinstallation close to existing footprint. International Conference on the Jack-up Platform, City University, London, UK. 17th & 18th Sep. 2013.


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