copyright by jiali han 2020

Copyright

by

Jiali Han

2020

The Thesis Committee for Jiali Han

Certifies that this is the approved version of the following Thesis:

A Study on Drained Residual Response of Axially Loaded Flowlines on

Gulf of Mexico Clay

APPROVED BY

SUPERVISING COMMITTEE:

Robert B. Gilbert, Supervisor

Chadi El Mohtar


Gulf of Mexico Clay

by

Jiali Han

Thesis

Presented to the Faculty of the Graduate School of

The University of Texas at Austin

in Partial Fulfillment

of the Requirements

for the Degree of

Master of Science in Engineering

The University of Texas at Austin

August 2020

Dedication

To my parents

Han Tingyi and Lu Maowen

my girlfriend

Sun Mengnan

and my loyal friend

Guo Yalin

v

Acknowledgements

I would like to express my sincere gratitude to my Master Supervisor, Dr. Robert

B. Gilbert, for his trust, support, and guidance. He provided me with the opportunities to

work in exciting research projects to broaden my horizon in Offshore Geotechnical

Engineering. He inspired me to be a thoughtful learner and to sharpen my decision

making based on knowledge. I would like to also extend my thanks to Dr. Chadi El

Mohtar for his support and comments to improve this thesis.

I would like to thank Shell Global Solutions U.S. Inc. for contributing data and

sponsoring this project. I would like to also thank Dr Jinbo Chen for his comments and

professional guidance throughout this study.

I would like to thank other geotechnical engineering faculty members, Dr Jorge

Zornberg, Dr Ellen Rathje, and Dr Brady Cox for their excellent teaching efforts that

helped me acquire this degree.

I am extremely grateful for being a member of an amazing research team. I would

like to thank Ahmed Hussien, Dr. Yunhan Huang, Chihun Sung, and Lucas Melo

Monteiro for their technical support, kindness, and friendship. I could never ask for a

better team, and this thesis is the medal of our collaborative efforts.

Living in Austin has been one of the best experiences in my life, and it would not

be possible without loyal friendships. I would like to thank Yalin Guo, Benchen Zhang,

Yunfei Shang, Meibai Li, Chuan Shi, Zhongze Xu, Thiago Araujo, Shradha Panda,

Behdad Mofarraj, Jodie Crocker, Mohamad Hallal, Kai Feng, Lorenzo Peve, Ugur

Arslan, Shiyuan Li, Qiuyu Wang, and Susana Orellana for being my awesome friends.

vi

I would like to thank Mengnan Sun for her patience, encouragement, and

company during the last 9 years. Life is a one-way journey, and I will never regret riding

along with you.

Finally, I would like to express my love and gratitude to my father Tingyi Han

and my mother Maowen Lu. They defined happiness in my life, and for that I will forever

be grateful.

vii

Abstract


Gulf of Mexico Clay

Jiali Han, M.S.E.

The University of Texas at Austin, 2020

Supervisor: Robert B. Gilbert

Lateral buckling and axial walking are the main design issues for offshore

pipelines and flowlines. Accurate pipe-soil interaction estimation can lead to optimizing

design with significant cost reduction. The objective of this study is to characterize the

interface interaction between pipelines and Gulf of Mexico (GOM) clay by carrying out

physical model tests using 4.25-inch diameter pipes. T-bar tests were conducted after

reconstitution of the clay test bed to capture the undrained shear strength, which allows

the sensitivity of the clay to be interpreted. Pipe embedment tests and consolidation tests

were performed to estimate the pore pressure dissipation rate of the GOM clay after the

placement of the pipeline. After the test beds are mixed to the desired water content,

model pipes were placed and allowed to settle for about three weeks. Then, four series of

axial load tests were carried out to simulate pipe walking under various loading

conditions. The soil-pipe interface axial resistance was measured, and factors including

consolidation time, loading conditions, loading sequence, and effective overburden stress

were investigated. Drained response of the soil-pipe interface in this study appeared to be

viii

mobilized by a “slow” axial motion rate at 1×10-5 in/s. Also, it was observed that the

drained residual state could be reached by performing numerous “fast” sweeps. Once the

drained residual state is reached, the axial shear resistance can stay constant with the

motion rate increased by 100 times. Same as suggested by other literature, the axial

resistance will increase if the test is paused for days. The axial resistance will then drop

back to the drained residual axial resistance with continued shearing. For test beds with

large water contents, consolidation of the pipe-soil interface may take longer, and the

measured axial resistance will keep increasing with time. Also, possible stress

concentration at the invert of the pipe might lead to a decreased axial resistance due to

increased effective contact stress between the pipe and the clay. Eventually, axial load

test results are compared with previously performed tilt table tests results. Both tests

indicate that the drained residual interface friction coefficient decreases with increasing

effective overburden stress.

ix

Table of Contents

List of Tables .................................................................................................................... xii

List of Figures .................................................................................................................. xiii

Chapter 1 Introduction .........................................................................................................1

1.1 Motivation .............................................................................................................1

1.2 Objectives .............................................................................................................3

1.3 Structure of Thesis ................................................................................................4

Chapter 2 Literature Review ................................................................................................5

2.1 Introduction and Objectives ..................................................................................5

2.2 Geotechnical Engineering Aspects of Offshore Flowline/Pipeline Design ..........5

2.3 Tests to Measure the Undrained Shear Strength of Marine Clays ........................6

2.4 T-Bar Test Parameters ..........................................................................................8

2.5 Tilt Table Tests ...................................................................................................13

2.6 Axial Load Model Tests .....................................................................................15

Chapter 3 Test Bed Preparation and Testing Equipment ...................................................18

3.1 Gulf of Mexico Clay ...........................................................................................18

3.1.1 General Characteristics ........................................................................18

3.1.2 Undrained Shear Strength Measurement Using T-bar Penetrometer

Tests ........................................................................................................19

3.1.2.1 T-bar Test Setup .......................................................................19

3.1.2.2 T-bar Test Results ....................................................................21

3.1.3 Drained Residual Shear Strength between Soil-Polypropylene

Interface ..................................................................................................29

3.2 Test Bed ..............................................................................................................30

x

3.3 Flowlines .............................................................................................................31

3.4 Electric Motor .....................................................................................................31

3.5 Load Frames .......................................................................................................32

3.6 Load Cell.............................................................................................................33

3.7 Linear Motion Transducer (LMT) ......................................................................34

3.8 Linear Variable Differential Transformer (LVDT) ............................................35

3.9 Data Acquisition and Motion Control Programs ................................................36

3.10 Axial Load Model Test Setup ...........................................................................37

3.10.1 Test Bed Mixing ................................................................................37

3.10.2 Load Conveyance System ..................................................................38

3.10.3 Instrumentation ..................................................................................40

Chapter 4 T-bar Tests and Embedment Tests ....................................................................44

4.1 Pipe Embedment Tests........................................................................................44

4.1.1 Initial Embedment ................................................................................44

4.1.2 Consolidation after Initial Embedment ................................................47

4.1.3 Torsional Spinning Test .......................................................................52

4.2 Embedment Tests Conclusions ...........................................................................54

Chapter 5 Axial Loading Tests ..........................................................................................56

5.1 Test Plan .............................................................................................................56

5.2 Test Procedures ...................................................................................................58

5.2.1 System Load-up ...................................................................................58

5.2.2 Load Sweep towards the Motor ...........................................................58

5.2.3 Load Sweep towards the Counterweight/Spring..................................61

xi

5.2.4 Pulley Friction Correction....................................................................61

5.2.5 Normalized Axial Resistance Calculation ...........................................64

5.3 Test Series 1 – 9-ft Pipe in Test Bed 1 ...............................................................64




5.7 Axial Load tests Conclusions .............................................................................92

Chapter 6 Conclusions .......................................................................................................96

References ..........................................................................................................................98

xii

List of Tables

Table 3.1 Gulf of Mexico Clay properties .........................................................................19

Table 3.2 T-bar Test Results Summary .............................................................................22

Table 3.3 T-bar Test Groups ..............................................................................................26

Table 4.1 Test Beds information ........................................................................................45

Table 4.2 Initial embedment estimation for Pipes 1-4 .......................................................45

Table 4.3 Consolidation parameters using Square-Root-of-Time Method ........................48

Table 5.1 Axial load testing plan .......................................................................................57

Table 5.2 Axial Load Test results of 9-ft 27-lb Pipe with counterweight .........................65

Table 5.3 Axial Load Test results of 6-ft 16.4-lb Pipe with counterweight ......................70

Table 5.4 “Very fast” and “fast” axial sweep results of 9-ft 53.7-lb Pipe with

counterweight ................................................................................................79

Table 5.5 Axial Load Test results of 6-ft 16.4-lb Pipe with counterweight ......................84

Table 5.6 Test plan for Pipe 4 in Test Bed 2 .....................................................................89

xiii

List of Figures

Figure 1.1 Example of offshore operation layout (Jayson et al., 2008) ...............................2

Figure 1.2 An engineering lateral buckle in a deepwater pipeline (Jayson et al., 2008) .....2

Figure 2.1 T-bar and ball full-flow penetrometers (Randolph et al., 2011).........................8

Figure 2.2 LDFE analysis of soil deformation patterns during T-bar tests (White et al.,

2010) .............................................................................................................10

Figure 2.3 Undrained shear strength profile measured by T-bar tests on the GOM

Clay at different penetration rates (Lai, 2017) ..............................................12

Figure 2.4 Undrained shear strength measured by T-bar tests on the GOM Clay at

different cycles (Gerkus, 2016) .....................................................................13

Figure 2.5 Schematic of tilt table test (Bae, 2009) ............................................................14

Figure 2.6 Pore pressure generation during shearing vs. OCR (Ballard et al., 2013)........16

Figure 2.7 Axial resistance peak after long set-up time.....................................................17

Figure 3.1 e vs. logσv .........................................................................................................18

Figure 3.2 T-bar and Penetrometer Rod (Gilbert et al. 2012) ............................................20

Figure 3.3 Nc as a function of depth (Melo Monteiro, 2019) ............................................21

Figure 3.4 Undrained Shear Strength vs. Depth - Test Bed 1............................................23

Figure 3.5 Undrained Shear Strength vs. Depth - Test Bed 2............................................24

Figure 3.6 Comparison of remolded undrained shear strengths Sur ...................................25

Figure 3.7 Undrained shear strength Su vs. time (Melo Monteiro, 2019) .........................27

Figure 3.8 Remolded undrained shear strength Sur vs. time (Melo Monteiro, 2019) ........27

Figure 3.9 Sensitivity (Su/Sur) vs. time (Melo Monteiro, 2019) ........................................28

Figure 3.10 Sur vs. water content (Melo Monteiro, 2019) .................................................28

xiv

Figure 3.11 Drained residual friction coefficient from tilt table tests (after Melo

Monteiro, 2019) ............................................................................................30

Figure 3.12 Powered soil mixer .........................................................................................31

Figure 3.13 Stepper motor .................................................................................................32

Figure 3.14 Load frames ....................................................................................................33

Figure 3.15 Load cell (Huang, 2015) .................................................................................34

Figure 3.16 Linear motion transducer (Melo Monteiro, 2019) ..........................................35

Figure 3.17 LVDT connection to (left) the tank and (right) the pipe ................................36

Figure 3.18 LabVIEW user interface (Huang, 2015) ........................................................37

Figure 3.19 Immediately after mixing and smoothing of Test Bed 2 ................................38

Figure 3.20 Test setup of axial load tests with linear actuator and (a) counterweight or

(b) spring (Hussien, 2020) ............................................................................39

Figure 3.21 (a) Counterweight (b) Spring..........................................................................40

Figure 3.22 (a) Motor-side loadcell (b) Counterweight-side loadcell ...............................41

Figure 3.23 LVDT installation overview ...........................................................................42

Figure 3.24 LVDT .............................................................................................................42

Figure 4.1 Measured pipe embedment (After Hussien, 2020) ...........................................47

Figure 4.2 Consolidation of Pipe 1 ....................................................................................49




Figure 4.6 Predicted consolidation for Pipe 1 and 2 ..........................................................51

Figure 4.7 Torsional Spinning Test Schematic ..................................................................52

Figure 4.8 Torsional Spinning Test being performed ........................................................53

Figure 4.9 Scaled Torsional Spinning Test results vs Pipe 4 embedment time history .....54

xv

Figure 5.1 Motor-side loadcell measurements of sweep 9 of Pipe 1 .................................59

Figure 5.2 Counterweight loadcell measurements of sweep 9 of Pipe 1 ...........................60

Figure 5.3 Axial force of sweep 9 of Pipe 1 ......................................................................60

Figure 5.4 Pulley friction measurement test (pulling towards the motor) .........................62

Figure 5.5 Pulley friction vs counterweight .......................................................................62

Figure 5.6 Example of applying pulley friction (towards motor) ......................................63

Figure 5.7 Axial force vs Displacement from cyclic axial load tests performed at

"slow" and "intermediate” speeds with counterweight, Pipe 1 in Test

Bed 1 .............................................................................................................66

Figure 5.8 Friction coefficients vs displacement from cyclic axial load tests

performed at "slow" and "intermediate” speeds with counterweight, Pipe

1 in Test Bed 1 ..............................................................................................67

Figure 5.9 Friction coefficient, μ vs sweep - Pipe 1 ..........................................................67

Figure 5.10 Friction coefficient, μ vs time elapsed after pipe placement - Pipe 1 ............68

Figure 5.11 Pipe transient speeds vs displacement, Pipe 1 - sweep 4 and 7 ......................69

Figure 5.12 Axial force vs Displacement from cyclic axial load tests performed at

"fast” speeds with spring, Pipe 2 in Test Bed 1 (After Hussien, 2020) ........72

Figure 5.13 Friction Coefficient vs Displacement from cyclic axial load tests

performed at "fast” speeds with spring, Pipe 2 in Test Bed 1 (After

Hussien, 2020) ..............................................................................................73

Figure 5.14 Friction coefficient, μ vs sweep at “fast” speeds - Pipe 2 ..............................74

Figure 5.15 Friction coefficient vs displacement from cyclic axial load tests

performed at "slow” speeds with spring, Pipe 2 in Test Bed 1 (After

Hussien, 2020) ..............................................................................................75


xvi


performed at "slow” speeds with counterweight, Pipe 2 in Test Bed 1

(After Hussien, 2020)....................................................................................77

Figure 5.18 Friction coefficient vs sweep - Pipe 2 ............................................................78

Figure 5.19 Friction coefficient vs time after pipe placement - Pipe 2 ..............................78

Figure 5.20 Friction coefficient vs sweep - Pipe 3 at “very fast” and “fast” speeds .........81


performed at "very fast” and “fast” speeds with counterweight, Pipe 3 in

Test Bed 2 .....................................................................................................83

Figure 5.22 Gaps between the model pipe and clay trench ...............................................86

Figure 5.23 Friction coefficient vs sweep - Pipe 3 ............................................................87

Figure 5.24 Friction coefficient vs time after pipe placement - Pipe 3 ..............................87

Figure 5.25 Friction coefficient vs displacement, Pipe 3 - sweeps 57 to 61 with

increasing speed (After Hussien, 2020) ........................................................88


performed at "slow” speeds with counterweight, Pipe 4 in Test Bed 2 ........90


Figure 5.28 Large-displacement drained residual friction coefficient obtained from

axial load tests and tilt table tests (After Hussien, 2020) ..............................94

Figure 5.29 Friction coefficient vs time after pipe placement – Test Bed 1 ......................94

Figure 5.30 Friction coefficient vs time after pipe placement – Test Bed 2 ......................95

1

Chapter 1 Introduction

1.1 MOTIVATION

Flowlines in offshore developments have been extensively used as the demand for

offshore oil and gas production and exploration keeps increasing over the years. An

example of in-field layout of offshore oil and gas production is shown in Figure 1.1.

These flowlines are often required to be operated at high temperatures and pressures.

Variations in temperatures and pressures due to start-up or shut-down of oil and gas

transmission lead to thermal gradient, which induces expansion of the flowlines. While

the thermal cycles of the flowlines continue, some portion of the expansion may not be

recovered, and a ratcheting response in the flowline axial displacement may occur.

Depending on the start-up and shut-down cycles, the flowline could exhibit varying rates

of displacement along the flowline (Senthilkumar et al., 2016). In addition, steep seabed

slopes, steel catenary riser tension, and liquid hold-up also create pipeline walking

potential, which requires design considerations (Simpson et al., 2015). In order to

properly design for axial walking and buckling, an accurate assessment of the axial pipe-

soil interaction is essential. (Randolph et al., 2010). The drained residual axial resistance

between a flowline and seabed clay is a crucial factor which affects the types and

magnitudes of the movement. Figure 1.2 shows an illustration of lateral buckling of a

flowline. Therefore, it is essential to choose a suitable range of axial resistance during the

design stage (Carr and Bruton, 2006).

2

Figure 1.1 Example of offshore operation layout (Jayson et al., 2008)

Figure 1.2 An engineering lateral buckle in a deepwater pipeline (Jayson et al., 2008)

3

Tilt table tests have been performed to measure the drained residual shear

resistance of the soil-pipe interface for its capability of measuring under very low

effective normal stress compared with direct shear test and ring shear test (Melo

Monteiro, 2019). Although tilt table test has been proved as an effective and efficient tool

to capture the drained residual resistance along the interface, it has certain drawbacks.

For instance, tilt table test cannot control the shearing rate and may lead to load

eccentricity. Also, tilt table test only measures the shear stress along a flat interface,

while the flowline-soil interface is curved. Therefore, large-scale model tests were

proposed and conducted in this study to investigate factors influencing the axial

resistance. In particular, the pipe-soil interface response against axial loading may change

significantly as time passes.

Over the last decade, several large-scale model tests were performed using marine

clays from Onsøy and West Africa, while no attempt has been made to conduct model

testing on GOM Clay. This study took the initiative and carried out axial load model test

on GOM Clay to obtain a better understanding of possible pipe-soil interface response in

this region.

1.2 OBJECTIVES

The main objectives of this study are:

1. Estimate the primary consolidation time of test beds by performing embedment

tests of model pipes and consolidation tests of the GOM Clay. Trace the pore pressure

dissipation history. Correlate the degree of consolidation of the test beds to the soil-pipe

interface drained residual shear strength. Develop an understanding of how the pipe-soil

interface behaviors under axial loading changes with time after placement.

4

2. Perform axial load model tests on GOM Clay test beds to simulate various

flowline loading conditions in the field including walking. Measure the soil-pipe interface

shear strength under different load conditions. Trace the timeline of the development of

drained residual state throughout the tests. Identify the effects of variables such as pipe

weights, soil strengths, elapsed time after pipe laydown, axial loading conditions, and

sequence of axial loading.

1.3 STRUCTURE OF THESIS

This thesis consists of six chapters. Chapter 1 is the introduction to the

motivations and objectives of this study. Chapter 2 focuses on literature review, which

outlines previous studies on geotechnical aspects of offshore flowline/pipeline design, T-

bar tests, flowline placement and axial load model tests. Test materials and equipment are

introduced in Chapter 3. Chapter 4 presents the embedment tests and torsional spinning

resistance tests results. Chapter 5 illustrates the testing methodology of axial load model

tests of this study. Four series of axial load tests are discussed in chronological order. A

summary of major conclusions from this study can be found in Chapter 6.

5

Chapter 2 Literature Review

2.1 INTRODUCTION AND OBJECTIVES

This chapter introduces background information about marine clay properties

under axial loading of flowlines. The objective of this section is to summarize results

from previously performed tests on characterization of flowline-marine clay interaction.

To measure the undrained shear strength of marine clay, T-bar test is adopted for this

study, and previous studies on the T-bar tests are introduced. In addition, methods that

measure drained residual shear strength are also mentioned, including the tilt table test. In

the end, results of similar model testing of axially loaded flowlines are reviewed.

2.2 GEOTECHNICAL ENGINEERING ASPECTS OF OFFSHORE FLOWLINE/PIPELINE

DESIGN

As the demand for energy supply keep rising, the offshore oil and gas industry

have been expanding over the last decades. Field operation of offshore oil and gas

extraction involves various design aspects. Challenges of geotechnical design for deep

water applications as such are different from those encountered at shallow water, which

are caused by the floating/subsea nature of the facilities and the properties of seabed

sediments (Randolph and Gourvenec, 2011). Among all the components, pipeline

interaction with the seabed is an important technical aspect that has been studied

extensively over the years.

One of the major objectives of this study is to simulate the soil-pipe interaction

during pipeline walking in the offshore environment. Flowline walking is a phenomenon

in which axial displacement of “short”, high temperature pipelines due to

thermal/pressure induced axial loading. The term “short” stands for pipelines that do not

reach full constraint in the middle. These changes in pressure and temperature are due to

6

start-up/shut-down of oil and gas transportation of the flowlines. Three different

conditions can potentially facilitate pipeline walking, including (1) tension at the end of

the flowline, associated with a steel catenary riser, (2) Seabed slope along the pipeline,

and (3) thermal gradients along the flowline during operation changes such as start-

up/shut-down. Numerous cycles of this ratcheting response may cause large global axial

displacements with unfavorable consequences, such as overstressing of

spoolpieces/jumpers, loss of tension in a steel catenary riser, and a need for restraint

using anchors (Carr et al., 2006).

Analogies have been made between piles and pipelines. However, pipeline

operation differs from pile operation in several aspects. Stress level between pipelines

and underlying soils is significantly lower compared with pile-soil interface stress. Also,

axial expansions of flowlines sometimes happen as slow as a thousandth of an inch per

second but may end up inducing several feet of displacement (Randolph et al., 2011).

In summary, flowlines under operation are susceptible to flowline walking and

lateral buckling issues. Failure mechanisms including local buckling, fracture and low-

frequency fatigue damage require thorough analyses to satisfy design limit states. In

particular, the pipe-soil interface response under such conditions is the largest uncertainty

of the design (Bruton et al., 2009). As a result, to properly design the pipelines, an

accurate assessment of the axial soil-pipe interface shear resistance is required (Randolph

et al., 2011).

2.3 TESTS TO MEASURE THE UNDRAINED SHEAR STRENGTH OF MARINE CLAYS

Deep water operation sites typically consist of soft marine clay that is usually

slowly deposited with strength increasing with depth (Randolph et al., 2011). Due to the

extreme difficulty of handling and retrieving undisturbed soil samples in deep-water

7

environment, in-situ testing is preferred. In-situ testing is performed using a seabed frame

equipped with built-in penetrometer rods. Characterization of soft clay often involves

using traditional cone penetration test (CPT) for stratigraphy profiling, followed by vane

shear tests or undisturbed sampling (Stewart and Randolph, 1994). Empirical

relationships are used to correlate the CPT tip resistance to the undrained shear strength

of the site. Relationship such as 𝑆𝑢 = 𝑞𝑐/𝑁𝑐 requires knowledge of the bearing capacity

factor 𝑁𝑐 . 𝑆𝑢 is the converted undrained shear strength of the soil, and 𝑞𝑐 is the

measured tip resistance. Without any prior knowledge of the site condition, the selection

of an appropriate bearing capacity factor can be difficult, as 𝑁𝑐 is determined by factors

including stiffness of the soil, stress level, and stress history (Stewart and Randolph,

1994). However, T-bar test has become an increasingly popular option to obtain

undrained shear strength of soils. The advantage of the T-bar test compared with

traditional tests such as vane shear and CPT is that soil can flow around the T-bar

cylinder. Not only does a T-bar test produces a continuous undrained shear strength

profile, the “flow around” nature of the T-bar also leads to fewer corrections for

overburden stress, which can be significant for the CPT. (White et al., 2010). As shown

in Figure 2.1, the projected area of the T-bar cylinder is 5-10 times larger than the cross-

sectional area of the shaft. Therefore, correction for overburden pressure can be

neglected, which is considered crucial for tests performed at sites with low c/p ratio

In addition to the T-bar test, the ball penetrometer is another testing method with

similar characters as the T-bar penetrometer. Compared with the T-bar test, the ball

penetrometer has a simpler geometry with pore pressure measurements (Randolph et al.,

2011).

8

Figure 2.1 T-bar and ball full-flow penetrometers (Randolph et al., 2011)

In summary, the T-bar penetrometer test has proven to be a simple and cheap

alternative to conventional methods of in-situ shear strength measurement. For this

reason, T-bar test was adopted and performed in this study to measure the undrained

shear strength of the GOM Clay.

2.4 T-BAR TEST PARAMETERS

One of the major considerations for analysis of T-bar test results is the choice of a

proper bearing capacity factor. Randolph and Houlsby (1984) originally proposed a

bearing capacity factor of 10.5 for the correlation between the measured penetration

resistance and the undrained shear strength using a simple equation (Stewart and

Randolph, 1994),

𝑃

𝑆𝑢𝑑= 𝑁𝑏 (Eq. 2.1)

9

where 𝑃 is the measured force per unit length acting on the T-bar cylinder, 𝑑 is

the diameter of the cylinder, and 𝑁𝑏 is the bar factor (bearing capacity factor).

The value of 𝑁𝑏 is dependent on the roughness of the bar, ranging from about 9

to 12 from a smooth bar transitioning to a rough bar (Stewart and Randolph, 1994), The

value of 10.5 for the bearing capacity factor was developed and calibrated by experiments

with various soil types, stress levels, and stress histories (House et al., 2001).

However, the bearing capacity factor of 10.5 for T-bar tests only applies when the

T-bar penetrometer was pushed deep enough and a full “flow-around” failure surface is

formed. Therefore, an accurate estimation of undrained shear strength at shallow depths

using T-bar tests requires appropriate adjustment of the bearing capacity factor. White et

al. (2010) performed Large Deformation Finite Element (LDFE) analysis to estimate the

shallow bearing capacity factors (Figure 2.2) which illustrates the mechanism of a

reduced bearing capacity at shallow depths with a shorter failure plane.

10

Figure 2.2 LDFE analysis of soil deformation patterns during T-bar tests (White et al.,

2010)

Then, an empirical expression is established to fit the trend of bearing capacity

factor profile obtained from LDFE analysis,

𝑁𝑇−𝑠ℎ𝑎𝑙𝑙𝑜𝑤 = 2 + (𝑁𝑇−𝑑𝑒𝑒𝑝 − 2) (�̂�

�̂�𝑑𝑒𝑒𝑝)

𝑃

(Eq. 2.2)

𝑝 = 0.61 (𝑆𝑢

𝛾′𝐷)

−0.31

where 𝑁𝑇−𝑠ℎ𝑎𝑙𝑙𝑜𝑤 is the corrected shallow bearing capacity factor, 𝑁𝑇−𝑑𝑒𝑒𝑝 is

the bearing capacity at deep depths, �̂� is the penetration depth of the invert of the

cylinder normalized by the diameter of the cylinder, �̂�𝑑𝑒𝑒𝑝 is the transition normalized

11

depth, 𝛾′ is the effective unit weight of the soil, and 𝐷 is the diameter of the acrylic

cylinder.

In addition to the calibration of the bearing capacity factor, the penetration rate is

another important aspect that requires careful selection. The conventional full-flow

penetrometer test often uses a penetration rate of 0.8 in/s (Dejong et al., 2011). Due to

viscous effects, the penetration resistance increases as the penetration rate increases.

Also, if the penetration rate is decreased, partial drainage and consolidation may occur

along the way, leading to an increase in penetration resistance. Therefore, the minimum

penetration resistance occurs at the rate at which the loading conditions transition from

undrained to partially drained (Bemben and Myers 1974) (Chung et al., 2006). As a

result, the principle of selecting the penetration rate should be to ensure the undrained

condition is reached, while limiting the viscous effects. For this study, the standard

penetration rate of 0.8 in/s is adopted. Lai (2017) performed T-bar tests on the same

GOM Clay used in this study to investigate the rate effect. Penetration rates of 0.8 in/s,

1.6 in/s, 3.2 in/s, 4.8 in/s, 6.4 in/s and 8 in/s are used. Compared with the baseline case

with a penetration rate of 0.8 in/s, it was observed that the measured undrained shear

strength will be overestimated once the loading rate becomes larger than 3.2 in/s. Beyond

3.2 in/s, the measured undrained shear strength increases with increasing penetration rate.

12

Figure 2.3 Undrained shear strength profile measured by T-bar tests on the GOM Clay at

different penetration rates (Lai, 2017)

The T-bar test can also be used to measure the remolded undrained shear strength

of clay by performing multiple penetration and extraction cycles (Randolph and

Anderson, 2006). Yafrate et al. (2009) proposed that remolded undrained shear strength

is typically reached after 10 cycles. Gerkus (2016) performed T-bar tests on the GOM

Clay and showed that the remolded shear strength was reached after 4 to 6 cycles. For

this study, since the T-bar tests were performed not long after clay mixing, not much

thixotropy effect was able to be developed for the clay. As a result, the remolded

undrained shear strength was assumed to be reached after three cycles of penetration and

extraction for this study.

13

Figure 2.4 Undrained shear strength measured by T-bar tests on the GOM Clay at

different cycles (Gerkus, 2016)

2.5 TILT TABLE TESTS

Several methods have been developed to measure the drained residual shear

strength at the interface between soils and solid interfaces. Direct shear and ring shear

tests have been extensively used to measure the interface shear strength between different

soils and material surfaces. However, as discussed, one of the major differences between

pipelines and conventional driven piles is that the effective normal stress between the

pipe-soil interface is very low. Since the effective normal stress is small, even small

amount of friction from the testing device may lead to skewed results (Najjar et al.,

14

2009). Direct shear tests and ring shear tests generally involve use of a large normal

stress, which usually cannot be achieved for flowline/soil interface.

Under low effective normal stress, tilt table tests can be performed to measure the

interface shear strength. A schematic of tilt table test is shown in Figure 2.5.

Figure 2.5 Schematic of tilt table test (Bae, 2009)

Compared with ring shear and direct shear tests, tilt table test eliminates the

mechanical friction by applying the normal and shear stresses to the soil-plate interface.

Also, different from direct shear or ring shear test, the failure surface is not forced to

occur along the interface. Disadvantages of tilt table tests include inability to control the

displacement and to measure the post-failure response. Also, since the loading surface is

inclined, loading eccentricity may cause ununiform stress distribution along the contact

surface and adversely affect the results (Pederson et al. 2003).

Najjar et al. (2003 and 2007) performed tilt-table tests on Gulf of Mexico clay

extracted from several different sites. Four solid interfaces with different properties were

15

tested at a range of effective normal stress below 104 psf. The measured drained residual

friction coefficient rapidly decreases at low effective normal stress, and the curve

becomes flatter as the effective normal stress keeps increasing.

In addition, Bae (2009) performed the tilt table tests on GOM Clay with similar

effective normal stress. From these tests, it is verified that the drained residual interface

shear strength is dependent on the chemical composition of the clay, the type of solid

materials, and the surface roughness.

2.6 AXIAL LOAD MODEL TESTS

Geotechnical model testing of axially loaded flowlines was increasingly used to

characterize the pipe-soil interaction at deep marine environments, as in-situ testing and

sampling can be difficult (Langford et al., 2007). Over the last decade, several groups of

axial load model tests attempting to characterize the interface between pipeline and

marine clay were reported, and many of which tests used Onsøy and West African marine

clay.

Bruton et al. (2009) performed axial load tests using three cubic meters (106 cubic

ft) of reconstituted clay. Consolidation of the test bed was facilitated by vacuum

consolidation to an undrained shear strength profile that matches the field conditions. The

pipe was pushed to the desired embedment depth and held in place until pore pressure

was fully dissipated. Axial sweeps were performed at speeds from 4×10-5 in/s to 4×10-3

in/s. When the pipe was moving at 4×10-3 in/s, pore pressure generation was recorded. As

a result, fast “undrained” and slow “drained” movements led to different axial resistance.

Hill et al. (2012) confirmed the existence of rate effects, stating that a change to

higher velocity could cause an increase in shear strength due to large negative excess

16

pore pressure generation. Also, it was observed that heavier pipes ended up with greater

embedment depth.

Ballard et al. (2013) investigated the effects of stress state of the test bed on the

recorded axial friction factor. As shown in Figure 2.6, when the unloading ratio for the

pipes were small, positive pore pressure was generated during shearing, while a slight

dilative behavior was obtained when the unloading ratio for the pipe was 1.4 and the pipe

was loaded at a fast moving rate. With this negative pore pressure generation, the axial

resistance is expected to be increased.

Figure 2.6 Pore pressure generation during shearing vs. OCR (Ballard et al., 2013)

Undrained fast speeds can lead to large breakout peaks in axial resistance if the

axial sweeps had pauses in between (Langford et al., 2007) (White et al., 2011).

17

Figure 2.7 Axial resistance peak after long set-up time

18

Chapter 3 Test Bed Preparation and Testing Equipment

3.1 GULF OF MEXICO CLAY

3.1.1 General Characteristics

The marine clay that was used to reconstitute the test bed is extracted from

multiple deep-water project sites from the Gulf of Mexico. The specific gravity and

plasticity indices were measured by Gerkus (2016). In addition, A consolidation test (In

accordance with ASTM 2435) was performed on a reconstituted sample that was

retrieved from the test bed. Figure 3.1 shows the void ratio e vs effective overburden

stress σv on a semi-log scale. The properties of the marine clay are summarized in Table

3.1.

Figure 3.1 e vs. logσv

0.0

0.5

1.0

1.5

2.0

2.5

3.0

100 1,000 10,000 100,000

Void

rat

io, e

Effective vertical stress (psf)

19

Table 3.1 Gulf of Mexico Clay properties

Parameter Unit Value

Specific Gravity, SG - 2.75

Liquid Limit, LL (%) 105

Plasticity Index, PI - 62

Compression Index, Cc - 0.6

Recompression Index, Cr - 0.18

Coefficient of Consolidation, Cv (ft2/year) 3-8

3.1.2 Undrained Shear Strength Measurement Using T-bar Penetrometer Tests

3.1.2.1 T-bar Test Setup

T-bar tests were performed at the beginning of each axial loading test series to

characterize the profile of the remolded undrained shear strength of the test bed. As

shown in Figure 3.2, the apparatus consists of a 4-inch-long cylindrical bar with 1-inch

diameter, which is attached to a long penetrometer rod. The test is performed by

vertically penetrating the T-bar into the soil at a constant rate at 0.8 in/sec (20mm/sec).

Load pieces are mounted to the rod to help the T-bar penetrate the soil. Due to thixotropy

effects, the initial T-bar penetration test may result in larger resistance. To account for the

strength gain for the initial penetration, after the initial penetration test, the test is

repeated for two more times to fully disturb the clay bed. Studies showed that the

measured resistance of penetration normalized by the resistance of the initial penetration

does not change significantly after the third penetration. Therefore, the remolded

undrained shear strength of the soil is taken as the measurements of the third penetration.

In the end, the cylindrical bar is removed, and the penetrometer rod is pushed into the soil

20

to measure the rod resistance. The rod resistance needs to be subtracted to obtain the

ultimate bearing capacity of the clay.

Figure 3.2 T-bar and Penetrometer Rod (Gilbert et al. 2012)

Similar as CPT and SPT, The T-bar penetrometer test essentially causes bearing

capacity failure of the clay under the T-bar. Therefore, the undrained shear strength can

be related to the measurements using the following equation,

𝑆𝑢 =(𝐹𝑡𝑜𝑡𝑎𝑙−𝐹𝑟𝑜𝑑)

𝑁𝑐∗𝐴 (Eq.3.1)

where Ftotal is the total resistance measured during the T-bar penetration tests, Frod

is the rod resistance, A is the projected area of the T-bar, and Nc is the bearing capacity

factor. Generally, at deep depths, the bearing capacity factor Nc varies from 9.14 for a

fully smooth interface to a fully rough interface (Randolph and Houlsby 1984; Martin

21

and Randolph 2006). For this study, a Nc value of 10.5 is adopted. On top of this, as the

depth gets shallower, Nc needs to be reduced as the failure path for T-bar penetration is

shorter. White et al. (2010) suggests that Nc should be reduced at shallower depth.

Therefore, Nc is set to 5 at the soil surface and linearly increases to 10.5 till the depth of

2.5D, where D is the diameter of the T-bar. Figure 3.3 illustrates the Nc at different

depths.

Figure 3.3 Nc as a function of depth (Melo Monteiro, 2019)

3.1.2.2 T-bar Test Results

T-bar tests were performed to characterize the undrained shear strength profile of

the test bed right after mixing of the clay before axial loading tests. The test results are

summarized in Table 3.2.

22

Table 3.2 T-bar Test Results Summary

Test Beds Test Bed 1 Test Bed 2

Water Content (%) 138 124

Remolded Undrained Shear

Strength, Sur (psf) 1.6 2.1

The undrained shear strength profiles show an increasing trend at shallow depths

till 1 inch. Since not enough time had passed to allow primary consolidation of the test

beds at deeper depths to finish, the profiles then stay constant with the rest of the depth.

According Figure 3.4 and Figure 3.5, the recorded constant remolded undrained shear

strength Sur are 1.6 psf and 2.1 psf for Test Bed 1 and Test Bed 2, respectively.

23

Figure 3.4 Undrained Shear Strength vs. Depth - Test Bed 1

0

1

2

3

4

5

6

0 1 2 3 4 5

Dep

th (

in)

Undrained shear strength (psf)

Cycle 1

Cycle 2

Cycle 3

24

Figure 3.5 Undrained Shear Strength vs. Depth - Test Bed 2

0

1

2

3

4

5

6

0 1 2 3 4 5

Dep

th (

in)

Undrained shear strength (psf)

Cycle 1

Cycle 2

Cycle 3

25

Figure 3.6 Comparison of remolded undrained shear strengths Sur

0

1

2

3

4

5

6

7

8

0 0.5 1 1.5 2 2.5 3

Dep

th (

in)

Remolded undrained shear strength (psf)

Test Bed 1

Test Bed 2

26

Prior to the T-bar tests performed on the test beds, additional T-bar tests were

carried out on a separate smaller clay bed to investigate the effects of changing water

contents on the undrained shear strength, as well as to correlate the thixotropy effects

with increasing time. In total, three groups of T-bar tests at different water contents were

performed. Usually, water content was measured after each T-bar test, except for a few

tests that the water content was not measured. The average water contents of the clay

penetrated by the T-bar tests for each test group are summarized in Table 3.3. The test

results are shown in Figure 3.7, Figure 3.8, and Figure 3.9.

Table 3.3 T-bar Test Groups

Test groups Average water

content (%) Number of Tests

A 125 7

B 124 4

C 140 2

27

Figure 3.7 Undrained shear strength Su vs. time (Melo Monteiro, 2019)

Figure 3.8 Remolded undrained shear strength Sur vs. time (Melo Monteiro, 2019)

28

Figure 3.9 Sensitivity (Su/Sur) vs. time (Melo Monteiro, 2019)

Figure 3.10 Sur vs. water content (Melo Monteiro, 2019)

29

Based on the previous results of T-bar tests, Su stayed unchanged for the first 1-2

hours, which was followed by an increasing trend until 72 hours after remixing. After 72

hours, except for one outlier from Test Group A, no significant increase was observed.

Also, Melo Monteiro (2019) shows that water content plays a huge part in the remolded

undrained shear strength Sur. On a semi-log scale, logSur decreases linearly with

increasing water content. Compare Figure 3.10 with the results of the test beds shown in

Table 3.2, the measured Sur values of the test beds at specified water contents are very

close to the trendline in Figure 3.10. The lower the water content is, the stronger the test

bed will be, which requires an accurate estimation of the pipe unit weight in need to reach

the desired embedment ratio.

In addition, Figure 3.9 shows that sensitivity of the clay increased with time after

mixing. Clay sensitivity represents the thixotropy phenomenon. Thixotropy is a reversible

and time-dependent process where the soil particles and structured solutes rearrange

themselves, which leads to more thick and viscous mixture in time, hence the increased

undrained shear strength Su. In addition, as the soil is disturbed and remixed, the

thixotropic soil/fluid will return to an equilibrium viscosity where the remolded

undrained shear strength Sur is defined. Thixotropy causes increased axial load resistance

after a long pause between loading sweeps, which will be discussed in detail in Chapter

5.

3.1.3 Drained Residual Shear Strength between Soil-Polypropylene Interface

Melo Monteiro (2019) performed tilt table tests on the GOM Clay used in this

study to develop the Mohr-Coulomb failure envelope for the drained residual interface

shear strength at different effective normal stress. Just like the results reported by Najjar

30

(2003 & 2007), the curved decreasing trend of drained residual shear resistance with

increasing effective normal stress was also obtained from these tests.

Figure 3.11 Drained residual friction coefficient from tilt table tests (after Melo Monteiro,

2019)

3.2 TEST BED

Two test beds were constructed at difference specified water content. The Gulf of

Mexico Clay was placed as a single layer in a 10-foot long by 4-foot wide open tank with

1-foot depth. Salt water (35 grams of salt per liter of water) was used to mix with the clay

to represent the salinity level of the sea water of the Gulf of Mexico. Desired water

contents were achieved by mixing the clay with adequate amount of water using a power

mixer. To examine the effects of change in effective overburden pressure on axial

resistance, test bed 1 was mixed to a water content of around 140 % (simulating a marine

clay bed with no “crust”) while test bed 2 was prepared at a water content of around 120

% (simulating a clay bed with a “crust” which was able to support heavier pipes). After

0.00

0.10

0.20

0.30

0.40

0.50

0.60

0.70

0.80

0.90

0 20 40 60 80 100 120

Dra

ined r

esid

ual fr

iction c

oeff

icie

nt

Effective normal stress (psf)

31

thoroughly being thoroughly mixed, the clay was covered by additional free salt water

with the same salinity to ensure full submersion of pipelines to be placed. The test bed

soil then consolidates at its own weight for several weeks before any axial loading tests.

Figure 3.12 Powered soil mixer

3.3 FLOWLINES

The model flowlines are 4.25-inch diameter polypropylene pipes with an outer

surface simulating the flowline coating used in practice. The lengths of the pipes are 6 ft

and 9 ft, respectively. End effects are neglected as the length to diameter ratio is larger

than 15 for both pipes. The unit weight of the pipes is adjusted by placing ballast inside

the pipes and varies from 3 lb/ft to 6 lb/ft when submerged in seawater used in the test

bed.

3.4 ELECTRIC MOTOR

A stepper motor (powered by a Superior Electric SLO-SYN MH112-ff-206) was

used to control the displacement rate of the loading cables for the T-bar tests and the axial

32

loading tests. As shown in Figure 3.13, the motor is mounted on a loading table, which is

then stabilized by load pieces to prevent slipping along the floor. The electric motor

system contains two linear actuators along two orthogonal directions, two stepper motors,

two translator drivers, and a computer controller card (Gerkus, 2016). The horizontal

moving range of the stepper motor is limited to 12.5 inches.

Figure 3.13 Stepper motor

3.5 LOAD FRAMES

Three load frames are used in this test program. To apply axial loading to the

flowlines, two wooden frames with pulleys were fabricated to establish a conveyance

system. Cables are connected to both ends of the pipeline through the wooden frames to

apply axial loads.

33

Figure 3.14 Load frames

In addition, a 4.7 ft tall by 5 ft wide aluminum load frame is exclusively used for

T-bar tests. The function of the load frames will be further discussed in the later sections.

3.6 LOAD CELL

Two load cells manufactured by Lebow Products Inc. were used to measure the

loads during the tests. The maximum load capacities of the load cells are 100 lbs and 200

lbs, respectively. Only tensile forces were measured during the tests. The measured

output voltage from the load cells were converted to the actual tensile forces by

subtracting the zero-force voltage and multiplied by the calibration factors. Regular

34

recalibrations were performed throughout the tests to ensure accurate conversion of the

results.

Figure 3.15 Load cell (Huang, 2015)

3.7 LINEAR MOTION TRANSDUCER (LMT)

The linear motion transducer (LMT) is used to measure displacements during the

T-bar tests. The LMT is a RayecoTM model P-50. With a range of 50 inches, the LMT is

attached to a track system on the aluminum load frame as shown in Figure 3.16. For a T-

bar test, the penetrometer rod is connected to the LMT through a plastic wire. Friction of

the system was determined negligible for the T-bar test, which requires no friction

correction (Melo Monteiro, 2019).

35

Figure 3.16 Linear motion transducer (Melo Monteiro, 2019)

3.8 LINEAR VARIABLE DIFFERENTIAL TRANSFORMER (LVDT)

A linear variable differential transformer (LVDT) with a range of 6 inches is used

to measure the displacement of the pipeline during axial loading tests. The needle of the

LVDT is clamped to the pipe to move along the pipe, while the LVDT body is fixed to

the tank through wood pieces. Same as the load cells, the LVDT is also calibrated on a

regular basis to ensure the accuracy of the recorded data.

36

Figure 3.17 LVDT connection to (left) the tank and (right) the pipe

3.9 DATA ACQUISITION AND MOTION CONTROL PROGRAMS

Huang (2015) first developed the data acquisition and motion control system used

in this study on LabVIEW platform. Data is recorded by the Data Acquisition (DAQ)

hardware (Figure 3.18) and motion control card manufactured by National Instruments.

The DAQ system records the output voltage data from all sensors and exports the signals

as csv files.

37

Figure 3.18 LabVIEW user interface (Huang, 2015)

The LabVIEW user interface (Figure 3.18) consists of a controlling panel, which

prompts the operator to start or finish a recording task, and a calibration area, which

allows the user to type in the calibration factors. Additionally, the recorded sensor

voltages can be instantaneously displayed as the test proceeds, which can be seen on the

right side of the interface.

3.10 AXIAL LOAD MODEL TEST SETUP

3.10.1 Test Bed Mixing

Test beds were prepared by thoroughly mixing the GOM Clay with salt water to

the desired water content using a powered soil mixer. After mixing, the water content of

the test bed was measured at different locations. If the water content was too high, the test

38

bed is left drying for days until the specified water content is reached. If the measured

water content was too low, additional salty water was added and mixing is repeated.

Once the desired water content was reached, the clay surface was smoothed to

prevent uneven surfaces. Figure 3.19 shows Test Bed 2 after mixing and surface

smoothing. Salt water was then carefully added into the tank to cover the clay and to

ensure the pipes were submerged later.

Figure 3.19 Immediately after mixing and smoothing of Test Bed 2

3.10.2 Load Conveyance System

Axial loads were applied to the pipes by cables through a load conveyance system

consisting load frames with pulleys. The linear actuator and the motor were connected to

39

one end of the pipe. Meanwhile, a counterweight was attached to the other end of the pipe

to maintain a constant load. Otherwise, the counterweight was replaced by a spring to

simulate the stiffness of flowlines during pipe walking (Figure 3.20).

A schematic of the pipe loading system is shown in Figure 3.20.

Figure 3.20 Test setup of axial load tests with linear actuator and (a) counterweight or (b)

spring (Hussien, 2020)

40

Figure 3.21 (a) Counterweight (b) Spring

3.10.3 Instrumentation

Load cells:

As illustrated in Figure 3.22, two load cells were used to measure the tensile force

of the cables on the motor side and the counterweight side. In addition, pulleys were used

between the loadcells and the ends of the pipe on both sides. As a result, friction forces

induced by the pulleys need to be accounted for. Pulley friction correction is explained in

detail in test procedure section.

41

Figure 3.22 (a) Motor-side loadcell (b) Counterweight-side loadcell

LVDT:

As shown in Figure 3.23, a LVDT was attached to an outpost fixed on the crown

of the pipe to measure the horizontal displacement. The needle of the LVDT was

clamped to the pipe outpost, and the LVDT was mounted to a wood frame that is further

fixed to the clay tank.

42

Figure 3.23 LVDT installation overview

Figure 3.24 LVDT

43

When an axial load test was taking place, the pipe dragged the needle along with

its movement, and the LVDT could measure and export the output voltage change

induced by the relative movement.

44

Chapter 4 T-bar Tests and Embedment Tests

4.1 PIPE EMBEDMENT TESTS

4.1.1 Initial Embedment

The target of pipe laydown is to have a final embedment ratio of around half of

the pipe diameter (0.5D). To achieve the desired embedment ratio, the embedment

estimation method from SAFEBUCK JIP is used (Simpson et al., 2015). Equation 4.1

estimates the pipe embedment,

𝑉

𝐷∙𝑆𝑢= min [6 (

𝑧

𝐷)

0.25

; 3.4 (10∙𝑧

𝐷)

0.5

] + 1.5𝛾′∙𝐴𝑏𝑚

𝐷∙𝑆𝑢 (Eq. 4.1)

where V is the pipe submerged unit weight, Su is the undrained shear strength at

pipe invert, D is the outer diameter of the pipe, z is the estimated pipe embedment, 𝛾’ is

the soil submerged unit weight, and Abm is the pipe submerged cross-sectional area.

Iteration is required to calculate the embedment depth of the pipe.

45

Based on the water content measurements and T-bar test results, the unit weights

of the pipes are determined and adjusted. The submerged unit weights of the pipes are

summarized in Table 4.1.

Table 4.1 Test Beds information

Test Beds

Remolded

Undrained Shear

Strength (psf)

Average water

content (%)

Submerged Unit Weights (lb/ft)

6 ft pipe 9 ft pipe

1 1.6 138 2.7 (Pipe 2) 3 (Pipe 1)

2 2.1 124 6 (Pipe 4) 6 (Pipe 3)

Using Equation 4.1, the initial embedment depth for the four pipes were

calculated (Table 4.2).

Table 4.2 Initial embedment estimation for Pipes 1-4

Pipes Estimated initial

embedment (in)

Estimated initial

embedment normalized by

pipe diameter (D)

Measured initial

embedment normalized

by pipe diameter (D)

1 1.01 0.26 0.24

2 0.81 0.22 0.28

3 0.53 0.47 0.47

4 0.56 0.31 0.31

46

After the test beds were thoroughly mixed, pipes were placed on top the clay by

hand. The target of pipe laydown was to place the pipe on the test bed surface as gentle as

possible, which allows the pipes to settle only by their self-weights. Unfortunately, the 9-

ft pipe for Test Bed 2 was accidently dropped about one pipe diameter above the test bed

surface, leading to an additional dropping force and excessive settlement compared with

the prediction. The settlements of the pipes were constantly monitored for a period of

time after placement. Figure 4.1 shows the pipe embedment measurements. As shown in

Table 4.2, the measured initial embedment ratios for Pipe 3 and Pipe 4 were much

smaller than the estimation. Even for Pipe 3, the measured embedment was still smaller

than the estimation with the dynamic loading effect. One of the possible explanations for

this discrepancy might be related to inconsistent water contents across Test Bed 2. As the

Test Bed was gigantic, mixing the clay thoroughly and evenly became hard. It was

possible that the water content was measured at a stronger region of the test bed, while

the pipes were placed at regions that were weaker with potentially higher water contents.

The initial embedment ratio varies from about 0.25D to 0.3D for gently placed pipes,

while the dynamically placed pipe had an initial embedment ratio of 0.5D. The final

embedment ratios measured before axial load tests range from 0.4D to 0.65D.

47

Figure 4.1 Measured pipe embedment (After Hussien, 2020)

4.1.2 Consolidation after Initial Embedment

Because the pipes were placed by hand, additional time was required to setup the

instrumentation to measure the embedment of the pipes after placement. As a result, the

embedment measurements only started minutes after the placement of the pipe.

Therefore, identifying the initial embedment ratio became hard. It was decided that the

initial embedment was extrapolated using the linear trend of a classical Square-Root-of-

Time method consolidation curve based on the Terzaghi 1D consolidation theory. The

settlement curves for the four pipes were fitted, and Cv values were calculated using

Equation 4.2,

𝑐𝑣 =0.159×(𝐻𝑑𝑟)2

𝑡45 (Eq. 4.2)

Initial Embedment

48

where Cv is the coefficient of consolidation, 𝐻𝑑𝑟 is the drainage length which is

taken as the diameter of the pipe based on the assumption from Krost et al. (2010), and t45

is the time for 45% primary consolidation. Table 4.3 shows the calculated Cv values. The

calculated Cv values are 1 to 2 magnitudes larger than those measured by the oedometer

test. White et al. (2019) explained that drainage is about 10 times faster for a curved soil-

pipe interface (e.g. a pipe sitting on soil) than that of a strip due to three-dimensional

effects. Because of factors such as anisotropy of stiffness and permeability, the operative

Cv value may be 1 to 2 magnitude higher than that obtained by an oedometer test.

Table 4.3 Consolidation parameters using Square-Root-of-Time Method

Pipes t45 (hr) Cv (ft2/year)

1 1.96 89

2 2.25 76

3 0.64 273

4 1.21 141

49

Figure 4.2 Consolidation of Pipe 1


50



51

Comparing consolidation curves of Pipe 2 and Pipe 3, it seemed like Pipe 3 was

consolidating at a much faster rate. Therefore, pore pressure dissipation along the pipe-

soil interface for Pipe 3 probably happened faster and essentially finished earlier than

Pipe 2. Consequently, it was expected that continued pore pressure dissipation would last

longer for Pipe 2 than for Pipe 3. According to Figure 4.6, based on the measured

consolidation trends, it was possible for Pipe 1 and Pipe 2 to further consolidate after the

start of the axial load tests, which potentially last around 100 days for Pipe 2. This

observation is further elaborated in Chapter 5.

Figure 4.6 Predicted consolidation for Pipe 1 and 2

52

4.1.3 Torsional Spinning Test

A simple torsional spinning test was performed to indirectly measure the

embedding and pore pressure dissipation process within a week after pipe placement. The

test was performed in a smaller test bed using a 3-ft pipe with the same coating and

diameter as the other model pipes. The clay was mixed to a water content of 123%, and

the torsional resistance force was measured using a digital pound meter.

The test only lasted for 7 days, and it was then suspended due to the coronavirus

pandemic.

Figure 4.7 Torsional Spinning Test Schematic

53

Figure 4.8 Torsional Spinning Test being performed

The torsional spinning test was performed on a test bed with similar water content

as Test Bed 2, and the measured embedment depth of the 3-ft pipe was also similar to

Pipe 4. As shown in Figure 4.9, the measured torsional resistance force was scaled to fit

the consolidation curve of Pipe 4. Basically, right after the pipe was placed, due to large

amount of pore pressure generation, the pipe was “floated” by the pore pressure and span

almost freely. As time passed, the amount of excessive pore pressure was reduced, and

the pipe embedded more. As a result, the torsional resistance also increased.

54

Figure 4.9 Scaled Torsional Spinning Test results vs Pipe 4 embedment time history

4.2 EMBEDMENT TESTS CONCLUSIONS

1. The initial embedment estimation equation introduced by SAFEBUCK JIP

Merged Guideline provided decent starting points to predict the actual initial embedment

of the model pipes.

2. Based on the consolidation measurements of the pipes, the consolidation rate at

the soil-pipe interface for Pipe 2 (Test Bed 1) was seemingly slower than that for Pipe 3

(Test Bed 2). According to the embedment measurements, it was possible for the soil-

pipe interface of Pipe 2 to experience pore pressure dissipation for up to 100 days.

55

Available drained shear strength along the soil-pipe interface for Pipe 2 was probably

affected by this trend.

3. Torsional Spinning Test results confirmed the continued process of pore

pressure dissipation along the soil-pipe interface after a pipe was placed for up to seven

days. Torsional resistance was observed to increase with time, which was likely

associated with the pore pressure dissipation process.

56

Chapter 5 Axial Loading Tests

5.1 TEST PLAN

The axial load tests schedule is shown in Table 5.1.

57

Table 5.1 Axial load testing plan

Test Beds Pipe

Number

Pipe

Length (ft)

Submerged Unit

Weights (lb/ft)

Test Specifics

No. of sweeps Intended

Speeds (in/s)

1

1 9 3

3 2×10-5

4 1×10-5

2 2×10-5

4 1×10-4

2 6 2.7

36 2×10-3

2 1×10-5

4 2×10-3

2

3 9 6

4 3.5×10-2

4 2.9×10-3

37

3.5×10-2 &

2.9×10-3

(Alternating)

15 1×10-5

4 6 6 2 1×10-5

58

5.2 TEST PROCEDURES

5.2.1 System Load-up

After the pipes were connected to the loading system, the cables on both sides of

the pipe were tensioned up by moving the motor and slowly increasing the

counterweight. The goal was to have a large enough counterweight to pull the pipe away

from the motor, and cautionary steps were taken to prevent any shock loads and sudden

movements. For example, if the estimated axial force resistance by the soil-pipe interface

was about 20 lb, then the deadweight should be larger than 20 lb. In addition, during the

load-up phase of the test, tensile forces of cables on both sides should be increased

carefully with small increments. The load difference between the motor force and

counterweight should never exceed 10 lb during load-up. After the targeted

counterweight was reached, the system was monitored for at least 24 hours before axial

load testing started. Even though the pipes tested were deemed long enough so end

effects could be ignored, clay at both ends of the pipe was cleared away manually before

axial load tests to minimize end resistance.

5.2.2 Load Sweep towards the Motor

To start an axial load sweep towards the motor, the operator sets the motor to

move away from the tank to pull the cable towards on the motor. The tensile force builds

up in the cable, and a load difference is induced between the cables across the pipe. The

net axial force is then pulling the pipe towards the motor, and the soil-pipe interface shear

strength is gradually mobilized. The net axial force keeps increasing until shear failure is

reached. The pipe then moves towards the motor, and the axial resistance is measured and

recorded by the sensors. When the pipe displacement reaches the target, the motor is

stopped. The net axial force drops below the available shear strength, and the load sweep

59

is finished. Figure 5.1 and Figure 5.2 illustrate the mobilization process of the axial

resistance.

Figure 5.1 Motor-side loadcell measurements of sweep 9 of Pipe 1

60

Figure 5.2 Counterweight loadcell measurements of sweep 9 of Pipe 1

Figure 5.3 Axial force of sweep 9 of Pipe 1

61

In addition, to investigate the rate effects, the axial moving speeds were divided

into four categories based on the speed magnitude: “Very fast” - 1×10-2 in/s, “Fast” -

1×10-3 in/s, “Intermediate” - 1×10-4 in/s, and “Slow” - 1×10-5 in/s.

5.2.3 Load Sweep towards the Counterweight/Spring

To initiate a load sweep towards the counterweight/spring, the motor is set to

move towards the tank. Subsequently, motor-side tensile force drops, and the net axial

force now acts towards the counterweight/spring. Same as load sweeps towards the

motor, the pipe will start moving when the motor-side cable tension drops to the point

when the net axial force is larger than the shear strength. For the test setup using a

counterweight, the sweep is stopped by terminating the motor. For the test setup using a

spring, as the pipe moves towards the spring, the spring force keeps decreasing. The

sweep automatically stops when the spring force drops below the axial resistance.

5.2.4 Pulley Friction Correction

As shown in Figure 3.20, two pulleys are used between the loadcells and the pipe.

If the pulley friction is not accounted for, the measured axial force will be overestimated.

The setup of pulley friction measurement test is illustrated in Figure 5.4. The only

difference between the pulley friction measurement test and an axial load test is that the

pipe is removed from the system. In this way, the difference between the measurements

of the loadcells is the pulley friction induced from the two pulleys (P3 and P4).

62

Figure 5.4 Pulley friction measurement test (pulling towards the motor)

According to the friction measurement results, pulley friction increases with

increasing deadweight. The correlation between the loadcell measurements difference

and the counterweight is plotted in Figure 5.5.

Figure 5.5 Pulley friction vs counterweight

63

An important assumption for pulley friction correction is that the two pulleys are

assumed to have identical friction properties. In other words, the pulley friction

measurement result is assumed to be the sum of two identical friction force from the two

pulleys. With the correlation established, pulley friction correction is performed

separately for the two pulleys based on the measurements of the loadcells.

For example, Figure 5.6 shows the process of applying pulley friction to an axial

sweep towards the motor. The equation for corrected net axial force is,

𝐴𝐹𝑐 = 𝑀𝐿𝐶 − 𝐶𝐿𝐶 − 𝐹𝑃3 − 𝐹𝑃4 (Eq. 5.1)

where AFc is the corrected axial force, MLC is the motor-side loadcell

measurement, CLC is the counterweight-side loadcell measurement, FP3 is the friction

from pulley P3, and FP4 is the friction from pulley P4.

Figure 5.6 Example of applying pulley friction (towards motor)

For this example, the estimated pulley friction forces from P3 and P4 are

64

𝐹𝑃3 = 𝐶𝐿𝐶 × 0.01255 = 58.065 𝑙𝑏 × 0.01255 = 0.73 𝑙𝑏

𝐹𝑃4 = 𝑀𝐿𝐶 × 0.01255 = 85.787 𝑙𝑏 × 0.01255 = 1.08 𝑙𝑏

Therefore, the corrected axial force is

𝐴𝐹𝑐 = 𝑀𝐿𝐶 − 𝐶𝐿𝐶 − 𝐹𝑃3 − 𝐹𝑃4 = 26 𝑙𝑏

5.2.5 Normalized Axial Resistance Calculation

The axial pipe-soil shear resistance is represented by a normalized friction

coefficient. As shown in Eq. 5.2, the friction coefficient is defined as the ratio of the

measured net axial force over the submerged weight of the pipe in salt water,

𝜇 =𝐴𝐹𝑐

𝑉 (Eq. 5.2)

where, μ is the axial friction coefficient, AF is the corrected axial force, and V is

the submerged pipe weight.

5.3 TEST SERIES 1 – 9-FT PIPE IN TEST BED 1

The first pipe was a 9-ft pipe with a submerged unit weight of 3 lb/ft on Test Bed

1. Before the first sweep of the axial test, 25 days were waited to allow the test bed to

consolidate. The test results are summarized in Table 5.2.

65

Table 5.2 Axial Load Test results of 9-ft 27-lb Pipe with counterweight

Sweep Direction

Pulled

Intended

Speed (in/s)

Large Displacement Speed

(in/s)

Coefficient of

Friction

1 Deadweight 2×10-5 2.50×10-5 0.547

2 Deadweight 2×10-5 2.00×10-4 0.611

3 Motor 2×10-5 2.00×10-5 0.468

4 Motor 1×10-5 1.75×10-5 0.547

5 Deadweight 1×10-5 3.00×10-4 0.655

6 Motor 1×10-5 1.50×10-5 0.517

7 Deadweight 1×10-5 3.00×10-4 0.684

8 Deadweight 2×10-5 2.50×10-5 0.683

9 Motor 2×10-5 2.25×10-5 0.510

10 Deadweight 1×10-4 1.10×10-4 0.669

11 Motor 1×10-4 1.10×10-4 0.502

12 Motor 1×10-4 1.10×10-4 0.613

13 Deadweight 1×10-4 1.10×10-4 0.652

For axial load tests on Pipe 1, the adopted speeds were relatively slow. Under

such conditions, the drained response of the soil-pipe interface was expected. The

measured axial force versus displacement for each sweep are plotted in Figure 5.7.

Positive axial forces and friction coefficients stand for sweeps towards the motor, and

negative axial forces and friction coefficients are for sweeps towards the counterweight.

This sign convention is consistent throughout the tests on the four pipes.

The first feature to notice from Figure 5.7 is that a breakout peak was observed

when the axial load test is performed after a 5-day pause in testing. This observation is

consistent with conclusions of other literatures. The additional axial resistance induced by

a pause decreased with shearing.

66

Figure 5.9 shows the measured friction coefficient of each axial sweep. Overall,

the measured residual friction coefficient ranges from 0.5 to 0.68 with a slightly

increasing trend. Since only 13 sweeps were performed on Pipe 1, it is unlikely that the

drained residual state was reached, and the friction coefficient might keep increasing

should more sweeps were carried out. In addition, the slight increasing trend of the axial

resistance might serve as evidence of continuing consolidation of the soil-pipe interface.

If the soil under Pipe 1 was allowed to fully consolidate, the measured axial resistance

would probably increase as the effective contact stress decreases. A better way of

illustrating the effects of test bed consolidation is to plot the friction coefficient against

time elapsed since pipe placement (Figure 5.10).

Figure 5.7 Axial force vs Displacement from cyclic axial load tests performed at "slow"

and "intermediate” speeds with counterweight, Pipe 1 in Test Bed 1

67

Figure 5.8 Friction coefficients vs displacement from cyclic axial load tests performed at

"slow" and "intermediate” speeds with counterweight, Pipe 1 in Test Bed 1

Figure 5.9 Friction coefficient, μ vs sweep - Pipe 1

68

Figure 5.10 Friction coefficient, μ vs time elapsed after pipe placement - Pipe 1

Although the motor was set to move at intended speeds, the measured actual

speeds of the pipe sometimes deviated from the intended speeds. This was due to the

inability of cables to timely adjust to sudden movement of the pipe, which often led to a

jagged shaped displacement plot such as sweep 4 and 11 of Pipe 1 (Figure 5.11). Before

the pipe moved, the axial force kept building up as the cables were tightened. Right after

the axial force surpassed the axial resistance, the pipe started moving, and the axial force

started to drop as the specified speed of the motor was slower than the actual speed of the

pipe. As a result, the recorded axial force dropped to the point where the axial force was

no longer able to pull the pipe, which led to a complete stop of the pipe movement. Then,

the tensile force in the cable started to build up again, and the cycle repeated until the test

was finished. This issue also happened to Pipe 3, which was then mitigated by using

heavier counterweights to stiffen the loading system.

69

Figure 5.11 Pipe transient speeds vs displacement, Pipe 1 - sweep 4 and 7


Pipe 2 is a 9-ft pipe with a submerged unit weight of 3 lb/ft on Test Bed 1. The

test results are summarized in Table 5.3.

70

Table 5.3 Axial Load Test results of 6-ft 16.4-lb Pipe with counterweight

Sweep Direction

Pulled


(in/s)

Coefficient of

Friction

1 Motor 1.5×10-3 0.321

2 Spring 1.5×10-3 0.358

3 Motor 1.5×10-3 0.447

4 Spring 1.5×10-3 0.507

5 Motor 1.5×10-3 0.549

6 Spring 1.5×10-3 0.564

7 Motor 1.5×10-3 0.595

8 Spring 1.5×10-3 0.592

9 Motor 1.5×10-3 0.646

10 Spring 1.5×10-3 0.610

11 Motor 1.5×10-3 0.666

12 Spring 1.5×10-3 0.621

13 Motor 1.5×10-3 0.673

14 Spring 1.5×10-3 0.666

15 Data lost

16 Spring 1.5×10-3 0.626

17 Motor 1.5×10-3 0.741

18 Spring 1.5×10-3 0.672

19 Motor 1.5×10-3 0.753

20 Spring 1.5×10-3 0.651

21 Motor 1.5×10-3 0.741

22 Spring 1.5×10-3 0.646

23 Motor 1.5×10-3 0.792

24 Spring 1.5×10-3 0.717

25 Motor 1.5×10-3 0.816

71

Table 5.3 Continued

26 Spring 1.5×10-3 0.690

27 Motor 1.5×10-3 0.829

28 Spring 1.5×10-3 0.723

29 Motor 1.5×10-3 0.834

30 Spring 1.5×10-3 0.721

31 Motor 1.5×10-3 0.986

32 Spring 1.5×10-3 0.768

33 Motor 1.5×10-3 1.118

34 Spring 1.5×10-3 0.860

35 Motor 1.5×10-3 0.947

36 Spring 1.5×10-3 0.805

37* Motor 1.3×10-5 0.940

38* Spring 5×10-6 to 2×10-3 0.854

39 Deadweight 1.8×10-3 0.956

40 Data lost

41 Deadweight 1.8×10-3 0.849

42 Deadweight 1.8×10-3 1.003

* Sweeps were performed at slow speeds.

For Pipe 2, the counterweight used in Pipe 1 was replaced by a spring as a

surrogate for flowline stiffness in the field during flowline walking (Figure 3.21 (b)).

To simulate walking, the first 36 axial sweeps of Pipe 2 were performed at a

“fast” loading speed at 1.5×10-3 in/s. Undrained response from the soil-pipe interface was

expected, and excess pore pressure was generated. After the pipe was pulled towards the

motor, the motor was set to move back, and the spring pulled the pipe towards the other

72

end. However, the spring force kept changing as the sweep progressed, the speed was not

maintained at a constant rate and the pipe was observed to move in a stick-slip manner.

Due to pore pressure generation at the “fast” speed, the test was paused for at least 8

hours to allow pore pressure dissipation. The test results of the first 36 sweeps are plotted

in Figure 5.12 and Figure 5.13.

Figure 5.12 Axial force vs Displacement from cyclic axial load tests performed at "fast”

speeds with spring, Pipe 2 in Test Bed 1 (After Hussien, 2020)

73

Figure 5.13 Friction Coefficient vs Displacement from cyclic axial load tests performed

at "fast” speeds with spring, Pipe 2 in Test Bed 1 (After Hussien, 2020)

Same as Pipe 1, when an axial sweep was performed after a long pause, a

breakout peak was observed. Similarly, the initial sweep also has a big peak at the

beginning, which is reasonable as the initial sweep of Pipe 2 was performed 54 days after

the pipe was placed.

As the resisting axial strength increased through shearing, the spring could not

pull the pipe back to its original position before the previous sweep towards the motor.

Therefore, the final position of the pipe after a sweep towards the spring gradually moved

towards the motor.

74

The measured large-displacement friction coefficients of each sweep at “fast”

speed are shown in Figure 5.14.

Figure 5.14 Friction coefficient, μ vs sweep at “fast” speeds - Pipe 2

According to Figure 5.14, the axial resistance steadily increased from about 0.3 to

0.9 with the undrained load cycles. For initial sweeps, large amount of positive pore

pressure was generated during shearing, which led to undrained shear resistance smaller

than the drained shear resistance. As the number of sweeps increased, soil under the pipe

got stronger and stronger with less pore pressure generated. Theoretically, the axial

resistance should reach a maximum value in the end when the drained residual state is

reached. Once the drained residual state was reached, a fast sweep would no longer

induce significant positive pore pressure, hence the expected friction coefficient should

be close to that obtained from a slow sweep.

75

In order to confirm that the drained residual state was reached, two “slow” sweeps

were performed after the “fast” sweeps. The results of the “slow” sweeps are shown in

Figure 5.15.

Figure 5.15 Friction coefficient vs displacement from cyclic axial load tests performed at

"slow” speeds with spring, Pipe 2 in Test Bed 1 (After Hussien, 2020)

As shown in Figure 5.15, the measured friction coefficients of the “slow” sweeps

were around 0.9. It was evident that the drained residual state had been reached.

During sweep 38, possibly due to the same reason that caused jagged movement

curves for Pipe 1, the measured speed jumped to 2×10-3 in/s and returned to the intended

speed (Figure 5.16). However, with the movement rate of the pipe varying from 5×10-

76

6 in/s to 2×10-3 in/s, the recorded friction coefficient did not change drastically. This

serves as another piece of evidence that the drained residual state was reached.


At the end of the “slow” sweeps, additional fast sweeps were performed with a

counterweight after a 6-day pause. As expected, the recorded large-displacement friction

coefficients of these fast sweeps remained close to 0.9 (Figure 5.17).

77


"slow” speeds with counterweight, Pipe 2 in Test Bed 1 (After Hussien,

2020)

From sweep 39, after a 6-day pause, the measured friction coefficient shows a

peak initially and rapidly decreases to the drained residual friction coefficient. The

drained residual friction coefficient remained at 0.9.

The measured normalized axial resistance of Pipe 2 are shown in Figure 5.18.

78

Figure 5.18 Friction coefficient vs sweep - Pipe 2

Figure 5.19 Friction coefficient vs time after pipe placement - Pipe 2

79


After tests on Pipe 2 were finished, the pipes were removed from Test Bed 1, and

the clay was dried out and mixed again to a decreased water content of about 120 %.

Hence, Test Bed 2 was stronger with larger undrained shear strength compared with Test

Bed 1, which required Pipe 3 and Pipe 4 to be heavier to embed 0.5 D.

The testing sequence for Pipe 3 was determined based on the same rationale as

that for Pipe 2. For the first 45 sweeps of Pipe 3, “Very fast” and “fast” speeds were

carried out to simulate flowline walking. Pipe 3 has a submerged unit weight of 6 lb/ft,

and the “very fast” and “fast” test results are summarized in Table 5.4.

Table 5.4 “Very fast” and “fast” axial sweep results of 9-ft 53.7-lb Pipe with

counterweight

Sweep Direction

Pulled


(in/s)

Coefficient of

Friction

1 Motor 3.3×10-2 0.382

2 Deadweight 3.3×10-2 0.359

3 Motor 3.3×10-2 0.366

4 Deadweight 3.3×10-2 0.383

5 Motor 2.9×10-3 0.345

6 Deadweight 2.9×10-3 0.309

6a* Motor 2.9×10-3 0.316

7 Motor 2.9×10-3 0.338

8 Deadweight 2.9×10-3 0.386

9 Motor 3.3×10-2 0.542

10 Deadweight 2.9×10-3 0.443

11 Motor 3.3×10-2 0.620

12 Deadweight 2.9×10-3 0.471

13 Deadweight 2.9×10-3 0.474

80

Table 5.4 Continued

14 Motor 3.3×10-2 0.654

15 Deadweight 2.9×10-3 0.485

16 Motor 3.3×10-2 0.628

17 Deadweight 2.9×10-3 0.478

18 Motor 3.3×10-2 0.627

19 Deadweight 2.9×10-3 0.478

20 Motor 3.3×10-2 0.631

21 Deadweight 2.9×10-3 0.470

22 Motor 3.3×10-2 0.633

23 Deadweight 2.9×10-3 0.467

24 Motor 3.3×10-2 0.622

25 Deadweight 2.9×10-3 0.474

26 Motor 3.3×10-2 0.626

27 Deadweight 2.9×10-3 0.477

28 Motor 3.3×10-2 0.639

29 Deadweight 2.9×10-3 0.476

30 Motor 3.3×10-2 0.641

31 Deadweight 2.9×10-3 0.489

32 Motor 3.3×10-2 0.653

33 Deadweight 2.9×10-3 0.498

34 Motor 3.3×10-2 0.697

35 Deadweight 2.9×10-3 0.497

36 Motor 3.3×10-2 0.652

37 Deadweight 2.9×10-3 0.502

38 Motor 3.3×10-2 0.670

39 Deadweight 2.9×10-3 0.501

81

Table 5.4 Continued

40 Motor 3.3×10-2 0.668

41 Deadweight 2.9×10-3 0.499

42 Motor 3.3×10-2 0.614

43 Deadweight 2.9×10-3 0.496

44 Motor 2.9×10-3 0.433

45 Deadweight 3.3×10-2 0.663

* After cable snapped during sweep 6, Pipe 3 was repositioned in sweep 6a.

Figure 5.20 Friction coefficient vs sweep - Pipe 3 at “very fast” and “fast” speeds

For sweeps performed at the “very fast” speed, the friction coefficient increased

from 0.36 to 0.65, while for sweeps at the “fast” speed, the friction coefficient varied

from 0.31 to 0.5. The alternating “very fast” and “fast” sweeps showed a rate effect, in

82

which the “very fast” speed sweep induced a larger friction coefficient than that from a

“fast” sweep. This difference in axial resistance could be caused by negative pore

pressure generation during the “very fast” sweeps, which led to increased effective

contact stress and increased shear strength.

At the end of the “fast” sweeps, an outlier with its friction coefficient lower than

the trend was measured. This was caused by directional difference. Due to the complexity

of the load conveyance system, it is hard to calibrate and correct the data to have identical

friction coefficients towards both directions. Similar results were also observed from Pipe

2, where sweeps towards one direction constantly produced larger calculated friction

coefficients than the other. For this study, the observed difference between directions is

not significant. However, if the difference between the two directions gets bigger, it is

recommended that the sensors should be recalibrated, and the pulley friction forces

should be measured again.

After 45 “very fast” and “fast” sweeps, additional “slow” sweeps were performed

to confirm that the drained residual state was reached. Table 5.5 summarizes the test

sequence.

83


"very fast” and “fast” speeds with counterweight, Pipe 3 in Test Bed 2

84

Table 5.5 Axial Load Test results of 6-ft 16.4-lb Pipe with counterweight

Sweep Direction

Pulled

Large Displacement

Speed (in/s) Speed Type

Coefficient of

Friction

46 Motor 1.3×10-5 Slow 0.470

47 Deadweight Results discarded due to insufficient counterweight

48 Motor 1.3×10-5 Slow 0.434


50 Motor 1.3×10-5 Slow 0.425


52* Motor 1.3×10-5 Slow 0.431

53 Motor Results discarded due to insufficient counterweight


55 Motor 1.3×10-5 Slow 0.502


57a Motor 1.3×10-5 Slow 0.479

57b Motor 1×10-4 Intermediate 0.479

57c Motor 1×10-3 Fast 0.491

58a Deadweight 1.3×10-5 Slow 0.442

58b Deadweight 1×10-4 Intermediate 0.457

58c Deadweight 1×10-3 Fast 0.470

59a Motor 1.3×10-5 Slow 0.474


59c Motor 1×10-3 Fast 0.485

60a Deadweight 1.3×10-5 Slow 0.424

60b Deadweight 1×10-4 Intermediate 0.410

60c Deadweight 1×10-3 Fast 0.440

61a Motor 1.3×10-5 Slow 0.386


85

Table 5.5 Continued

61c Motor 1×10-3 Fast 0.423

* Testing was suspended for 160 days due to lab construction after sweep 52

After closer examination of the data, results from sweeps 47, 49, 51, 53, 54 and

56 were deemed not presentable due to a lack of stiffness of the load conveyance system,

which was then mitigated by increasing the amount of counterweight for sweeps 57 to 61.

The axial resistance of the “slow” sweeps was similar to that measured from the

final “fast” sweeps, facilitating the conclusion that the soil interface had reached drained

residual state. In addition, axial resistance of the interface gradually dropped from sweep

46 to 52 during continued “slow” shearing. After a 160-day pause, the axial resistance

increased to 0.5, and started to decrease again from sweeps 55 to 61. A possible theory

that could explain this reduction in drained axial resistance is stress concentration at the

invert of the pipe. As “slow” shearing continued, due to the high curvature of the pipe

and inevitable lateral disturbance during shearing, the contact area between the pipe and

the test bed might be decreasing. Subsequently, the effective stress on the clay was

increased, and the available friction resistance was decreased (Figure 5.23 and Figure

5.24). Also, gaps were formed between the sides of the pipe and the trench. This is a

piece of visual evidence that supports the possibility of stress concentration under the

pipe invert (Figure 5.22).

86

Figure 5.22 Gaps between the model pipe and clay trench

Another change in the procedure of the third series of testing is an addition of

sweeps with speeds increasing from “slow” speed to “fast” speed. At the final stage of

axial load testing on Pipe 3, it was assumed that the drained residual state was already

reached. Therefore, it was expected that not only “slow” speeds, but also “intermediate”

and “fast” speeds would mobilize the same drained shear resistance along the soil-pipe

interface. The measured results plotted in Figure 5.25 demonstrate that the axial

resistance stayed constant with the moving speed increasing by 100 times from 1.3×10-

5 in/s to 1×10-3 in/s.

87

Figure 5.23 Friction coefficient vs sweep - Pipe 3

Figure 5.24 Friction coefficient vs time after pipe placement - Pipe 3

88

Figure 5.25 Friction coefficient vs displacement, Pipe 3 - sweeps 57 to 61 with increasing

speed (After Hussien, 2020)

89


Pipe 4 is a 6-ft pipe with a submerged unit weight of 6 lb/ft on Test Bed 2.

Currently, axial tests on Pipe 4 have not been finished and were unavoidably suspended

due to the coronavirus pandemic.

With lessons learned from the first three test series, the testing procedures for Pipe

4 are designed to better track the strengthening process of the soil-pipe interface through

undrained load sweeps. The revised axial test plan is shown in Table 5.6.

Table 5.6 Test plan for Pipe 4 in Test Bed 2

Test no. Number of

cycles Speed (in/s) Speed type Remarks

1 1

(2 sweeps) 1×10-5 Slow

• 6-12 hours between the two

sweeps, and this rule applies

to all the tests.

2 1 3.5×10-2 Very fast

• Wait for 3 days between the

slow and very fast tests to

dissipate any excess pore

pressure

3 2 3.0×10-3 Fast

• Wait for 3 days between the

very fast and fast tests to

dissipate the excess pore

pressure

• Wait for one week after test #

1 to 3

• Repeat test # 1 to 3 following

the same procedure until the

drained residual state is

reached.

Instead of performing “very fast” or “fast” cycles at the beginning, Pipe 4 was

first pulled at the “slow” speed. The purpose of the initial “slow” sweeps is to measure

the initial drained friction coefficient, which will be used as a benchmark to compare

against the expected increasing trend of the friction coefficients of the undrained load

sweeps at “very fast” and “fast” speeds. In this way, any changes between the final

90

drained residual shear resistance and the initial drained resistance will be captured. Also,

as the soil-pipe interface gets stronger with continued “very fast” and “fast” tests, the

difference between the initial friction coefficient and the undrained friction coefficient

can be regarded as an indirect indicator of the amount of pore pressure (positive or

negative) generation.

So far, two sweeps at “slow” speeds have been performed. The recorded large-

displacement friction coefficient of these two sweeps were around 0.6 (Figure 5.26).


"slow” speeds with counterweight, Pipe 4 in Test Bed 2

According to Figure 5.26, a small peak was found from sweep 1. Also, the friction

coefficient measured at the end of sweep 2 also increased from 0.6 to 0.7. Therefore, the

91

local axial resistance for the pipe maybe larger when the pipe is at 0-inch displacement.

This could be caused by local soil conditions of the clay near the end of the pipe towards

the counterweight. Further testing is required to verify this assumption.

Another observation is that the curves show sawtooth pattern. As discussed

before, this pattern is not favorable, and it usually indicates insufficient stiffness of the

load conveyance system. Fortunately, after a closer examination of the measured

transient speeds of Pipe 4 (), the measured speeds remain under 7×10-3 in/s most of the

time, which should be slow enough for the test results to be satisfactory. Future

improvements are required to stiffen the load conveyance system. Feasible measures

consist of increasing the counterweight and switch the cable connections to more rigid

options.


92

5.7 AXIAL LOAD TESTS CONCLUSIONS

Four series of axial load model tests were performed to simulate offshore flowline

behaviors under different loading conditions. The following conclusions can be

summarized from the test results:

1. Similar as the tilt table test results, the drained residual axial resistance of the soil-

pipe interface decreases with increasing effective overburden stress. In Figure

5.28 the large-displacement drained shear resistance of the axial load tests has a

decreasing trend as the effective normal stress increases.

2. Although the final large-displacement friction coefficient of Pipe 1 is below that

of Pipe 2 and the tilt table results, it is likely that the soil-pipe interface was still

consolidating at the end of testing for Pipe 1. Figure 5.29 shows a seemingly

increasing trend of the normalized axial resistance against time elapsed after pipe

placement. If the soil-pipe interface under Pipe 1 was allowed to consolidate for

more time, the effective contact area would be increased, and the effective contact

stress would decrease, which then led to a larger axial resistance.

3. The normalized axial resistance obtained from Pipe 3 are lower than those from

Pipe 4, which was possibly caused by stress concentration at the invert of the

pipe. As the effective normal stress used in Figure 5.28 is calculated using the

projected area of the pipe, if the contact area between Pipe 3 and Test Bed 2 was

actually smaller, the actual effective normal stress would be larger. Visual

evidence also confirms this assumption by showing that gaps were formed

between Pipe 3 and the clay. Subsequently, the green points on Figure 5.28 will

be shifted to the right and brought closer to the tilt table results.

4. Comparing Test Bed 1 and Test Bed 2, the consolidation time for the pipe-soil

system in Test Bed 1 seemed to be longer. According to Figure 5.29, the

93

increasing trend for normalized axial resistance extends until at least 90 days after

pipe placement. Figure 5.30 shows the normalized axial resistance time history of

Test Bed 2, and it appears that the consolidation at the pipe-soil interface was

finished before 50 days. Comparing the two test beds, Pipe-soil interfaces in Test

Bed 1 took longer to consolidate with a larger water content and lighter model

pipes.

5. For axial load model tests performed on GOM Clay at 120% to 140% water

content with a 4.25-in diameter polypropylene pipe, drained shear resistance at

the pipe-soil interface can seemingly be mobilized at “slow” speed of 1×10-5 in/s.

Also, the drained residual state can be achieved by performing a number of

“undrained” loading cycles with fast speeds, which leads to hardening of the

interface.

6. Once the soil-pipe interface reaches the drained residual state, the measured axial

resistance can remain constant with increasing the loading speed by 100 times

from “slow” speed to “fast” speed. However, as shown in Figure 5.30, when the

speed is “very fast”, a larger axial resistance was observed probably due to

negative pore pressure generation.

7. A breakout peak would be observed if a fast axial load sweep was performed after

a long pause due to thixotropy. This increase in axial resistance would go away

with continued shearing. However, slow sweep does not seem to generate a peak

after a pause.

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Figure 5.28 Large-displacement drained residual friction coefficient obtained from axial

load tests and tilt table tests (After Hussien, 2020)

Figure 5.29 Friction coefficient vs time after pipe placement – Test Bed 1

95

Figure 5.30 Friction coefficient vs time after pipe placement – Test Bed 2

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Chapter 6 Conclusions

T-bar tests, embedment tests and axial load tests were performed to characterize

the offshore flowline behavior under axial loading conditions. With T-bar tests, the

undrained shear strength of the test beds was measured. Thixotropy effects of GOM Clay

were captured, and parameters for embedment estimation equation were obtained.

Embedment tests monitored the settlement of the model pipes after placement, and the

measured settlements matched very well with the embedment prediction equation from

the SAFEBUCK JIP Merged Guideline. Eventually, four series of axial load tests have

been performed on two test beds with the fourth series currently suspended. The drained

residual friction resistance is obtained, and the process of reaching the drained residual

state for the two test beds are analyzed. The following conclusions are summarized from

these tests:

1. The embedment estimation method introduced by SAFEBUCK JIP Merged

Guideline provided decent match with the actual measured embedment of the

statically placed model pipes.

2. Same as the tilt table results on GOM Clay by Melo Monteiro (2019), the drained

residual axial resistance of the soil-pipe interface decreases with increasing

effective normal stress.

3. Consolidation of pipe-soil interface of test beds with a thickness of 1 ft and a

water content between 120 % and 140 % may take 25 to 100 days to complete.

During the consolidation process of Pipe 1 and Pipe 2, the drained residual

friction resistance of the soil-pipe interface seemingly keeps increasing, which

was supported by the interpretation of the embedment test results.

97

4. For axial load model tests performed on GOM Clay at 120% to 140% water

content with a 4.25-in diameter polypropylene pipe, drained shear resistance at

the pipe-soil interface can seemingly be mobilized at “slow” speed of 1×10-5 in/s.

Also, the drained residual state can be achieved by performing numerous sweeps

with fast loading speeds.

5. When the drained residual state is reached, the axial resistance can remain

constant with increasing the loading speed by 100 times from “slow” speed to

“fast” speed. However, a larger axial resistance might be obtained at “very fast”

speed due to negative pore pressure generation.

6. A stress concentration at the pipe invert possibly happened to Pipe 3. Under this

assumption, the effective normal stress between the pipe and clay trench was

increased, and the mobilized drained shear strength was reduced. If a model pipe

with a larger diameter is used, this effect might be mitigated.

7. A breakout peak would be observed if there is a long pause before a fast axial

load sweep, which is due to thixotropy. This increase in axial resistance would

reduce with continued axial shearing. However, slow sweep does not seem to

generate a peak after a pause.

98

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