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HSEHealth & Safety

Executive

Development of design guidance

for neoprene-lined clamps for

offshore application

Phase II

Prepared by MSL Engineering Limited

for the Health and Safety Executive 2002

RESEARCH REPORT 031

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Crown copyright 2002

First published 2002

ISBN 0 7176 2577 X

All rights reserved. No part of this publication may bereproduced, stored in a retrieval system, or transmitted inany form or by any means (electronic, mechanical,photocopying, recording or otherwise) without the priorwritten permission of the copyright owner.

Applications for reproduction should be made in writing to:Licensing Division, Her Majesty's Stationery Office,St Clements House, 2-16 Colegate, Norwich NR3 1BQor by e-mail to [email protected]

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FOREWORD

This document has been prepared by MSL for three sponsoring organisations:

Health and Safety Executive

ExxonMobil

Shell U.K. Exploration and Production.

In addition, MSL themselves partly funded the work described herein.

The document is concerned with a test programme investigating the slip capacity of

neoprene-lined clamps. In Phase I of the project, which is reported separately, a total of

sixteen full-scale tests were conducted at Memorial University of Newfoundland, Canada.

Based on the results of the Phase I tests, interim recommendations were made for the

estimation of frictional coefficients. The results indicated some surprising effects, and furthertests were recommended. The further tests have now been conducted under Phase II of the

project and are reported herein.

A project steering committee including representatives of the sponsoring organisations

oversaw the work and contributed to the development of this document. The following

individuals served on the committee:

Mr P Bailey

Mr J Bucknell

Dr A Dier

Mr D Galbraith (Chairman)

Mr M Lalani

Mr B McCullough

The Project Manager at MSL was Mr J Bucknell who carried out the work with guidance and

support from Dr A Dier and Dr K Chen. The tests were conducted at Memorial University,

Newfoundland.

The recommendations presented in this document are based upon the knowledge available at

the time of publication. However, no responsibility of any kind for injury, death, loss,

damage or delay, however caused, resulting from the use of the recommendations can be

accepted by MSL Engineering or others associated with its preparation.

The participants do not necessarily accept all the recommendations given in this document.

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JIP DEVELOPMENT OF DESIGN GUIDANCE FOR

NEOPRENE LINED CLAMPS FOR OFFSHORE

APPLICATION PHASE II

FINAL REPORT

CONTENTS

Page No

FOREWORD .............................................................................................................................3

1. INTRODUCTION..........................................................................................................6

1.1 General ...............................................................................................................6

1.2 Summary of Phase I Programme........................................................................7

2. OBJECTIVE AND SCOPE OF TESTS ......................................................................10

3. DESCRIPTION OF TEST CONFIGURATION AND PROCEDURES.....................123.1 Clamp Specimen ..............................................................................................12

3.2 Test Rig ............................................................................................................12

3.3 Test Instrumentation.........................................................................................15

3.4 Test Procedures ................................................................................................17

3.4.1 Pre-Testing Procedures ........................................................................17

3.4.2 Testing Procedure.................................................................................18

3.4.3 Post-Test Procedures ............................................................................19

3.5 Loading Schedule .............................................................................................19

4. RESULTS.....................................................................................................................234.1 Failure Criterion Quasi-Static Load

(T1A &T1B; T1, T4 & T4A) ...........................................................................23

4.2 Bolt Pre-Load (T1A & T1B;T1, T2 & T3) ......................................................23

4.3 Neoprene Hardness (T1A, T18 & T18A).........................................................25

4.4 Cyclic Loading (T4C & T1C) ..........................................................................26

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5. DISCUSSION AND GUIDELINES FOR DESIGN/ASSESSMENT.........................30

5.1 Discussion ........................................................................................................30

5.2 Design/Assessment Guidelines ........................................................................33

6. CONCLUDING REMARKS .......................................................................................36

REFERENCES

FIGURES FOR SECTION 4

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1. INTRODUCTION

1.1 General

This report is concerned with a second phase of a Joint Industry Project (JIP)

investigating the experimental slip capacity of neoprene-lined clamps.

The use of such clamps in the offshore industry has been, and continues to be,

widespread throughout the world. The following applications of such clamps may be

given:

attachment of retrofitted appurtenances such as conductors and risers temporary attachments for lifting purposes, especially pipelines repair of damaged members (dented or corroded) attachment of new structural members.

Neoprene-lined clamps contain a liner that lies between the clamp steelwork and the

enclosed member. The liner provides tolerance against lack of fit of the clamp saddle

around the tubular brace. In general, the linear is made of polychloroprene (neoprene)

sheet that is bonded to the inner surface of the clamp saddle plates. The neoprene

liner is usually plain for structural connections designed to transmit axial or rotational

loads, although ribbed linings are sometimes used to accommodate potentially large

lack of fit tolerances.

Stressed neoprene-lined clamps rely on applied stud bolt pre-loads to generatecompressive forces normal to the interface between the clamp liner and the surface of

the clamped brace. The strength is considered to be dependent on the magnitude of

the normal force, the relative stiffness of the steel and liner and the effective

coefficient of friction at the liner/brace interface.

Despite the widespread use of neoprene-lined clamps through the world, there were

only limited data, and no data in the public domain, on the slip capacity of these

clamps. As part of a Joint Industry Project conducted by MSL entitled

Demonstration Trials of Diverless Strengthening and Repair Techniques, a static

slip test on a neoprene-lined clamp exhibited a slip capacity significantly less than

that expected from the guidance available at that time.

It was against the above background that MSL launched this current Joint Industry

Project. The project was intended to generate test data so that more reliable design

guidance could be formulated, both for the rational assessment of the reliability of

neoprene-lined clamps currently in service and for the safe design of such clamps in

future applications.

With the support of HSE and two major North Sea operators, Phase I of this current)JIP was concluded in May 1999 with the issue of a final report (1 to the sponsoring

organisations. Phase I covered a programme of 16 full-scale neoprene-lined clamp

tests. The tests in Phase I encompassed both axial and torsional loading and weredesigned to investigate the influence of a variety of parameters, including the bolt

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load, test repeatability, neoprene thickness, pipe surface condition, clamp

length/diameter ratio and pipe radial stiffness. Interim guidance was prepared on the

basis of the test results generated in Phase I. Further tests were recommended in the

following three areas where new data would lead to a substantial and significant

enhancement of the interim guidance created in Phase I:

Clamp tests with imposed interface pressures lower than those used in thePhase I programme.

Clamp tests with neoprene liners having hardness values different to thatadopted in Phase I.

Clamp tests with loading rates different to those used in Phase I.In light of the extensive use of neoprene-lined clamps, and the benefits that will result

through generation of data in the above three areas, this Phase II of the subject JIP

was instigated.

1.2 Summary of Phase I Programme

The Phase I testing programme involved a total of 16 tests. The programme was

designed to investigate the influence of the following parameters on the slip strength

of neoprene-lined clamps:

stud bolt preload (Tests T1, T2 and T3) repeatability (Tests T1 and T4) failure criterion (Tests T1 and T4A) surface conditions of the clamped tubular member (Tests T1, T7 and T8) brace stiffness (D/T) ratio (Tests T1, T9 and T10) clamp length to diameter (L/D) ratio (Tests T1, T11 and T12) torsional v. axial slip (Tests T1, T13, T14 and T15) thickness of the neoprene liner (Tests T1 and T16) clamp brace pinching (liner/brace interference) (Tests T16 and T17).The Phase I test programme is summarised on the pullout table at the back of this

document. The primary Phase I finding is that coefficient of friction for neoprene

lined clamps is substantially below the range of values adopted in practice. Hence,

some existing structural neoprene-lined clamps potentially have capacities that may

be less than the design intent. In addition to the primary findings, the following

results were also achieved during the Phase I tests:

The failure load for all axial tests was defined as the position at which theload-slip curve was seen to deviate substantially from the trendline defining its7

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initial slope. The failure torque was defined as the position with a relative

rotation of 0.45 between the clamp and tubular for pure torsional load orwhen an axial displacement of 1.25 mm was reached under the combination of

tensile and torsional loads.

The axial tests were repeatable and similar failure loads could be derived fromtests with identical conditions according to the definition of failure. The increase in the applied stud bolt load did not lead to a corresponding

increase in the clamp axial capacity, at least for a preload level of 40% - 60%

of the stud bold yield strength. For a given stud bolt pre-load level, significant

drop in bolt load with the increase of the slip was shown, although such

reduction was not so marked for the initial 4 to 5 mm slip.

The clamp axial slip capacity varied little with either the pipe surfacecondition or the pipe radial stiffness (i.e. pipe diameter to thickness ratio)

according to the definition of failure criterion.

As the clamp length was reduced, with a corresponding reduction in totalapplied bolt load, the axial capacity of the clamp was also seen to reduce. The

reduction in axial capacity was proportional to the reduction in total applied

bolt load.

No increase in clamp slip capacity was obtained by increasing the neoprenethickness. Reducing the circumferential length of the pipe/neoprene interface

could lead to an increase of the clamp slip capacity.

Application of the torsional load would reduce the clamp slip capacity.With the above observations achieved in the Phase I test programme, interim

guidance was formulated for clamp slip capacities under axial load alone, torsional

moment alone and combined axial and torsional loadings respectively.

The developed guidance had to be considered as being of an interim nature until

further data became available due to the unexpected slip behaviour of the clamp,

particularly with regard to the relationship between applied bolt load and clamp

capacity. There were insufficient test data to permit a proper clarification of the role

of bolt load.

The interim guidance can be considered conservative, particular for small bolt pre

loads where no data exists in the Phase I test programme. For combined axial and

torsional loadings, it could also lead to a rather conservative prediction of the

torsional capacity at high values of co-existing axial load.

The combination of lower bolt pre-loads, applied loading rate effects typical of those

due to wave action, and possibly liners of greater Shore hardness may give higher

apparent coefficients of friction. It is for this purpose that Phase II of this current JIP

was launched.

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The remainder of this document presents the Phase II JIP test programme in detail,

viz:

Section 2 Objective and Scope of Tests Section 3 Description of Test Configuration and Procedures Section 4 Test Results Section 5 Design/Assessment Guidance Section 6 Concluding Remarks.

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2. OBJECTIVE AND SCOPE OF TESTS

Objective

The objective of Phase II of the JIP is to conduct a programme of tests that, in

combination with the Phase I results, will enable cost-effective, robust and safeguidance for the design of neoprene-lined clamps to be established.

Testing Programme

The testing programme involves a total of 6 tests, as summarised in Table 2.1. The

programme has been designed to investigate the influence of the following parameters

on the slip strength of neoprene-lined clamps:

Interface pressure - Tests T1A and T1B (along with Phase I Tests T1, T2 &T3)

Neoprene hardness - Tests T18 and T18A (along with Test T1A) Cyclic loading effect (Test T4C) Failure definition (full cyclic and half cyclic loading) (Test T1C)Phase I test rig was utilised for the first two tests on T1A and T1B. In order to apply

a sinusoidal-type loading in both tensile and compressive direction (tests on T4C and

load steps 1 through 5 of tests on T1C), the Phase I test rig was slightly modified to

remove slack in the bearings.

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CH10010R006Rev1January2001

Test

No.

Nature

OfTest

Loading

C

lamp

L

ength

(mm)

Brace

D/T

(mm)

No.Of

Bolts

Bolt

Size

(nom.)

Bolt

Load

(%fy)

Brac

e

Surfa

ce

Condition

Neoprene

Thickness

(mm)

Neop

rene

Hardness

(IRH

D)

T1A

Incremental

800

324/17

8

M36

20%

Blacko

xide

10

60

T1B

Incremental

800

324/17

8

M36

10%

Blacko

xide

10

60

T4C

FullCyclic&

HalfCyclic

800

324/17

8

M36

20%

Blacko

xide

10

60

T18

Incremental

800

324/17

8

M36

20%

Blacko

xide

10

70

T18A

Incremental

800

324/17

8

M36

20%

Blacko

xide

10

50

T1C

HalfCyclic

800

324/17

8

M36

20%

Blacko

xide

10

60

Notes:IRHDInternationalRu

bberHardnessDegree

Table2.1:

Summ

aryofTestProgrammeforP

haseII

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CH10010R006 Rev 1 January 2001

3. DESCRIPTION OF TEST CONFIGURATION AND PROCEDURES

A complete description of the test configurations and procedures is provided in the

Annex. Here, a summary of the more salient points is given.

3.1 Clamp Specimen

The clamp test specimen, as used in the Phase I trials, is illustrated in Figure 3.1. The

clamp was structurally typical of many clamps used for the retrofitting of risers to

existing installations and for the handling of pipe spools.

bolt

800 mm

600

460mm

saddle platesside plates

neoprene

lining

stiffeners

stud bolt centrelines (8 No. total)mm

Figure 3.1: Clamp Test Specimen

3.2 Test Rig

Phase II tests were restricted to pure axial loading of the clamp along the longitudinal

axis of the pipe. The same test rig configuration used for application in axial tensile

loading in the Phase I tests, as shown in Figure 3.2 and illustrated in Figure 3.3, was

utilised in the Phase II tests. Modifications to the end connections have been made topermit load reversal. The modified end connection and load cell tubular interface

(flange joint) are illustrated in Figure 3.4.

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cell

tubularsection

clamptest

specimen

hydraulic

actuator

load

cell

tubularsection

load

hydraulic

actuator

Figure3.2:

TestRig

ConfigurationforAxialLoad

ingOnly

1

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Figure 3.3: Photograph of Axial Load Rig

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Figure 3.4: Modified End Connection and Load Cell Tubular Interface

3.3 Test Instrumentation

Strain Measurement

Strains were measured by means of linear 120-Ohm electrical resistance straingauges.

A total of six strain gauges were mounted on the pipe to verify loading of thespecimen. The locations and identification numbers for each of strain gauges

are shown in Figure 3.5. The identification numbers of each of the stud bolts

are also shown in Figure 3.5.

Each bolt was instrumented with two strain gauges in a half bridgeconfiguration to monitor the total load on each bolt.

Dummy gauges were provided on blocks of steel to allow for temperaturecompensation of measured values.

Displacement Measurement

Displacements were measured by means of temperature compensating LinearVoltage Displacement Transducers (LVDTs).

A total of five LVDTs were positioned about the specimen to record therelative displacement of the clamp with respect to the pipe. Figure 3.6

presents the locations of the LVDTs.

LVDTs were aligned to ensure that displacement were measuredperpendicular to specimen or reaction surfaces.

Sufficient travel lengths were specified for the LVDTs such that the entireloading regime could be recorded.

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ensuring that the split line on each side of the clamp remained even along the

length of the clamp and approximately equal either side of the clamp. A

sufficient length of bolt protruded above the top plate of the upper clamp half

to accommodate the hydraulic stud bolt tensioning system.

5. The leads from the strain gauges on each of the bolts and on the tubular wereconnected to the appropriate channels of the data acquisition system.

6. The operation of all instrumentation and data acquisition system was checked.

7. Initial datum (zeroed) readings of all instrumentation were taken.

8. The bolts were hand-torque, evenly, using a standard wrench.

Tensioning of Stud Bolts

The stud bolts were simultaneously tensioned using the Hydratight hydraulic

tensioning system. The procedure for the tensioning involved a three-stage pressure

application. A qualified Hydratight technician supervised the tension operation.

Strain gauge readings from each stud bolt were continuously monitored throughout

the tensioning procedure to confirm:

(a) the desired average bolt load had been achieved to within 5%(b) maximum variation of load between bolts did not exceed 10% of the target

load.

Installation of the Test Rig

The specimen was installed into the appropriate test rig and the LVDTinstrumentation was set up.

The operation of all instrumentation and data acquisition system was checked. Initial datum (zeroed) readings of all instrumentation was again checked. The specimen was bedded down by applying load cycles not greater than 5%

of the estimated failure load.

3.4.2 Testing Procedure

After completion of the steps described in Section 3.4.1, the application of loading

proceeded in accordance with the following procedure:

Throughout all loading and unloading operations, data from eachinstrumentation point, including all strain and displacement gauges, were

continuously recorded, calibrated and logged by the computerised data

acquisition system. Visible and/or audible events were manually recorded and

photographed as appropriate.

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20

Details of loadings are given in the Annex in the form of load-time plots. Here, the

loadings are summarised as follows:

Tests T1A, T1B, T18 & T18A

These four tests were comprised of two loading cycles each.

Loading cycle one was incremental loading of 5-minute durations commencing at 50

kN, proceeding to 75 kN, then 100 kN. After this, the load increments were reduced

to intervals of 10 kN, continuing until a relative displacement of 4 mm of the tubular

in the clamp was measured.

Loading cycle two involved a linear increase of the load, at a rate of 700 lbs

(approximately 31.2 kN) per minute, from zero load to a point where 20 mm slip was

measured between the clamp and the tubular, after which the load was gradually

reduced back to zero.

Test T4C

A total of 7 loading steps were comprised in this test. All loading steps were similar,

sinusoidal loading in tensile and compressive directions about a mean load of zero.

The only variations between load steps were frequency and amplitude of the cyclic

load and the total number of times each load was applied (cycle numbers). Figure 3.7

presents a typical loading cycle in Test T4C. Specifications of each of the 10 loading

cycles are summarised in Table 3.2. The amplitudes, periods and the steepness of

1/16 are explained below.

-40000

-30000

-20000

-10000

0

1 0 0 0 0

2 0 0 0 0

3 0 0 0 0

4 0 0 0 0

167

133

199

265

331

397

463

529

595

661

727

793

859

925

991

1057

1123

1189

1255

1321

1387

1453

1519

1585

1651

T ime (s )

Load(lbs)

-2.5

-2

-1.5

-1

-0.5

0

0. 5

1

1. 5

2

Displacement(mm)

L O A D L B S

A V E R A G E B O LT L O A D L BS

L V D T 1 M M

Figure 3.7: Typical Load Step in Test T4C

Test T1C

Test T1C included a total of 10 load steps, the first five being sinusoidal and the latter

being the tensile load, i.e. positive portion only, of similar cycles as shown in Figure

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21

3.8. The variations in the frequency, amplitude and number of cycles are presented in

Table 3.2.

-5000

5000

15000

25000

35000

45000

55000

156

111

166

221

276

331

386

441

496

551

606

661

716

771

826

881

936

991

1046

1101

1156

1211

1266

1321

1376

Time (s)

Load(lbs)

-1

-0.5

0

0 .5

1

1 .5

2

Displacement(mm)

LO A D LB S

AV ER AG E BO LT LO AD LB S

L V D T 1 M M

Figure 3.8: The 10th Loading Step of Test T1C

Test T4C

Loading

Cycle

No. of

Cycles

Amplitude

(kN)

Period

(seconds)

Comment

1 10 24 7.0 Steepness of 1/16

2 10 51 9.0 Steepness of 1/16

3 5 99 10.5 1-year return wave

4 5 150 13.2 Steepness of 1/16

5 5 188 12.0 100-year return wave

6 5 220 12.0 Steepness of 1/16

7 5 220 10.5 Steepness of 1/16

Test T1C (Loading Cycles 15: TensionCompression; 610 Tension only)

Loading

Cycle

No. of

Cycles

Amplitude

(kN)

Period

(seconds)

Comment

1 & 6 10 24 7.0 Steepness of 1/16

2 & 7 10 51 9.0 Steepness of 1/16

3 & 8 5 99 10.5 1-year return wave

4 & 9 5 150 13.2 Steepness of 1/16

5 & 10 5 188 12.0 100-year return wave

Table 3.2: Specifications of Loading Steps in Tests T4C and T1C

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The amplitudes, periods and steepness values in Table 3.2 were proposed, in

consultation with the Project Steering Committee, to be similar to what existing

clamps may experience. For this purpose, it was assumed that the test specimen had

been used to attach a 26retrofit riser to a typical UK Southern North Sea platform.The clamp was assumed to be located at an elevation close to the first horizontal

frame below the waterline (-8.0 m). The water depth was taken as 35.4 m. It wasassumed that the clamp had been designed with an interface pressure of 3.2 MPa (i.e.

consistent with Test T1A). It was further assumed that, in the design of the clamp, a

value of 0.2 was used for the coefficient of friction and a factor of safety of 1.7 was

applied to the extreme event load. On this basis, the clamp notional design axial slip

capacity is 188 kN.

The design wave height and period used for the 100-year event were 15.1 m and 12

seconds respectively. At 8 m water depth the lateral wave load on the riser was 25.9

kN/m. The clamp was therefore designed to resist the wave load on approximately

7.3 m of riser. The wave height and period used for the 1-year event were 10.8 m and

10.5 seconds respectively. At 8 m water depth, the load on the riser was 13.6 kN/m,therefore, the load on the clamp during the 1-year return period wave was 99 kN. For

the determination of the wave periods associated with the intermediate load steps a

wave steepness of 1/16 has been assumed.

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4. RESULTS

This section presents the main results of the various tests conducted in Phase II. The

results are grouped, according to the parameter under investigation, in the following

subsections. Further details may be found in the Annex, especially of the condition of

the liner following each test.

4.1 Failure Criterion Quasi-Static Load (T1A &T1B; T1, T4 & T4A)

It has been observed from the tensile tests in Phase I that the slip behaviour of the

clamp is extremely ductile. Typical load-slip curves obtained from Tests T1 and T4

in Phase I are reproduced in Figure 4.1. In order to provide greater accuracy in the

definition of the failure criterion, Test T4A was also carried out in Phase I. Tests T1,

T4 and T4A had a similar test programme except for the load application rate.

Compared with Tests T1 and T4, where the load was gradually and continuously

ramped at a slow rate, the load in Test T4A was applied in increments with a period of5 minutes between each load increment. It was observed during Test T4A that sliding

did not stop during the hold periods and equilibrium was never established, see Figure

4.2. Due to the creep effect observed in Test T4A, the failure load could not be

defined by selecting a certain amount of limiting slip.

Based on the observations from Tests T1, T4 and T4A, a failure criterion was defined

in Phase I for clamps under axial loading. The failure load was defined as the

position at which the load-slip curve was seen to deviate substantially from the

trendline defining its initial slope.

Tests T1A and T1B in Phase II have an identical test frame and specimenconfigurations to those of T1, T4 and T4A in Phase I. However, in Tests T1A and

T1B, the total stud bold loads were lower. The load-slip curves of Tests T1A and

T1B are presented in Figures 4.3 and 4.4 respectively. The figures show the mean

slip of the two clamp halves for each test. The initial trendlines are also shown. Once

again, a ductile form of slip can be seen in each of the tests.

In loading cycle one of Tests T1A and T1B, relatively low axial loads were applied

with a low loading rate. These tests called for incremental loading until 4 mm of

displacement was measured between the tubular and the clamp, whereupon the load

was reduced to zero. It can be observed from Figures 4.3 and 4.4 that the unloading

paths are approximately parallel to the initial slope.

On the basis of the above observations, the failure criterion defined in Phase I for

axial quasi-static load is retained herein.

4.2 Bolt Pre-Load (T1A & T1B;T1, T2 & T3)

Tests T1A and T1B, together with Tests T1, T2 and T3 conducted in Phase I,

represent an investigation of the influence of bolt pre-load on clamp axial slip

capacity. For Test T1A the bolts were tensioned to a nominal pre-load of 20% of

tensile yield. In Test T1B the nominal pre-load of each bolt was 10% of tensile yield.

As reported in the Phase I final report, the pre-load of each bolt in Tests T1, T2 andT3 were 40%, 50% and 60% of tensile yield, respectively. It was observed in Phase I

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tests that a pre-load of 60% of tensile yield caused excessive bulging of the liner

during bolt preload application and damage of the neoprene liner during slip. Hence,

50% of tensile yield was taken as the design limit of the bolt pre-load level.

Bolt Pre-load

The bolts were simultaneously tensioned using Hydratight hydraulic tensioning tools.

The loads in the bolts were continuously monitored during the tensioning operation by

means of the attached strain gauges, each bolt having been previously calibrated to

80% of yield. The average applied pre-loads at the start of Tests T1A and T1B are

presented in the table below, where the pre-loads at the start of Tests T1 and T2 in

Phase I are also included for reference.

Test IDApplied Bolt Pre-Load at Start of Test (KN)

Average Total % of tensile yield

T1A 106 848 20%

T1B 55 440 10%

T1 212 1696 40%

T2 274 2192 50%

Bolt Load Variation

The variation of the pre-load in each bolt for Test T1A is shown in Figure 4.5.

Similar variations in bolt loads were observed for Test T1B. The variation of the

average bolt load for each of Tests T1A and T1B is shown in Figure 4.6. The plotsshown a significant drop in bolt load over the duration of the tests, however, for the

first 4-mm of slip, the reduction is not so marked. The reduction of bolt load was also

observed in Phase I tests.

In all quasi-static tensile tests the variation in the bolt loads followed a similar pattern.

The bolts at the end of the clamp from which the pipe was pulled, numbers 1 and 8

(see diagram below), experienced an immediate fall-off in load. At the other end of

the clamp bolts 4 and 5, initially, see a small increase in load, until about 3 mm of

clamp displacement relative to the pipe. They then see a similar rate of load loss as

the other bolts. The central bolts, numbers 2, 3, 6, and 7, see little or no loss in load

until about 3 mm of relative displacement, at which time a similar rate of loadreduction to the end bolts occurs. A similar bolt load variation pattern was also

observed in Phase I tests.

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Slip

The load-slip behaviour recorded in Tests T1A and T1B are shown in Figures 4.3 and

4.4 respectively. The load-slip curves for Tests T1 and T2 in Phase I are reproduced

in Figure 4.1. The figures show the mean slip of the two clamp halves for each test.

A ductile form of slip can be seen in each of the test, as already mentioned. Thelinear trend line through the initial slope of the curves has been plotted on Figures 4.3

and 4.4. The failure load for each test, based on the definition discussed in Section

4.1, is given in the table below, together with the corresponding apparent friction

coefficient. Those obtained in Phase I for Tests T1 and T2 are also included for

reference. The apparent friction coefficient is defined as the failure load divided by

the total applied bolt pre-load per clamp half. As reported in Phase I final report, the

design capacities for Tests T1 and T2 were estimated as 441 kN and 552 kN

respectively (= 0.2, factor of safety = 1.7). The corresponding estimated designcapacities for Tests T1A and T1B are 220 kN and 110 kN respectively.

Test ID Failure Load Apparent Friction Coefficient

T1A (Fb= 0.2 Fy) 150 0.088

T1B (Fb= 0.1 Fy) 115 0.131

T1 (Fb= 0.4 Fy) 150 0.044

T2 (Fb= 0.5 Fy) 150 0.034

During the failure load determination for Tests T1A and T1B, the load cycle 2 curves

in Figures 4.3 and 4.4, which more actually represent the practical load application

rate, were utilised. The load-slip behaviour of each of the Tests T1A, T1B, T1 andT2 are shown, for comparison, in Figure 4.7. Given the feature that all these fours

tests behaved very similar shown in Figure 4.7, the failure load of Test T1B presented

in the above table may be considered to be conservative.

It can be observed that the increase in the applied stud bolt load does not necessarily

lead to a corresponding increase in the clamp axial capacity. For a bolt load of 20%

of tensile yield and above, the nature and magnitude of clamp failure remain similar.

This, in turn, results in a reduction in the apparent friction coefficient for each of the

Tests T1A, T1 and T2.

4.3 Neoprene Hardness (T1A, T18 & T18A)

Tests T18 and T18A were carried out to assess the effect of neoprene hardness on

clamp displacement by comparison with Test T1A of Phase II. For Test T1A the

neoprene liner had a hardness of 60 IRHD. The neoprene hardness in Tests T18 and

T18A were 70 IRHD and 50 IRHD respectively.

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Bolt Pre-Load

The bolts were simultaneously tensioned, using Hydratight hydraulic tensioning tools.

The load in the bolts was continuously monitored during the tensioning operation by

means of the attached strain gauges, each bolt having been previously calibrated to

80% of yield. The applied bolt pre-loads at the start of Tests T1A, T18 and T18A aregiven in the table below.

Test Neoprene Applied Bolt Pre-Load at Start of Test (KN)

ID Hardness Average Total % of tensile yield

T1A 60 106 848 20%

T18 70 100 800 20%

T18A 50 105 840 20%

Bolt Load Variation

Variation of the bolt pre-load for Tests T18 and T18A followed a similar pattern to

that for Test T1A shown in Figure 4.5. The description of the bolt pre-load variation

and observations therefrom are discussed in Section 4.2, above.

Slip

The load-slip behaviour recorded in Tests T1A, T18 and T18A are shown in Figure

4.8. The curves show the mean slip of the two clamp halves for each test. Once

again, a ductile form of slip can be seen in each of the tests. The linear trend lines

through the initial slope of each curve have been plotted on the figure and the slopescan be seen to correlate with neoprene hardness. The failure load for each test, based

on the definition discussed in Section 4.1, is tabulated below, together with the

corresponding apparent friction coefficient. The apparent friction coefficient is

defined as the failure load divided by the total applied bolt pre-load per clamp half.

The estimated design capacity for all three tests is 220 kN.

Test ID Neoprene

Hardness

Failure Load

(kN)

Apparent Friction

Coefficient

T1A 60 150 0.088

T18 70 160 0.100

T18A 50 130 0.077

The correlation between neoprene hardness and the apparent friction coefficient is

presented in Figure 4.9. It can be seen that the apparent friction coefficient increases

with neoprene hardness.

4.4 Cyclic Loading (T4C & T1C)

Tests T4C and T1C represent an application of cyclic loading on the neoprene-lined

clamp to determine slip due to simulated wave action on a riser fitted to a platform inthe UK Southern North Sea (see Section 3.5). Seven loading steps were conducted in

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cycle and a load loss during the compressive half-cycle were observed. The rates of

load increase and loss were similar and approximately equalled to those of bolts 1 and

8. The central bolts, numbers 2, 3, 6 and 7 see little or no loss/increase in load during

the entire tensile-compressive cycle. The above observations are similar to the

findings of the quasi-static tests as already described in Section 4.2. The variation of

the bolt pre-load for Test T1C followed a similar pattern to that for Test T4C.

Figure 4.11 presents the average bolt pre-load variations for each of the loading steps

in Tests T4C. Comparisons of the average bolt pre-load variations among the loading

steps in Test T4C, full-cycle steps in Test T1C and half-cycle steps in Test T1C are

shown in Figure 4.12. As can be seen, there is little difference in response for full- or

half-cycle loading.

Slip

As seen in the quasi-static axial loading tests, a ductile form of slip appeared in each

of the load steps of Tests T4C and T1C. The load-slip response for step 5 of TestT4C is shown in Figure 4.13. It can be seen from Figure 4.13 that the load-slip

response forms a closed hysteresis loop. Similar load-slip behaviour was observed for

all load steps in Tests T4C.

Figure 4.14 presents a single cycle of each step in Test T4C. It can be observed that

the slope and area of the load-slip hysteresis loop depend on the frequency and

amplitude of the applied cyclic load. Generally, higher load amplitude results in a

larger loop area.

The first five loading steps of Test T1C were carried out to assess the repeatability of

the clamp slip behaviour under cyclic load. As addressed in the test procedures, theneoprene liner was to be replaced whenever visual damage became evident or, in any

case, following three successive tests. Test T1C was conducted using a new neoprene

liner with the same hardness as that in Test T4C. The comparative load-slip

behaviour of Tests T4C and T1C is shown in Figure 4.15. Figure 4.15 reveals that the

clamp slip response in Tests T4C and T1C are similar.

As mentioned above, Test T4C had the identical specimen configurations and similar

bolt pre-load levels to Test T1A. The peak load-slip responses of each loading step in

Test T4C are compared with that of Test T1A in Figure 4.16. It can be seen from

Figure 4.16 that the clamp has similar slip behaviour under cyclic loading and quasi

static loading.

Full-cycle Tests T4C and T1C showed that the load-slip response formed a stable

hysteresis loop. The amplitudes of Steps 1 to 4 in these tests did not exceed their

static axial load capacity. It can be seen from Figure 4.15 that a residual displacement

of clamp relative to pipe is within 0.4 mm for Steps 1 to 4 when the applied loadreduced to zero. It was observed during tests that even this residual displacement was

recovered within a very short period of time.

In order to determine whether the observed residual displacements of the clamp are

purely from the shear deformation of the neoprene lining or if they represent true slip

movement (ie. sliding at interface) of the clamp, 5 half-cycle loading steps wereconducted in Test T1C. The five pulse loading steps simulate the tensile half of the

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five full-cycle steps respectively. Figure 4.17 presents one loading cycle of Step 5 in

Test T1C.

The clamp load-slip response under the tensile half-cycle loading steps was again

observed to form a stable hysteresis loop. A typical example is shown in Figure 4.18

for loading Step 10 of Test 1C.

Comparisons of the clamp load-slip response among the full-cycles tests in T4C and

T1C and the half-cycle tests in T1C are shown in Figure 4.19. The neoprene liner

under the half-cycle load is slightly stiffer than when experiencing a full-cycle load.

In part, this may be because the loading rate in the half-cycle is slightly higher.

It can be seen from Figures 4.18 and 4.19 that the residual displacement of the clamp

under the half-cycle loading is so small that it can be neglected. The residual

displacement observed in the full-cycle tests can be taken as the pure shear of the

neoprene.

Given the above observations, the following conclusions can be drawn:

1) A clamp under cyclic load has a similar deformation response to that when

subjected to static load.

2) The clamp displacement, at least up to an amount of 2 mm, results purely from

the neoprene undergoing shear deformation.

3) With the application of axial load less than the failure load, the shear

deformation of the neoprene results in a small amount of residual

displacement, which is recoverable within a short period of time.

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5. DISCUSSION AND GUIDELINES FOR DESIGN/ASSESSMENT

This section is concerned with the development of guidance for the slip capacity of

stressed neoprene-lined clamps based on the results of the Phase I tests and of the test

programme described in Section 4.

5.1 Discussion

It is appropriate to begin with the tests that were not subject to cyclic loading. A most

illuminating, albeit short, discussion of friction behaviour, as pertains to natural

rubbers, can be found in Reference (2). It is not known how applicable it is to

synthetic rubbers such as neoprene but it would seem to explain the results of most of

the tensile tests in both Phase I and Phase II. Quoting from Reference (2) (underline

inserted):

The coefficient of friction is defined by = WF , where F is the/tangential friction force and W the applied normal load. For rubber,the coefficient of dry friction is not constant, but falls with increasing

normal load. At light loads the dependence is weak, but it becomes

more pronounced at high loads. The friction forceFis proportional to

the real surface contact area, which for normally rough surfaces under

light loads is much less than the geometric area of contact. At very

high loads the relatively low modulus of rubber results in the real

contact area approaching the geometric area, andFtends to a limiting,

maximum value. For dry contacts, the constant of proportionality

between Fand the real contact area is of the same order as the shear

modulus, but it is reduced by surface contamination.

The above passage suggests that the bolts loads in the majority of tests, although

typical of offshore practice, were sufficiently high so that the limiting value of slip

loadFwas reached. The evidence of liner extrusion due to preloading the bolts tends

to confirm that liner was highly stressed, and the real contact area was approaching

the geometric area. Figure 4.7 illustrates the similarity of clamp slip loads over a

wide range of bolt loads.

However, rather than taking the ultimate slip load (ie. the load occurring at a slip of

15 mm or more), a more conservative failure criterion has been used herein. The

failure load has been taken as the load when significant departure from the initial

linear elastic behaviour occurs. This corresponds to when relative displacement

occurs under sustained loads (see Test T4A curve in Figure 4.2). With this failure

criterion, most specimens again give a similar failure load, and hence approximately

similar factors of safety against true slip. In only one test (Test T1B) were bolt loads

so low that a lower failure load was inferred.

To assist in the development of design guidance, reference is made to Figure 5.1,

which is a plot of the interface shear capacity against the interface pressure. The

figure shows the results of Tests T1A, T1B, T1, T2 and T3 in which the preload was

the parameter under investigation. Superimposed are lines corresponding to the

apparent coefficients of friction inferred from the tests. Also shown is the linecorresponding to =0.8 which is typical of values suggested by liner manufacturers.

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The results from Tests T18 and T18A for different neoprene hardness are also

presented.

InterfaceShearCapacity,

=P

c/DL(N/m

m2)

0.35

0.3

0.25

0.2

0.15

0.1

0.05

0

=0.8

18=0.100 T1A, =0.088 T1, =0.044 T2, =0.034

1=0.131 18

=0

.077T3, =0.024

0 1 2 3 4 5 6 7 8 9 10 11

Interface Pressure, q=FB/DL (N/mm2)

Figure 5.1: Plot of Interface Shear against Interface Pressure

The design guidance given in Section 5.2 is formulated in terms of a limiting interface

shear capacity (this is 0.29 N/mm2for the results of Tests T1A, T1 and T2 all having

neoprene hardness IRHD = 60). A limit of 8.5 N/mm2is put on the interface pressureas the liner of specimen T3 in Phase I was extruded when the stud bolt preload was

applied and it also suffered damage during the slip test. For interface pressures less

than that corresponding to about that in Test T1A (q = 3 N/mm2in fact), the limiting

interface shear capacity is ramped down in a parabola form, effectively ending with a

slope of = 0.19 at the origin. Although it is conservative compared with a slope of0.8, uncertainties exist in the small interface pressure region, and there is no test data

available for an interface pressure less than that of Test T1B (1.7 N/mm 2). The

parabola is given by:

q

2 = 0.29

q2 5.19 3

The limiting interface shear capacity was observed to be a function of the neoprene

hardness, measured in International Rubber Hardness Degrees (IRHD), see TestsT18A, T1A and T18 (IRHD values of 50,60 and 70 respectively) in Figure 5.1.

Assuming that these test results are indicative of their respective plateau regions, the

limiting interface shear capacity can be plotted against IRHD as shown in Figure 5.2.

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Interface

Shear,

(N/mm

2)

0.34

0.32

0.3

0.28

0.26

0.24

0.22

0.2

T18

T1A

T18A

40 50 60 70 80

IRHD

Figure 5.2: Effect of Neoprene Hardness on Interface Shear Capacity

The curve shown in the figure is a suitably simple yet accurate approximation to the

data; its equation is given by:

= 0.29 * 1.15 * sin(IRHD) sin (IRHD) 5.23

in which the sine function is based on degrees angular measure.

Equations 5.1.and 5.2 are combined and the coefficients rounded off for the purposes

of the design/assessment provisions in Section 5.2.

Two other sets of results, not shown on Figure 5.1, are worth mentioning. Firstly, the

results of the tests conducted at Karlsruhe fall in the region of an interface pressure of

2.5 N/mm2at a slightly higher shear capacity than Tests T1 and T2. Secondly, some

ad hoc tests on flat plate specimens confirmed the =0.8 line for low interfacepressures. Both sets of results indicate that the design guidance is conservative.

The design/assessment provisions include a factor of safety, . For long term loads

(eg. Gravity loads) applied to the clamp, should notbe taken as less than unity assuch values could lead to creep of the liner material (recall Test T4A in Figure 4.2).

Indeed, it should not be forgotten that the design/assessment provisions are essentially

based on mean values of a relatively small data sample and therefore are not even

characteristic or lower bound. However, as discussed below, environmental loads are

short term in nature, and this allows a less onerous interpretation to be assigned to .

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Cyclic loading tests, especially those of half-cycle tests in T1C, reveal that the relative

displacement of the clamp is almost entirely due to neoprene shear deformation. The

response of the clamp took the form of closed hysteresis loops, with only negligible

deformation following load removal, see Figure 4.19(e).

Rubber-like material is highly sensitive to creep, during which the material continuesto deform under a given load. Figure 5.3, reproduced from Reference (2), shows that

for a certain types of rubber, creep varies approximately linearly with the logarithm of

time under load. It would appear that the durations of the cyclic tests were such that

no substantial creep occurred, and that this is the essential difference between the

cyclic and quasi-static tests. In the design/assessment provisions, it is therefore

recommended that the factor of safety may be taken as unity for designing clampssubject to environmental loads. For assessment purposes of existing clamps, a factor

lower than unity may be justified for the storm event. This is because the storm event

occurs infrequently and very minor slippage (certainly less than 0.1 mm) does not

have any significant structural consequence. The data presented in Figure 4.19(e)would suggest that a clamp load of 200 kN should be perfectly acceptable which,

when compared with the quasi-static limit of 150 kN, leads to an allowable of150/200 = 0.75.

Figure 5.3: Shear Creeping Curves for Different Rubber Materials

The provisions in Section 5.2 for torsional and combined axial/ torsional loads follow

the Phase 1 findings and recommendations.

5.2 Design/Assessment Guidelines

Base on the above observations, design guidance can be formulated as follows. The

clamp slip capacity under axial load alone has been updated with the test results in

Phase II. The clamp slip capacities under either torsional moment or combined axial

and torsional loading are reproduced from Phase I.

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(i) Slip capacity of one clamp half under axial load alone:

P =DL when 3.0 N/mm2 q 8.5 N/mm2c

2

when q < 3.0 N/mm2DL q2 q

3 9Pc =

In the above:

is a limiting stress to be taken as = sin (IRHD) (degree angular measure).3

D and L are respectively the tubular diameter and length of the clamp, both to

be expressed in units of millimetres to give Pcin unit of Newtons.

q is the radial pressure at the neoprene liner/tubular interface, to be calculated

as:

9 =FB , where FBis the total stud bolt load.DL

is a factor of safety which is selected according to the following:

i. For long term (gravity) loads, should not be taken as less than unity.

ii. For designing new clamps for environmental loads, = 1.0.

iii. For the assessment of existing clamps for environmental loads, lessthan unity may be used with caution. On the basis of the project

results, = 0.75 may be acceptable.

The total axial capacity of a clamp is thus 2Pc, but note that in many situations

the axial load is transferred to only one clamp half in the first instance.

The above formulation assumes that there is no interference, i.e. that the

tubular outside diameter is not greater than the inside diameter of the neoprene

liner. If there is interference, a lower capacity may result.

(ii) Slip capacity of clamp under torsional moment:

Mc = PcDWhere Pcis defined in item (i) and D the tubular diameter.

Note that the value of Mc above is the total torsional capacity. Because the

axial stiffness of stud bolts is much greater than the circumferential shear

stiffness of neoprene, the applied torsion is effectively resisted by both halves

of the clamps even where the torsional loading is applied to only one half inthe first instance.

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6. CONCLUDING REMARKS

A programme of slip tests have been carried out on neoprene-lined clamps as used in

offshore applications. The programme of Phase II consisted of six axial slip tests

under either quasi-static loadings or cyclic loadings that simulate wave action in the

UK Southern North Sea. The following parameters were investigated in the Phase IItest programme:

Bolt pre-load Neoprene hardness Cyclic loading effectsThe Phase II tests, with lower bolt loads than those in Phase I, have allowed the

conservatism of the Phase I design guidelines to be removed. This is important as

many existing clamps have neoprene/steel interface pressures corresponding to lowerbolt loads.

The tests with clamps having different neoprene hardness (IRHD value) have

confirmed that hardness affects capacity.

The cyclic loading tests indicate that at the design capacity, the relative displacement

of the clamp and member is recoverable. In other words the displacement is largely

due to neoprene shear deformation as opposed to true slip. It is only when the loads

are applied statically that time dependent phenomena such as creep are manifested.

Design guidance has been formulated based on the results of both Phase I and PhaseII test programmes, see Section 5.2. The provisions encapsulate the above

observations. It is recommended that the factor of safety be adjusted depending on

whether quasi-static or dynamic loading is being considered. A further relaxation

may be used if the clamp is existing, as opposed to a new design.

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REFERENCES

1. MSL Engineering Limited. Development of Design Guidance for Neoprene-Lined

Clamps for Offshore Application. JIP Phase I Final Report, Doc. Ref.

CH10010R005, Rev 1, May 1999.

2. The Malaysian Rubber Produces Research Association, Engineering Design with

Natural Rubber, NR Technical Bulletin, 5thEdition, ISSN-0956-3856, 1992.

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Printed and published by the Health and Safety ExecutiveC30 1/98

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FIGURES FOR SECTION 4

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Figure4.1:

Load-SlipResponseforTests

T1andT4(Reproducedfrom

Phase1FinalReport)

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Figure4.2:

L

oad-SlipResponseforTestsT

1andT4A(ReproducedfromP

hase1FinalReport)

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050

100

150

200

250

300

350

400

450

0

5

10

15

20

25

Clam

pslip(mm)

Appliedaxialload(kN)

Loa

ding

Cyc

le2(higher

loa

dingra

te)

Loa

ding

Cyc

le1(lower

loa

dingra

te)

Linear(

Initials

lope

)

Figure4.3:

Load-SlipResponseforTestT1A

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050

100

150

200

250

300

350

400

0

5

10

15

20

25

30

Clampslip(mm)


Loa

din

gCyc

le2(higher

loa

dingra

te)

Loa

din

gCyc

le1(lower

loa

dingra

te)

Linear

(In

itials

lope

)

Linear

(In

itials

lope

)

Figure4.4:

Load

-slipResponseforTestT1B

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40

50

60

70

80

90

100

110

120

130

-5

0

5

10

15

20

25

Cla

mpslip(mm)

Boltload(kN)

Figure4.5:

BoltPre-LoadVariationduringTestT

1A

3

2

4

6

7

1

8

5

1

5

8

4

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020

40

60

80

100

120

-5

0

5

10

15

20

25

30

Clampslip(mm)

Boltload(kN)

Figure4.6:

AverageBoltL

oadVariationforTestsT1A

andT1B

TestT1A

(F

b=0.2

Fy

)

TestT1B

(Fb=0.1

Fy

)

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-1000

100

200

300

400

500

-5

0

5

10

15

20

25

AxialDisplacemen

tofClampRelativetoPipe(mm)

AppliedAxialTestLoad(kN)

Figure4.7:

Load-SlipResponseforTestsT1,

T2,

T1Aa

ndT1B

TestT2(F

b=

0.5

Fy

)

TestT1(F

b=

0.4

Fy

)

TestT1B(F

b=0.1

Fy

)

TestT1A

(F

b=0.2

Fy

)

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0

100

200

300

400

500

600

0

5

10

15

20

25

30

35

AxialDisplacementofClampRelativetoPipe(mm)


TestT

18

(IRHD70)

TestT18A

(IRHD50)

TestT1A

(IRHD60)

Figure4.8:

Load-SlipResponseforTestsT1A,

T18andT18A

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0

0.0

2

0.0

4

0.0

6

0.0

80.1

0.1

2

ApparentFrictionCoefficient

40

45

50

55

60

65

70

75

NeopreneHardness(IRHD)

Figure4.9:

Correlationbetweenneoprenehardnessandtheapparentfrictioncoefficient

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90

95

100

105

110

115

120

125

130

-3

-2

-1

0

1

2

3

Clampslip(mm)

Boltload(kN)

Figure4.1

0:

BoltLoadVariationinTestT4C(Step

5)

1

5

8

4

18

4

5

3

7

26

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110

111

112

113

114

115

116

117

118

-4

-3

-2

-1

0

1

2

3

4

AxialDisplacementof

ClampRelativetoPipe(mm)

Boltload(kN)

Figure4.1

1:

AverageBoltPre-loadVariationofOneCycleforEachStepinTestT4C

Step1

Step2

Step3

Step4

Step5 S

tep6

Step7

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118

BoltLoad(kN)

110

111

112

113

114

115

116

117

TestT4C

(Step1)

TestT1C

(Step6)

TestT1C

(Step1)

-0.3

-0.2

-0.1

0

0.1

0.2

0.3

AxialDisplacement

ofClampRelativetoPipe(mm)

Figure

4.1

2(a):

AverageBoltP

re-loadVariationof24kNA

mplitudeCycle

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112

113

114

115

116

117

118

TestT4C

(Step2)

TestT1C

(Step2)

Tes

tT1C

(Step7)

BoltLoad(kN)-0.6

-0.4

-0.2

0

0.2

0.4

0.6

Ax

ialDisp

lacement

ofClamp

Re

lative

toPipe

(mm

)

Figure

4.1

2(b):

AverageBoltP


mplitudeCycle

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118

BoltLoad(kN)

112

113

114

115

116

117

TestT4C

(Step3)

TestT1C

Ste

3

TestT1C

(Step8)

-1.5

-1

-0.5

0

0.5

1

1.5

Ax

ialDisp

lacement

ofClamp

Re

lative

toPipe

(mm

)

Figure

4.1

2(c):

AverageBoltP


mplitudeCycle

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TestT4C

Ste

4

TestT1

4

TestT1C

(Step9)

112

1

12

.5

113

1

13

.5

114

1

14

.5

115

1

15

.5

116

1

16

.5

BoltLoad(kN)

C

Ste

-2

-1.5

-1

-0.5

0

0.5

1

1.5

2

Ax

ialDisp

lacement

ofClamp

Re

lative

toPipe

(mm

)

Figure4.1

2(d):

AverageBoltPre-loadVariationof150kNA

mplitudeCycle

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112

113

114

115

BoltLoad(kN)

TestT4C

(Step5)

TestT1C

(Step5)

TestT1C

(Step10)

1

11

.5

1

12

.5

1

13

.5

1

14

.5

1

15

.5

-3

-2

-1

0

1

2

3

Ax

ialDisp

lacement

ofClamp

Re

lative

toPipe

(mm

)

Figure4.1

2(e):

AverageBoltPre-loadVariationof188kNA

mplitudeCycle

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-250

-200

-150

-100

-500

50

100

150

200

250

-3

-2

-1

0

1

2

3

Clamps

lip

(mm

)


Figure4.1

3:

Load-SlipResponseofStep5forTestT4C(188kNA

mplitudeCycle)

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0

100

200

300

-4

-3

-1

0

1

2

3

4

l

Step1

Step4

-300

-200

-100

-2

Appiedaxialload(kN)

Step2

S

tep3

Step5

Step7

Step6

AxialDisplacement


Figure4.1

4:

Load-SlipResponseof

OneCycleforVariousTestT4CLoadSteps

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-200

-1000

100

200

-2.5

-2

-1.5

-1

-0.5

0

0.5

1

1.5

2

2.5


AppliedAxialLoad(kN)

T4C

T1C

Figure4.1

5(c):

Load-SlipR

esponseforTestsT4CandT

1C(Step3)

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-200

-1000

100

200

-2.5

-2

-1.5

-1

-0.5

0

0.5

1

1.5

2

2.5



T4C

T1C

Figure4.1

5(d):

Load-SlipR


1C(Step4)

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-200

-1000

100

200

-2.5

-2

-1.5

-1

-0.5

0

0.5

1

1.5

2

2.5



T4C

T1C

Figure4.1

5(e):

Load-SlipR


1C(Step5)

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050

0

1

2

3

4

5

i

il

il

l

l

l

i

-50

100

150

200

250

300

-0.5

0.5

1.5

2.5

3.5

4.5

AppliedAxalTestLoad(kN)

T4C

-TensePeak

T4C

-CompressvePeak

T1A-LoadCyce1(ow

oadingrate)

T1A-LoadCyce2(hghloadingra

te)

AxialDisplacementofclampRelativetoPipe(mm)

Figure4.1

6:

Load-SlipResponseforTestsT1AandT4C

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250

AppliedAxialTestLoad(kN)

200

150

100

50 0

-50

-100

-150

-200

-250

0

1

2

3

4

5

6

7

8

9

10

11

12

13

14

Step10ofTest1C

Step5ofTest1C

Time(Second)

F

igure4.1

7:

AppliedAxialL

oadofStep5inTestT1C(OneCycle)

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-1000

100

200

-0.5

0

0.5

1

1.5

2

2.5



Figure4.1

8:

Load-Slip

ResponseofStep10inTestT

1C

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50

-0.5

0.5

-50

AxialDisplacement


25 0

-0.2

5

-25

0

0.2

5

T4C-Step

1

T1C-Step

1

T1C-Step

6

Figure4.1

9(a):

Load-SlipResponseofStep1

inTestsT4CandT1CandStep6inTestT1C(OneCycle

)

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-100

-500

50

100

0

1

-1

-0.5

0.5


T4C-S

tep

2

T1C-S

tep

2

T1C-S

tep

7

Ax

ialDisp

lacement

ofClamp

Re

lative

toPipe

(mm

)

Figure4.1

9(b):



)

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150

-1

1

-150

Ax

ialDisp

lacement

ofClamp

Re

lative

toPipe

(mm

)

75 0

-75

-0.5

0

0.5

T4C-S

te

p3

T1C-S

te

p3

T1C-S

te

p8

Figure4.1

9(c):


inTestsT4CandT1CandS

tep8inTestT1C(OneCycle

)

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0

-2

-1.5

-1

-0.5

0

1

2

-200

-100

100

200

0.5

1.5


T4C-S

tep

4

T1C-S

tep

4

T1C-S

tep

9

Ax

ialDisp

lacement

ofClamp

Re

lative

toPipe

(mm

)

Figure4.1

9(d):



)

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-200

-1000

100

200

0

1

2

-2.5

-2

-1

.5

-1

-0.5

0.5

1.5

2.5


T4C-S

tep

5

T1C-S

tep

5

T1C-S

tep

10

Ax

ialDisp

lacement

ofClamp

Re

lative

toPipe

(mm

)

Figure4.1

9(e):

L

oad-SlipResponseofStep5

inTestsT4CandT1CandSt

ep10inTestT1C(OneCycle)

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