DnV-RP-E302 - Design and Installation of Plate Anchors in Clay [2000]

DET NORSKE VERITAS

RECOMMENDED PRACTICERP-E302

DESIGN AND INSTALLATION OFDRAG-IN PLATE ANCHORS

IN CLAY

2000

FOREWORD

DET NORSKE VERITAS (DNV) is an autonomousand independent Foundation with the objectives ofsafeguarding life, property and the environment, atsea and onshore. DNV undertakes classification,certification, and other verification and consultancyservices relating to quality of ships, offshore unitsand installations, and onshore industries worldwide,and carries out research in relation to these functions.DNV publishes various documents related to theoffshore industry, aimed at promoting quality andsafety on offshore units and installations.

The Recommended Practice publications (RP-series)cover proven technology and solutions which havebeen found by DNV to represent good practice, andwhich represent one alternative for satisfying therequirements stipulated in the DNV OffshoreStandards or other codes and standards cited byDNV. The DNV RP-series is divided into 6 parts, asfollows.

A. Quality and Safety MethodologyB. Materials TechnologyC. StructuresD. SystemsE. Special FacilitiesF. Pipelines & Risers

As well as forming the technical basis for DNVverification services, the Offshore Standards andRecommended Practices are offered as DNV’sinterpretation of safe engineering practice for generaluse by the offshore industry.

ACKNOWLEDGEMENTSThis Recommended Practice is based upon a designprocedure developed within the Joint Industry Project

"Design Procedures for Deep Water Anchors" /1/, /2/and /4/. The following companies sponsored this JIP:

BP Exploration Operating Company Ltd.; BruceAnchor Ltd.; Det Norske Veritas; Health & SafetyExecutive; Minerals Management Service; NorskHydro ASA; Norske Conoco AS; Petrobras; SagaPetroleum ASA; Shell Internationale PetroleumMaatschappij B.V. (Part 1 only); SOFEC Inc.(Part 1 only); and Statoil.

DNV is grateful for valuable co-operations anddiscussions with these companies. Their individualsare hereby acknowledged for their contribution.

As part of the publication of this RP, a draft copy wassent for hearing to several companies. Significant,valuable and concrete comments were providedwithin the resulting feedback. The followingorganisations, which actively participated, arespecially acknowledged:

Elf Exploration Production; and Naval FacilitiesEngineering Service Center of US Navy.

THIS REVISIONThis revision of RP-E302 is made in order toharmonise the document with other relateddocuments. The only significant revision in thisdocument relates to the partial safety factors on linetension for the ALS case.

Comments may be sent by e-mail to [email protected] . For subscription orders or information about subscription terms, please [email protected] .

Comprehensive information regarding DNV services, research and publications can be found at http://www.dnv.com , or can be obtained fromDNV, Veritasveien 1, N-1322 Høvik, Norway; Tel +47 67 57 99 00, Fax +47 67 57 99 11.

© 2000 DET NORSKE VERITAS. All rights reserved. No part of this publication may be reproduced or transmitted in any form or by anymeans, including photocopying and recording, without the prior written consent of DET NORSKE VERITAS.

Printed in Norway by Det Norske Veritas AS

DET NORSKE VERITAS

CONTENTS1. General ..............................................................................11.1 Introduction....................................................................... 11.2 Objective ............................................................................ 11.3 Scope of application......................................................... 11.4 Structure of the RP........................................................... 11.5 Abbreviations.................................................................... 11.6 Symbols and explanation of terms ................................. 12. Drag-in plate anchor components..............................43. General behaviour of drag-in plate anchors ...........43.1 Introduction....................................................................... 43.2 Penetration phase.............................................................. 53.3 Principles of the Stevmanta anchor............................... 63.4 Principles of the Denla anchor....................................... 64. Methodology for design of drag-in plate anchors..74.1 General................................................................................ 74.2 Analytical tools ................................................................. 75. Recommended design procedure................................75.1 General................................................................................ 75.2 Contributions to anchor resistance................................ 85.3 Step-by-step description of procedure ........................ 115.4 Tentative safety requirements ...................................... 115.5 Installation measurements............................................. 146. Requirements to soil investigation ...........................147. References.......................................................................15

Appendix A: Analysis tool for fluke anchor designAppendix B: Drag-in plate anchors in layered clayAppendix C: Installation and testing of drag-in plate anchors

Appendix D: Consolidation effectAppendix E: Cyclic loading effectsAppendix F: Uplift angle at the seabed

Appendix G: General requirements to soil investigation

Recommended Practice No. RP-E302 1May 2000

DET NORSKE VERITAS

1. General

1.1 IntroductionThis Recommended Practice features a substantial part of thedesign procedure developed in Part 2 /2/ of the joint industryproject on Design procedures for deep water anchors. Animportant basis for the recommended design procedure wasestablished in Part 1 /1/ of the JIP, which lead to a DNVRecommended Practice for design and installation of flukeanchors in clay /3/ . An overview of the JIP is given in /4/.

1.2 ObjectiveThe objective is that this Recommended Practice shall berecognised as an international guideline for geotechnicaldesign and installation of drag-in plate anchors in clay.

1.3 Scope of applicationThis Recommended Practice applies to the geotechnicaldesign and installation of drag-in plate anchors in clay fortaut mooring systems (TMS).

The recommendations herein are in principle applicable toboth long term (permanent) and temporary mooring.

The design procedure outlined is a recipe for how drag-inplate anchors in clay can be designed to satisfy therequirements by DNV.

Until the design rule presented herein has been calibratedbased on reliability analysis, the partial safety factors will betentative.

1.4 Structure of the RPDefinition of the main components of a drag-in plate anchoris given in Chapter 2, followed by a description of thegeneral behaviour of drag-in plate anchors in clay in Chapter3.

In Chapter 4 a design methodology is outlined, which isbased on the use of a calibrated and validated analytical tool.

The recommended procedure for design and installation ofdrag-in plate anchors is presented in Chapter 5. The closeand important relationship between the assumptions fordesign and the consequential requirements for the installationof drag-in plate anchors is emphasised.

General requirements to soil investigations are given inChapter 6.

The intention has been to make the procedure as concise aspossible, but still detailed enough to avoid misinterpretationor misuse. For transparency, details related to the variousdesign aspects are therefore given in the appendices.

A number of Guidance notes have been included as an aidin modelling of the anchor line, the anchor and the soil. Theguidance notes have been written on the basis of theexperience gained through the mentioned joint industryproject, see /1/ and /2/.

1.5 AbbreviationsAHV Anchor Handling Vessel

ALS Accidental damage Limit State

CAU Anisotropically Consolidated Undrained

DSS Direct Simple Shear

BTL Break Test Load

ULS Ultimate Limit State

UU Unconsolidated Undrained

1.6 Symbols and explanation of termsSymbol Term Explanation of term

α Uplift angle Line angle with the horizontalat the dip-down point

αi Installationuplift angle

Uplift angle during anchorinstallation

αmax Maximumpossibleuplift angle

Installation uplift angle, whichmakes the anchor drag atconstant tension withoutfurther penetration at the actualdepth

α Anchoradhesionfactor

Accounts for remoulding ofthe clay in the calculation ofthe sliding resistance at theanchor members

αmin Minimumadhesion

Set equal to the inverse of thesensitivity, αmin = 1/St

AF Anchorfluke aspectratio

Typically AF ≈ 1 for theanchors considered herein

αsoil Lineadhesionfactor

Used to calculate unit frictionin clay of embedded anchorline

Afluke Anchorfluke area

Based on manufacturer's datasheet.

β Anchorpenetrationdirection

Angle of the fluke plane withthe horizontal

AB Effectivebearing area

Related to anchor line segmentin the soil

AS Effectivesurface area

Related to anchor line segmentin the soil

d Nominaldiameter

Diameter of wire, rope orchain

ds Elementlength

Related to embedded anchorline

e Lever arm Between shackle and the lineof action of the resultingnormal resistance at the fluke

η Empiricalfactor

Reduction factor, related to Nc,derived from field tests

2 Recommended Practice No. RP-E302May 2000

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Symbol Term Explanation of term

f Unit friction Sliding resistance, bothfrictional and cohesive, ofembedded part of anchor lineand anchor members

Φ Mooringline upliftangle

Angle between mooring lineand seabed

γm,1 Partialsafety factoron RS(zi)

Accounts for the uncertainty insu(zi) and su,r(zi) as it affectsRS(zi), Ur and reference strainrate vref, the prediction methodand the analytical model

γm,2 Partialsafety factoron ∆Rcy(zi)

Accounts for the uncertainty insu(zi) as it affects ∆Rcy(zi),thecyclic test data used and Ucy,the prediction method and theanalytical model

γm,i Partialsafety factoroninstallationseabedfriction

Accounts for the uncertainty inthe predicted seabed friction tobe overcome during anchorinstallation

γmean Partialsafety factoron TC-mean

Accounts for the uncertainty inthe mean line tension

γdyn Partialsafety factoron TC-dyn

Accounts for the uncertainty inthe dynamic line tension

κ Anchorgeometryfactor

Related to WF,typically κ ≈ 1.00

k Undrainedshearstrengthgradient

Average gradient betweenseabed intercept su,0 and shearstrength at installation depthsu(zi)

k c Empiricalfactor

Used to estimate the cyclicdegradation effect

λ Ultimatedepth factor

Relates to zult, function of soilstrength, forerunner, etc.,typically λ ≈ 7-9

Ls,i Line lengthon seabed atanchorinstallation

For the anchor installation lineconfiguration and theinstallation tension Tmin

µ Coefficientof seabedfriction

Average friction coefficient(also of cohesive nature) overline length Ls,i

M Relativedepth ofpenetration

M = zi/zult (applicable to claywith constant gradient k only)

n Exponent Used in empirical formula forloading rate effect


Nc Bearingcapacityfactor forclay

Corrected as relevant for theeffects of shape, orientationand embedment of respectiveresistance members

Neqv Equivalentnumber ofcycles tofailure

The number of cycles at theconstant cyclic shear stress thatwill give the same effect as theactual cyclic load history

OCR Overconsoli-dation ratio

Ratio between maximum pastand present effective verticalstress on a soil element

q Normalstress

Used in equilibrium equationsfor anchor and embeddedanchor line

θ Orientationof anchorline element

θ = 0 for a horizontal element

Q1, Q2 Pileresistance

Pile resistance at loading ratesv1 and v2, respectively

R Anchorresistance

Resistance in the line directionin the line dip-down point withreference to the anchorpenetration depth z

RC Characteris-tic anchorpulloutresistance

RC = RS(zi) + ∆Rcy(zi)

Rd Designanchorpulloutresistance

Rd = RS(zi)/γm,1 + ∆Rcy(zi)/γm,2

∆Rcons Consolida-tion effect

At restart after stoppage duringanchor installation

Rcons Consolidated anchorresistance

Anchor installation resistanceincluding ∆Rcons at restart afterstoppage.

Rp,cr Creeppulloutresistance

Threshold for development ofsignificant creep, whichdepends on actual line tensionhistory, duration of operation,soil characteristics, etc.

Rp,i Anchorinstallationpulloutresistance

The pulloutresistance'immediately' afteranchor installation andtriggering in the dip-downpoint

Rp,u Ultimatepulloutresistance

The resistance at ultimatepenetration depth zult.

∆Rcy Cyclicloadingeffect

Depends on extreme linetension history and soilcharacteristics, added to RS

Rp,cy Cyclicpulloutresistance

Pullout resistance including∆Rcy (Rp,cy = RC)


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RS Staticpulloutresistance

Installation pullout resistanceRp,i corrected for loading rateeffect

Ri Installationanchorresistance

May be set equal to Ti, if Ti isproperly verified anddocumented at installation

Rult Ultimateanchorresistance

The resistance reached at z =zult, the anchor drags withoutfurther increase in theresistance during continuouspulling

Rai Sum of soilresistance atanchorcomponents

Excluding soil resistance at thefluke

RFN Soil normalresistance

At the fluke

RFS Soil slidingresistance

At the fluke

Rmai Momentcontribution

From Rai

RmFS Momentcontribution

From RFS

RTIP Tipresistance

At anchor members

ρ Creep factor Related to Rp,cr

sc Shape factor Related to Nc

St Soilsensitivity

The ratio between su and su,r,as determined e.g. by fall-coneor UU triaxial tests

su Intactstrength

For drag-in plate anchoranalysis the DSS strength orthe UU triaxial strength isassumed to be mostrepresentative

su,r Remouldedshearstrength

The undrained shear strengthmeasured e.g. in a fall-cone ora UU triaxial test after havingremoulded the clay completely

τf,cy Cyclic shearstrength

Accounts for both loading rateand cyclic degradation effectson su.

tcons Consolida-tion time

Time elapsed from anchorinstallation to time of loading

tcy Time tofailure

Rise time of line tension frommean to peak level during thedesign storm (= 1/4 load cycleperiod)

thold Installationtensionholdingperiod

Period of holding Tmin at theend of anchor installation


tsu Time tofailure

Time to failure in a laboratorytest for determination of theintact undrained shear strength(typically 0.5 − 2 hours)

T Line tension Line tension model followingsuggestion in /5/

Tv, Th Componentsof linetension atthe shackle

Vertical and horizontalcomponent of the line tensionat the anchor shackle for theactual anchor and forerunner

TC Characteris-tic linetension

Split into a mean and dynamiccomponent, TC-mean andTC-dyn

TC-mean Characteris-tic meanline tension

Due to pretension and theeffect of mean environmentalloads in the environmentalstate

TC-dyn Characteris-tic dynamicline tension

The increase in tension due tooscillatory low-frequency andwave-frequency effects

Td Design linetension

With specified partial safetyfactors included

Td-mean Designmean linetension

= TC-mean ⋅ γmean

Td-dyn Designdynamicline tension

= TC-dyn ⋅ γdyn

Ti Targetinstallationtension

Specified installation tensionat the dip-down point, seeFigure 4

Tmin Minimuminstallationtension

Specified installation tensionat the touch-down point (if Ls,i

> 0), for Ls,i = 0, Tmin = Ti

∆Tmin Oscillationofinstallationtension

Double amplitude tensionoscillation around Tmin duringperiod thold

Tpre Pretensionin mooringline

As specified for the mooringsystem

Ucons Soilconsolida-tion factor

Ucons = (1+∆Rcons/Ri), whereratio ∆Rcons/Ri expresses theeffect of consolidation on Ri

Ucy Cyclicloadingfactor

Ucy = (1+∆Rcy/RS), where ratio∆Rcy/RS expresses the effect ofloading rate and cyclicdegradation on RS

Ur Loading ratefactor

Ur = (vi/v2)n

Ur = (v/vref)n

v Loading (orstrain) rate

Actual pullout rate

v1 Loading (orstrain) rate

Actual installation rate


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v2 Loading (orstrain) rate

Reference rate at the end ofinstallation

vref Loading (orstrain) rate

Reference rate for pulloutresistance

WF Equivalentfluke width

Equal to κ* (Afluke)0.5,

typically κ=1.00Wa' Submerged

anchorweight

Taken as 0.87 ⋅ anchor weightin air

Wm Momentcontribution

From anchor weight W

Wl' Submergedweight ofanchor line

Per unit length of actual linesegment

ψ Inverseloading ratefactor

Adjustment for strain rateeffect on the pullout resistanceRp,i measured in an offshoretest when calculating RS

z Penetrationdepth

Depth below seabed of thefluke tip.

zi Installationpenetrationdepth

For R = Ri (at end ofpenetration).For the final normal loadposition, zi refers to flukecentre.

zlab.test Depthreflected bylab. tests

Refers to cyclic soil test datagiven in Appendix E

zmin Minimumpenetrationdepth

To ensure deep embedment(deep failure)

zult Ultimatepenetrationdepth

Relates to Rult, and isexpressed in number of flukewidths, λ⋅WF

2. Drag-in plate anchor componentsThe main components of a drag-in plate anchor (Figure 1)are:

− the shank− the fluke− the shackle− the forerunner

Shackle

Shank

Fluke (plate)

Normalloading

Flukeangle

Installation

ForerunnerInver

se catenary

Figure 1 Main components of a drag-in plate anchor.

The fluke angle is the angle arbitrarily defined by the flukeplane and a line passing through the rear of the fluke and theanchor shackle. It is important to have a clear definition(although arbitrary) of how the fluke angle is beingmeasured.

Normally the fluke angle is fixed within the range 30° to 50°,the lower angle used for sand and hard/stiff clay, the higherfor soft normally consolidated clays. Intermediate anglesmay be more appropriate for certain soil conditions (layeredsoils, e.g. stiff clay above softer clay). The advantage ofusing the larger angle in soft normally consolidated clay isthat the anchor penetrates deeper, where the soil strength andthe normal component on the fluke is higher, giving anincreased resistance.

The forerunner is the line segment attached to the anchorshackle, which will embed together with the anchor duringinstallation. The anchor penetration path and the ultimatedepth/resistance of the anchor are significantly affected bythe type (wire or chain) and size of the forerunner, seeFigure 2.

The inverse catenary of the anchor line is the curvature ofthe embedded part of the anchor line, see Figure 1.

3. General behaviour of drag-in plateanchors

3.1 IntroductionDrag-in plate anchors have been developed for use incombination with taut mooring systems (TMS) and they canresist both the vertical and the horizontal loads transferred tothe anchors in such a system. This anchor is installed as aconventional fluke anchor, see /3/, and when the targetinstallation load Ti has been reached it is triggered to createnormal loading against the fluke (plate). In this normalloading mode the anchor acts as an embedded plate with ahigh pullout resistance Rp.

Examples of drag-in plate anchors are the Denla anchor /7/and the Stevmanta anchor /8/. Use of the Bruce AnchorTracker on both types of anchor, see description of trackerand testing experience in /9/ and /10/, has significantly


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R(z)

Rult=R(zult)

Z

zult

θ

T

Th

Tv

Wire forerunner

(with wire)

T(z)

z

αi

Pen

etra

tion

dept

h

Drag

Chain forerunner

zult

(with chain)

Figure 2 Installation behaviour of drag-in plate anchors and definition of Rult.

improved the quality of the data base for drag-in plateanchors.

3.2 Penetration phaseThe pullout resistance of a drag-in plate anchor depends onthe ability of the anchor to penetrate and to reach the targetinstallation tension (Ti).

The penetration path for a drag-in plate anchor and theultimate penetration depth zult depends on

− the soil conditions (soil layering, variation in intactand remoulded undrained shear strength)

− the type and size of anchor,− the anchor’s fluke angle,− the type and size of the anchor forerunner (chain or

wire and nominal diameter), and− the installation uplift angle αi at the seabed level.

The predicted ultimate penetration zult of the anchor iscrucial for sizing the anchor, given Ti and the shearstrength profile.

In clay without significant layering a fluke anchornormally penetrates along a path, where the ratio betweenincremental penetration and drag decreases with depth, seeFigure 2 with the Denla anchor as example. At the ultimatepenetration depth zult the anchor is not penetrating anyfurther. The anchor is “dragging” with a near horizontalfluke, and the tension in the line is constant. At theultimate penetration depth the anchor reaches its ultimatepenetration resistance Rult. It is important not tooverestimate zult. In the worst case the target installationtension Ti will not be reached before the anchor startsdragging without further increase in the anchor resistance.To avoid this, the design (sizing) of the anchor shouldassume that the anchor is installed to a depth zi, which issignificantly less than zult.

An additional requirement is that the minimum penetrationdepth zmin should be 4.5 fluke widths (WF) to ensure thatthe boundary conditions for assuming deep failure aresatisfied in the computation of the anchor pulloutresistance.

It is recommended as a target that the relative depth ofanchor penetration M = zi/zult is in the range 0.50 to 0.80for drag-in plate anchors installed in normally consolidatedclay, with 0.75 as a tentative default value. This will alsolead to more predictable drag.


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The cutting resistance of a chain forerunner will be greaterthan the resistance of a steel wire, with the result that achain forerunner will have a steeper curvature (inversecatenary) at the anchor shackle than a wire forerunner, i.e.the angle θ at the shackle is larger, see Figure 2. Thisincreases the upward vertical component Tv of the linetension T at the shackle with the consequence that a drag-in plate anchor with a chain forerunner penetrates less thanone with a wire forerunner, and mobilises less resistancefor a given drag distance. Increasing the diameter of a steelwire forerunner will also reduce the anchor penetration forthe same reason.

It has been demonstrated in /1/ that a non-zero installationuplift angle αi, see Figure 2, can be acceptable undercertain conditions as discussed in Appendix F and appliedin /3/. Drag-in plate anchors in deep water will normallybe installed under an uplift angle, in which case there willbe no line on the seabed. A high uplift angle reduces thecontribution from the anchor line to the anchor resistanceR(z), which leads toan increased penetration of the anchorfor a given installation tension Tmin. If the uplift anglebecomes excessive during installation the ultimatepenetration depth may, however, be reduced, see AppendixF.

3.3 Principles of the Stevmanta anchorThe Stevmanta anchor has no rigid shank, but a system ofwires connected to a fixed plate angle adjuster. Forillustration of the principles of the Stevmanta anchor, thedouble-line installation method as illustrated in Figure 3will be used, see also /8/. The installation line is attachedto the front-shackle of the angle adjuster, whereas theactual mooring line is attached to the back-shackle.

When pulling in the installation line, normally towards thecentre of the mooring pattern, the fluke angle adjusts to theangle set for deep penetration in the actual soil conditions.This angle is determined by the length of the wiresattached to the front of the fluke relative to the length ofthe wires attached to the rear of the fluke.

When pulling in the mooring line the angle adjuster willrotate and create the normal loading mode and themaximum pullout resistance for the actual fluke area andsoil strength at the depth of penetration.

When the anchor is installed by the double-line installationmethod the anchor is retrieved by again pulling in theinstallation line as shown in Figure 3

The performance ratio Pr expressing the relationshipbetween the pullout resistance Rp in the loading directionnormal to the fluke plane and the target installation tensionTi is explained in Figure 5 with the Denla anchor as anexample.

Figure 3 Installation, normal loading and retrieval ofthe Stevmanta anchor.

3.4 Principles of the Denla anchorThe principles for installing the Denla anchor andpreparing it for hook-up to the floater are illustrated inFigure 4 and Figure 5, see also Figure 1.

The anchor handling vessel (AHV), while steaming awayfrom the centre of the mooring pattern, installs the Denlato the target installation tension Ti. Then the AHV steamsback, towards the centre of the mooring pattern, over theburied Denla, applying a vertical/backward pull on theshank.

The shear pin controlling the installation fluke anglebreaks at a certain predefined load and the shank rotatesand locks in a position, which creates normal loading onthe fluke (plate). The AHV then continues steamingtowards the centre and rotates the anchor into a positionwhich gives approximately normal loading for thespecified mooring line angle Φ with the seabed, seeFigure 5b).


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4. Methodology for design of drag-inplate anchors

4.1 GeneralThe methods used to design drag-in plate anchors have sofar been highly empirical, using the experience from smallto full-scale offshore tests, see /7/, /8/, and /9/. In mostcases there are significant uncertainties attached to thereported data from offshore tests, e.g. soil stratigraphy, soilstrengths, anchor installation tension, possible contributionfrom sliding resistance along the anchor line segment onthe seabed, depth of anchor penetration, possible effect ofanchor roll during penetration, etc. Well instrumentedonshore tests, as reported in /10/, have given moreconclusive results, and together with the best documentedoffshore tests they have provided the experimentalevidence behind the recommended design procedureoutlined in Chapter 5.

The introduction of the anchor tracker system as describedin /9/ and /10/ has improved the quality of the test datasignificantly. To further improve the anchor test database,and as a means to verify the design of commercial anchors,measurements during anchor installation and testingshould be carried out on a routine basis. For planning andexecution of such measurements, see guidance inAppendix C.

A drag-in plate anchor, in its intended operational mode,orients itself such that the fluke plane (plate) is normal tothe direction of line tension at the anchor shackle. Sincethe soil disturbance during penetration of the anchor iscaused by penetration in a direction parallel to the flukeplane this disturbance will have only marginal effect onthe pullout resistance. It is therefore conservativelyassumed in the present design methodology that anyconsolidation effects on the pullout resistance can bedisregarded, although they should be considered for thepenetration phase, see Section 5.5 and Appendix D.

Sound engineering judgement should always be exercisedin the assessment of the characteristic resistance of thechosen anchor, giving due consideration to the reliabilityof the analytical tool and the uncertainty in the designparameters provided for the site. Extrapolation from testresults with small to medium size anchors to prototype sizeanchors should be made with due consideration of possiblescale effects. The development of design methods basedon theoretical models and geotechnical principles shouldbe encouraged. Based on the experience gathered throughthe JIP on deepwater anchors /1/ and /2/, requirements toanalytical tools for design of drag-in plate anchors arepresented below.

4.2 Analytical tools

4.2.1 General

The analytical tool should be based on geotechnicalprinciples, be calibrated against high quality anchor tests,and validated.

With an analytical tool the designer should be able tocalculate:

− the relationship between line tension, anchorpenetration depth and drag for the actual anchor andline configuration in the prevailing soil conditions

− how this relationship is affected by changing the typeand/or size of the anchor, the type and/or size of theforerunner, or the soil conditions

− the effect of soil consolidation at restart after stoppageduring anchor installation, see guidance in AppendixD

− the effects of cyclic loading in terms of loading rateand cyclic degradation, see guidance in Appendix E,which also discusses in some detail the basis for theproposed static pullout resistance RS and the creeppullout resistance Rp,cr

− the effect on the penetration trajectory of changing theuplift angle at the seabed, see guidance in Appendix F.

The analytical tool must satisfy the equilibrium equationsboth for the embedded anchor line and for the drag-in plateanchor.

4.2.2 Equilibrium equations for drag-in plateanchor analysis

The inverse catenary of the embedded anchor line duringanchor installation is resolved iteratively such thatequilibrium is obtained between the applied line tensionand the resistance from the surrounding soil, see /11/. Forthe drag-in plate anchor both force and momentequilibrium is sought. The equilibrium equations for theanchor line and the anchor as included in an analytical tooldeveloped by DNV are given in Appendix A.

5. Recommended design procedure

5.1 GeneralDrag-in plate anchors should be designed based ongeotechnical calculations using a suitable analytical tool,as discussed in Section 4.2. A condensed step-wisedescription of the recommended procedure is given inSection 5.3.2. The basic nomenclature used in theprocedure and the main contributions to the anchorresistance both during installation and pullout aredescribed in Section 5.2. Further details about the anchorresistance contributions and calculation guidance are givenin the Appendices.

As an alternative, or as a supplement, to the recommendedprocedure, the design may be based on the results fromsite-specific drag-in plate anchor tests. This alternativedesign procedure is described in Section 5.3.3

In the description of the two alternatives it is assumed thatthe anchor installation line intersects the seabed under apositive uplift angle towards the final stage of installation,thus no need to correct for line seabed friction duringinstallation. If this assumption is invalidated, then theprocedure should be corrected to include these frictionforces, see Eq. (13).


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5.2 Contributions to anchor resistance

5.2.1 General

In the following, the basic contributions to the resistanceof drag-in plate anchors are presented and explained. Theparameters involved will be described as they appear inrelation to anchor installation, anchor performance ratio,and anchor design.

5.2.2 Anchor installation

The anchor should be installed by continuous pulling untilthe minimum installation tension Tmin has been reached.Stoppage of the anchor installation at a smaller line tensionshould be avoided, since reconsolidation of the remouldedclay around the anchor during the stoppage period mayincrease the penetration resistance such that the necessarypulling force to restart the anchor exceeds the capacity ofthe available installation equipment. The penetration depthof the anchor will then be limited to the depth reached withthe installation tension applied before the stoppage. Sincethe anchor pullout resistance, see Section 5.2.4.2, isdirectly related to the undrained shear strength at theinstallation depth, the reduced penetration depth will alsoreduce the safety of that particular anchor point such thatthe specified safety level cannot be achieved. Measuresshould be taken to avoid this situation in the planning andexecution of the anchor installation, see further AppendixC.

Figure 4a) illustrates the anchor installation phase for thecase without uplift (αi = 0, Ls,i > 0). The installation anchorresistance Ri(zi) in the dip-down point is assumed to beequal to the target installation line tension Ti(zi), bothreferring to the dip-down point and anchor penetrationdepth zi. The minimum installation tension Tmin, whichshall be measured and documented as discussed in Section5.5 and Appendix C, has to exceed Ti(zi) by the seabedfriction developed over the length Ls,i between the dip-down point and the touch-down point. The anchor isassumed to penetrate to a depth zi if Ti is mobilised in thedip-down point and held for a specified period of time thold,i.e.

( ) islii LWTzT ,'

min ⋅⋅−= µ (1)

where

µ = coefficient of seabed friction (also cohesivein nature)

Wl' = submerged weight of anchor line per unitlength

If Ls,i = 0, see Figure 4b), then Ri(zi) = Ti(zi) = Tmin.

The installation anchor resistance Ri(zi) is thus dependenton the correct assessment of length Ls,i and the resultingseabed friction.

Figure 4b) illustrates a situation where the anchor lineintersects the seabed in the dip-down point under an upliftangle (αi > 0, Ls,i = 0) during the final stage of anchorinstallation.

Ti(zi)

Ri(zi)

Tmin

Dip-down Touch-down

a) At installation:(no uplift)

Ls,i

µ W l' Ls,i

z = zi

Pen

etra

tion

dept

h

z

Drag

Ti (zi)= Tmin

Ri(zi)

Dip-down = Touch-down

b) At installation:(uplift angle αι)

αι

z

Drag

z = zi

Pen

etra

tion

dept

h

Figure 4 Basic nomenclature (1 of 2).

Ti (ztest )

1. Installation

Rpi (ztest)

3. Pull-out test

2. Triggering

Dip-down angle αι

Performance Ratio: P r = Rp,i (ztest)/ Ti(ztest)

ztest

a) Anchor test:

Φ test

Ti (zi )

Rd(zi) = RS(zi )/γm,1 + ∆Rcy(zi) /γm,2

zi

Td = TC-mean*γmean + TC-dyn*γdyn

b) At operation:

Φ

Figure 5 Basic nomenclature (2 of 2)


DET NORSKE VERITAS

5.2.3 Anchor performance ratio Pr

The assessment of the target installation tension Ti is acrucial design issue, which is directly related to the anchorperformance ratio Pr, which is illustrated in Figure 5a).This ratio can be predicted using recognised geotechnicalprinciples combined with empirical correction factorsderived from offshore and onshore field anchor tests.

The sketch in Figure 5a) shows a typical anchor testsituation, where the anchor, in this case a Denla anchor, ispulled in to a penetration depth ztest applying a line tensionTi(ztest) and an uplift angle αi. After triggering of theanchor, a pullout test is performed, which gives theinstallation pullout resistance Rp,i(ztest). The performanceratio Pr is defined as

( )( )testi

testipr zT

zRP ,= (2)

Note in Figure 5a) that an offshore pullout test may beperformed by applying either a vertical or an inclinedpulling force in the anchor line. An offshore test isnormally carried out with inclined loading in order toreduce the cyclic fluctuations in the line tension due tovessel motions, whereas vertical loading can be used in anonshore test. Anchor installation and installationmeasurements are discussed in Appendix C.

It should be appreciated that both the installation resistanceRi and the installation pullout resistance Rp,i, whenmeasured in an offshore drag-in plate anchor test, areweather (loading rate) dependent, which in the end willaffect the back-calculated performance ratio Pr for theactual test. Loading rate effects on the anchor performanceare further discussed in Appendix C and Appendix E, seealso /10/.

Assume that results as given by Eq. (2) are available froman instrumented offshore test, that the undrained shearstrength increases linearly with depth, and that theperformance ratio Pr derived for depth ztest isrepresentative also for depth zi, then the installation pulloutresistance Rp,i at depth zi may be estimated as follows

( ) ( )testitestipiip zzzRzR /)( ,, ⋅= (3)

As discussed in Section 5.2.4.2, and further explained inAppendix E, the installation pullout resistance measured inan offshore test includes a loading rate effect, which isexpressed by a loading rate factor Ur. By dividing themeasured Rp,i(ztest) by this loading rate factor one wouldobtain the static pullout resistance RS(ztest) at the test depthztest.

( ) ( ) ( )testitestipiS zzzRzR /, ⋅⋅= ψ (4)

where

ψ = inverse loading rate factor (=1/Ur), whichaccounts for the strain rate effect on su, whencomparing an offshore pullout test with astatic pullout test, see discussion inAppendix E (typically ψ = 0.8 for soft clay)

The cyclic loading effect on the static undrained shearstrength of soft to lightly overconsolidated clay, andconsequently on the static pullout resistance of an anchorin that clay, may be expressed by a cyclic loading factorUcy. This gives the following expression for prediction ofthe target installation tension Ti(zi), which is relying on theanchor performance ratio Pr obtained from site specificanchor tests

( ) ( ) ( ) rtestitestipcyii PzzzRUzT //, ⋅⋅⋅= ψ (5)

where

Ucy = cyclic loading factor, see details in Section5.2.4.2 and Appendix E

On the basis of a careful interpretation of available sitespecific and/or other relevant anchor tests the performanceratio Pr for design of drag-in plate anchors at the actual sitemay be assessed. Given the characteristic line tensions forthe ULS and ALS conditions, a target installation linetension Ti(zi) can then be assessed, which satisfies thespecified safety requirements and can be accomplishedwith the available installation equipment.

5.2.4 Anchor design

5.2.4.1 Design line tension Td

The design line tension Td is made up of the design meanline tension Td-mean and the design dynamic line tension Td-

dyn, which are obtained by multiplying the characteristicmean tension TC-mean and the dynamic tension TC-dyn bytheir respective partial safety factors γmean and γdyn.

dyndynCmeanmeanCd TTT γγ ⋅+⋅= −−(6)

For details regarding the two line tension components inEq. (6), see /5/.

5.2.4.2 Design anchor pullout resistance Rd

Consolidation effects on the pullout resistance aredisregarded since the remoulding of the clay due to theanchor installation occurs in a narrow zone adjacent to thepenetrating members, whereas the clay providing the mainresistance to the pullout resistance is virtually unaffectedby the anchor installation. The anchor pullout resistance isthen split into a static component RS (zi) and a cycliccomponent ∆Rcy(zi), see Eqs. (7) and (8) and Figure 5b).

( ) ( )( ) flukeioucciS AzkssNzR ⋅+⋅⋅⋅= ,η (7)

( ) ( ) ( )1−⋅=∆ cyiSicy UzRzR (8)

where


DET NORSKE VERITAS

Nc = bearing capacity factor (Nc = 12.5, planestrain)

sc = shape factor (sc = 1.1 for assumed flukewidth/length ratio Af = 0.5)

η = empirical reduction factor from fieldtests, which accounts for effects of soilremoulding, strain softening, out-of-normal and eccentric loading of the plate(typically η = 0.73)

su,0 = undrained shear strength seabed intercept

k = undrained shear strength gradient(s)between seabed and installation depth zi

Afluke = fluke area of selected anchor type

As a 'first guess' the necessary fluke area of the selectedanchor type may be set equal to

u

dfluke s.

TA

⋅=

258(9)

where

su = best estimate static, undrained, DSS shearstrength at the assumed installation depth

If results from offshore pullout tests are available, thestatic pullout resistance RS(zi) may be assessed from Eq.(4).

The cyclic loading component ∆Rcy accounts for theeffects of cyclic loading on the static pullout resistanceRS(zi). ∆Rcy consists of two parts, one is the loading (orstrain) rate effect, and the other is the cyclic degradationeffect on the undrained shear strength of the clay. Thesetwo effects are linked together and may be expressedthrough the cyclic shear strength τf,cy, see Appendix E.

The expression for the design pullout resistance thebecomes

( ) ( ) ( )

( )

−+

⋅

=∆

+=

2,1,

2,1,

11

m

cy

miS

m

icy

m

iSid

UzR

zRzRzR

γγ

γγ(10)

where γm,1 and γm,2 are partial safety factors.

As mentioned before, the effects of consolidation shouldnormally be disregarded in the assessment of the pulloutresistance of drag-in plate anchors, but it should be notedthat there are long term consolidation effects, which maylead to an increase in the pullout resistance. Such long-term effects are dependent on the method of anchorinstallation and the amount of remoulding associated withrotation of the plate into its normal loading mode, inaddition to the soil characteristics. If considered at all,these effects therefore need to be evaluated on a case-by-case basis.

The basis for calculation of the effects of consolidation,cyclic loading and uplift at the seabed are discussed inAppendix D, Appendix E and Appendix F, respectively.Although creep is not foreseen to become a problem withdrag-in plate anchors designed to satisfy the ULS and ALSconditions and the design procedure recommended herein,this is a design issue that needs to be addressed. Tentativeguidance for addressing this issue is given in Appendix E.

Of particular importance for the design of drag-in plateanchors is a realistic assessment of the anchor penetrationtrajectory and the ultimate depth of penetration zult, seeillustration in Figure 2. The penetration trajectory assumedfor the anchor has a direct bearing on the undrained shearstrength that may be assumed in the calculation of theinstallation pullout resistance Rp,i and correspondingly onthe static pullout resistance RS for a given installation linetension Ti.

The ultimate penetration depth zult can be estimated as:

Fult Wz ⋅= λ (11)

where

λ = ultimate depth factor, which varies with theundrained shear strength (e.g. seabedintercept, depth gradient), seabed installationuplift angle, type of anchor forerunner andthe anchor itself (typically λ = 7-9)

WF = fluke width, equal to κ⋅(Afluke)0.5

(typically κ = 1.00)

The recommended design procedure is presented step-by-step in Section 5.3.2 and tentative safety requirements aregiven in Section 5.4. Since there is a close relationshipbetween the actual anchor installation tension and theresulting design anchor resistance, the design procedureintegrates these items through an iterative process. Theassessment of the minimum installation tension Tminresulting from this process is addressed in Section 5.5.Finally, requirements to soil investigation are given inSection 6 and Appendix G.


DET NORSKE VERITAS

5.3 Step-by-step description of procedure

5.3.1 General

The applied line tension T and the anchor resistance R areboth related to the dip-down point, where the mooring lineenters down into the seabed. The anchor resistance willthen be a function of the installation depth zi, the soilconditions, and the nature and rate of loading.

The main steps in the design of drag-in plate anchorsaccording to the procedure recommended herein aresummarised in the following, see also flowchart in Figure6.

5.3.2 Step-by-step procedure

1) Select mooring pattern (position, number andconfiguration of the mooring lines).

2) Determine the design line tension Td in the dip-downpoint from Eq. (6).

3) Select anchor type and make a 'first guess' on the flukearea Afluke from Eq. (9), i.e.

u

dfluke s

TA

⋅=

258.

4) Compute the penetration path vs. line tension T(z)down to the ultimate penetration depth zult for thisanchor, see Figure 2 and Eq. (11) for guidance.

5) Compute the design anchor pullout resistanceaccording to Eq. (10) for a number of points along thepenetration path, concentrating on the range 50 to 80 %of zult.− Find the penetration depth range where Rd ≥ Td.− If necessary, increase the selected installation

depth to ensure that the minimum depth criterionzi ≥ (zi)min in Section 3.2 is also satisfied.

− If the design limit state cannot be satisfied withinan acceptable depth range return to Step 1 or toStep 3 and select another mooring pattern oranother anchor.

− Check that anchor long-term creep during theoperational period is acceptable, see guidance inAppendix E. If not, evaluate the possibility toincrease zi, and thus the minimum installationtension Tmin, see Step 6 below.

6) Compute the minimum installation tension Tminaccording to Eq. (13) for the smallest acceptable depth.− Check if Tmin is feasible with respect to cost and

availability of installation equipment.− The anchor is acceptable if Tmin is feasible.− If Tmin is excessive, return to Step 1 or Step 3 and

consider a different mooring pattern or anchor.− If anchor long-term creep is intolerable, return to

Step 1, Step 3 or Step 5.7) Estimate the anchor drop point based on the drag

length computed in Step 4 and the selected installationpenetration depth zi.

The iteration process is continued until a suitable anchor isfound, while also taking into account the combined costsof purchase of equipment, installation, and retrieval.

Note 1. In case of significant layering reference is made toguidance in Appendix B.

Note 2. The acceptable uplift angle during installation of adrag-in plate anchor may be evaluated based on theguidance in Appendix F.

Note 3. The proposed partial safety factors for design ofdrag-in plate anchors, see Section 5.4, are tentative untilthe design rule proposed herein has been calibrated basedon reliability analysis.

Note 4. Analytical tools used for prediction of anchorperformance during installation and operational conditionsshould be well documented and validated.

5.3.3 Alternative procedure - based primarily onanchor tests

If anchor tests are being planned, and the results areintended to become the basis for design and installation ofanchors in the same area, the tests should be documentedby measurements, see guidance in Appendix C.

High quality anchor tests should continue to be performed,since these are vital for further development of the anchordesign procedures outlined herein.

Properly executed and interpreted site specific anchor testswill provide installation tension Ti and pullout resistanceRp,i versus installation depth zi, which may partly replacethe use of a computer programme to predict thepenetration trajectory versus line tension, see Step 4 of therecommended procedure in Section 5.3. However, beforeusing the anchor test results as a basis for anchor design,corrections for the effect of pullout speed on the pulloutresistance need to be made, see Eq. (4) and Appendix E.The method of extrapolating the measured behaviour ofthe test anchor to the prototype anchor size, and theassessment of the required installation tension, see Eq. (5)and Step 6, are examples of important design issues thatneed to be discussed and documented.

5.4 Tentative safety requirements

5.4.1 General

The safety requirements are based on the limit statemethod of design, where the anchor is defined as a loadbearing structure. For geotechnical design of the anchorsthis method requires that the following two limit statecategories be satisfied by the design:

− the Ultimate Limit State (ULS) and− the Accidental damage Limit State (ALS)

The following inequality must be satisfied for both limitstates:

0≥− dd TR (12)

The design line tension Td and the design anchor pulloutresistance Rd are given by Eq. (6) and Eq. (10),respectively.


DET NORSKE VERITAS

Step 2DetermineTd for selected

mooring pattern, see Eq. (6)

Step 3Choose anchor type and size

Step 5Compute Rd (z) for

50 to 80 % of zult along that pathSee Eq. (10)

Rd (zi) >Td ?zi > zmin?

(Section 3.2)

Step 6Compute necessaryTmin topenetrate the anchor to zi

See Eq. (15)

Step 7Estimate anchor drop point

for drag length computed in Step 4for z = zi

Step 1Select mooring pattern

Yes

Return toStep 1 or Step 3

No

No

Yes

NoReturn toStep 1, Step 3

or Step 5

Tmin feasible?- Equipment available?

- Cost OK?

Installationuplift angle OK?

See App. F

Step 4Compute penetration path to zult

See Figure 2 and Eq. (11)

Return toStep 1 or Step 3

Yes

Anchor creeptolerable?

Anchor design OK?

No

Yes

Yes

Find means toreduce angle- Return to

Step 4

Figure 6 Design procedure - flowchart.


DET NORSKE VERITAS

The ULS is intended to ensure that the anchor canwithstand the loads arising in an intact mooring systemunder extreme environmental conditions. The ALS isintended to ensure that the mooring system retainsadequate capacity if one mooring line or anchor should failfor reasons outside the designer's control.

Two consequence classes are considered, both for the ULSand for the ALS, defined as follows:

1) Failure is unlikely to lead to unacceptableconsequences such as loss of life, collision with anadjacent platform, uncontrolled outflow of oil or gas,capsize or sinking,

2) Failure may well lead to unacceptable consequencesof these types.

The primary function of an anchor, in an offshore mooringsystem, is to hold the lower end of a mooring line in place,under all environmental conditions. Since extremeenvironmental conditions give rise to the highest mooringline tensions, the designer must focus attention on theseconditions.

The ULS and ALS are nominally defined for the inceptionof anchor displacement in a direction normal to the flukeplane. This is an idealisation. A small amount ofdisplacement is implied in the mobilisation of the designanchor pullout resistance. This displacement may bereferred to as the failure displacement. The failuredisplacement may typically be 5 to 10% of the equivalentwidth WF of the fluke (plate) area. This magnitude offailure displacement assumes that consolidation hasoccurred after rotation of the anchor into position fornormal loading.

The basic safety requirements are given by Eqs. (6), (10)and /12/. The present ULS and ALS criteria (with partialsafety factors) are intended to avoid long-term anchorcreep. However, it should be borne in mind that the anchorload-displacement relationship from onset of creep tofailure is non-linear. Therefore it is probably realistic toadmit that even if the drag-in plate anchors are designedaccording to the present requirements, some small amountof anchor displacement may occur during the design life.More detailed investigation of long-term creep may,however, be advisable for systems with sustained highloading, as discussed in Appendix E. For discussion of theproposed partial safety factors, see Section 5.4.4.

For drilling rigs operating under conditions where anchorfailure is unlikely to lead to unacceptable consequences, atolerable anchor displacement may be agreed upon. Thetolerable displacement should be set to a value that issignificantly less than the failure displacement.Furthermore, the tolerable displacement of any anchorshould not lead to a significant change in mooring linetensions or platform position. It may be necessary to checkthis by calculation of mooring system response.

5.4.2 Partial safety factors for anchor resistance inthe ULS

With an intact mooring system the partial safety factorsgiven in Table 1 are tentatively suggested to avoid anchorfailure and intolerable long-term anchor creep. Partialsafety factors are specified for the two alternative methods,dynamic or quasistatic, to calculate line tension. Formooring in deep water a dynamic analysis is required.

Table 1 Partial safety factors for the ULS.

Consequenceclass

Type ofanalysis

γmean γdyn γm,1 γm,2

1 Dynamic 1.10 1.50 Note 1 1.402 Dynamic 1.40 2.10 1.15 1.401 Quasi-static 1.70 1.152 Quasi-static 2.50 1.15

Note 1: γm,1 = 1.0 + 0.35⋅(Td-mean/Td) (γm,1)min = 1.15γm,1 = 1.15 for temporary mooring

The resistance factors γm,1 and γm,2 shall account also forthe uncertainty in the calculation method and analysismodel, and the intact undrained shear strength, as far as itaffects the calculation of the anchor resistancecontributions in Eq(10).

For the ULS consequence class 1, when long-termmooring is considered and the ratio Td-mean/Td is higherthan about 0.40, γm,1 is increased from the basic value of1.15 according to the expression in Note 1 under Table 1.This is tentatively proposed to account for the increasedrisk of anchor long-term creep when this ratio becomeshigh.

For temporary mooring long-term creep is not a concernand γm,1 = 1.15 for all ULS-cases.

The resistance factors are intended for use with anchorresistance calculated by the recommended proceduredescribed in Section 5.3.2. Higher factors may be requiredif the alternative procedure described in Section 5.3.3 isadopted.

5.4.3 Partial safety factors for anchor resistance inthe ALS

Subsequent to a one-line failure or anchor failureintolerable displacement of any anchor is conservativelyassumed to imply mooring system failure in the ALSconsequence class 2. This implies that the failure eventdefined for the ULS is retained for the ALS consequenceclass 2. It is also expected that an initial mooring linefailure in severe weather will lead to a more unevendistribution of line tensions in the remaining lines.

For the ALS consequence class 1, failure is defined as ananchor displacement, which would have a noticeable, buttolerable, effect on the load distribution between the lines.The partial safety factors given in Table 2 are tentativelysuggested, when the same definition of the characteristicanchor resistance is used as for the ULS:


DET NORSKE VERITAS

Table 2 Partial safety factors for the ALS.

Consequenceclass

Type ofanalysis

γmean γdyn γm,1 γm,2

1 Dynamic 1.00 1.10 1.00 1.152 Dynamic 1.00 1.25 1.00 1.301 Quasi-static 1.10 1.002 Quasi-static 1.35 1.00

The maximum acceptable displacement of the anchorduring the period in which consequence class 1 conditionsprevail needs to be estimated, and the consequencesassessed and accepted from an anchor load-displacementbehaviour point of view during the actual period.

5.4.4 The partial safety factors are tentative

The partial safety factors in Table 1 and Table 2 are fortemporary use only until a formal calibration of therecommended anchor design procedure has been carriedout.

The present partial safety factors on tension are taken fromthe calibration of a mooring line design procedure, whichis dominated by uncertainty in the loading and littleaffected by uncertainty in the line strength.

The same partial safety factors on tension are notnecessarily equally suitable for an anchor designprocedure, in which uncertainty in the anchor resistance isexpected to play a significant part; e.g. the occurrence ofsimultaneous high anchor resistance and low tensionestimates has a probability less than one, and this needs tobe taken into account.

Calibration of the anchor design procedure may lead tolower safety factors on the line tension. Alternatively,these factors may be retained for convenience, andcompensated by relatively low safety factors on the anchorresistance.

5.5 Installation measurements

5.5.1 Minimum installation tension

The prescribed minimum installation tension Tmin will to agreat extent determine the geotechnical safety of theanchor as installed. Tmin may be assessed from Eq. (13)below, which includes the seabed friction developed overline length Ls,i lying on the seabed, i.e. for the condition ofno uplift during the anchor installation phase, see Figure4a).

( ) imislii LWzTT ,,min ' γµ ⋅⋅⋅+= (13)

The first term in Eq. (13) Ti(zi) is the line tension thataccording to the design calculations needs to be applied atthe dip-down point during anchor installation to ensurethat the design anchor resistance is provided. The secondterm is the seabed friction that has to be overcome in orderto ensure that a line tension equal to Ti(zi) is transmitted tothe dip-down point.

To account for the uncertainty in the predicted seabedfriction the resistance factor is set to

γm,i = 1.3

The friction coefficient µ will vary with the seabed soilcomposition and the type of forerunner, chain or wire.

For deepwater installations it is, however, likely that thedrag-in plate anchors are installed under an uplift angle,see Figure 4b), and the length of line on the seabed willthen be set to zero (i.e. Ls,i = 0), which changes Eq. (13) toTmin = Ti(zi).

In practice, Tmin will have to be calculated through aniterative process following the step-by-step procedureoutlined in Section 5.3.2. The resulting Tmin will then beevaluated and compared with the installation tension thatcan be mobilised with the installation scenarios underconsiderations. Regarding measurements required forverification of Tmin, see Appendix E.

5.5.2 Anchor rotation into normal loading mode

When the anchor has reached the final penetration depth,the anchor needs to be rotated as close as possible into itsoperational normal loading position. The necessary linetension to achieve this varies with type and size of anchor,but tentatively the line tension should be at least equal toTi. In addition other means should be used to control thattriggering has taken place, e.g. breaking the shear pin,which eventually is used to control the installation flukeangle. See further guidance in Appendix C.

6. Requirements to soil investigationThe planning and execution of soil investigations fordesign of drag-in plate anchors should follow establishedand recognised offshore industry practice. As a generalguidance to achieve this quality of soil investigationreference is made to the NORSOK standard /12/, whichmakes extensive references to international standards.Some specific recommendations are given herein withrespect to soil investigations for drag-in plate anchors.

For design of drag-in plate anchors the soil investigationshould provide information about:

− Seafloor topography and sea bottom features− Soil stratification and soil classification parameters− Soil parameters of importance for all significant layers

within the depth of interest.

The most important soil parameters for design of drag-inplate anchors in clay are the intact undrained shearstrength su, the remoulded undrained shear strength su,r , theclay sensitivity St, and the cyclic shear strength τf,cy foreach layer of significance.


DET NORSKE VERITAS

As a minimum, the soil investigation should provide thebasis for specification of a representative soil profile andthe undrained shear strengths, su and su,r , respectively, foreach significant soil layer between the seabed and theultimate depth of anchor penetration zult. The number ofsoil borings/in situ tests required to map the soil conditionswithin the mooring area should be decided from case tocase.

The ultimate penetration depth zult of drag-in plate anchorsin clay varies with the size of the anchor and the undrainedshear strength of the clay. It is convenient to account forthe size of the anchor by expressing the ultimatepenetration depth in terms of fluke widths WF. In very softclay zult may be up to 12-15 WF decreasing with increasingstrength of the clay. In strong, overconsolidated clay thedepth of penetration will seldom be more than 1-2 WF forpractical reasons (limitation in pulling force).

Drag-in plate anchors are primarily of interest for softclay, such that a reasonable range of ultimate penetrationto consider from a soil investigation point of view wouldbe 7-15 WF. However, an anchor is never (or seldom)designed for full utilisation of the ultimate resistance,partly because of the associated large drag distance.

The necessary depth of a soil investigation in clay withoutsignificant layering will be a function of the size of theanchor, the shear strength of the clay and the expectedrelative depth of penetration M, typically 30 to 50 m. Theupper few metres of the soil profile is of interest particularfor assessment of the initial anchor penetration resistanceand related fluke angle. This part of the soil profile willalso have an influence on the penetration (cutting)resistance against the embedded part of the anchor line andthe resulting inverse catenary.

General requirements to the soil investigation for drag-inplate anchor foundations, in addition to therecommendations in /12/, are provided in Appendix G.

7. References/1/ Dahlberg, R., Eklund, T. and Strøm, P.J. (1996),

Project Summary - Part 1 , Joint Industry Projecton design procedures for deep water anchors,DNV Report No. 96-3673. Høvik.

/2/ Dahlberg, R, Strøm, P.J., Ronold, K.O., Cramer,E. Mathisen, J., Hørte, T. and Eklund, T. (1998),Project Summary - Part 2 , Joint Industry Projecton design procedures for deep water anchors,DNV Report No. 98-3591. Høvik.

/3/ DNV Recommended Practice RP-E301 (1999),Design and Installation of Fluke Anchors inClay, April 1999.

/4/ Dahlberg, R. (1998), Design Procedures forDeepwater Anchors in Clay, OffshoreTechnology Conference, Paper OTC 8837, pp.559-567. Houston.

/5/ Sogstad, B.E., Mathisen, J., Hørte, T. and Lie H.(1998), Modifications to DNV Mooring Code(POSMOOR) and their Consequences,Conference on Offshore Mechanics and ArcticEngineering (OMAE), Paper 1460. Lisbon.

/6/ DNV Rules for Classification of MobileOffshore Units (1996), Position Mooring(POSMOOR), Pt.6 Ch.2, January 1996.

/7/ Foxton, P., Latest Developments for VerticallyLoaded Anchors, 2nd Annual Conference onMooring & Anchoring, June 2-3, 1997,Aberdeen.

/8/ Agnevall, T., Installation and Performance ofP27 Stevmanta-VLA Anchors, paper presentedat the Conference on Moorings & Anchors,October 12-13, 1998. Aberdeen.

/9/ Foxton, P., Bruce, P. and Craine, G.A.,Development of the Bruce Anchor Tracker,Conference on Moorings & Anchors, October12-13, 1998. Aberdeen.

/10/ Dahlberg, R. and Strøm, P.J. (1999), UniqueOnshore tests of Deepwater Drag-in PlateAnchors, Offshore Technology Conference,Paper OTC 10989. Houston.

/11/ Vivitrat, V., Valent, P.J., and Ponteiro, A.A(1982), The Influence of Chain Friction onAnchor Pile Behaviour, Offshore TechnologyConference, Paper OTC 4178. Houston.

/12/ NORSOK standard (1996), CommonRequirements Marine Soil Investigations, G-CR-001, Rev. 1, dated May 1996.


DET NORSKE VERITAS

Appendix A Analysis tool for fluke anchor design

A1 GeneralAn analytical tool for design of drag-in plate anchorsshould be able to calculate anchor line catenary in soil aswell as the force and moment equilibrium of the drag-inplate anchor itself, both in the penetration and in thenormal loading mode. Further, the analytical tool shouldbe capable to predict the effect of consolidation on thepenetration resistance should there be a stoppage duringinstallation of the anchor and it has to be restarted. Thefollowing section describes in brief the principles for suchan analytical tool developed by DNV /A-1/

A2 Anchor line seabed frictionThe resistance due to seabed friction during anchorinstallation is given by the second term in Eq. (1), see alsosafety requirements in Section 5.5. The coefficient ofseabed friction µ to be used in the prediction of the seabedfriction is different for wire and chain.

Guidance NoteBased on the back-fitting analysis of data frommeasurements on chain segments reported in /A-2/ andestimated values for wire, the following coefficients ofseabed friction are recommended for clay1):

Table A-1 Coefficient of seabed friction

Wire Lower bound Default value Upper bound

µ 0.1 0.2 0.3

Chain Lower bound Default value Upper bound

µ 0.6 0.7 0.8

The unit friction f along the embedded part of the anchorline as required for calculation of anchor line contributionto the anchor resistance Ri is given by Eq. (A-4).

--- End of Guidance Note ---

A3 Equilibrium equations of embedded anchorlineThe equilibrium of the embedded part of the anchor linecan be solved approximately by closed form equations orexactly in any soil strength profiles by iterations /11/. Thenormal stress q and the unit soil friction f, which act on ananchor line element in the soil are shown schematically inFigure A-1.

q

f

θW l'

ds

dθT

Figure A-1. Soil stresses at an anchor line segment inthe soil.

The loss in line tension dT over one element length ds iscalculated from the following formula:

)sin(' θ⋅−⋅−= lWASfdsdT

(A-1)

where

T = anchor line tension

? = orientation of anchor line element (θ = 0 fora horizontal element)

AS = effective surface of anchor line per unitlength of line

ds = element length

The angular advance from one anchor line element to thenext is then solved by iterations from the followingformula:

TWABq

dsd l )cos(' θθ ⋅−⋅

= (A-2)

where

q = normal stress

AB = effective bearing area of anchor line per unitlength of line

Guidance NoteThe following default values are suggested for theeffective surface area AS and the effective bearing areaAB:


DET NORSKE VERITAS

Table A-2 Effective surface and bearing area

Type offorerunner

AS AB

Chain 11.3⋅d 2.5⋅d

Wire or rope π⋅d d

where

d = nominal diameter of the chain andactual diameter of the wire or rope.

-- End of Guidance Note ---

The normal stress q on the anchor line is calculated fromthe following equation:

uc sNq ⋅= (A-3)

where

Nc = bearing capacity factor

su = undrained shear strength (direct simple shearstrength suD is recommended)

Effect of embedment on the bearing capacity factor shouldbe included.

Guidance NoteBased on the back-fitting analyses reported in /A-2/ and/A-3/ the following bearing capacity factors arerecommended for the embedded part of the anchor line inclay1):

Table A-3 Bearing capacity factor for wire/chain 1)

Wire / Chain Lowerbound

Defaultvalue

Upperbound

Nc 9 11.5 14

1) See Table A-2 for values of the effective bearing areaAB, which is a pre-requisite for use of the bearingcapacity factors given here.


The unit friction f along the anchor line can be calculatedfrom the following formula:

usoil sf ⋅= α(A-4)

where

αsoil = adhesion factor for anchor line

Guidance NoteBased on the back-fitting analysis of data frommeasurements on chain segments reported in /A-2/, andestimated values for wire, the following coefficients ofseabed friction are recommended for the embedded partof the anchor line clay1):

Table A-4 Adhesion factor for wire and chain1)

Wire Lowerbound

Defaultvalue

Upperbound

αsoil 0.2 0.3 0.4

Chain Lowerbound

Defaultvalue

Upperbound

αsoil 0.4 0.5 0.6

1) See Table A-2 for values of the effective surface areaAS, which is a pre-requisite for use of the adhesionfactor given here.


A4 Equilibrium equations for drag-in plateanchorMoment equilibrium and force equilibrium can be solvedfor the drag-in plate anchor for two different failuremodes. One mode leading to further anchor penetration ina direction close to the fluke penetration direction, and asecond mode leading to reduced or no further penetration.In principle, the soil resistance contributions are the samefor the two failure modes, but in the first failure mode thesoil resistance normal to the fluke may not take on theultimate value. Using the symbols shown in Figure A-2 thenecessary equilibrium equations are defined and explainedin the following.

β

Penetrationdirection

θe

RFN

W'

RFS

T

Rai

RTIP

Figure A-2. Principal soil reaction forces on a drag-inplate anchor (penetration direction coincides with flukepenetration direction).

For the range of possible penetration directions, thehorizontal and vertical equilibrium should satisfy thefollowing equations:


DET NORSKE VERITAS

Horizontal equilibrium:

+⋅+⋅=⋅ ∑=

)cos()cos()cos(1

ββθ FS

N

iai RRT

)sin()cos( ββ ⋅+⋅ FNTIP RR(A-5)

Vertical equilibrium

−+⋅=⋅ ')cos()sin( aFN WRT βθ

⋅+⋅+⋅∑

=

)sin()sin()sin(1

βββ TIPFS

N

iai RRR

(A-6)

where

T, θ = tension and corresponding orientation ofanchor line at the shackle

RFN = soil normal resistance at the fluke

RFS = soil sliding resistance at the fluke

RTIP = tip resistance at the fluke

Rai = soil resistance at the remaining componentsof the anchor (separated through anchorgeometry)

Wa' = submerged anchor weight

β = penetration direction of fluke

The normal resistance will be the normal stress times thebearing area of the anchor part being considered, and mayneed to be decomposed in the three orthogonal directionsdefined (one vertical and two horizontal). The normalstress can be calculated from the following formula:

uc sNq ⋅= (A-7)

where

Nc = bearing capacity factor

Sliding resistance will be the unit friction f times theadhesion area of the anchor part being considered. The unitfriction f can be calculated from the following formula:

usf ⋅= α(A-8)

where

α = adhesion factor for anchor

su = undrained shear strength (the direct simpleshear strength suD is recommended)

The bearing and adhesion areas should in this case bemodelled with due consideration of the actual geometry ofthe anchor.

Guidance NoteBased on the back-fitting analysis performed in the JIP ondeepwater anchors /A-2/ the following tentatively valuesare recommended for the resistance towards the variousanchor members in clay:

Table A-5 Bearing and adhesion factors for anchor

Bearing capacity factor Nc 1) Adhesion factor α

RFN Rai RTIP RTIP RFS

12.5 2) 12.5 12.5 1 / St 1 / St

1) Effect of shape, orientation and embedment of thevarious resistance members on the anchor should beincluded as relevant.

2) Actual degree of mobilisation of this value asrequired to satisfy moment equilibrium.


Horizontal and vertical equilibrium for a certain flukepenetration direction can now be achieved for a number offluke orientations and line tensions at the shackle. In orderto determine the correct penetration direction and thecorresponding line tension, moment equilibrium must besatisfied (here taken with respect to the shackle point):

( ) 01

=⋅+−++∑=

eRWmRmRmRm FNTIPFS

N

iai

(A-9)

where

RmFS = moment contribution from soil slidingresistance at the fluke

RmTIP = moment contribution from tip resistance atthe fluke

Wm = moment contribution from anchor weight

RFN = soil normal resistance at the fluke

e = lever arm between shackle and the line ofaction of the normal resistance at the fluke

Rmai = moment contribution from soil resistanceat the remaining components of the anchor(separated through anchor geometry).

When the anchor penetrates in the same direction as thefluke, any possible lever arm (e) and normal resistance thatcan be replaced by a realistic stress distribution at the flukeshould be considered. When the anchor penetrates inanother direction than the fluke, the centre of normalresistance on the fluke should act in the centre of the flukearea.

A5 References/A-1/ Eklund T and Strøm, P.J. (1998), DIGIN

Users’s Manual ver. 5.3 , DNV Report No. 96-3637, Rev. 03, dated 20 April 1998. Høvik

/A-2/ Eklund T and Strøm, P.J. (1998), Back-fittingAnalysis of Fluke Anchor Tests in Clay, DNVReport No. 96-3385, Rev. 03, dated 16September 1997. Høvik

/A-3/ Strøm, P.J. and Dahlberg, R. (1998), Back-fitting Analysis of Drag-in Plate Anchors inClay, DNV Report No. 98-3585, Rev. 01,dated 13 January 1999. Høvik


DET NORSKE VERITAS

Appendix B Drag-in plate anchors in layered clayGuidance for assessment of the penetration ability of drag-in plate anchors in layered clay is given in the following.Layering is understood herein as a soil layer sequencecomprising a soft layer overlain and/or underlain by arelatively stiffer clay (or sand) layer.

Drag-in plate anchors are particularly suitable for softnormally consolidated clays and their general behaviour isaddressed in Chapter 3. Figure 2 relates to anchors in claywithout significant layering.

Experience has shown, however, that drag-in plate anchorsoften penetrate through an overlying layer of sand orstiffer clay as long as the thickness of this layer is less than30 to 50 % of the fluke width WF of the actual anchor.Penetration through the upper stiff layer may sometimesrequire a smaller fluke angle than desirable for penetrationthrough the underlying soft layer. This may be resolved bydesigning a shear pin, which fails for a shackle load that issufficient to penetrate the anchor through the surface layer.When this shear pin breaks the fluke angle opens up to anangle, which is found to be suitable for deep embedmentinto the underlying soft clay layer.

In a soft-stiff layer sequence the anchor should normallystay in the soft layer and avoid partly penetration into thestiff layer. Since the pullout resistance will be governed bythe undrained shear strength of the soft overlying clay, atarget installation load related to the penetration resistanceof the stiffer clay will be misleading. If predictions oranchor tests show that there is a risk that the targetinstallation load cannot be reached without penetrationinto the stiffer layer, changing to another type and/or sizeof anchor may improve the situation. If drag-in plateanchors at all should be used is dependent on the thicknessof the soft layer and the loads, which have to be resisted.

A stiff-soft-stiff layer sequence will in most circumstancesinvolve extra complications in that penetration through theupper stiff layer may require a smaller fluke angle thandesirable for penetration through the locked-in soft layer.Again, the drag-in plate anchor should be designed to staywithin the soft layer and avoid partial penetration into theunderlying stiff layer. If the strength of the locked-in softlayer is smaller than assumed in designing the anchor, thetarget installation tension Ti may not be reached, visualisedby continuous drag at constant tension. Designing theanchor for less than ultimate penetration as discussed inSection 3.2 may reduce this risk. In most cases, predictionsmay show that the penetration path improves in thatrespect, and becomes steeper for a given depth and a givenfluke angle, if the anchor is increased in size. In manycases it may be possible to find an optimal, non-standard,combination between anchor size and fluke angle, whichaccounts both for the overlying and the underlying stifflayer and ensures that the anchor stays within the soft claylayer in between. To consider drag-in plate anchors at allin layered soil the target clay layer must be reachable andhave a strength and thickness, which confidently can beutilised to provide a safe pullout resistance.

From the above it is evident that layer thickness, and depthto boundaries between layers, need to be documented forproper design of a drag-in plate anchor and to avoidunexpected behaviour of the anchor during the installationphase, see further about requirements to soil investigationin Chapter 6 and Appendix G.


DET NORSKE VERITAS

Appendix C Installation and testing of drag-in plate anchors

C1 GeneralBoth the offshore and the onshore testing of drag-in plateanchors have focussed a great deal on the performanceratio Pr. The Denla and Stevmanta anchors tested undercontrolled onshore conditions, see /10/ have givenperformance ratios in the range Pr = 1.5 - 2.5, which agreefairly well with the experience from offshore tests. Highervalues have been published and an explanation to thisdiscrepancy is offered below. It is desirable to continuetesting of these anchors and other plate anchors, since thedatabase is rather thin, although many tests are of a highquality.

From an installation point of view a drag-in plate anchorcan be compared with a fluke anchor, and it may betempting to use results from drag-in plate anchor tests todevelop charts similar to those found in /C-1/ for flukeanchors. In fact, in the Anchor Manual 2000 from Vryhof,see /C-2/, a design chart for the Stevmanta anchor isincluded. It should be noted, however, that theperformance ratio assumed in this chart is Pr = 3.0, whichis unconservative and not recommended for use until welldocumented test data can support such high values.

It may well be that one under certain offshore testingconditions can obtain a better performance ratio than 2.5,but an explanation to this may be found in the offshoretesting procedures. The final anchor installation tensionTmin is normally maintained over a specified period oftime, e.g. 15 to 30 minutes, which takes out part (but notall) of the loading rate effect included in the anchorinstallation resistance. The actual loading rate effectremaining in the line tension during this holding period isweather dependent and increases with the motions of theinstallation vessel. Also the pullout test is affected by thesea state prevailing during the time of testing. If theweather gets rougher the motions of the installation vesselincreases and the tension variation during the pullout testmay become quite significant and lead to higher thannormal pullout resistance.

With the performance ratio Pr being defined as the ratiobetween the installation pullout resistance Rp,i and thetarget installation resistance Ti it is therefore quite possibleto explain the large scatter in Pr from offshore tests.

In order to improve the basis for interpretation of anchortest results it is recommended to take records of theweather conditions during the period of anchor testing andto measure the pullout rate, see further guidance in SectionC3.2 and Appendix E.

However, use of design charts as presented in /C-2/ as abasis for final design of drag-in plate anchors is notrecommended due to the inherent uncertainties in suchcharts with reference to the above discussion.

As a general reference to installation procedures normallyadopted for both fluke anchors and drag-in plate anchors,as well as desciptions of associated installation equipment,the Anchor Manual from Vryhof /C-2/ is recommended asa useful supplement to this RP.

All reasonable efforts should therefore be made to ensurethat the measurements are reliable and include the mostcrucial test data for maximum usefulness of the results andimprovement of the database. This should be fullyappreciated when installing both test anchors andprototype anchors. Tentative guidelines for monitoring ofdrag-in plate anchor tests and commercial anchorinstallations are given in Section C3.

C2 Minimum installation tensionThe anchor installation should follow procedures, whichhave been presented and agreed to by all parties well aheadof the installation. By prescribing a minimum installationtension Tmin, see Section 5.4.2, the intention is to ensurethat the design assumptions are fulfilled during anchorinstallation. In other words, if the anchor is installed to Tminthe design anchor resistance Rd has implicitly also beenverified, unless the anchor penetrates partly into anunderlying stiffer layer, see Appendix B Tmin should beheld for a specified holding period, which period may besoil dependent. Any relaxation (drag) during this periodshould be compensated for, such that the required linetension is maintained as constant as possible. The anchorinstallation and testing log should document the events andthe measurements taken from start to end of theinstallation.

The possible extra resistance due to seabed friction duringanchor installation should be accounted for according toEq. (13), see Guidance Note in Appendix A.

C3 Installation measurementsC3.1 GeneralWhen installation of prototype, or test anchors, is beingplanned it is essential that the most essential boundaryconditions for the installation be taken into consideration.Well ahead of the installation such backgroundinformation should be compiled and documented.

If practical (e.g. if ROV assistance is available duringanchor installation) it is recommended to check theposition and orientation of the anchor, as well as thealignment, straightness and length on the seabed of the aslaid anchor line, before start of tensioning.

During the anchor installation a number of parametersneed to be measured to serve as a documentation of theinstallation. The more information that is recorded beyondthe minimum documentation requirements, the moreuseful the installation data will become in the end.


DET NORSKE VERITAS

Monitoring of the anchor installation should, as aminimum, provide data on

− line tension− line (pitch) angle at the stern roller− anchor drag

These items should be measured as a function of time fromstart to end of the installation using the clock on the PC asa reference time. A calibrated load ecll, being a segment ofthe installation line, should preferably be used to measurethe line tension.

If manual measurements are taken intermittently, seechecklist below, they should be stored into the PC log atthe time of the event.

The final installation measurements should at leastdocument that the minimum installation tension Tmin hasbeen achieved and maintained during the specified holdingtime.

After installation the triggering of the anchor should becontrolled by means as provided for each anchor type. Asa final preparation before the anchor is ready for hook-upto the floater the anchor is to be rotated into its normalloading mode. The line tension at deck level to achievethis should be at least equal to Tmin, and it should bemeasured and documented.

The checklist below indicates the type of information thatshould be focussed on before and during the installationand testing of drag-in plate anchors. This checklist can beused as a guidance for installation of both prototype andtest anchors.

C3.2 Checklist1) Before the installation.

a) Assessment of the most likely soil stratigraphy atthe anchor location and the soil strength ofsignificant layers (from soil investigation report),see Chapter 6, Appendix B and Appendix G forguidance.

b) Specification of the anchor and the installation lineconfiguration.

c) Specification of the fluke angle(s) to be used, andhow this angle is defined, see Section 2 and Figure1 for guidance.

d) Estimate of friction resistance at the stern roller.e) Equipment and procedures for anchor installation,

e.g. type and tensioning system of the vessel,method of laying and tensioning the anchors,availability of ROV, etc.

f) Type of measurements to be undertaken, andprocedures to be applied, from check list below.

2) During the installation.a) Line tension (at deck level)1

b) Drag (method of measurement, reference point)

1 The installation tension should be measured as accurately aspossible, e.g. by means of the TENTUNE method /C-3/.

c) Penetration depth (method of measurement, atleast the final depth)

d) Line angle with the horizontal outside the sternroller (at least for the final line tension)

e) Pull-in speed (vessel speed, drag and line angle atstern roller versus time)

3) Final installation measurementsa) Maintaining Tmin (during specified holding period,

thold = 15 to 30 minutes)b) Measure tension vs. time during holding period

(mean tension ≥ Tmin)c) Drag (corresponding to final penetration depth)d) Penetration depth (best estimate of final depth)

4) Triggering and rotation of anchora) In a one-line installation scenario verify that the

anchor has triggered through measurements (andobservation) of line tension variation, e.g. a suddendrop in line tension followed by a rapid increase,when the shear pin breaks.

b) In a two-line installation scenario the triggering isaccomplished by changing from the installationline to the mooring line.

c) After triggering verify that the anchor has rotatedinto its normal loading position, approximate lineangle Θ with the horizontal, by measuring anddocumenting a line tension at deck level at leastequal to Tmin.

5) Pullout test of anchor in normal loading modea) Line tension (at deck level) 1

b) Penetration depth (method of measurement, atleast the final depth)

c) Line angle with the horizontal outside the sternroller (at least for the final line tension)

d) Pullout speed (e.g. vessel or winch speed and lineangle at stern roller versus time)

The database for drag-in plate anchors loaded to theirultimate resistance Rult and ultimate penetration depth zult isstill very limited, and for practical reasons the size of thetest anchors is normally rather small, limited to theavailable pull force. The largest anchors tested inconnection with offshore projects have therefore notreached zult, but for the future it would be fruitful for theindustry if the most significant parameters (line tension,drag, final penetration depth, pullout line tension andpullout rate) are recorded during all installations, at least ina few locations out of many.

In this connection it is important that all reasonable effortsare made to make the recorded data as reliable as possible,since the assessment of the safety of the anchoring systemdepends on such installation data.


DET NORSKE VERITAS

C4 Anchor installation vesselsThe bollard pull of the most powerful new generationanchor handling vessels is in the range 2 to 2.5 MN.Depending on the required minimum installation tensionTmin at the touch-down point, one or two AHV's may berequired. If feasible, and practically possible, the anchortensioning can also be done from a special tensioningvessel/barge or from the floater itself. If two oppositeanchors are tensioned simultaneously line tensions up to 5to 6 MN or even 10 MN can be reached.

The chosen scenario for anchor installation shall ensurethat the specified minimum installation tension Tmin can bereached. The bollard pull, winch capacity and break testload (BTL) of the installation wire on the actual vessel(s)will have to be assessed on this basis. If Tmin cannot bereached due to pulling limitations set by the vessel(s), thedesign anchor resistance Rd according to Eq. /10/, and thusthe intended safety level of the anchors, will not beachieved.

It is essential that all parties involved in the decisionsrelated to the anchor design appreciate the relationshipbetween anchor resistance and installation tension. In deepwaters, unless lightweight anchor lines are used, theweight and sea bed friction of the anchor lines limits thenet line tension that can be used for anchor penetration,which must be considered when the requirements for theinstallation vessel are specified.

C5 References/C-1/ API Recommended Practice 2SK (1996),

Recommended Practice for Design andAnalysis of Stationkeeping Systems forFloating Structures, 2nd Edition, effectivefrom March 1997.

/C-2/ Vryhof (1999), Anchor Manual 2000 ,VryhofAnkers. Krimpen a/d Yssel, The Netherlands.

/C-3/ Handal, E. and Veland, N. (1998),Determination of tension in anchor lines, 7th

European Conference on Non-DestructiveTesting, Copenhagen, 26-29 May, 1998.


DET NORSKE VERITAS

Appendix D Consolidation effect

D1 GeneralThe zone of clay being remoulded during penetration of adrag-in plate anchor will reconsolidate after completion ofthe anchor installation. The consolidation of this volume ofclay will have little or no effect on the normal loadingresistance of the anchor, since this involves the resistanceof a volume of clay that was not remoulded during thepenetration phase. The effect of this consolidation should,however, be considered if the anchor penetration isdelayed and the anchor has to be restarted again. Theincrease in penetration resistance due to such temporarystoppage during installation depends on the anchorcharacteristics, the soil characteristics, the duration of thestoppage, etc. as discussed in the following.

After completion of the anchor penetration phase theanchor is triggered and prepared for normal loading andhook-up to the floater. This will normally involve avariable degree of rotation of the anchor fl, leading tomore significant soil disturbance than during penetrationof the anchor. This rotation may therefore affect (reduce)the anchor resistance that can be achieved in a pullout testperformed shortly after this rotation took place. In thisdesign procedure the possible increase in the pulloutresistance due to consolidation of this partly remouldedclay has been disregarded.

D2 Assessment of the effect of consolidationDuring continuous penetration of the anchor, the slidingresistance will be governed by the remoulded shearstrength, sur, in a narrow zone close to the anchor. In ananalytical model this may be accounted for through theadhesion factor, α, which will depend on the soilsensitivity, St, i.e. the ratio between the intact (in situ)undrained shear strength, su, and su,r

St = su / su,r (D-1)

The minimum α-value is tentatively set equal to theinverse of the sensitivity, i.e.

αmin = 1 / St (D-2)

After an anchor has been installed to a certain installationtension (and depth), the remoulded soil will graduallyreconsolidate and regain its intact shear strength. As aresult the resistance against further penetration willincrease. This effect is in the literature referred to assoaking, set-up or consolidation of the anchor and theanchor line.

The effect of soil consolidation is that the installationanchor resistance Ri will increase as a function of the timeelapsed since installation tcons to a maximum value, whichdepends on the soil sensitivity St. For a particular anchorand depth of penetration this increase may be describedthrough the consolidation factor Ucons, which is a functionof the soil consolidation characteristics, as well as thegeometry, depth and orientation of the anchor.

Ucons = f(tcons, St, etc.) (D-3)From a geotechnical point of view there should be nomajor difference between drag-in plate anchors and e.g.piles or the skirts of a gravity base structure, when theeffects of installation and subsequent reconsolidation onthe clay undrained shear strength are considered. Theconsolidated resistance Rcons is the installation resistancewith superimposed consolidation effect as shown in Eq.(D-4).

consiconscons RRURR ∆+=⋅= i (D-4)

The degree of consolidation that can be applied to thefrictional part of the resistance can be assessed by lookingat the drainage characteristics in a zone adjacent to theanchor, which is influenced (remoulded) due to the anchorpenetration. The length of this zone depends on the anchorgeometry and the actual soil characteristics. Guidance formodelling and calculation of the consolidation effect canbe obtained using the experience from e.g. tests on piles.

The consolidation factor Ucons related to the total anchorresistance will be much smaller than reflected by thesensitivity of the clay, since the sliding resistance onlycontributes to part of the total resistance. The relationbetween the consolidation factor Ucons and the increase inthe sliding resistance depends on the geometry of theanchor, and its final depth of penetration into the soilduring the installation phase. A reliable quantification ofthis effect can only be obtained by site-specific relevantfull-scale tests or by adequate analytical tools. Theanalytical tools should be able to predict both thepenetration part and the subsequent consolidatedcondition. It is essential that the analytical tool accountsfor full force and moment equilibrium that is compatiblewith the failure modes in question, see Appendix A.


DET NORSKE VERITAS

Caution is recommended in the assessment of the possibleconsolidation effect when the likely failure mode,following upon such consolidation, may either reduce orprevent further penetration. Overloading will in this caseinitiate anchor movement in the direction of the linetension, before the full effect from consolidation isutilised. When such movement has been initiated, the soilcloser to the flukes will loose the effect fromconsolidation, and the anchor will continue to drag inremoulded soil conditions. This can in particular beexpected close to the seabed, where the resistance in thedirection of the line tension is limited, but may also berelevant at larger depths, if the anchor has penetrated witha very large fluke angle, or in layered soil if the fluke tiphas penetrated partly into a stiffer layer underlying a softlayer. Therefore, the consolidation factor Ucons must beassessed on a case by case basis.

In practice, the anchor installation should be planned suchthat stoppage before the minimum installation tension Tmin

has been reached is avoided.

Guidance NoteValues for Ucons vs. typical soil sensitivity St, as applicable forassessment of the restart resistance including the consolidationeffect due to stoppage during anchor installation, are given inTable D-1.

Table D-1 Consolidation factor Ucons

UconsSoil sensitivitySt Lower bound Default value Upper bound

2 1.35 1.45 1.50

2.5 1.50 1.60 1.70

-----End of Guidance Note ----


DET NORSKE VERITAS

Appendix E Cyclic loading effectsFundamental work on the effects of cyclic loading on theundrained shear strength of clay and the cyclic response ofgravity base foundations has been published by Andersenand Lauritzen /E-1|/. This provides a basis forunderstanding also how cyclic loading may affect theresistance of drag-in plate anchors.

Cyclic loading affects the static undrained shear strength(su) in two ways:

1) During a storm, the rise time from mean to peak loadmay be about 3 - 5 seconds (1/4 of a wave frequencyload cycle), as compared to 0.5 to 2 hours in a staticconsolidated undrained triaxial test, and this higherloading (or strain) rate leads to an increase in theundrained shear strength

2) As a result of repeated cyclic loading during a storm,the undrained shear strength will decrease, thedegradation effect increasing with theoverconsolidation ratio (OCR) of the clay.

The most direct, and preferred, approach to account forboth the loading rate effect and the cyclic degradationeffect is to determine the cyclic shear strength τf,cy of theclay, following the strain accumulation proceduredescribed in /E-1/.

The strain accumulation method utilises so-called strain-contour diagrammes to describe the response of clay tovarious types, intensities and duration of cyclic loading:

Given a clay specimen with a certain su and OCR, which issubjected to a load history defined in terms of a sea stateand a storm duration, the intensity of that storm isgradually increased until calculations according to thestrain accumulation method show that the soil fails incyclic loading.

In a catenary mooring system the loads transmitted to theanchors through the mooring lines will always be intension (one-way), which has a less degrading effect on theshear strength than two-way cyclic loading (stressreversal). The failure criterion for one-way cyclic loadingis development of excessive accumulated permanentstrains. The maximum shear stress the soil can sustain atthat state of failure is equal to the cyclic shear strength τf,cy.

The load history for use in the calculations should accountfor the combination of wave-frequency load cyclessuperimposed on low-frequency, slowly varying, loadcycles, particularly the amplitude of cyclic loads relative tothe mean line tension Tmin, including the line pretensionTpre.

If cyclic soil data, applicable for the actual site, areavailable, the cyclic strength τf,cy may be determinedaccording to the procedure outlined in /E-1/. The cyclicstrength τf,cy as defined in /E-1/ incorporates effects of bothloading rate and cyclic degradation, provided that thecyclic load period is representative for the variation in linetension with time at the anchoring point. This would leadto a combined loading rate and cyclic degradation factor,or simply a cyclic loading factor Ucy as shown in Eq. (E-1)below.

Ucy = τf,cy/su = f [tsu/tcy, soil data, load history, etc] (E-1)where

τf,cy = cyclic shear strength with time to failure

tc y = (1/4)⋅(load period)

su =static undrained shear strength with time tofailure

tsu set equal to 1 hour

E2 Loading rate effectsThe loading rate plays a role both during installation andpullout of a drag-in plate anchor, which has beendemonstrated in the field tests described in /10/.

Important work on the effect of loading rate on axial pilecapacity has been published by Bea and Audibert /E-2/,followed by Kraft et al /E-3/, and later by Briaud andGarland /E-4/. The following relationship is suggested in/E-4/ for description of the effect of the loading rate, v , onpile capacity, Q

(Q1/Q2) = (v1/v2)n

(E-2)

where Q1 and Q2 represent the pile capacity at loading ratesv1 and v2, respectively

A loading rate factor Ur may be introduced, whichincorporates the loading rate effect on the drag-in plateanchor resistance, i.e.

Ur = (v1/v2)n (E-3)One practical problem with Eq. (E-3) is to determinerepresentative values for the loading rates v1 and v2 atextreme line tension and at end of installation,respectively. Another problem is to assess the value ofexponent n in Eq. (E-3). The effect of loading rate on theinstallation anchor resistance Ri may be significantlyreduced by holding the minimum installation tension Tminover a specified period of time thold while takingmeasurements as recommended in Appendix C.


DET NORSKE VERITAS

The effect of loading rate on the pullout resistance Rp,i

measured in offshore or onshore anchor tests may varyfrom one test to another. This has the consequence that themeasured pullout resistance of the test anchor includes acertain loading rate effect. Based on the results from theinstrumented drag-in plate anchor tests at Onsøy inNorway as reported in /10/ an approach has been proposedfor quantification of the loading rate effect andestablishment of a quasi-static pullout resistance RS. Thisapproach is explained in the following.

E3 Loading rate factor versus strain rateIt has been possible to establish a relationship betweenresults from the anchor tests in clay /10/ and extensivelaboratory tests on Drammen clay /E-1/ and Troll clay /E-5/. The following approach was used:

1) Find results from anisotropically consolidatedundrained compression (CAU) triaxial tests on clayswith an overconsolidation ratio OCR = 1.

2) Relate the static undrained shear strength su to theapplied strain rate v (in % per hour).

3) Find for comparison the peak cyclic shear stress in thefirst cycle (Neqv = 1), which gives a permanent axialstrain of 10 % (failure strain), using results from testswith 1-way cyclic loading. With a cycle period of 10seconds, this gives a strain rate of 14,400 %/hour forcomparison with 3 %/hour in the static test (failurestrain set to 3 %, reached after 60 minutes).

4) Compare the measured relationship between anchorpullout resistance and strain rate in the Onsøy testswith the behaviour observed in the triaxial tests.

Following this approach it was possible to express theloading rate factor Ur as a function of the strain rate v asshown in Figure E-1.

Since the triaxial test data suggest that the static undrainedshear strength varies exponentially with the strain (orloading) rate, it is convenient to relate the loading ratefactor Ur to the strain rate as shown in Eq. (E-4).

( )nrefr vvU /= (E-4)

where

v = actual strain rate (%/hour)

vref = reference strain rate, set to 3 %/hour

n = exponent, which is dependent on type of soiland method of testing

The exponent derived from the triaxial tests on Drammenclay and Troll clay was 0.040 and 0.041, respectively,when the reference strain rate was set to vref = 3 %/hour.Combining the results from the anchor tests at Onsøy withthe criterion that also this line must intersect the staticresistance line at a strain rate of 3 %/hour, an exponent n =0.054 is found for the anchor tests, assuming the failurestrain to be 5 %.

-3 -2 -1 0 1 2 30

0.2

0.4

0.6

0.8

1.0

1.2

1.4

1.6

Ur

Rp,i (typical)

RS = ψ . Rp,i

Ur = 1.25

ψ = 1/Ur =0.80

Rp,cr = ρ . RS

log10(v/vref)3.7

Figure E-3 Loading rate factor versus strain rate v,with vref = 3 %/hour (tentative).

Tentatively it is suggested that a typical offshore pullouttest corresponds to a strain rate of about 180 %/hour,which gives Ur = 1.25 according to Eq. (E-4), i.e. a 25 %increase in anchor resistance above the static pulloutresistance RS at vref = 3 %/hour.

This means that the measured values of Rp,i need to bemultiplied by a factor Ψ = 1/Ur = (1/1.25) = 0.80 to get thestatic pullout resistance RS. Until more test data becomeavailable Ψ = 0.80 is recommended as a default value, seealso Appendix C about guidance for measurements duringinstallation and testing of drag-in plate anchors.

Assuming further that the lines can be extrapolatedbackwards towards strain rates less than 3 %/hour, athreshold strain rate associated with sustained loading maybe established, which would give only negligible anchorcreep over the actual operational period. Anchor creep isdiscussed in Section E4 following.

E4 Assessment of creep pullout resistanceAnchors for deepwater mooring in taut mooring systemwill be subjected to significant mean line tensions Tmeanduring severe weather conditions. This makes anchor creepa design issue, which needs to be addressed. It should,however, be mentioned that for a plate embedded to some20 to 30 m depth below seabed, creep should not representa serious threat to the safety of the mooring system, if theanchors are designed to satisfy the ULS and ALSrequirements according to this procedure.

The creep pullout anchor resistance Rp,cr is defined suchthat anchor creep is avoided during the design life of thefloater, if the design mean line tension Td-mean, does notexceed Rp,cr, i.e.

crpmeand RT ,≤− (E-5)

With reference to Figure 1 Rp,cr is related to the staticanchor pullout resistance RS through the creep factor ρ

Scrp RR ⋅= ρ, (E-6)


DET NORSKE VERITAS

The value of the creep factor ρ should reflect the designlife of the floater on the actual location, the intensity andduration of the various sea states, the type of clay and itscharacteristic properties. Tentatively, the creep factor ρ isexpected to lie in the range 0.75 to 0.85, which accordingto Figure 1 would correspond to a maximum strain rate vof about 0.015 to 0.15 %/hour.

The loading rate factor Ur may be presented as a functionof time to failure Tf as indicated in Figure E-2. The basisfor this figure is taken from a comprehensive paper byBerre and Bjerrum /E-7/, where the experience from testson Drammen clay was presented.

1.6

1.4

1.2

1.0

0.8

0.6

0.4

0.2

0

(Tf)ref (60 minutes)

Time to failure Tf, minutes

UrThreshold for creep

Typical anchor test

-2 -1 0 1 2 3log10(Tf)

Figure E-2 Loading rate factor Ur versus time to failureTf, with (Tf)ref = 60 minutes.

The curve in Figure E-2 has been adjusted slightly to fitwith a time to failure Tf = 60 minutes instead of 140minutes as used in /E-7/. Using the times to failure in theanchor tests and the loading rate factor derived in FigureE-1 a straight-line slope representing the anchor tests hasbeen plotted in Figure E-2. The curved shape for Tf > 60minutes is roughly taken from /E-7/.

E5 Assessment of the cyclic pullout resistanceThe expression for the cyclic pullout resistance Rp,cy of adrag-in plate anchor is

cyScyScyp RRURR ∆+=⋅=, (E-7)

The expression for Ucy then becomes:

( )Scycy RRU /1 ∆+= (E-8)

If no relevant cyclic soil data exist for the site, andexperience from better documented sites with similar soilconditions cannot be drawn upon, a conservativeassessment of τf,cy may be made based on Eq. (E-4)corrected for the effect of cyclic strength degradation. Inorder to account for the possible strength degradation dueto one-way cyclic loading, the net effect of loading rate(Ur - 1) should therefore be multiplied by a cyclicdegradation factor k c. The expression for Ucy then changesto:

Ucy = 1 + k c⋅(Ur - 1) = 1 + k c⋅{(v/vref)n -1} (E-9)

k c is a function of the line tension load history through astorm and the characteristics of the clay. The load historyvaries with water depth, type of floater and mooring lineconfiguration. Therefore the value of k c should be assessedfrom case to case.

Guidance for assessment of both the loading rate factor Urand the cyclic loading factor Ucy can be found in thepublished information about cyclic behaviour of clay, e.g.tests on Drammen clay in /E-4/, on Troll clay in /E-5/ andon Marlin clay in /E-6/. It is noted based on the test resultspresented for the Marlin clay that carbonate content maysignificantly affect the cyclic response of clay. Caution istherefore warranted in the use of experience from tests onnon-carbonate clay, if the actual clay contains more than10 % carbonate.

Guidance NoteBasis for an approximate assessment of the effect of cyclicloading is provided in the following.

Loading rate factor UrAs outlined above the effect of cyclic loading is two-fold, theloading rate effect and the cyclic degradation effect.In a cyclic laboratory test on clay the cycle period is often setto 10 seconds, which means that the load rise time tcy frommean level to the first peak load is 2.5 seconds (= tcy). If thecycle amplitude is high enough to fail the clay specimen duringthat first quarter of the first load cycle (Neqv = 1), thecorresponding cyclic strength τf,cy of the clay divided by thestatic undrained shear strength suD is a measure of the loadingrate factor Ur for the actual clay, i.e.

Ur = τf,cy/su,D (for Neqv = 1).

Figure E-3 presents excerpts of published results from cyclicdirect simple shear tests on the Drammen clay /E-4/, on theTroll clay /E-5/ and on the Marlin clay /E-6/.Figure E-3a) shows the loading rate factor Ur as a function ofthe average shear stress level τa/suD during the test. It is worthnoting that the loading rate effect is most pronounced for τa/suDin the range 0.5 to 0.7, and that for higher shear stress levelsthe effect reduces at an accelerating rate, particularly for thecarbonate type Marlin clay (Unit IIb), which has a carbonatecontent of 15 - 20 % according to /E-6/.Based on the mooring analysis it will be possible to define themean, low-frequency and wave-frequency components of thecharacteristic line tension, such that a basis is obtained forassessment of a likely range for the parameter τa/suD. Typicallythe line tension in a taut mooring system may generate anaverage shear stress level τa/suD in the range 0.5 to 0.8. For thisrange Ur = 1.4 - 1.75 for four of the examples shown in FigureE-3a), but may be as low as 1.2 (or lower) as indicated by thecurve for the Marlin carbonate clay.


DET NORSKE VERITAS

Cyclic loading factor UcyFollowing the strain accumulation procedure as described indetail in /E-4/, and briefly summarised in this Appendix, thecyclic test data may be used for prediction of the cyclic loadingfactor Ucy.In Figure E-3b) and c) the Ucy-factor is plotted for Neqv = 3 andNeqv = 10, respectively. In the latter case this means that if thecalculations leads to failure in cyclic loading for a given cyclicload history, the same effect will be achieved if 10 cycles ofthe extreme load amplitude in the same load history is appliedto the clay.Experience has shown that the cyclic shear strength will oftenbe found for Neqv = 5 - 10, but unless site specific tests havebeen performed it is recommended to make conservativeassumptions about the cyclic loading effect. By conservative ismeant that the strength and plasticity properties of the clayshould be evaluated and compared with the database, that thestress history of the soil profile is assessed, that possiblecarbonate content is accounted for, etc. When looking at rangeof Ur and Ucy reported for the different clays in Figure E-3 it isevident that experience from testing of one clay will notnecessarily be representative of the behaviour of another clayin another geological environment. Unless a site specific cyclictesting programme has been designed and executed, theempirical data like those shown in the figure and elsewhere inthe literature should therefore be used with caution.As a further background for the results shown in Figure E-3Table E-1 gives some characteristics of the tested clay.

Other effectsThe cyclic laboratory tests behind Figure E-1 were carried outon normally consolidated clay (OCR = 1-1.5), but the effect ofOCR on the cyclic bahaviour for so-called one-way cyclicloading (no shear stress reversal), which is a relevantassumption when mooring line tension is considered, ismoderate. Typically Ur and Ucy will be reduced by up to 5 %when OCR increases from 1 to 4, by up to 15 % when OCRincrease from 1 to 7 and by 20 % when OCR increases from 1to 10.The cyclic response will also be affected by the frequency ofloading, e.g. low-frequency versus wave-frequency tensioncomponents. The low-frequency component has typically aperiod, which is about 10 times longer than the wave-frequency component represented in the test results plotted inFigure E-1. Recognising the effect of loading rate an increasein the load rise time tcy from 2.5 seconds to 25 seconds, i.e. onelog-cycle change, will give a reduction in the net cyclic loadingeffect by about 10 %, e.g. a reduction from Ucy = 1.3 to Ucy =1.27.

Loading rate factor U r (for Neqv=1)

0.8

1

1.2

1.4

1.6

1.8

0.5 0.6 0.7 0.8 0.9 1

Average shear stress level τa/suD

Lo

adin

g r

ate

fact

or

U r

Drammen clay

Troll clay (Unit 1)Troll clay (Unit 2)Marlin clay (Unit IIa)Marlin clay (Unit IIb)

Neqv = 1

(a)

Cyclic loading factor Ucy (for Neqv=3)

0.8

1

1.2

1.4

1.6

1.8

0.5 0.6 0.7 0.8 0.9 1

Average shear stress level τa /suD

Cyc

lic

load

ing

fac

tor

U cy

Drammen clayTroll clay (Unit 1)Troll clay (Unit 2)Marlin clay (Unit IIa)

Marlin clay (Unit IIb)

Neqv = 3

(b)

Cyclic loading factor Ucy (for Neqv=10)

0.8

1

1.2

1.4

1.6

1.8

0.5 0.6 0.7 0.8 0.9 1

Average shear stress level τa/suD

Cyc

lic lo

adin

g f

acto

r U cy

Drammen clayTroll clay (Unit 1)Troll clay (Unit 2)Marlin clay (Unit IIa)Marlin clay (Unit IIb)

Neqv = 10

(c)

Figure E-3. Example of cyclic direct simple shear testdata (from /E-4/, /E-5/ and /E-6/).

Table E-1 Characteristics of tested clay (ref. Figure E-3).

Parameter Drammen Troll (Unit 1) Troll (Unit 2) Marlin (Unit IIa) Marlin (Unit IIb)

suD [kPa] 8.6 ≈20 ≈90 ≈10 ≈30zlab.test [m] ≈ 50 10 - 20 20 - 60 5 - 15 15 - 35OCR [-] 1 1.45 1.45 1 1w [%] 52 47-70 18-26 60-90 40-65PI [%] 27 37 20 35-60 30-42

Note: Carbonate content in the Marlin clay (Unit IIa) in the range 8-10 %, in Unit IIb in the range 15-20 %.


+


DET NORSKE VERITAS

E6 Anchor pullout resistance versus soilstrengthThe described approach for assessment of a creep pulloutresistance Rp,cr may be followed up by using the strainaccumulation method to determine the cyclic pulloutresistance Rp,cy. To do this, direct simple shear (DSS)cyclic test data are used, but the anchor pullout resistance,with reference to Figure E-3, replaces the undrained shearstrength, i.e.

su,D = RS = ψ⋅ Rp,i

τf,cy = RC = Ucy ⋅ RS (∆Rcons =0)

τcrit = Rp,cr = ρ ⋅ RS

The factors ψ and ρ need to be established on a case bycase basis until sufficient understanding of theserelationships has been developed in the industry. Thecyclic loading factor Ucy is computed according to theprocedure outlined in Appendix E, Guidance Note, butwith RS replacing su and Rp,cy replacing τf,cy.

Currently ψ=0.80 is suggested as a default value fordetermination of RS from the pullout resistance measuredunder offshore testing conditions, see above.

The design mean line tension Td-mean should be less thanRp,cr

τcr

τ f,cy (Neqv=1)

τf,cy (Neqv=10)

su

0 0.25 0.50 0.75 1.0τa / su

0

0.25

0.50

0.75

1.0

1.25

1.50

τ f,cy

/ su

τ cy /

s u

Parallel to anchor design: RS (~ su) = ψ . Rp,iRC (~ τ f,cy) = Ucy

. RSRp,cr (~ τcr) = ρ . RS

EXAMPLE

Figure E-4 Anchor pullout resistance versus soilstrength.

E3 References/E-1/ Andersen, K. H. and Lauritzen, R. (1988), Bearing

capacity for foundations with cyclic loads, ASCEJournal of Geotechnical Engineering, Vol. 114,No. 5, May, 1988, pp. 540-555.

/E-2/ Bea, R.G. and Audibert, J.M.E. (1979),Performance of dynamically loaded pilefoundations, Proceedings from BOSS’79, PaperNo. 68, pp. 728-745. London.

/E-3/ Kraft, L.M., Cox, W.R. and Verner, E.A. (1981),Pile load tests: Cyclic loads at varying load rates,American Society of Civil Engineers, Vol. 107,No. GT1, January 1981, pp. 1-19.

/E-4/ Briaud, J-L and Garland, E. (1983), Loading ratemethod for pile response in clay, American Societyof Civil Engineers, Vol. 111, No. 3, March 1985,pp. 319-335.

/E-5/ By, T. and Skomedal, E. (1992), Soil parametersfor foundation design, Troll platform, Behaviour ofOffshore Structures BOSS'92, pp. 909-920.

/E-6/ Jeanjean. P, Andersen K.H. and Kalsnes B. (1998),Soil parameters for design of suction caissons forGulf of Mexico deepwater clays , OffshoreTechnology Conference, Paper OTC 8830, pp.505-519. Houston.

/E-7/ Berre, T and Bjerrum, L. (1973), Shear strength ofnormally consolidated clays , Proc. 8th InternationalConference on Soil Mechanics and FoundationEngineering, Vol. 1.1, pp. 75-80. Moscow.


DET NORSKE VERITAS

Appendix F Uplift angle at the seabed

F1 GeneralThe anchor line in a mooring system may be split intothree parts, one part embedded in the soil, a second partresting on the seabed, and a third part suspended in water.

The length of anchor line lying on the seabed at any timeduring anchor installation will be a function of at least thefollowing factors

− the configuration of the anchor line− the actual length of line between the anchor shackle

and the pulling source (stern roller)− the actual line tension− the anchor line catenary (suspended part)− the inverse catenary of the line (embedded part)− the penetration trajectory of the anchor (position of

the shackle)

At some point the length of the seabed part becomes zeroand a further increase in the line tension or decrease indistance will result in a situation where the anchor lineintersects the seabed under an uplift angle (α), seeFigure 2.

It has been a recognised understanding for some time thatboth fluke anchors and drag-in plate anchors can beinstalled with a uplift angle, although only a small numberof tests have been performed with a non-zero uplift angle.There is a potential for significant cost savings if a safeinstallation uplift angle can be documented and agreedupon. In the following, guidelines are given for achievingthis.

F2 Assessment of a safe uplift angleNon-zero uplift angles during installation typically occurwhen anchors are installed using a short scope of line,either by bollard pull (and blocked line) or by winch pull(from a stationary vessel).

An anchor should under no circumstances be set with ananchor line giving an initial non-zero uplift angle fromstart of the installation. This would reduce the possibilityfor the anchor to enter the soil. As a minimum, theembedment of the fluke should be 3 fluke widths (WF)before uplift is applied. This will also limit the possiblemaximum uplift angle for all practical means consideringthe path reaching an ultimate depth. An uplift angleexceeding 10° should not be expected during installationof a drag-in plate anchor according to this procedure, evenif the anchor approaches its ultimate depth.

The penetration path is only slightly affected by the upliftangles following upon the adoption of the installationprocedure described above. If the anchor were to beinstalled to the ultimate depth using this procedure, theultimate depth reached would be reduced only by a fewpercent as a result of the increased uplift angle at theseabed. Considering that the anchor resistance is mainly afunction of the penetration depth, this means that thechange in anchor resistance for most installation cases isnegligible.

The anchor line may have either a wire or a chainforerunner, and the effect of using one type of line or theother affects the behaviour of the anchor. An anchorpenetrated with a wire will reach a larger ultimate depththan an anchor with a chain, since the soil cuttingresistance is less for a wire than for a chain, see sketch inFigure 2. The maximum acceptable uplift angle for ananchor installed to the ultimate depth with a wireforerunner therefore becomes larger than with a chainforerunner.

The penetration path becomes shallower the higher theuplift angle at the seabed is. The maximum possible upliftangle (αmax) is the angle, which makes the anchor drag at aconstant depth, and gradually pulls the anchor out of thesoil for higher angles. Tentatively, a safe α-angle may beset to 50% of αmax, although limited to α = 10.

The effect on the installation anchor resistance Ri ofincreasing the uplift angle from 0°to θ/2 may be assumedto vary linearly according to the following simpleexpression

)/1( max0,, αααα −= =LL RR

(valid for α<αmax/2 and α<10°)

(F-1)

where RL is the contribution to the installation anchorresistance Ri from the embedded part of the anchor line.


DET NORSKE VERITAS

Appendix G General requirements to soil investigation

G1 Geophysical surveysThe depth of sub-bottom profiling should correspond tothe depth of rock or the expected depth of anchorpenetration. The seismic profiles should preferably be tiedin to geotechnical borings within the mooring area, whichwill improve the basis for interpretation of the results fromthe geophysical survey. Emphasis should be on very highresolution tools for surveying the top 50 metres ofsediments.

G2 Geotechnical surveysThe soil investigation should be planned and executed insuch a way that the soil stratigraphy can be described insufficient detail for both the anchor and the anchor lineanalysis. The required depth coverage will vary from caseto case, see Chapter 6.

The extent of the soil investigation, sampling frequencyand depth of sampling/testing, will depend on a number ofproject specific factors, e.g. the number of anchorlocations, soil stratigraphy and variability in soilconditions with depth and between the potential anchoringpoints, water depth, sea floor bathymetry, etc.

The challenge to secure soil samples of sufficient qualityto determine realistic strength parameters increases withthe water depth, and the continuous efforts to improve theexisting, and develop new, sampling procedures should befollowed. Nevertheless, in situ testing will becomeincreasingly important for mapping of the soil conditionsin deep waters.

If soil layering is such that the layer sequence and thevariation of thickness and layer boundaries will become animportant anchor design and installation consideration, itmay be necessary to document the soil layer sequence ateach anchor location. The thickness of all significantlayers, and the thickness variation between the anchoringlocations, should be known with reasonable accuracy priorto the design of the anchor foundation.

Piezocone penetration testing (PCPT) normally bringsvaluable and useful information about soil stratigraphy, butthe undrained shear strength derived from such tests willbe uncertain if the PCPT results are not calibrated againstlaboratory strength tests on recovered soil samples. Ifgenerally adopted correlation factors are used theundrained shear strength derived will be affected by theuncertainty in this correlation factor.

For the anchor design, most weight should be given to theundrained shear strength derived from direct simple shear(DSS) and unconsolidated undrained (UU) triaxial tests.These types of test are considered to give the mostrepresentative estimates of the intact undrained shearstrength of the clay. Clay sensitivity (St) is also asignificant soil parameter in the anchor design, whichrequires companion determinations (on the same soilspecimen) of intact and remoulded shear strengths, eitherby UU triaxial tests or by fall-cone tests.

One possibility to acquire soil information, in addition tothe data obtained from a standard soil investigation asaddressed above, may be to perform a few instrumentedanchor tests at the actual location. The objective with thesetests would be to obtain data on anchor penetrability,depth-drag relationships, etc as relevant for the design ofdrag-in plate anchors at the actual location.

The test anchor should be sized such that it can bepenetrated to the ultimate penetration depth in the actualsoil, and then retrieved with the available installationequipment. Such anchor tests should provide relationshipsbetween line tension (preferably at the anchor shackle),depth and drag versus time. Anchor pullout tests fordetermination of the performance ratio Pr may also beincluded in the test programme. Regarding considerationsrelated to planning and performance of anchor testsreference is made to Appendix C.

For calculation of the effect of cyclic loading on thelongterm anchor resistance, it is recommended to carry outstatic and cyclic undrained DSS tests. These tests shouldbe carried out on representative soil samples of highquality, which shall be subjected to stress conditions,which simulate the in situ conditions as closely as possible.A combined static/cyclic test programme should allowdetermination of the strength of the soil under the range ofloading conditions to be covered by the anchor design, e.g.cyclic tests with a representative combination of averageand cyclic shear stresses. The test programme should allowthe construction of a strain contour diagram, as requiredfor calculation of the cyclic shear strength (τf,cy), see /E-4/and Appendix E for details. If site specific soil data are notprovided for assessment of the cyclic loading effect, aconservative assessment of this effect is warranted.

Documents

DnV-RP-E302 - Design and Installation of Plate Anchors in Clay [2000]