FACULTY OF SCIENCE AND TECHNOLOGY MASTER'S THESIS

FACULTY OF SCIENCE AND TECHNOLOGY

MASTER’S THESIS

Study program/specialization:

Offshore Technology - Marine and Subsea

Technology

Spring semester, 2014

Open

Author:

Dawit Asefaw Ghebreghiorghis

………………………………………… (Authour’s signature)

Faculty Supervisor: Prof. Kenneth A. Macdonald – UiS

External Supervisor: -

Title of Master Thesis:

On-Bottom Stability Analysis of Subsea Pipelines According to DNV-RP-F109

Credits (ECTS): 30

Key words:

Subsea pipelines

Design codes and standards

DNV-RP-F109

Waves and hydrodynamic loads

Parametric analysis

SESAM: StableLines V1.4-01

Pages: 72

+ Enclosure/others: 58

Stavanger, 15.06.2014

ABSTRACT

Subsea pipelines are exposed to wave and current generated forces and other internal andexternal loads. The stability of the pipeline can be checked by performing on-bottomstability analysis. The primary requirement for the pipe is that no lateral movement ofthe pipe is accepted, or that certain limited movements that do not cause interferencewith adjacent structures or objects are allowed depending on the method of analysis used.

The objective of this Thesis is to make a sensitivity analysis of the on-bottom analysis,where the effect of the different parameters on the stability of pipelines resting on seabedsis checked. Data from a case study project from Subsea 7 is utilized in the analysis.

The lateral and vertical stability analyses of the pipeline for different seabed types ischecked. In the lateral stability analysis, all the forces which will force the pipe to movesideways are taken into consideration. The main lateral forces are the drag and inertiaforces and the resisting forces. These forces are checked against each other. If the drivingforce is larger than the resisting force, the pipe will move. The requirement is to makesure that the pipe does not move at all or more than it is allowed or expected to move.

In the vertical stability analysis, the main objective is to make sure that the pipe sinksinto the water. Depending on the type of soil of the seabed, the pipe may be required tosit on the seabed, be buried or anchored. Hence, the objective is to make sure that thepipe is stable under these conditions.

The historical development of some of the most common offshore pipeline codes are alsodiscussed in the literature. Besides, other published approaches from different conferencepapers and journals are referred to perform the analysis.

In the analysis process, after performing the desired calculations using the different meth-ods, a sensitivity analysis is carried out to identify the critical parameters which can affectthe stability.

The Absolute and the Generalized Lateral Stability methods are used in the case study.As the results show, concrete and steel wall thicknesses are the most critical parameters.In the Absolute Lateral Stability method, to make the pipeline absolutely stable a verythick concrete coating and/or steel thickness is required. This makes it uneconomical anddifficult to install. The Generalized method which allows a lateral displacement of up to10 pipe diameters is therefore, identified to be a more economical analysis method.

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ACKNOWLEDGEMENTS

I would like to first thank my advisor, Prof. Kenneth A. Macdonald, for his guidance andconstructive comments, and for the efforts he made in negotiating with DNV-GL to getme the SESAM: StableLines software on time.

My gratitude also goes to the Faculty of IKM at UiS for facilitating the authentication ofthe software.

Finally, my warmest gratitude goes to my caring and loving sister, Semhar and her familyfor their continuous support and encouragement.

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DEDICATION

I dedicated this work to my beloved parents, Hiwet Tuquabo and Asefaw Ghebreghiorghis.

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Intentionally left blank.

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Table of Contents

Abstract . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . i

Acknowledgements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ii

Dedication . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . iii

Table of Contents . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . vii

List of Figures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ix

List of Tables . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . xi

List of Abbreviations & Symbols . . . . . . . . . . . . . . . . . . . . . . . . . . xii

Preface . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . xvi

1 INTRODUCTION 1

1.1 General . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1

1.2 Background . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2

1.3 Scope . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3

2 FUNDAMENTALS OF SUBMARINE PIPELINE DESIGN 4

2.1 General . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4

2.2 Design Loads . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4

2.2.1 Functional Loads . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5

2.2.2 Environmental Loads . . . . . . . . . . . . . . . . . . . . . . . . . . 6

2.2.3 Accidental Loads . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6

2.2.4 Construction/ Installation Loads . . . . . . . . . . . . . . . . . . . 6

2.2.5 Combination of Loads . . . . . . . . . . . . . . . . . . . . . . . . . 6

2.3 Design Limit States . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6

2.3.1 Serviceability Limit State (SLS) . . . . . . . . . . . . . . . . . . . . 7

2.3.2 Accidental Limit State (ALS) . . . . . . . . . . . . . . . . . . . . . 7

2.3.3 Ultimate Limit State (ULS) . . . . . . . . . . . . . . . . . . . . . . 7

2.4 Wall Thickness . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8

2.5 Pipeline Installation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9

2.5.1 S - Lay . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9

2.5.2 J - Lay . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9

2.5.3 Reel - Lay . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10

2.5.4 Tow-in Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10

2.6 Hydrodynamic On-Bottom Stability . . . . . . . . . . . . . . . . . . . . . . 11

2.7 Free Spans . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13

2.8 Expansion and Buckling . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13

2.9 Fatigue . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14

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3 DESIGN CODES AND STANDARDS 15

3.1 Background . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15

3.2 Det Norske Veritas (DNV) . . . . . . . . . . . . . . . . . . . . . . . . . . . 16

3.2.1 General . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16

3.2.2 DNV-OS-F101 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16

3.2.3 DNV-RP-E305 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16

3.2.4 DNV-RP-F109 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18

3.3 British Standards - BS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19

3.3.1 Background . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19

3.3.2 Stability Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . 20

3.4 International Standardization Organization (ISO) . . . . . . . . . . . . . . 21

3.4.1 General . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21

3.4.2 ISO-13623: Petroleum and Natural Gas Industries - Pipeline Trans-portation Systems . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21

3.5 American Petroleum Institute (API) . . . . . . . . . . . . . . . . . . . . . 22

3.6 American Gas Association (AGA) . . . . . . . . . . . . . . . . . . . . . . . 23

3.7 Comparisons . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24

3.7.1 Similarities . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24

3.7.2 Differences . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24

4 WAVE THEORIES AND HYDRODYNAMIC LOADS 26

4.1 General . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 26

4.2 Design Waves and Currents . . . . . . . . . . . . . . . . . . . . . . . . . . 26

4.2.1 Waves . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 26

4.2.2 Currents . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 33

4.3 Hydrodynamic Loads and Seabed Resistance . . . . . . . . . . . . . . . . . 37

4.3.1 General . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 37

4.3.2 Loads and Load Coefficients . . . . . . . . . . . . . . . . . . . . . . 37

4.3.3 Pipe-Soil Interaction . . . . . . . . . . . . . . . . . . . . . . . . . . 39

4.3.4 Seabed Resistance . . . . . . . . . . . . . . . . . . . . . . . . . . . 40

5 ON-BOTTOM STABILITY DESIGN METHODOLOGIES (ACCORD-ING TO DNV-RP-F109) 44

5.1 Load Combinations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 44

5.1.1 Temporary (Empty) Conditions . . . . . . . . . . . . . . . . . . . . 44

5.1.2 Permanent (Operational) Conditions . . . . . . . . . . . . . . . . . 44

5.2 Analysis Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 45

A. Lateral Stability Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . 45

5.2.1 Absolute Lateral Static Stability Method . . . . . . . . . . . . . . . 45

5.2.2 Generalized Lateral Stability Method . . . . . . . . . . . . . . . . . 47

5.2.3 Dynamic Lateral Stability Method . . . . . . . . . . . . . . . . . . 51

B. Vertical Stability Analysis . . . . . . . . . . . . . . . . . . . . . . . . . 51

5.3 Stability Enhancement Methods . . . . . . . . . . . . . . . . . . . . . . . . 52

5.4 Analysis Software: SESAM (StableLines) . . . . . . . . . . . . . . . . . . . 53

5.4.1 General . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53

5.4.2 Features . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53

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6 ANALYSIS, RESULTS AND DISCUSSIONS 556.1 Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55

6.1.1 Design Data . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 566.2 Results and Discussions . . . . . . . . . . . . . . . . . . . . . . . . . . . . 56

6.2.1 Absolute Lateral Stability . . . . . . . . . . . . . . . . . . . . . . . 566.2.2 Generalized Lateral Stability . . . . . . . . . . . . . . . . . . . . . . 596.2.3 Vertical Stability . . . . . . . . . . . . . . . . . . . . . . . . . . . . 63

6.3 Sensitivity Analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 636.3.1 Case I: Concrete Thickness, tconc . . . . . . . . . . . . . . . . . . . 636.3.2 Case II: Pipe Diameter . . . . . . . . . . . . . . . . . . . . . . . . . 686.3.3 Case III: Water Depth . . . . . . . . . . . . . . . . . . . . . . . . . 68

7 CONCLUSIONS & RECOMMENDATIONS 697.1 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 697.2 Recommendations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 69

REFERENCES 72

APPENDICES 73Appendix A: Manual Calculation . . . . . . . . . . . . . . . . . . . . . . . . . . 73Appendix B: SESAM: StableLines Output . . . . . . . . . . . . . . . . . . . . . 79

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List of Figures

1.1 Illustration of the forces acting on a submarine pipeline [28] . . . . . . . . 2

2.1 Uses of offshore pipelines. [13] . . . . . . . . . . . . . . . . . . . . . . . . . 5

2.2 S-lay method [21] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9

2.3 J-lay barge method [13] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10

2.4 Surface-tow method of pipeline installation. [13] . . . . . . . . . . . . . . . 11

2.5 Mid-depth tow method of pipeline installation. [13] . . . . . . . . . . . . . 11

2.6 Off-bottom method of pipeline installation. [13] . . . . . . . . . . . . . . . 12

2.7 Bottow-tow method of pipeline installation. [13] . . . . . . . . . . . . . . . 12

2.8 Illustration of pipeline design procedures [1] . . . . . . . . . . . . . . . . . 14

3.1 Three dimensional stability analysis . . . . . . . . . . . . . . . . . . . . . . 21

4.1 Regions of application of wave theories [9] . . . . . . . . . . . . . . . . . . 27

4.2 Water particles motion. [23] . . . . . . . . . . . . . . . . . . . . . . . . . . 28

4.3 Effect of significant wave height on wave energy. [12] . . . . . . . . . . . . 30

4.4 Effect of peak period on wave energy. [12] . . . . . . . . . . . . . . . . . . 31

4.5 JONSWAP Spectrum (Peakedness) [10] . . . . . . . . . . . . . . . . . . . . 31

4.6 Significant flow velocity amplitude Us at seabed. [12] . . . . . . . . . . . . 32

4.7 Mean zero up-crossing period of oscillating flow Tu at seabed. [12] . . . . . 33

4.8 Large and short period waves. [23] . . . . . . . . . . . . . . . . . . . . . . 34

4.9 Reduction factor due to wave spreading and directionality. [12] . . . . . . . 35

4.10 Keulegan - Carpenter number, K, relative to pipe diameter, D. [12] . . . . 36

4.11 The influence of current velocity on a pipeline on seabed [23] . . . . . . . . 36

4.12 Illustration of penetration [12] . . . . . . . . . . . . . . . . . . . . . . . . . 40

4.13 Peak load reduction due to penetration [12] . . . . . . . . . . . . . . . . . 41

4.14 Trench parameters [12] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 41

4.15 Peak load reduction due to trenching [12] . . . . . . . . . . . . . . . . . . . 42

4.16 Soil embedment [23] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 42

5.1 Peak load coefficients [12] . . . . . . . . . . . . . . . . . . . . . . . . . . . 47

5.2 Minimum weight, Lstable/(2 +M)2, for pipe on sand. [12] . . . . . . . . . . 50

5.3 Minimum weight, L10/(2 +M)2, for pipe on sand. [12] . . . . . . . . . . . 50

6.1 SESAM:StableLines Output: Necessary weight vs. direction presentation . 57



6.4 SESAM: StableLines Output: Necessary weight vs concrete thickness, pipediameter and water depth from top to bottom . . . . . . . . . . . . . . . . 65

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6.5 SESAM:StableLines Output: Concrete thickness vs. water depth presen-tation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 67

1 SESAM:StableLines Output: Necessary weight vs. direction presentation . 832 SESAM:StableLines Output: Concrete thickness vs. density presentation . 833 SESAM:StableLines Output: Necessary weight vs concrete thickness pre-

sentation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 844 SESAM:StableLines Output: Necessary weight vs pipe diameter presentation 855 SESAM:StableLines Output: Necessary weight vs water depth presentation 866 SESAM:StableLines Output: Necessary weight vs current velocity presen-

tation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 887 SESAM:StableLines Output: L vs. K for three values of M presentation . . 888 SESAM:StableLines Output: Necessary weight vs. direction presentation . 959 SESAM:StableLines Output: Concrete thickness vs. density presentation . 9610 SESAM:StableLines Output: Necessary weight vs concrete thickness pre-

sentation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9711 SESAM:StableLines Output: Necessary weight vs pipe diameter presentation 9712 SESAM:StableLines Output: Necessary weight vs water depth presentation 9813 SESAM:StableLines Output: Necessary weight vs current velocity presen-

tation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9814 SESAM:StableLines Output: L vs. K for three values of M presentation . . 9915 SESAM:StableLines Output: Necessary weight vs. direction presentation . 10416 SESAM:StableLines Output: Concrete thickness vs. density presentation . 10517 SESAM:StableLines Output: Necessary weight vs concrete thickness pre-

sentation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10618 SESAM:StableLines Output: Necessary weight vs pipe diameter presentation10719 SESAM:StableLines Output: Necessary weight vs water depth presentation 10720 SESAM:StableLines Output: Necessary weight vs current velocity presen-

tation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10821 SESAM:StableLines Output: L vs. K for three values of M presentation . . 10922 SESAM:StableLines Output: Concrete thickness vs. water depth presen-

tation, Absolute Stability . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11223 SESAM:StableLines Output: Concrete thickness vs. water depth presen-

tation, Generalized Stability, 0.5D . . . . . . . . . . . . . . . . . . . . . . . 11324 SESAM:StableLines Output: Concrete thickness vs. water depth presen-

tation, Generalized Stability, 10D . . . . . . . . . . . . . . . . . . . . . . . 114

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List of Tables

2.1 Typical link between scenarios and limit states. [9] . . . . . . . . . . . . . 8

4.1 Seabed roughness of different types of soils. [9] . . . . . . . . . . . . . . . . 36

5.1 Peak horizontal load coefficients. [12] . . . . . . . . . . . . . . . . . . . . . 465.2 Peak vertical load coefficients. [12] . . . . . . . . . . . . . . . . . . . . . . 465.3 Safety factors, winter storms in North Sea. [12] . . . . . . . . . . . . . . . 465.4 Minimum weight, Lstable/(2 +M)2, for pipe on sand, K > 10. [12] . . . . . 485.5 Minimum weight, Lstable/(2 +M)2, for pipe on sand, K 6 5. [12] . . . . . . 495.6 Minimum weight, L10/(2 +M)2, for pipe on sand. [12] . . . . . . . . . . . 49

6.1 StableLines: Absolute Lateral Stability main page . . . . . . . . . . . . . . 586.2 StableLines: Generalized Stability, Displacement=0.5D main page . . . . . 606.3 StableLines: Generalized Stability, Displacement=10D main page . . . . . 62

1 Pipe and soil data. [25] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 792 Material densities. [25] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 793 Environmental data. [25] . . . . . . . . . . . . . . . . . . . . . . . . . . . . 794 StableLines: Absolute Lateral Stability main page . . . . . . . . . . . . . . 815 SESAM:StableLines Output: Necessary weight vs. Direction . . . . . . . . 826 SESAM:StableLines Output: Concrete thickness vs. density . . . . . . . . 827 SESAM:StableLines Output: Necessary weight vs. concrete thickness . . . 848 SESAM:StableLines Output: Necessary weight vs. pipe diameter . . . . . . 859 SESAM:StableLines Output: Necessary weight vs. water depth . . . . . . . 8610 SESAM:StableLines Output: Necessary weight vs. current velocity . . . . . 8711 SESAM:StableLines Output: L vs. K for three values of M . . . . . . . . . 8912 SESAM:StableLines Output: Iteration (Empty condition) . . . . . . . . . . 9013 SESAM:StableLines Output: Iteration (Operational condition) . . . . . . . 9114 SESAM:StableLines Output: Results: All RPV combinaions . . . . . . . . 9115 SESAM:StableLines Output: General output . . . . . . . . . . . . . . . . . 9216 StableLines: Generalized Stability, Displacement=0.5D main page . . . . . 9417 SESAM:StableLines Output: Necessary weight vs. Direction . . . . . . . . 9518 SESAM:StableLines Output: Concrete thickness vs. density . . . . . . . . 9619 SESAM:StableLines Output: Necessary weight vs. concrete thickness . . . 9620 SESAM:StableLines Output: Necessary weight vs. pipe diameter . . . . . . 9721 SESAM:StableLines Output: Necessary weight vs. water depth . . . . . . . 9822 SESAM:StableLines Output: Necessary weight vs. current velocity . . . . . 9923 SESAM:StableLines Output: L vs. K for three values of M . . . . . . . . . 10024 SESAM:StableLines Output: Results: All RPV combinaions . . . . . . . . 10025 SESAM:StableLines Output: General output . . . . . . . . . . . . . . . . . 101

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26 StableLines: Generalized Stability, Displacement=10D main page . . . . . 10327 SESAM:StableLines Output: Necessary weight vs. Direction . . . . . . . . 10428 SESAM:StableLines Output: Concrete thickness vs. density . . . . . . . . 10529 SESAM:StableLines Output: Necessary weight vs. concrete thickness . . . 10530 SESAM:StableLines Output: Necessary weight vs. pipe diameter . . . . . . 10631 SESAM:StableLines Output: Necessary weight vs. water depth . . . . . . . 10632 SESAM:StableLines Output: Necessary weight vs. current velocity . . . . . 10833 SESAM:StableLines Output: L vs. K for three values of M . . . . . . . . . 10934 SESAM:StableLines Output: Results: All RPV combinaions . . . . . . . . 11035 SESAM:StableLines Output: General output . . . . . . . . . . . . . . . . . 11136 SESAM:StableLines Output: Concrete thickness vs. water depth, Absolute

Stability . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11237 SESAM: StableLines Output: Concrete thickness vs. water depth, Gener-

alized Stability, 0.5D . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11338 SESAM: StableLines Output: Concrete thickness vs. water depth, Gener-

alized Stability, 10D . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 114

xi

List of Abbreviations & Symbols

As Acceleration of water particles

C1, C2, C3 Coefficients for calculating minimum required pipe weight

CD Drag coefficient

CL Lift coefficient

CM Inertia coefficient

C∗Y Peak horizontal load coefficient

C∗Z Peak vertical load coefficient

D Outer diameter of pipeline including all the coatings

FC Vertical contact force between pipe and soil

FD Drag force

FI Inertia force

FL Lift force

FR Passive soil resistance

F ∗Y Peak horizontal load

FZ Vertical hydrodynamic (lift) load

F ∗Z Peak vertical load

G(ω) Wave transfer function

GC Soil (clay) strength parameter

H Wave height

Hs Significant wave height

K Keulegan-Carpenter number

Kb Equivalent sand roughness parameter

L Significant weight parameter

M Steady to oscillatory velocity ratio for design spectrum

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Mn Spectral moment of order n

N Spectral acceleration factor

R Resistance

RD Reduction factor

SF Safety factor

Sg Specific density of pipe

Su Un-drained shear strength of clay

SUU(ω) Wave induced velocity spectrum at seabed

Sηη Wave spectral density

T Wave period

T ∗ The period associated with the single design oscillation velocity

Tn Wave induced velocity spectrum at seabed

Tu Mean zero up-crossing period

U∗ Oscillatory velocity amplitude for single design oscillation

Uc Current velocity at the top of the pipeline

Ur Current velocity at the reference height

Us Significant flow velocity amplitude at pipe level

V ∗ Steady current velocity associated with design oscillation

Ws Submerged weight

Y Dimensionless lateral pipe displacement

α Generalized Phillip’s constant

γ Peakedness

γ′ Submerged unit weight of soil

γw Safety factor

γsc Safety factor

κc Initial penetration in clay

κs Initial penetration in sand

µ Coefficient of friction of seabed

ω Wave frequency = (2π)/T

xiii

ωp Peak wave frequency = (2π)/Tp

φ Phase angle of wave

ρw Density of seawater

σ Spectral width parameter

τ Number of waves in the storm

θ Angle made by trenching with the horizontal

θc Angle between current velocity and pipeline axis

b Buoyancy of pipeline per unit length

d Water depth

g Acceleration due to gravity

k Wave number

kt The ratio of the periods of the single design oscillation and design spectrum

rpen Load reduction factor due to penetration into seabed

rperm Load reduction factor due to permeability of seabed

rtot Total load reduction factor

rtrench Load reduction factor due to trenching

z0 Seabed roughness

zp Depth of penetration

zr Reference measurement height

zt Depth of trench

AGA American Gas Association

ALS Accidental Limit State

API American Petroleum Institute

BS-PD British Standard - Published Document

BSI British Standards Institution

CWC Concrete Weight Coating

DNV Det Norske Veritas

DNV-GL Det Norske Veritas - Germanischer Lloyd

DNV-OS Det Norske Veritas - Offshore Standard

xiv

DNV-RP Det Norske Veritas - Recommended Practice

ISO International Standardization Organization

JONSWAP Joint North Sea Wave Project

LRFD Load Resistance Factor Design

SLS Serviceability Limit State

SMYS Specified Minimum Yield Stress

TDP Touch Down Point

ULS Ultimate Limit State

xv

PREFACE

On-bottom stability design and analysis of subsea pipelines had been done using the oldDNV-RP-E305 (1988) code until it has been replaced by the new DNV-RP-F109 (2007,updated 2010) code recently. The new design and analysis methods have significantlymaintained, revised and updated the old methods. This paper is therefore preparedbased on the new DNV Recommended Practice (DNV-RP-F109).

The organization of the Thesis can be summarized as follows.

Chapter 1 discusses the historical development and a general overview on why on-bottomstability analysis is required for subsea pipelines and which methods to use for the analysis.

Chapter 2 covers the fundamental principles of pipeline design procedures. Here, anoverview of offshore pipelines’ design loads, limit states, the installation methods, thehydrodynamic loads, free span, etc. will be discussed.

Chapter 3 introduces the different types of design codes and standards used in the de-sign of offshore pipelines in general and the codes and standards used in the on-bottomsstability analysis of subsea pipelines in particular.

Chapter 4 discusses about design waves and currents, on how to predict extreme wavesand currents based on the data available and/or the hind-casted data. It also discusseswhich types of wave theories and wave spectrum to use which can represent the sea-state.It also discusses about the hydrodynamic loads which are created by the waves and cur-rents.

Chapter 5 discusses about the design methodologies used for the on-bottom stabil-ity analysis according to the DNV-RP-F109 code. It also discusses about the types ofloads and load combinations considered and the design criteria for the temporary andoperational conditions. A short introduction to the analysis software package, SESAM(StableLines) is also made. In case of instability, the stability enhancing methods arediscussed at the end of the chapter.

Chapter 6 focuses on the analysis of the on-bottom stability design procedures using areal data found from Subsea 7, and the results found will be analyzed and discussed. Asensitivity analysis will also be made by changing the governing parameters, hence, themost sensitive parameters will be identified.

Chapter 7 covers the conclusions drawn and the recommendations for further works.

Appendices will include a manual sample calculation, numerical results, output graphsand tables from the SESAM: StableLines software.

xvi

Chapter 1

INTRODUCTION

1.1 General

A standard engineering task when designing subsea pipelines is to ensure that the pipelineis stable on the seabed under the action of hydrodynamic loads induced by waves andsteady currents. If it is too light, it will slide sideways under the action of hydrodynamicforces. If it is too heavy, it will be difficult and expensive to construct. This movementof the pipeline will cause bending stresses on the pipe, which may then cause the pipeto fatigue and fail. Simultaneously, it may cause damage to the coatings, for example,cracking of concrete. Conventionally a subsea pipeline has been considered stable if it hasgot sufficient submerged weight so the lateral soil resistance is sufficiently high to restrainthe pipeline from deflecting sideways.

On-bottom stability calculations are performed to establish requirements for pipeline sub-merged mass. The required pipeline submerged mass will have a direct impact on therequired pipe-lay tensions, installation stresses and the pipe configuration on the sea bot-tom. Subsea pipeline stability is governed by the fundamental balance of forces betweenloads and resistances. From the installation viewpoint, especially where spans are not aconcern, the priority is to minimize the required pipeline submerged weight. [22]

On-bottom stability calculations shall be performed for the operational phase and for theinstallation phase. The pipeline is filled with content at the expected lowest density whenconsidering the operational phase. [3]

The primary stabilization method used traditionally by designers has been to apply suffi-cient amount of external concrete weight coating over the anticorrosion coating to achievethe on-bottom stability as it will not be cost efficient to increase the steel wall thickness inorder to increase the submerged weight. The concrete weight coating can then also protectthe anticorrosion coating against mechanical damage. But there are limits in using con-crete weight coating depending on the type of installation (pipe-laying) method to be used.

But there is a practical limit to how much concrete weight coating can be applied topipeline, for example, due to the pipe-lay vessel tension capacity limitation or limitationto the practical thickness that can be applied to a pipeline with a given diameter or tothe pipe joint weight that can be practically handled in the coating plant. A secondarystabilization method may have to be adopted to overcome these limitations, such as

1

1.2. BACKGROUND

Figure 1.1: Illustration of the forces acting on a submarine pipeline [28]

through lowering the pipeline into the seabed by pre-lay dredging or post-lay trenchingor by using on-seabed restraints. The latter can involve covering the pipeline by crushedrock. However, where trenching and backfilling or rock dumping is not technically feasibleor cost efficient, the designer may have to resort to more costly and technically challengingsolutions such as anchoring the pipeline to the seabed.

1.2 Background

On-bottom stability of submarine pipelines is governed by the fundamental balance offorces between loads and resistances. This approach to stability design of pipelines wasincorporated by various pipeline codes and standards, such as, DNV ’76 (Rules for Sub-marine Pipeline Systems), the American Gas Association, AGA - Level 1, etc., which usethis model as a basis for design.

Predicting on-bottom pipeline stability is complex with many disciplines requiring in-tegration, including soil constitutive modeling, seabed liquefaction, scour and sedimenttransport, structural mechanics and prediction of ocean waves and hydrodynamic loads.Considering this complexity, most pipelines are designed using very simplistic methods.

The Coulomb friction model describes pipe-soil behavior and adopt force balance meth-ods to ensure the pipe does not displace horizontally, i.e., the stability failure criterion isdefined such that the hydrodynamic loading does not exceed the soil resistance. Until theearly 1970s, the Coulomb theory of soil friction was the one mainly employed to estimatesoil resistance to lateral displacement of subsea pipes subject to hydrodynamic forces. Inthis approach, the effects of passive resistance due to the soil heap created in the processof pipe movement were ignored. [20]

More recent updates such as DNV-RP-F109 (2010), retain simplistic stability charts, buthighlight the advantages of dynamic time domain and three-dimensional stability analy-sis. A lateral displacement limit-state is set as a “failure” criterion in this latter analysis

2 On-Bottom Stability Analysis of Subsea Pipelines According to DNV-RP-F109

1.3. SCOPE

methodology preferred by DNV. [26]

The Morison’s Equation was the most widely used equation for pipeline on-bottom sta-bility analysis. But some limitations were identified on this equation recently. DNV haveidentified this problem and have presented a more experimental approach to modify theparameters in the latest and state-of-the-art code, DNV-RP-F109, on which this Thesisis based on.

1.3 Scope

Here are the main topics to be discussed in this Thesis;

• A literature review about the design of offshore pipelines, in general and the designprocedures of the on-bottom stability analysis of subsea pipelines, in particular willbe discussed,

• Identification and explanation of the different standards and codes of practice willbe made,

• The difference between these codes will be addressed and the state-of-the-art ofon-bottom stability analysis will be identified as well,

• A case study will be performed on a specific project, based on the data from Subsea7 for the analysis,

• A case study will be performed for different sensitive parameters which affect theon-bottoms stability of subsea pipelines, and finally,

• The most critical parameters will be identified and addressed.

The on-bottom stability design analysis will be carried out using the SESAM: StableLinessoftware package. This software is developed by DNV according to the DNV-RP-F109code.

On-Bottom Stability Analysis of Subsea Pipelines According to DNV-RP-F109 3

Chapter 2

FUNDAMENTALS OFSUBMARINE PIPELINE DESIGN

2.1 General

The establishment of marine pipelines is a more recent development of the twentieth cen-tury. The fuel line installed across the English Channel in 1944 to supply the allied troopsWorld War II is cited as the first example. The first oil-producing well in the Mexican Gulfwas drilled in 1947, the first pipe-lay barge commissioned in 1952, and the first pipelinewas laid on the seabed in 1954 [5]. The petroleum industry has proven that pipelinesare by far the most economical means of large scale transportation of crude oil, naturalgas and their products. Transporting petroleum fluids with pipelines is a continuous andreliable operation. Pipelines have demonstrated an ability to adapt to a wide variety ofenvironments including remote areas and hostile environments.

Offshore pipelines, as illustrated in Figure 2.1, can be classified as follows:

• flow-lines transporting oil and/or gas from satellite subsea wells to subsea manifolds;

• flow-lines transporting oil and/or gas from subsea manifolds to production facilityplatforms;

• infield flow-lines transporting oil and/or gas between production facility platforms;

• export pipelines transporting oil and/gas from production facility platforms to shore;

• flow-lines transporting water or chemicals from production facility platforms, throughsubsea injection manifolds, or injection wellheads.

The pipelines are sized to handle the expected pressure and fluid flow. To ensure thedesired flow rate of product, pipeline size varies significantly form project to project. [13]

2.2 Design Loads

Loads on a marine pipeline have been classified in a lot different ways by different nationaland international standards. But they can generally be classified as:

• Functional loads;

4

2.2. DESIGN LOADS

Figure 2.1: Uses of offshore pipelines. [13]

• Environmental loads;

• Accidental loads;

• Construction/ Installation loads;

• Combination of loads.

2.2.1 Functional Loads

Functional loads are defined as actions that result from the operation of the pipeline. Theweight of the pipeline, including components and fluid, and loads due to pressure andtemperature are examples of functional loads arising from the intended use of the system.Pre-stressing, residual stresses from installation, soil cover, external hydrostatic pressure,marine growth, subsidence and differential settlement, frost heave and thaw settlement,and sustained loads from icing are examples of functional loads from other sources. Re-action forces at supports from functional loads and loads due to sustained displacements,rotations of supports or impact by changes in flow direction are also functional. Functionalloads include the effects of:

• internal static pressure;

• pressure surge; and

• operational temperature.


2.3. DESIGN LIMIT STATES

2.2.2 Environmental Loads

Environmental loads are defined as actions resulting from the interaction of the pipelineswith its environment. Environmental loads are primarily generated by wave and currentaction. Gravity forces (self-weight, buoyancy and hydrostatic pressure) and soil pressurecan also be characterized as environmental loads. In general, loads arising from theenvironment shall be classified as environmental, except where it is necessary that theybe considered as functional or when, due to a low probability of occurrence, as accidental.

2.2.3 Accidental Loads

Loads imposed on the pipeline under unplanned but plausible circumstances shall beconsidered as accidental. Both the probability of occurrence and the likely consequenceof an accidental load should be considered when determining whether the pipeline shouldbe designed for an accidental load. For subsea pipelines, such loads may be grouped intothe following:

• Natural hazards : earthquakes, mudslides, iceberg scouring.

• Third party hazards : dropped objects, fishing activities, shipping and military ac-tivities.

2.2.4 Construction/ Installation Loads

Installation of marine pipelines is to a great extent weather dependent, and part of theinstallation engineering is the determination of the acceptable limits (wind, wave height,current, etc.) for the installation to take place. Loads necessary for installation andcommissioning shall be classified as construction loads. Apart from the pipeline self-weight and the normal environmental loads, the specific actions during installation willmostly impose static and dynamic deformations.[5]

2.2.5 Combination of Loads

When calculating equivalent stresses or strains, the most unfavorable combination offunctional, environmental, construction and accidental loads that can be predicted tooccur simultaneously shall be considered.

2.3 Design Limit States

The DNV standard, DNV-OS-F101 is based on a Load Resistance Factor Design (LRFD)format, i.e. load factors are applied to the loads and a resistance factor is applied to thematerial strength. The safety requirement, as defined in the pipeline offshore standardDNV-OS-F101, is based on the Limit State Design approach. The on-bottom stabilityanalysis recommended practice DNV-RP-F109 is also based on the same safety require-ment as F101. In this regard, F109 states that for other than Absolute Lateral StabilityMethod, “the recommended safety level is based on engineering judgment in order toobtain a safety level equivalent to modern industry practice”.

The following Limit States may be considered in relation to on-bottom stability analysis:


2.3. DESIGN LIMIT STATES

• Serviceability Limit State;

• Accidental Limit State;

• Ultimate Limit State.

2.3.1 Serviceability Limit State (SLS)

For a marine pipeline it shall be ensured that during its installation and operation it willnot be unsuitable for its intended purpose. The SLS refers to a given load condition that,if exceeded, can cause the pipeline to be unsuitable for continued operation. The SLSare defined for all the relevant loading conditions that can be formulated. The followingissues are normally considered:

• Deformation and movements due to waves and currents (hydrodynamic stability);

• Longitudinal deformations due to temperature and pressure variations (pipelineexpansion);

• Lateral deformations due to restrained temperature and pressure expansion (up-heaval buckling)

• Blockage of the pipeline, due to hydrate formation or wax deposition (flow assur-ance).

In the case of permanent local damage or permanent unacceptable deformations the SLSis not the appropriate formulation and the Ultimate Limit State (ULS) design factorsshall be introduced.

2.3.2 Accidental Limit State (ALS)

Considering higher environmental return periods than ULS, local buckling limit state maybe adopted to capture non-linear structural response effects. It shall be ensured that thepipeline has the required safety against accidental loads.

2.3.3 Ultimate Limit State (ULS)

As pipelines on the seabed are allowed to displace laterally to some extent, a local bucklinglimit state must be checked for which very large bending moments can be encountered.It shall be ensured that the pipeline has the required safety against failure. ULS can bedefined in terms of:

• Plastic deformations (yield);

• Local instability (buckling);

• Crack instability (bursting);

• Repeated loading (fatigue).


2.4. WALL THICKNESS

It shall be ensured that the pipeline has the required safety against accidental loads. Itshould also be noted that fatigue or other time dependent deterioration mechanisms re-duce the strength of the structure, and may initiate ULS. In this relation it is useful todistinguish between damage tolerant structures (SLS) and damage intolerant structures(ULS).

The different types of scenarios and their relationships to the limit states are illustratedin Table 2.1.

Table 2.1: Typical link between scenarios and limit states. [9]

2.4 Wall Thickness

The primary objective of the line-pipe design is to determine the optimal wall thicknessand steel grade of the pipeline. Determination of pipeline wall thickness is based on thedesign internal pressure or the external hydrostatic pressure and performing a simplehoop stress calculation. A usage factor applied to the Specified Minimum Yield Stress(SMYS) defines the allowable stress which, when inserted into the hoop stress formulae,determines the minimum required thickness of the pipe wall. The engineer would thenselect the nearest standard, for example API, wall thickness above the required minimum.[5]

Increasing the grade of line-pipe will correspondingly decrease the wall thickness andtherefore provide cost benefits. In addition to this, a thinner wall thickness will alsohave various impacts on construction activities. A thinner wall thickness will require lessfield welding and therefore, in theory, has the potential to reduce construction or lay time.

Increasing the material grade and strength of line-pipe is beneficial to laying pipe in deeperwaters. Furthermore, certain projects can only be implemented with pipe having reducedweight and optimized strength and toughness. [4]


2.5. PIPELINE INSTALLATION

2.5 Pipeline Installation

Once design is finalized, pipeline is ordered for pipe construction and coating and/orinsulation fabrication. Upon shipping to the site, pipeline can be installed. There areseveral methods for pipeline installation including S-lay, J-lay, reel barge, and tow-inmethods. [13]

2.5.1 S - Lay

This method is the most frequently used technique for marine pipeline installation. Itwas originally developed to lay pipe in shallow water near the shore Gulf of Mexico in1940 and 1950s. [21]

Figure 2.2: S-lay method [21]

As shown in Figure 2.2, the S-lay requires a laying barge to have on its deck severalwelding stations where the crew welds together 12- to 24-meter lengths of insulated pipein a dry environment away from wind and rain. As the barge moves forward, the pipe iseased off the stern, curving downward through the water as it leaves until it reaches thetouchdown point (TDP). After the touchdown, as more pipe is played out, it assumes thenormal S-shape.

A stringer is used to support the pipe as it leaves the barge which reduces the bendingstress in the pipe. To avoid buckling of the pipe, a tensioning roller and controlled forwardthrust must be used to provide appropriate tensile load to the pipeline. This method isused for pipeline installations in the range of water depths from shallow to deep.

2.5.2 J - Lay

The J-lay method, as shown in Figure 2.3, avoids some of the difficulties of S-laying suchas tensile load and forward thrust. J-lay barges drop the pipe down almost vertically untilit reaches touchdown. After that, the pipe assumes the normal J-shape, J-lay barges havea tall tower on the stern to weld and slip pre-welded pipe sections of lengths up to 75meters. With the simpler pipeline shape, J-lay can be used in deeper water than S-lay.


2.5. PIPELINE INSTALLATION

Figure 2.3: J-lay barge method [13]

2.5.3 Reel - Lay

Small-diameter pipelines can be installed with reel barges. The pipe is welded and coatedand fabricated in long strings, called ’stocks’ at an onshore facility. These stocks are thenlaid from the reel offshore. Vertical reels most commonly do J-lay, so can S-lay too.

2.5.4 Tow-in Method

There are four variations of the tow-in method: surface tow, mid-depth tow, off-bottomtow and bottom tow.

In the Surface-tow approach as shown in Figure 2.4, buoyancy modules are added tothe pipeline so that it floats at the surface. Once the pipeline is towed on site by the twotowboats, the buoyancy modules are moved or flooded, and the pipeline settles to the seafloor.

The Mid-depth tow which is illustrated in Figure 2.5 requires fewer buoyancy modules.The pipeline settles to the bottom on its own when the forward progression ceases.

The Off-bottom tow, as illustrated in Figure 2.6 involves both buoyancy modules andadded weight in the form of chains. Once in location, the buoyancy is removed, and thepipeline settles to the sea floor.

In the Bottom tow, The pipeline is allowed to sink to the bottom and then towed alongthe sea floor, as illustrated in Figure 2.7. It is primarily used for soft and flat sea floor inshallow water.


2.6. HYDRODYNAMIC ON-BOTTOM STABILITY

Figure 2.4: Surface-tow method of pipeline installation. [13]

Figure 2.5: Mid-depth tow method of pipeline installation. [13]

2.6 Hydrodynamic On-Bottom Stability

Pipelines shall be designed to prevent horizontal and vertical movement, or shall be de-signed with sufficient flexibility to allow predicted movements within the strength criteriaof a standard. [18]

This Thesis addresses stability analysis of marine pipelines on the seabed under the in-fluence of hydrodynamic loads and provides guidelines for the pipeline stabilization usingconcrete. Other stabilization methods such as trenching techniques, mattress covers, rockdumping, etc. are not included here. Stability is checked for the installation (1-year re-turn period) and for operation (100-year storm) phases.

Regardless of the computer program selected, the general hydrodynamic stability analysisinvolves the following procedures:

Step 1 : data gathering for the 1-year and 100-year environmental conditions, which in-


2.6. HYDRODYNAMIC ON-BOTTOM STABILITY

Figure 2.6: Off-bottom method of pipeline installation. [13]

Figure 2.7: Bottow-tow method of pipeline installation. [13]

cludes:

• Water depth

• Wave spectrum

• Current characteristics

• Soil properties

• Seabed condition

Step 2 : Determination of the hydrodynamic coefficients for the traditional analysis: drag(CD), lift (CL) and inertia (CM) coefficients. These coefficients may be adjustedfor Reynolds Number, Keulegan Number, steady current to wave ratio, and em-bedment in the latest design procedures.

Step 3 : Calculation of the hydrodynamic forces, i.e. drag (FD), lift (FL) and inertia (FI)forces.


2.7. FREE SPANS

Step 4 : Performing static force balance at time step increments and assessing stabilityand calculating the concrete coating thickness for the worst combination of thehydrodynamic forces. [13]

2.7 Free Spans

The loss of contact between the pipeline and seabed over a significant distance on a roughseabed can lead to pipeline spanning. An evaluation of an allowable free-span length isrequired in pipeline design. Should actual span lengths exceed the allowable length, cor-rection is then necessary to reduce the span to avoid pipeline damage. The flow of waveand current around a pipeline span can result in the generation of sheet vortices in thewake. These vortices are shed alternately from top to bottom of the pipeline resulting inan oscillatory force exerted on the span.

Free span can result in failure of pipelines due to excessive yielding and fatigue. It mayalso cause interference with human activities such as fishing. Free span can occur dueto unsupported weight of the pipeline section and dynamic loads from waves and currents.

The presence of bottom currents can cause significant dynamic stresses, if fluid structureinteraction (vortex shedding) in these free-span areas causes the pipeline to oscillate.These oscillations can result in fatigue of the pipeline welds, which can reduce pipelinelife.[5]

2.8 Expansion and Buckling

A pipeline will elongate due to the operating temperature and the operating pressure,which is normally higher than the installation temperature and pressure. However, dueto the restraint offered by the seabed friction, such pipeline expansions only occur at theends, i.e. at the tie-in points to fixed structures. At undisturbed sections of the pipeline,the restraint against thermal and pressure induced expansion may cause a compressivepipeline force, which could result in a global buckling mechanism. [5]

The three main reasons contributing to the end force and expansion leading to the lateralor upheaval buckling are:

• Temperature

• Pressure

• Poisson contraction – associated with pressure effects. [21]

At the undisturbed sections of the pipeline the restraint against thermal and pressureinduced expansion may cause a compressive pipeline force, which could result in a globalbuckling mechanism.

Pipeline buckling may occur due to the axial compressive loads caused by thermal andinternal pressure actions on pipelines in operation. If the pipeline is resting on the seabeda lateral deflection mode will prevail (snaking), whereas a vertical deflection mode will


2.9. FATIGUE

Figure 2.8: Illustration of pipeline design procedures [1]

take place when the pipeline is trenched and buried (upheaval buckling).

Free spanning pipeline may buckle in a downward mode. When buckling occurs, part ofthe constrained thermal expansion is released, thus reducing the compression force in thebuckled section. The resulting buckling configuration depends to a large extent upon thefrictional resistance between the pipe and the soil. An external force, for example, trawlboard impact or anchor hooking, may initiate the buckling event. [5]

2.9 Fatigue

Fatigue analysis shall be performed for pipeline sections that may be exposed to cyclicloading in order to:

• demonstrate that initiation of cracking does not occur;

• define requirements for inspection of fatigue.

The analysis shall include the prediction of load cycles during construction/installationand operation and a method for the translation of these loads into normal stress andstrain cycles.The selection of safety factors shall take into account the inaccuracy of fatigue-resistancepredictions and access for inspection for fatigue damage. It can be necessary to moni-tor the parameters causing fatigue and to control possible fatigue damage accordingly. [18]

The general pipeline design procedure is illustrated in Figure 2.8.


Chapter 3

DESIGN CODES ANDSTANDARDS

3.1 Background

Modern codes go back to the third quarter of the 19th century. There was a huge ex-pansion of infrastructure, particularly railway bridges and rapid industrial development,above all in boilers and ships. An intolerable number of failures occured. For exampleGies’s book on bridges claims that 40 bridges a year failed in the United State in the1870s. Codes and standards are products of mistakes and errors made by human beingsand different countries have different codes. In Europe, for example, there used to bemany pipeline codes. In the United Kingdom there was an institute of Petroleum PI6,which was very old-fashioned. About 30 years ago, there came a wave of new develop-ment, prompted by the offshore industry and led by the Norwegian organization, DNV,which was originally a ship classification society.

Other countries have followed. The Dutch made the NEN3650, which is more famousin the Netherlands. The Germans produced the Germanischer Lloyd code. The Britishrewrote BS8010. The Americans wrote API 1111. There is also International StandardsOrganization (ISO) code, ISO 13623, and Eruocode, etc. [22]

The client and authorities in a country where a pipeline is to be installed shall endorsethe codes and standards used by the designer. The widely used offshore pipeline designcodes include:

• DNV: DNV-OS-F101, DNV-RP-E305, DNV-RP-F109

• BSI: BS-PD-8010:2

• ISO: ISO-13623

• API: API-1111

• AGA: AGA (Level-1,2,3)

15

3.2. DET NORSKE VERITAS (DNV)

3.2 Det Norske Veritas (DNV)

3.2.1 General

DNV has been merged with Germanischer Lloyd (GL) as recently as September 2013 andform one of the world’s leading classification and certification companies, DNV-GL.

In its 150 years of experience, DNV has published a number of codes and recommendedpractices for the offshore industry. The first offshore design standards were intended as’rules’ to supplement national regulations and were aimed mainly at North Sea oil andgas developments.

The 1976 version was replaced by 1981 version, which was not universally effective particu-larly with respect to guidance on pipeline stability. Extensive revisions and developmentswere made and it was published in January 2000 as DNV-OS-F101, ”Submarine PipelineSystems”. It was further reprinted in 2003, amended in 2005. The newest publicationwas made in October 2007 as DNV-RP-F109. It was then updated in 2010.

DNV is a much respected and detailed design guidance and a state-of-the-art in the off-shore industry which is referenced in BS 8010 and other standards.

The main DNV codes relevant for the design of offshore pipelines are discussed in thefollowing sections.

3.2.2 DNV-OS-F101

Force Balance Method

Conventionally, a subsea pipeline has been considered stable if it has got sufficient sub-merged weight so that the lateral soil resistance is sufficiently high to restrain the pipelinefrom deflecting sideways. The main stabilizing method has traditionally been to applysufficient amount of concrete weight coating (CWC).

The traditional design approach for subsea pipelines which is expressed in the early designcodes, such as “Rules for Submarine Pipeline Systems, DNV (1976, updated in 1981)”,was to not to allow for any horizontal movement when a pipe is exposed to the environ-mental conditions associated with an extreme return period, i.e., a 100 year Return Period.

Even if this method has been widely replaced by the empirical or calibrated methods, theforce balance method is still in common use in cases where a pipeline is exposed to purecurrent. [27]

3.2.3 DNV-RP-E305

The design method presented in the code relates to a pipeline resting on the seabedthroughout its lifetime, or prior to some other form of stabilization (eg. trenching, burial,self-burial). The stability of the pipeline is then directly related to the submerged weightof the pipelines, the environmental forces and the resistance developed by the seabed soil.Consequently, the aim of stability design is to verify that the submerged weight of the



pipeline is sufficient to meet the required stability criteria.

Three different analysis methods were used in this code. The Dynamic Stability Analysismethod and two simplified or calibrated methods that do not require full dynamic FEanalysis, namely, the Simplified and Generalized Stability Analyses methods. [11]

Dynamic Stability Design Method

This method is based on a time-domain solution of the pipeline stability and incorporatesthree-dimensional effects, surface wave spectra and nonlinear soil resistance.

Generalized Stability Design Method

The Generalized method comprises a set of design response curves which have been de-veloped based on a large number of dynamic FE simulations. The background for themethodology is based on the assumption that the pipeline lateral displacement is to alarge extent a function of a relative small number of non-dimensional parameters. Thecurves for use in sandy soils are for net pipe movements of up to 40 pipe diameters in DNVZone 1 (more than 500 m away from a platform) or 0 m in DNV Zone 2 (less than 500m from a platform). No displacement is allowed in clays. The curves are based on piperoughness and sea state spectrum (JONSWAP), which helps determine the wave-inducedvelocity.

Simplified Stability Design Method

DNV recommends this method based on a link between the traditional stability designprocedure and the generalized stability analysis. The results from the two distinct meth-ods are made consistent with each other through the use of two calibration factors: onebased on the soil conditions, and one based on the Keulegan-Carpenter number and ratioof wave to current velocity. The calibration factors ensure that the results of the simplifiedanalysis tie in with the generalized analysis. [22]

A calibration factor is multiplied by the submerged weight determined from the tradi-tional type design to arrive at the new design submerged weight. The factor varies from1.0 to 1.6. The calibration gives pipe weights that produce a conservative envelope to thegeneralized procedure.

For this method, simplified design is characterized as;

• a hydrodynamic force (Morison type) formulation with force coefficients 0.7, 0,9,and 3.29 for CD, CL, and CM respectively and,

• a simple frictional soil resistance model with coefficients of 0.7 for sand and 0.15 to1.3 for clays.

The wave induced water particle velocity is taken as the significant bottom velocity. Thisdeviates from traditional designs where the wave induced water velocity is normally takenas that associated with the significant wave or sometimes the maximum wave. This anal-ysis is again based on pipelines designed with an allowable lateral displacement of up to



20 m in sandy soil and 0 m in clayey soil. As a result of extensive research and develop-ments performed in the 1980s in the area of pipeline stability, it was accepted that somemovement can be allowed during extreme sea states provided that the lateral displace-ments were kept within defined limits. This is reflected in the two calibrated methods,the Simplified and the Generalized Methods, which do not require FE analyses.

Limitations of DNV-RP-E305

The following shortcomings are identified in this code:

• The code does not provide guidance on pipe-seabed interaction forces for pipelineson carbonate soils;

• It does not allow for the effect of pipeline embedment on soil resistance and hydro-dynamic loading; and

• It does not consider the effects of seabed instability within the response of thepipeline during storm loading. [19]

This code has been replaced by the new DNV-RP-F109.

3.2.4 DNV-RP-F109

This code was first published in 2007 and updated in 2010, and it has superseded the oldDNV-RP-E305. This updated code does allow for some effects of pipeline embedment.However it does not consider asymmetrical embedment levels, and also does not providequantitative guidance for carbonate soils. Three stability analysis methods are includedin this code:

Absolute Lateral Static Stability Method

The Force Balance Method in DNV-OS-F101 (2000) is not included in the modified andrevised DNV-OS-F101 (2007) and again the Simplified method of DNV-RP-E305 (1988)is not included in the new DNV-RP-F109 (2010). The Absolute Lateral Static StabilityMethod in F109 replaces both these methods and it will ensure that no pipe motionwill occur when exposed to maximum load during a sea state. This method appears tobe significantly more conservative than the more traditional Force Balance Method. Inmost situations some minor pipeline movements (< 1m) can safely be allowed, whichwill significantly reduce the required pipeline submerged weight. [27] As this methodhas evolved, the criteria for defining pipeline stability have loosened, and now extend toallowing the pipeline to displace a significant predefined lateral distance laterally under agiven loading conditions. This allowance leads to the next two types of stability analysismethods. [24]

Generalized Lateral Stability Method

As opposed to E305 which presented design curves for lateral displacements ranging from0 to 40 times the external pipe diameter of the pipe, the new design code, F109 is basedon a lateral displacement limited to up to 10 pipe diameters during a given sea state. Butit follows similar design principles as the respective method in E305.


3.3. BRITISH STANDARDS - BS

Dynamic Lateral Analysis Method

To predict the displacement of a pipeline during a design storm event a full dynamicanalysis of the structural response is required. Although the use of dynamic FE analysesto calculate pipeline structural response is the most comprehensive method available toassess pipeline stability, the method has not been widely used by pipeline engineers forseveral reasons. Firstly, in many locations around the world where stability can readilybe mitigated by applying a minimal amount of Concrete Weight Coating (CWC), therehas not been a strong motivation for replacing the calibrated (Simplified and Generalized)methods with a more advanced FE-based method. Secondly, the design tools based onthis method are not easily available. [27]

Limitations of DNV-RP-F109

• This code focus on pipeline on-bottom stability with relatively low embedment lev-els, and is not suited to the assessment of highly embedded pipeline sections,

• This updated code allows for some effects of pipeline embedment; however it doesnot consider asymmetrical embedment levels, and also does not provide quantitativeguidance for carbonate soils,

• It does not consider the effects of the changes in seabed bathymetry (seabed topog-raphy) during a storm event on pipeline stability. [19]

3.3 British Standards - BS

BS-PD-8010-2:2004 - Code of Practice for Pipelines(Part 2: Subsea Pipelines)

3.3.1 Background

This Published Document has been prepared to take into account the publications ofBS-EN-1461, which is based on ISO-13623 (see section 3.3). It covers a number of issuesthat are outside the scope of BS-EN-1461, which has superseded BS-8010. The BS 8010was withdrawn and a more comprehensive approach to pipeline design is achieved by re-placing it using the standard in association with BS-PD-8010 which is an updated versionof BS-8010.

The Published Document is intended for use by designers, manufacturers and operatorsof pipelines and it is published in two parts:

• BS-PD-8010-Part 1: Steel Pipelines on Land

• BS-PD-8010-Part 2: Subsea Pipelines

This Published Document reviews the contents of BS-8010-1, BS-8010-2.8 and BS-8010-3and takes into account the recent publication and/or revision of a number of European,International and industry standards, including BS-EN-1594 and IGE/TD/1. [6]


3.3. BRITISH STANDARDS - BS

3.3.2 Stability Analysis

The pipeline should be designed to be stable during construction and operation. Wheremovement is permitted, it should not adversely affect the integrity of the pipeline.

The stability analysis should take into account:

a. Hydrodynamic forces resulting from the action of near-seabed, wave-induced andsteady currents on the pipeline;

b. Lateral soil forces;

c. Vertical stability;

d. Historic stability of seabed;

e. Axial forces in the pipeline, where appropriate.

Two-dimensional Analysis Method

If the two-dimensional analysis method is used, the lateral soil friction coefficient, should be greater than or equal to the stability safety factor, as shown in equation (3.1), and the illustration is identical as the DNV’s Force-Balance method.

µ(Ws − FL) 6 SF (FD + FI) (3.1)

NOTE 1 : SF is commonly 1.1. In special conditions, where environmental conditions areknown to a high degree of certainty, a lower value can be applied.

When considering the most unfavorable conditions on the pipelines, the effect of wavephase angle on the hydrodynamic loads should be taken into consideration.

This method is similar to the Force Balance method of stability is the DNV code. To sup-plement the above methodology, the approaches defined in DNV codes may be adopted,so long as all the criteria are met. [6]

Three-dimensional Analysis Method

In the three-dimensional case,as illustrated in Figure 3.1, the allowable movements shouldbe based on acceptable stress levels and fatigue. Possible loss of concrete due to movementand damage to the pipeline by a third-party activity should be considered. Note also thatthe approaches defined in DNV and AGA may also be considered. [6]In most cases, a two-dimensional method is acceptable for determining the stability of apipeline. In area where excessive stability is required, a more complex three-dimensionalanalysis may be used.

Limitations of BS-PD-8010-2

The application of this standard to deep water, and the guidance and applicability tohigh-pressure and high-temperature pipelines are the main limitations of this standard.


3.4. INTERNATIONAL STANDARDIZATION ORGANIZATION (ISO)

(a) Force due to steady cur-rent

(b) Force due to crestedwaves (c) Combined force

Key :1 Pipeline2 Wave− crest− length

Figure 3.1: Three dimensional stability analysis

3.4 International Standardization Organization (ISO)

3.4.1 General

International Standards give state-of-the-art specifications for products, services and goodpractice, helping to make industry more efficient and effective. ISO is the world’s largestdeveloper of voluntary International Standards. Since its foundation in 1947, ISO havepublished more than 19500 standards covering almost all aspects of technology and busi-ness. ISO are networks of national standards. These national standard bodies make upthe ISO membership and they represent ISO in their country. [17]

The ISO standard relevant for the offshore pipelines is the ISO-13623.

3.4.2 ISO-13623: Petroleum and Natural Gas Industries - PipelineTransportation Systems

ISO-13623 was prepared by Technical Committee, TC 67, Materials, equipment and off-shore structures for petroleum, petrochemical and natural gas industries, Subcommittee,SC 2, Pipeline transportation systems.

This second edition cancels and replaces the first edition, (ISO 13623:2000), which hasbeen technically revised.Pipelines shall be designed to prevent horizontal and vertical movement, or shall be de-signed with sufficient flexibility to allow predicted movements within the strength criteriaof this International Standard. This standard recommends the following factors whichshould be considered in the stability design;

• hydrodynamic and wind loads;

• axial compressive forces at pipeline bends and lateral forces at branch connections;

• lateral deflection due to axial compression loads in the pipelines;

• exposure due to general erosion or local scour;

• geotechnical conditions including soil instability due to, for example, seismic activity,slope failures, frost heave, thaw settlement and groundwater level;


3.5. AMERICAN PETROLEUM INSTITUTE (API)

• construction method, including bundled or piggybacked lines;

• trenching and/or backfilling techniques.

Limitation of ISO-13623

This standard does not gives a detailed overview of pipeline stability design as compared tostandards like DNV-RP-F109, which give a very detailed rational way of pipeline stabilityanalysis procedures.

3.5 American Petroleum Institute (API)

An offshore pipeline is subject to wave-induced and current-induced forces. For a pipelineresting on the seabed, lift and drag forces will be created. For that portion of a pipelinesuspended between seabed irregularities, oscillation due to vortex shedding can occur.Evaluations of these forces should be made by alternately assuming:

• The pipe is empty (construction condition), and

• It is full of transported fluid (operating condition).

The lift and drag forces created by the current-induced and wave-induced flow of wateron the sea bottom can result in excessive strains, fatigue from repeated lateral move-ments, encroachment on other pipelines, structures, bottom features, etc. of an offshorepipeline if not countered by a restraining force. Generally, a restraining force is suppliedby on-bottom weight of the pipeline. Wall thickness of the pipe, thickness and densityof the weight coating, or both are commonly used to control on-bottom weight. Wherebottom conditions and water depths permit, anchors or weights may be viable alternatives.

The AGA Level 2 and Level 3 Analysis for Submarine Pipeline On-Bottom Stability maybe used for assessing on-bottom stability requirements.

Specific geographic locations are subject to natural phenomena that can expose an off-shore pipeline to unusual forces. The design of an offshore pipeline should consider suchforces regarding stability and safety of the pipeline.Example of natural phenomena and their effect on offshore pipelines follow:

a. Earthquakes can liquefy some sea-bottom sediments : As a result, a pipeline couldtend to either sink or float, depending on specific gravity relative to the liquefiedbottom.

b. Hurricanes, cyclones and typhoons : Hurricanes, cyclones and typhoons can causehigh currents and large cyclic wave action, which together or individually can causeliquefaction or weakening of some sea bottom sediments. As a result, a pipelinemay tend to sink, float or move laterally.

c. Gross sea bottom movement (such as mudslides or sea bottom subsidence) maysubject a pipeline to large lateral forces. As a result, a pipeline may tend to sink,float or move laterally as the moving sediment is effectively liquefied.


3.6. AMERICAN GAS ASSOCIATION (AGA)

d. Sediment transport or scour of susceptible soils due to bottom currents and/or waveaction may result in exposure of a buried or partially buried pipeline, loss of soilrestraint or increase in free spans.

It may not be possible to quantify the effect of these natural phenomena for a specificoffshore pipeline and location. Consideration should be given to modifying an otherwiseoptimum design to reroute around a potential sea-bottom movement zone. In thoserare conditions where weight-coating or trenching methods may not represent a suitablesolution – such as on a solid rock surface or in shallow water zones or extremely highcurrents – the use of anchors or pipeline weights may be a viable addition or alternative.[2]

3.6 American Gas Association (AGA)

Three design procedures (Levels 1, 2, and 3) are presented in the A.G.A. design guide-lines and software. The first procedure (Level 1) is based on traditional stability analysismethods (Morison type hydrodynamic forces and frictional soil resistance). It is intendedonly as a reference to the type of static analysis which has been done in the past.

The most detailed procedure uses finite element time domain simulation software (Level3). The software provides detailed information regarding pipe movement and stressesduring design events. Pipeline safety is then assessed based on these results.

The Level 2 procedure assumes no net movement of the pipe, and is based on a quasi-staticcalculation which simulates the embedment process modeled in the Level 3 software. Theprocess modeled is that of a pipe embedding itself into the soil during small amplitudedisplacement oscillations caused by wave loadings. The resulting lateral soil resistance iscalculated and compared to the expected hydrodynamic forces to determine pipe stability.

The Level 2 and Level 3 analyses differ from the traditional method (Level 1) in severalimportant areas:

1. In oscillatory flow conditions, the peak hydrodynamic forces predicted by the Level2 and Level 3 software are typically much larger than those calculated using thetraditional, Level 1 approach. The Level 2 and Level 3 softwares are accurate,and hydrodynamic forces have typically been under-predicted in the traditionalapproach.

2. Levels 2 and Level 3 account for the effect that past cyclic wave loading historyhas on pipe/soil interaction. As a result, soil resistance forces are typically largerthan those considered for a Level 1 analysis. The Level 2 and 3 predictions moreaccurately represent actual soil resistance.

3. Levels 2 and 3 allow a much better representation of the sea state than the “designwave” approach used in Level 1. [14]

Limitation of AGA

• The AGA design code does not allow pipeline displacement;


3.7. COMPARISONS

• The AGA hydrodynamic model is not based on field measurement data, it’s basedsolely on laboratory model tests;

• The AGA design code does not assume a design sea spectrum, but allows the userto specify the sea state representation.

3.7 Comparisons

The offshore pipeline standards have their own advantages and disadvantages dependingon the type of offshore pipeline analysis they refer to. The BS and API standards presenta general design overview of on-bottom stability design on offshore pipelines. The designapproaches they use are very limited and not detailed. Whereas, the DNV and AGAdesign tools give a very detailed design procedures. Hence, the comparison in this sectionis mainly focused between these two codes to which most pipeline engineers refer.

3.7.1 Similarities

1. Both methods are based on very similar dynamic simulation tools which use a FEpipeline model subject to irregular waves plus steady current loadings.

2. Both methods include history dependent pipe-soil interaction models, where soilresistance and pipe embedment are based on recent pipe movement history.

3. Both are based on the assumption that the duration of the design sea-state is 3hours;

4. The hydrodynamic force formulations are quite different between the two designtools. The resulting forces produced by both models are very similar though. Thisdemonstrates the accuracy of the hydrodynamic forces estimation.

3.7.2 Differences

1. The AGA’s Level 2 and DNV’s Generalized method have been developed based onvery different approaches;

• The AGA tool performs a simplified analysis by estimating wave content of thedesign event. It then simulates the embedment process checks static stabilityafter some embedement, i.e., zero net pipe movement.

• The DNV’s Recommended Practice non-dimensionalizes results of numerousdynamic analyses and presents these results in general form. These resultsanticipate limited pipe movements between zero to 10 pipe diameters.

2. DNV’s code was developed based on field measurement data and laboratory tests.The AGA’s hydrodynamic model is based solely on laboratory model tests.

3. DNV’s design procedure assumes that the pipe has no initial embedment at the timethe design sea state occurs. The AGA’s Level 2 design procedure uses a conservativeestimate of the pipe embedment expected during storm build up to the design sea-state.


3.7. COMPARISONS

4. DNV assumes a JONSWAP type sea state spectra to generate the design charts. However, the AGA tools make no assumption regarding sea state type, but allows the user to define the sea state representation.

The analysis methods in DNV-RP-F109 range from an absolute stability where no move-ment of the pipeline is allowed to a lateral displacement of up to 10 pipe diameters. Thisrange makes it a very versatile and flexible standard which can be used for different typesof purposes to which the pipeline is intended for. For the on-bottom stability analysisprocedures, it can be seen that, the DNV standard, which is based on empirical designparameters, has a very detailed approach compared to all the other codes and standards.Besides, DNV standard is regarded as a well respected code to which the majority of theother standards refer to.

Therefore, the design procedures and sensitivity analyses in this Thesis are based on thestate-of-the-art DNV-RP-F109 standard, and it is explained deeply in Chapter5 .


Chapter 4

WAVE THEORIES ANDHYDRODYNAMIC LOADS

4.1 General

Submarine pipelines have been widely used to conveniently transmit crude oil and gasfrom offshore drilling wells to onshore facilities. In many offshore pipeline projects, thepipes are laid un-trenched on the seabed rather than fully embedded into the seabed, forcost efficiency and simplicity of installation. Despite its advantages, especially for oil andgas development in deep water,this option presents some design challenges. Its shallowplacement makes it vulnerable to current and wave induced instability due to the severeenvironmental conditions in the sea. [20]

4.2 Design Waves and Currents

4.2.1 Waves

The selection of design wave depends on the data available. Wave data sets are useful forextreme wave prediction only if they extend for at least 5 years, preferably more. Manyof them don not meet this criterion. Direct extrapolation from data covering 1 or 2 yearsis so unreliable as to be dangerously misleading.

When satisfactory wave data are available, the methods for predicting extreme waves arebased on the standard methods of extreme statistics. When there are no adequate wavedata, or when an independent check is warranted the design has to be based on winddata. Numerical modeling of wave surges and currents has reached a level of accuracyat which hind-cast data (obtained by numerical modeling of past events) are probably amore reliable guide to future extreme events than measured data. The primary value ofmeasured data is to validate numerical models. [22]

Wave Theories

Ocean waves are, generally, random in nature. However, larger waves in a random waveseries may be given the form of a regular wave that may be described by a deterministictheory. Even though these wave theories are idealistic, they are very useful in the design

26

4.2. DESIGN WAVES AND CURRENTS

of an offshore structure and its structural members. To determine wave particle velocity,the theory used depends on wave height (H), water depth (d) and wave period (T).

Figure 4.1: Regions of application of wave theories [9]

Wave theories can be classified as:

• Linear (Airy)wave theory;

• Solitary wave theory;

• Stoke’s 2nd/5th order wave theories;

• Cnoidal wave theory;



Figure 4.2: Water particles motion. [23]

• Stream function wave theory.

For most situations, linear theory is adequate as bottom velocities and accelerations donot vary significantly between theories. However, as the wave height to water depth ratioincreases, Stoke’s fifth order theory becomes appropriate. For shallow water or very highwave heights, a solitary theory should be used to predict particle velocity and accelera-tions. For breaking waves, or large diameter pipe that may affect the flow regime, otheranalysis methods may be appropriate. [7]

The specific wave theory to be used in the analysis is selected by using the graph in Figure4.1. In this Thesis, the Linear/ Airy’s wave theory is used.

Surface wave that generates on-bottom water particle motions, which may affect thepipeline on the seabed, is illustrated in Figure 4.2.

Wave Spectrum

Short term stationary irregular sea states may be described by a wave spectrum; that is,the power spectral density-function of the vertical sea surface displacement.

Wave spectra can be given in table form, as measured spectra, or by a parameterizedanalytic formula. The most appropriate wave spectrum depends on the geographicalarea with local bathymetry and the severity of the sea state. The Pierson-Moskowitz(PM) spectrum and JONSWAP spectrum are frequently applied for wind seas. The PMspectrum was originally proposed for fully-developed sea. The JONSWAP spectrum isformulated as a modification of the Pierson-Moskowitz spectrum for a developing sea statein a fetch limited situation. Both spectra describe wind sea conditions that often occurfor the most severe sea-states. The JONSWAP spectrum, which is wind-generated and



developed for the North Sea conditions, is often used and its spectral density function is:

Sηη(ω) = α.g2.ω−5. exp

(−5

4.

(ω

ωp

)−4).γ

exp(−0.5(ω−ωp

σ.ωp)2)

(4.1)

The Generalized Phillips’ constant is given by:

α =5

16.H2s .ω

4p

g2.(1− 0.287.lnγ) (4.2)

The spectral width parameter is given by:

σ =

{0.07 if ω ≤ ωp

0.09 if else(4.3)

The peakedness may be taken as:

γ =

5 if φ ≤ 3.6

exp(5.75− 1.15φ) if 3.6 ≤ φ ≤ 5.0, φ =Tp√Hs

1 if φ ≥ 5.0

(4.4)

The Pierson - Moskowitz spectrum appears for γ = 1.0. [12]

The main parameters in the wave spectrum are:

• Significant wave height,

• Peak period,

• Peaked-ness,

• Direction

• Spreading

Significant wave height (Hs) is only a statistical measure for the wave elevation processapproximately the average of (1/3)rd largest waves. Figure 4.3 shows that increasingsignificant wave height increases wave energy and the potential for instability.

Peak period (Tp) is the wave period for which the maximum energy density appears. Fig-ure 4.4 shows that increasing the peak period shifts the wave energy towards larger waveswhich increases the potential for stability.

Peaked-ness determines the sharpness of the spectrum and its effects depend on peakperiod and water depth. This dependency is shown in Figure 4.5.The wave induced velocity spectrum at the seabed, SUU(ω) may be obtained through aspectral transformation of the waves at sea level using a first order wave theory.

SUU(ω) = G2(ω).Sηη(ω) (4.5)



Figure 4.3: Effect of significant wave height on wave energy. [12]

The transfer function G transforms sea surface elevation to wave induced flow velocitiesat seabed and is given by:

G(ω) =ω

sin(k.d)(4.6)

where d is the water depth and k is the wave number established by iteration from theequation:

ω2

g= k. tanh(k.d) (4.7)

The spectral moment of order n is defined by:

Mn =

∫ ∞0

ωn.SUU(ω) dω (4.8)

The significant flow velocity amplitude at pipe level is:

Us = 2√M0 (4.9)

It is recommended to consider no boundary layer effect on the wave induced velocity. Themean zero up-crossing period of oscillating flow at pipe level is:

Tu = 2π

√M0

M2

(4.10)

Assuming linear wave theory, Us may be taken from Figure 4-5 and Tu from Figure 4.6,induced from surface wave (Hs and Tp), in which,

Tn =

√d

g(4.11)



Figure 4.4: Effect of peak period on wave energy. [12]

Figure 4.5: JONSWAP Spectrum (Peakedness) [10]

If the pipeline is designed to move significantly, the displacement can be assumed to beproportional with the number of waves τ , in the storm:

τ =TstormTu

(4.12)

The maximum wave height during a storm during τ waves:

Hmax =Hs

2.

(√2 ln τ +

0.5772√2 ln τ

)(4.13)

The maximum wave induced particle velocity is given by a single design oscillation derivedfrom the design spectrum using the equation:



Figure 4.6: Significant flow velocity amplitude Us at seabed. [12]

U∗ = 0.5.

(√2 ln τ +

0.5772√2 ln τ

).Us.RD (4.14)

The period associated with the single design oscillation velocity is given by:

T ∗ =

(

(kt − 5.(kt − 1).(TnTu

)

).Tu for

TnTu

6 0.2

1 forTnTu≥ 0.2

(4.15)

where, kt =

1.25 for γ = 1.0

1.21 for γ = 3.3

1.17 for γ = 5.0

(4.16)

The relationship between the mean zero crossing period and the period associated withthe single design oscillation velocity, T ∗, is site specific. In the absence of other data, theabove expression can be used.

Wave Period

Wave induced particle flow spectrum at seabed can be characterized by:

• Increasing depth: velocity decays with depth;

• Short period waves decay quicker than large period waves.

The difference between large period waves and short wave periods are illustrated in Figure4.8



Figure 4.7: Mean zero up-crossing period of oscillating flow Tu at seabed. [12]

Wave Directionality and Spreading

The direction of the waves has a significant influence on the on-bottom stability designof subsea pipelines. It is common for on-bottom stability design procedures that only theflow and acceleration components perpendicular to the pipe axis will try to move the pipe.

The effect of main wave directionality and wave spreading is introduced in the form of areduction factor on the significant flow velocity, i.e. projection onto the velocity normalto the pipe and effect of wave spreading. The graph in Figure 4.9 is used to get thereduction factor.

Normally, s is taken between 2 and 8. If no information is available, the most conservativevalue between 2 and 8 shall be taken. In the North Sea, a value range between 6 and 8may be used.

Wave Size

The size of waves can be described by the Keulegan - Carpenter number, K. As waterparticle move in an ellipse, the K value tells us large this ellipse as compared to the pipediameter. This phenomenon is illustrated in Figure 4.10.

K =Us.TuD

(4.17)

4.2.2 Currents

Design currents should be determined from statistical analysis of recorded data (assumingthese are of sufficient duration) in combination with numerical model simulations.

The steady current flow at the pipe level may have components including:



Figure 4.8: Large and short period waves. [23]

• tidal current;

• wind-induced current;

• density-driven currents;

• storm surge;

• ocean oscillations;

• long-shore currents; and

• river discharge. [6]

Steady currents develop boundary layer due to the viscous forces in the water and theboundary flow condition of zero flow at the seabed, which is illustrated in Figure 4.11.The location of the pipe in the velocity boundary layer lowers the effective velocity seenby the pipe.

The current velocity at the pipe can be shown by two approaches; the (1/7)th power lawand the method used by DNV-RP-F109.

• The (1/7)th power law predicts the current at a height z based on the readings onthe current meter (i.e. a reference velocity Ur at a height, zr). This is often fed onthe stability calculation as the predicted velocity at the level of the top of the pipe.

• The second formula which is based on DNV-RP-F109, and it is the average current(perpendicular to the pipe) over the height of the pipe which is modified to take



Figure 4.9: Reduction factor due to wave spreading and directionality. [12]

account of the effect of the seabed roughness, z0. The rougher the seabed thethicker the boundary layer and the slower the average velocity over the pipe height.Different seabed types with their respective roughness values are shown in Table4.1.

The current velocity may be reduced to take account for the effect of the bottom boundarylayer and directionality:

U(Z) = Ur(zr).

(ln (z + z0)− ln z0ln (zr + z0)− ln z0

). sin θc (4.18)

The mean perpendicular current velocity over a pipe diameter applies the expression:

Uc(z0) = Ur(zr).

(1 +z0D

). ln (D

z0+ 1)− 1

ln (zrz0

+ 1)

. sin θc (4.19)

where, θc is the angle between current velocity and the pipeline axis. In case directionalityinformation is unavailable, the current should be assumed perpendicular to the pipeline.[12]

The total flow velocity acting on the pipeline on the seabed is therefore the combination ofthe wave velocity and current velocity. Waves by nature are oscillating and do not have aspecific direction, whereas currents are more or less unidirectional. Hence, the criticality ofthe pipeline depends on whether the sea state is wave-dominated or current-dominated. Ifthe sea state transforms from wave dominated to current dominated, then the pipeline willbe under a constant pressure from one direction which will significantly affect its stability.



Figure 4.10: Keulegan - Carpenter number, K, relative to pipe diameter, D. [12]

Figure 4.11: The influence of current velocity on a pipeline on seabed [23]

Table 4.1: Seabed roughness of different types of soils. [9]

Seabed Grain size d50 [mm] Roughness z0 [mm]Silt and clay 0.0625 ≈ 5.10−6

Fine sand 0.25 ≈ 1.10−5

Medium sand 0.5 ≈ 4.10−5

Coarse sand 1.0 ≈ 1.10−4

Gravel 4.0 ≈ 3.10−4

Pebble 25 ≈ 2.10−3

Cobble 125 ≈ 1.10−2

Boulder 500 ≈ 4.10−2


4.3. HYDRODYNAMIC LOADS AND SEABED RESISTANCE

4.3 Hydrodynamic Loads and Seabed Resistance

4.3.1 General

On-bottom stability analysis is performed to ensure the stability of the pipeline, whenexposed to wave and current forces and other internal or external loads (e.g. bucklingloads in curved pipeline sections). The requirement to the pipeline is that no lateralmovements at all are accepted, or alternatively that certain limited movements, that donot cause interference with adjacent objects or over-stressing of the pipe, are allowed.

Conventionally a subsea pipeline has been considered stable if it has got sufficient sub-merged weight so that the lateral soil resistance is sufficiently high to restrain the pipelinefrom deflecting sideways. Since increasing the steel wall thickness in order to increasethe submerged weight will not be cost efficient, the primary stabilization method hastraditionally been to apply sufficient amount of concrete weight coating to achieve theon-bottom stability. [27]

A pipeline near the seabed is exposed to hydrodynamic forces from wave and currentactions as illustrated in Figure 1.1, in Chapter 1. Hydrodynamic stability is determinedusing Morison’s equation, which relates hydraulic lift, drag and inertial forces to localwater particle velocity and acceleration. The force variation is in general quite complex,and simple analytical expressions can only describe the force variation in an approximatemanner. [5]

4.3.2 Loads and Load Coefficients

Drag ForceThe drag force is mainly the result of the high pressure in front of the pipe and the lowpressure region in the wake behind the pipe. The drag is influenced by the width of thewake and the wave action. The effect of the wave is that the wake from the previous 50%wave cycle is swept back on the pipe again.

Inertial ForceWaves produce cyclic loadings on the water particles in the water column. The cyclicloads accelerate and decelerate the water particles in both the horizontal and verticaldirection. Where a body sits within the water flow it experiences the loads which wouldhave been exerted on the water that would have occupied the volume of the body.

Lift ForceLift is produced in the same way as flow over an airfoil. The presence of the seabed pro-duces an asymmetry between the flow over the top of the pipe and the flow underneath.This causes a slow or no flow underneath the pipeline (high pressure) and higher velocitiesover the top (low pressure), resulting in a lift.

The hydrodynamic loads resulting from the wave and current actions are given by theMorison Equations:



Drag force per unit length:

FD =1

2.ρw.D.CD(Us. cosφ+ Uc).|Us cosφ+ Uc| (4.20)

Inertia force per unit length:

FI =(πD2)

4.ρw.Cm.As. sinφ (4.21)

Lift force per unit length:

FL =1

2.ρw.D.CL(Us. cosφ+ Uc)

2 (4.22)

The water particle acceleration normal to the pipe:

As =(2π.Us)

Tu(4.23)

[12]

There is a phase difference of 900 between the maximum water particle velocity and ac-celeration. The maximum lift and drag forces occur when the inertia load is zero, and themaximum inertial load occurs when the lift force is at a minimum.

The drag, inertia and lift coefficients are empirically determined, and they vary dependingon the flow conditions. The magnitudes of the hydrodynamic loads depend on the flowboundary layer and the level of turbulence.

Lift and drag coefficients are affected by:

• The Reynold’s number of the flow,

• The pipe roughness (bare steel, concrete coated, Marine growth),

• The Keulegan-Carpenter number of the waves,

• Any embedment of the pipe into the seabed,

• Properties of the seabed (soil).

Typical values of drag, inertial and lift force coefficients for a pipe on the seabed are givenas:

CD = 0.7, CM = 3.29, and CL = 0.9 [12]

The hydrodynamic load experienced by the pipeline is part of the flow induced forces“caused by the relative motion between the pipe and surrounding water” [9]. Morison’sequation has traditionally been used to calculate the loads on the pipe. The submergedweight is then increased to ensure the resistance of the system avoids lateral displacement.

According to the new DNV’s recommended practice, DNV-RP-F109, 2010, experimentshave shown that the standard Morison type of force calculations based on ambient flow



velocity and with time invariant coefficients have proved inadequate for calculating lateraldisplacement, lead to an overestimation of total displacement. This is due to the develop-ment of a wake that is swept back over the pipe in succeeding half wave cycles [15]. Use oftime-dependent coefficients and near pipe velocity produces a significant improvement inthe calculation of forces under oscillatory wave loads. Hence, load reduction coefficientsare introduced and the design analysis procedures are presented in detail in Chapter 5.

4.3.3 Pipe-Soil Interaction

To avoid the occurrence of pipeline lateral instability, the surrounding soil must provideenough lateral soil resistance to balance the hydrodynamic loads. Before 1970s, the tra-ditional Coulomb friction theory was employed to calculate the lateral soil resistance tothe pipeline in waves. However, the model tests by (Lyonsetal.1973) showed that theCoulomb friction theory is not appropriate for describing the complicated pipe-soil inter-action.

Until now, several empirical pipe-soil interaction models have been proposed to predictthe ultimate lateral soil resistance to the pipeline in waves. In the pipe-soil interactionmodel by Wagner et al. 1989, it was assumed that the ultimate soil lateral resistance isthe sum of the sliding resistance component and the soil passive resistance component.[16]

According to F109, the reduction of the hydrodynamic loads may occur due to:

• Permeability of the seabed; rperm,i

• Pipeline penetration into the seabed; and rpen,i

• Trenching, rtrench,i

The subscripts y and z stand for the reduction factors in the horizontal and verticaldirections, respectively.

Permeable Seabed

A permeable seabed will allow flow in the seabed underneath the pipe and thus reducethe vertical load, and rperm,z can be taken as 0.7. Figure 4.12 and Figure 4.13 illustratethe relationship between zp and D.

Penetration

The penetration factors in the horizontal and vertical directions are given, equations 4.24and 4.25 :

rpen,y = 1− 1.4 ∗ zpD≥ 0.3 (4.24)

rpen,z = 1− 1.3 ∗ (zpD− 0.1) ≥ 0 (4.25)



Figure 4.12: Illustration of penetration [12]

Trenching

The load reduction factors due to trenching, shown in Figure 4.14 and Figure 4.15 in thehorizontal and vertical directions are given, respectively, as:

rtrench,y = 1.0− 0.18 ∗ (θ − 5)0.25 ∗ (ztD

)0.42, 5 6 θ ≥ 5 (4.26)

rtrench,z = 1.0− 0.14 ∗ (θ − 5)0.43 ∗ (ztD

)0.46, 5 6 θ ≥ 45 (4.27)

The total reduction is then found as:

rtot,y = rpen,y ∗ rtrench,y (4.28)

rtot,z = rperm,z ∗ rpen,z ∗ rtrench,z (4.29)

4.3.4 Seabed Resistance

Seabed resistance can generally be represented by two parts: a pure Coulomb friction anda passive resistance, due to the build-up of soil penetration as the pipe moves laterally.The resistance can be expressed as:

R = µ(Ws − FL)− FR (4.30)

where: Ws = (self weight – buoyancy), and FL is the lift force.

The self-weight includes the pipe content, steel, concrete coating, corrosion coating andmarine growth. And the buoyancy is based on the overall pipe diameter. As per [12], themain seabed soil types are: sand, clay and rock and they can be defined as:

Sand is a soil that is permeable with negligible cohesive effects.Clay is a soil that is not permeable with significant cohesive effects.Rock are crushed rocks with 50% diameter fractile larger than 50 mm.

The coefficients of friction, µ, for concrete coated pipe, can normally be taken as 0.6 onsand, 0.2 on clay and 0.8 on rock. As the lift force fluctuates through the wave cycle, the



Figure 4.13: Peak load reduction due to penetration [12]

Figure 4.14: Trench parameters [12]

resistance will also fluctuate. [12]

When the pipe is first laid on the seabed, a small amount of settling or embedment willoccur. Each wave half-cycle pushes the pipe against the small soil berm created by thepipe resting on the seabed. As wave loads gradually increase during a storm build up, theforces displace the pipe back and forth against the soil berms, gradually pushing themand enlarging them. As a consequence, the pipe moves further down into the seabed. Thephenomenon is illustrated in Figure 4.16.

Passive Resistance

Sand



Figure 4.15: Peak load reduction due to trenching [12]

Figure 4.16: Soil embedment [23]

FRFC

=

(5.0 ∗ κ2s) ∗ (

zpD

)1.25 if κs ≤ 26.7

κ ∗ (zpD

)1.25 if κs ≥ 26.7(4.31)

κs =(γ′ ∗D2)

Ws − FZ=

(γ′ ∗D2)

FC(4.32)

where, FC = Ws − FZ

Clay



The passive resistance on clay can be taken as:

FRFC

= 4.1 ∗ κc(GC)0.39

∗(zpD

)1.31(4.33)

where, GC =Su

(D ∗ γs), and κc =

(Su ∗D)

FC


Chapter 5

ON-BOTTOM STABILITYDESIGN METHODOLOGIES(ACCORDING TO DNV-RP-F109)

5.1 Load Combinations

The characteristic load conditions to be identified shall reflect the most probable extremeresponse over a specified period of design time. Before being put to their main purpose,i.e. to transport oil and gas, pipelines need to be transported and installed on site. Hence,the load combinations to be considered in the design process are divided in two conditions:the temporary and permanent conditions. [12]

5.1.1 Temporary (Empty) Conditions

Temporary conditions are the conditions considered during the installation phases, whenthe pipeline is empty, and can be divided into two parts.

• For durations less than 12 months but in excess of 3 days: Here a 10-year returnperiod for the actual seasonal environmental conditions applies. An approximationto this condition is to use the severe condition among the following two combinations:

– A 10-year wave with a 1-year current,

– A 1-year wave with a 10-year current.

• For durations less than three days, an extreme load condition may be specified basedon reliable weather forecasts, For durations over 12 months, the condition can betaken as permanent. [12]

5.1.2 Permanent (Operational) Conditions

For permanent operational conditions and temporary conditions in excess of 12 months, a100-year return period applies, i.e., the characteristic load condition is the load conditionwith 10-2 annual exceedance probability. When detailed information about the jointprobability of waves and currents is not available, this condition may be approximatedby the most severe condition among the following two combinations:

44

5.2. ANALYSIS METHODS

• A 100-year return period for waves combined with a 10-year current,

• A 10-year return period for waves combined with a 100-year current. [12]

5.2 Analysis Methods

When performing pipeline on-bottom stability analysis, the pipe has to be stable bothvertically and laterally. Hence, these two methods are explained thoroughly in the comingsections.

A. Lateral Stability Analysis

The drag and inertia forces of the hydrodynamic loads tend to move the pipeline laterally.The soil resistance force also tends to resist this movement. The requirement is to ensurethat the pipe does not move laterally or moves within the allowed displacement limits. Thethree types of lateral stability analysis methods adopted by DNV-RP-F-109 are discussedhere. As the Dynamic Stability Method uses a Finite Element Analysis and has not beenwidely used by pipeline engineers, we will concentrate on the two methods: Absolute andGeneralized Stability Methods. [12]

5.2.1 Absolute Lateral Static Stability Method

The Absolute Stability Method gives an absolute static requirement for lateral on-bottompipelines based on static equilibrium of forces that ensures that the resistance of thepipe against motion is sufficient to withstand maximum hydrodynamic loads during a seastate, i.e. the pipe will experience no lateral displacement under the design extreme singlewave induced oscillatory cycle in the sea state considered. It is further based on a LoadResistance Factor Design (LRFD) approach with additional partial safety factors whichis said to satisfy the target safety level in the DNV code, [9].

In this method the lift, drag and inertia coefficients are replaced with horizontal andvertical load coefficients. These coefficients are based on experimental data.

Design Loads

The peak horizontal and vertical loads are:

F ∗Y = rtot,y.1

2.ρw.D.C

∗Y (U∗ + V ∗)2 (5.1)

F ∗Z = rtot,z.1

2.ρw.D.C

∗Z(U∗ + V ∗)2 (5.2)

U∗ can be taken from equations (4.15) and (4.16) and V* can be taken from Section 4.2.2.(Currents).

The peak load coefficients C∗Y and C∗Z are taken by combining Table 5.1 Table 5.2. Theintermediate values can be interpolated between the curves.



Load reductions due to a permeable seabed, soil penetration and trenching, can be cal-culated according to Section 4.3.3

The Keulegan-Carpenter number for single design oscillation and the associated steadycurrent to oscillatory wave velocity ratio, which are needed to determine the peak loadcoefficients, are calculated using equations 5-3 and 5-4 in reference to Tables 5-1 and 5-2.

K∗ =U∗.T ∗

D(5.3)

M∗ =V ∗

U∗(5.4)

Table 5.1: Peak horizontal load coefficients. [12]

Table 5.2: Peak vertical load coefficients. [12]

Table 5.3: Safety factors, winter storms in North Sea. [12]

For other regions and sea states, it can be seen in the DNV code, [12], section 3.6.3



(a) Horizontal (b) Vertical

Figure 5.1: Peak load coefficients [12]

Design Criteria

A pipeline can be considered to satisfy absolute stability requirement if:

γsc.F ∗Y + µ.FZµ.Ws + FR

6 1.0 (5.5)

and,

γsc.F ∗ZWs

6 1.0 (5.6)

The safety factors in winter storms in the North Sea are tabulated in Table 5.3.

5.2.2 Generalized Lateral Stability Method

This method limits the pipeline displacement in a design sea state such that the associ-ated strain remains within the acceptable limits. These limits range from a displacementof half of the pipeline diameter up to an allowable displacement of 10 pipeline diameters.However, the effects of axial loading due to high operating temperatures and pressuresand restraints at pipe ends are neglected and care should be taken to avoid buckling ofthe pipeline.

The dimensionless lateral pipe displacement, Y, is to a large extent governed by a set ofnon-dimensional parameters:

Y = f(L,K,M,N, τ,Gs, Gc) (5.7)

The specific weight of a pipe can be calculated by:

Sg = 1 +2

π∗ (N ∗K ∗ L), 1.05<Sg ≤ 3 (5.8)



where;

N =Us ∗RD

G ∗ Tu, K =

Us ∗RD ∗ TuD

, L =ws

0.5 ∗ ρw ∗D ∗ U2s

(5.9)

The Generalized stability method provides for an on-bottom stability design with anallowed lateral displacement within the range of less than half a pipe diameter (virtuallystable) up to a significant displacement of 10 diameters during the given sea state.Hence,Lstable is the weight required for achieving a virtually stable pipe; and

L10 is the weight required for achieving a displacement of 10 pipe diameters.

Lstable is independent of sea state duration whereas L10 is valid for 1000 waves and canbe assumed to be proportional to the number of waves τ in the sea state. If L<Lstable,then displacement should conservatively be regarded as varying linearly with number ofwaves in the sea state:

Yτ = 0.5 + (10− 0.5) ∗ τ

1000= 0.5 + 0.0095 ∗ τ (5.10)

It should be noted that the values are obtained from a large number of one dimensionaldynamic analyses, i.e. on flat seabed and neglecting bending and axial deformation of thepipe. [12]The design procedures for sand and clay soil can be carried out as follows:

Stability on Sand

For virtually stable pipe, the minimum pipe weight required can found from the designpoints independent of N as shown in Table 5.4 :

Table 5.4: Minimum weight, Lstable/(2 +M)2, for pipe on sand, K > 10. [12]

And the minimum pipe weight required for a virtually stable pipe dependent of N, hasthe design points shown in Table 5.5.

For lateral displacement up to 10 pipe diameters, the minimum pipe weight required canbe found from the design points shown in Table 5.6.



Table 5.5: Minimum weight, Lstable/(2 +M)2, for pipe on sand, K 6 5. [12]

Table 5.6: Minimum weight, L10/(2 +M)2, for pipe on sand. [12]

The minimum weights for pipe on sand can also be found from Figures 5.2 and 5.3.



Figure 5.2: Minimum weight, Lstable/(2 +M)2, for pipe on sand. [12]

Figure 5.3: Minimum weight, L10/(2 +M)2, for pipe on sand. [12]



Stability on Clay

To limit the maximum relative displacement Y to less than 0.5 on a clayey seabed, theminimum required pipe weight can be calculated by the following formula:

Lstable = 90

√Gc

N0.67.K∗ f(M) (5.11)

where,

f(M) = [0.58(logM)2 + 0.60 ∗ (logM) + 0.47]1.1 6 1.0

To limit the maximum relative displacement Y to (10 ∗ τ1000

), the minimum required pipeweight can be calculated by the following formula:

L10

(2 +M)2=

C1 +

C2

KC3for K ≥ Kb

C1 +C2

KC3for K<Kb

(5.12)

The coefficients can be found tabulated in Appendix A in the DNV’s On-Bottom StabilityRecommended Practice. [12].

5.2.3 Dynamic Lateral Stability Method

In the Dynamic Stability Analysis, the surface wave spectrum must be transformed to atime series for the wave induced particle velocity at the pipe position at the seabed.

The objective of this method is to calculate the lateral displacement of a pipeline sub-jected to hydrodynamic loads from a given combination of waves and currents during adesign sea state. [12]

This method requires the use of Finite Element Analysis (FEA) procedures and it has notbeen widely used by pipeline engineers for several reasons. One of the main reasons hasbeen that in many locations around the world where stability can be readily mitigatedby applying a minimal amount of concrete weight coating, there has not been a strongmotivation for replacing the simplified force model or the calibrated methods with a moreadvanced FE based method. Secondly design tools based on these methods are not easilyavailable. [27]

B. Vertical Stability Analysis

The vertical component of the hydrodynamic force (lift force) tends to lift the pipe up-wards. The requirement is to ensure that the pipe sinks into the water, floatation in watershould be avoided and it should not sink into the soil.

In order to avoid floatation in water, the submerged weight of the pipeline should fulfilthe criteria shown in equation equation 7.2.

γw ∗b

(Ws + b)=γwSg≤ 1.00 (5.13)


5.3. STABILITY ENHANCEMENT METHODS

The factors that need to be considered when assessing the vertical stability of a pipelineinclude:

• the specific weight of the soil;

• shear strength of the soil; and

• seabed liquefaction.

Effects of vertical instability

In case sinking of the pipe occurs, the adverse effects that may occur include:

• over-stressing of the pipe due to uneven sinking;

• obstruction to future access to the pipe.

5.3 Stability Enhancement Methods

After performing the analysis, if the pipe’s stability still needs to be enhanced, one ormore of the following methods should be considered:

Stability enhancement options:

• increasing the wall thickness;

• concrete weight coating (CWC);

• trench and backfill;

• rock dumping;

• flexible mattresses/ concrete saddles;

• fronds;

• anchors/suction anchors;

• grout bags/saddles;

• artificial seaweed mats;

• piling anchors;

• pipe-in-pipe.

When pipelines need to be designed to extremely high hydrodynamic loads, burial of theline by trenching and back-filling is the best option. [12, 6, 1]


5.4. ANALYSIS SOFTWARE: SESAM (STABLELINES)

5.4 Analysis Software: SESAM (StableLines)

5.4.1 General

STABLELINES is a Microsoft Excel VBA spreadsheet developed by DNV for the on-bottom stability design and assessment of submarine pipelines in compliance with DNV-RP-F109 “On-Bottom Stability Design of Submarine Pipelines”, issued October 2010.DNV-RP-F109 has replaced the design code DNV-RP-E305, which has been used exten-sively by the industry since it was issued in 1988.

The main objectives of this new revision which are found in the DNV code [12] are pre-sented below:

• harmonise the design approach for on-bottom stability with the design method inDNV-OS-F101, “Submarine Pipeline Systems”;

• present a design method with calibrated safety factors that ensure absolute stability;

• expand the window of applicability for sandy and clayey seabed allowing a displace-ment of up to 10 pipe diameters:

• present requirements for on-bottom stability design based on full dynamic analysisin more detail. [8]

5.4.2 Features

The following features are included in StableLines:

• two design methods for stability are used:

– Absolute lateral stability;

– Generalized lateral stability for 0.5D and 10D

• four different pre-defined soil models (Sand, Clay, Rock and Coulomb friction) areused;

• sea-states are modeled based on JONSWAP spectrum;

• wave induced particle velocities at seabed level are automatically calculated basedon user inputs, significant wave height (Hs) and peak period4 (Tp);

• reduction of both current and wave induced velocities is made based on directionalitywith independent specification of wave and current directions;

• there is a possibility of user-defined reduction factor due to added seabed penetrationof Pipeline/Umbilical;

• simultaneous calculation of pipeline in empty condition and in operational condi-tions;

• suggestions of added steel and concrete/armour thicknesses to stabilize the unstablepipelines/umbilicals:


5.4. ANALYSIS SOFTWARE: SESAM (STABLELINES)

• iteration calculation of optimal necessary pipe weight and concrete thickness for theAbsolute stability method, (Not for the Generalized method);

• graphical presentations of:

– required submerged weight as function of direction;

– required concrete/armour thickness as function of concrete/armour density;

– required submerged weight as function of current velocity;

– dimensionless weight parameter L as function of K (Keulegan-Carpenter num-ber) and sensitivity to M (ratio of current velocity to wave induced particlevelocity).

• report of final and intermediate results;

• powerful tool to analyse stability of a large number of cases simultaneously (Para-metric Runs). [8]


Chapter 6

ANALYSIS, RESULTS ANDDISCUSSIONS

6.1 Analysis

The design analysis in this Thesis will be carried out based on the new and state-of-the-art design code, DNV-RP-F109. This code is considered the most appropriate standardfor future design and it has an international approbation [5]. The software package to beused in this analysis (SESAM: StableLines) is also developed according to this code byDNV (see section 5.3 ).

The design procedure will be made as follows:

The design analysis procedures are explained in detail in Chapters 4 and 5. Here, onlythe results of the analysis are discussed. The manual calculations and the the outputs ofthe SESAM: StableLines software are attached in Appendices A and B respectively.

The analysis is made for different wave and current directions (00 to 3300) and the criticaldirection (usually perpendicular to the pipeline axis) is taken as the critical direction forthe design loads. The temporary (empty) and permanent (operational) conditions will beanalysed simultaneously.

The analysis is made for both the lateral and vertical stability methods. The AbsoluteLateral Stability method is first used to check the stability of pipeline where no movementof the pipeline is allowed. In case instability of the pipeline encountered, the General-ized Lateral Stability method is used, where pipeline displacement between 0.5 and 10pipe diameter is allowed. The result is presented in tables and respective graphs. Thevertical stability checks will also be made afterwards. Finally, sensitivity analysis will beperformed for the selected parameters, to see the response of the pipeline’s stability withrespect to each parameter.

As discussed in section 5.2.3, the Dynamic Lateral Stability Method has not been widelyused by pipeline engineers for several reasons, hence, it is not considered in the analysis.

55

6.2. RESULTS AND DISCUSSIONS

6.1.1 Design Data

The data used for the analysis represents a North Sea project from Subsea 7. The datais attached in Appendix B

For the temporary (empty) conditions, the SESAM:StableLines software considers theworst of the following two wave-current combinations as explained in detail in section 5.1

• A 10-year wave with a 1-year current,

• A 1-year wave with a 10-year current.

The operational condition is also carried out in the same manner by the software as inthe installation case, taking the worst of the following wave-current combinations.

• A 100-year return period for waves combined with a 10-year current,

• A 10-year return period for waves combined with a 100-year current. [12]

6.2 Results and Discussions

The design analysis procedures are explained in detail in Chapters 4 and 5. Here, only theresults and discussions of the analysis are shown in tables and graphs. The manual calcu-lations and the the results of the SESAM:StableLines software are attached AppendicesA and B.

6.2.1 Absolute Lateral Stability

In this method a similar approach to the force-balance method is adopted. However, thelift, drag and inertia coefficients are replaced with horizontal and vertical load coefficientsbased on experimental data. The pipeline is not allowed to move.

The analysis result shows that the pipeline does not fulfill the Absolute Lateral Stabilitycriteria for both the empty and operational conditions. To obtain the optimal necessarypipe weight and concrete thickness, an iteration is made as shown in Tables 12 and 13.Increasing the concrete thickness is the main option to stabilize the pipeline, hence, thethickness is increased from 55mm to 123mm in the empty condition and to 257mm in theoperational condition, which is a significant amount in each case as shown in Table 6.1.The thickness of the steel pipelines is also added by 17 mm and 37 mm for the emptyand the operational phases respectively. Therefore, the stability should be checked usingthe Generalized Lateral Stability method. The required submerged weights for both theempty and operational conditions are shown in Figure 6.1.



Figure 6.1: SESAM:StableLines Output: Necessary weight vs. direction presentation


6.2.R

ES

UL

TS

AN

DD

ISC

US

SIO

NS

Table 6.1: StableLines: Absolute Lateral Stability main page

58O

n-B

ottomS

tabilityA

nalysis

ofS

ubsea

Pipelin

esA

ccording

toD

NV

-RP

-F109


6.2.2 Generalized Lateral Stability

The Generalized Lateral Stability method adopts a quasi-static approach, which limitsthe pipeline displacement such that the it remains within the acceptable limits. Theselimits range from a displacement of a half of the diameter for a virtually stable pipe up toan allowable displacement of 10 pipe diameters. Design curves generated from dynamicstability assessments (performed on a flat seabed, neglecting the effect of axial loads) areused to determine the minimum pipe weight required to satisfy these stability criteria.The analysis is made for two lateral displacement conditions, 0.5 and 10 pipe diameters.

Generalized 0.5D

A lateral displacement of 0.5 pipe diameters is allowed in this method. The analysisresult shows that the pipeline is stable for the empty conditions and unstable for theoperational conditions. The results are displayed in Table 6.2. The necessary pipelinesubmerged weights are illustrated in Table 17 and Fig 8. As the pipeline is not stablein the operational condition. The thickness of concrete required is increased from 55mmto 83mm, a relatively smaller amount compared to the Absolute Stability. Besides, thethickness of the steel is kept constant in the empty phase and it is increased by 8 mmin the operational phase. Hence, a further analysis is made by increasing the lateraldisplacement criteria to a maximum of 10 pipe diameters.



6.2.R

ES

UL

TS

AN

DD

ISC

US

SIO

NS

Table 6.2: StableLines: Generalized Stability, Displacement=0.5D main page

60O

n-B

ottomS

tabilityA

nalysis

ofS

ubsea

Pipelin

esA

ccording

toD

NV

-RP

-F109


Generalized 10D

The pipeline in both the above methods is not completely stable. Hence a maximumlateral displacement criteria of 10 pipe diameters is allowed in this method. As can beseen in Table 6.3, the pipeline is stable in the empty condition, but it is still unstable inthe operational condition. But the required concrete thickness has reduced as comparedto the above method (Generalized 0.5D. An additional concrete thickness of 14mm isrequired to stabilize the pipeline which a relatively small value. Whereas, no increment ofthe steel thickness is required. Therefore, this method is chosen to be the analysis designmethod for this pipeline in the North Sea.



6.2.R

ES

UL

TS

AN

DD

ISC

US

SIO

NS

Table 6.3: StableLines: Generalized Stability, Displacement=10D main page

62O

n-B

ottomS

tabilityA

nalysis

ofS

ubsea

Pipelin

esA

ccording

toD

NV

-RP

-F109

6.3. SENSITIVITY ANALYSIS

6.2.3 Vertical Stability

In the vertical stability analysis, the main requirement is to make the pipeline sink intothe water. As can be seen in the analyses results for both the Absolute and GeneralizedStability methods, the required submerged weight for the vertical stability is very smallcompared to the actual submerged weight.

For the temporary (empty) condition;

Ws(= 2088N/m) ≥ Ws,vertical(= 421N/m)

and;

For the permanent (operational) condition;

Ws(= 2426N/m) ≥ Ws,vertical(= 421N/m)

Hence, the pipeline is stable vertically for both the empty and operational conditions.

6.3 Sensitivity Analysis

Sensitivity analysis is done by performing a number of parametric analyses, which is usedto analyse different cases and observe the behaviour of the on-bottom stability of thepipeline under consideration.

This sensitivity analysis is done for both the Absolute and Generalized Lateral StabilityMethods, and the effects of all the parameters are plotted, which are attached in AppendixB.

In this case study, the following four input parameters are utilized;

• Case I: concrete thickness,

• Case II: pipe diameter,

• Case III: water depth and

The response of the stability of the pipeline based on the required submerged weight(Ws,required) will be evaluated with the varying values of each parameter keeping the restof the parameters constant.

6.3.1 Case I: Concrete Thickness, tconc

Concrete thickness is the main parameter which directly affects the submerged weightof the pipeline. Concrete thickness to be added on the steel pipeline depends largely onthe purpose for which the pipeline is intended. A larger thickness may be required forpipelines where no movement is allowed. For those pipelines where lateral movement isallowed, the thickness of concrete required can be smaller as compared to the absolutelystable pipelines.



In the analysis, the thickness of the concrete is being varied while keeping the rest of theparameters constant. The response of the pipeline’s on-bottom stability to this changeis illustrated in terms of the necessary submerged weight against the change of concretethickness in Figure 5.

From the output, it can be seen that as the thickness of the concrete is increased signifi-cantly, the necessary submerged weight shows a gradual increase for both the empty andoperational conditions.


6.3.S

EN

SIT

IVIT

YA

NA

LY

SIS

Figure 6.4: SESAM: StableLines Output: Necessary weight vs concrete thickness, pipe diameter and water depth from top to bottom

On

-Bottom

Stability

An

alysisof

Su

bseaP

ipelines

Accordin

gto

DN

V-R

P-F

10965


The effect of the difference in water depth on the added concrete thickness is also analyzed,for both the empty and operational conditions. The illustration can be seen in Figure 6.5.


6.3.S

EN

SIT

IVIT

YA

NA

LY

SIS

Figure 6.5: SESAM:StableLines Output: Concrete thickness vs. water depth presentation

The result shows, as the water depth increases from the mean sea level downwards, the required added concrete thickness decreases, andbecomes zero at some specific depth. This is because when the water depth increases, the wave length (L) increases. This reduces thehorizontal and vertical loads on the pipeline, reducing the submerged weight, in turn decrease the concrete thickness. For example;

• Absolute Lateral Stability : No additional concrete thickness is required beyond the water depth of 140 meters for both the emptyand operational conditions;

• Generalized Lateral Stability (0.5D): No additional concrete thickness is required beyond the design water depth of 104 meters forthe empty condition. And no added concrete thickness is required beyond the depth of 140.75 meters for the operational condition;

• For the Generalized Lateral Stability (10D) method, no additional concrete thickness is required beyond the water depth of 158.75meters for the empty condition. And no added concrete thickness is required for the operational condition beyond the depth of140.75 meters.

On

-Bottom

Stability

An

alysisof

Su

bseaP

ipelines

Accordin

gto

DN

V-R

P-F

10967


6.3.2 Case II: Pipe Diameter

Pipeline diameter is one of the main parameters which affects the submerged weights to agreat extent. In the analysis the effect of the pipeline diameter to the necessary submergedweight is analyzed. Pipeline diameter is not directly involved in the calculation of waterparticle velocities. But it affects the hydrodynamic forces directly. The forces increasewith the increment of pipeline diameter. Hence, as the pipeline diameter increases, thepipeline submerged weight also increases as illustrated in Figure 5

6.3.3 Case III: Water Depth

The water depth affects also stability analysis to a great extent. i.e. as the depth getsdeeper, the wave length increases and the wave velocity becomes almost negligible atsea bed level and correspondingly wave velocity contribution to the hydrodynamic loadsbecomes very small. The submerged weight of the pipeline decreases with the increment ofwater depth. Hence the effect om the on-bottom stability of submarine pipeline is very less.This reduces the hydrodynamic forces because the water particle kinematics decreases,which in turn contributes to the decrement of the submerged weight. In the result as shownin Figure 5, the increase of the water depth decreases the required submerged weight closeto zero, meaning, the pipeline is stable at deeper water depths. This conclusion can also beproved by another result in Figure 6.5 which shows the added concrete thickness reducingto zero as the water depth increases.


Chapter 7

CONCLUSIONS &RECOMMENDATIONS

7.1 Conclusions

On-Bottom Stability of Subsea pipelines is governed by a fundamental balance of forcesbetween loads and resistances. To do this analysis, a governing standard or code is nec-essary for the pipeline engineer. These codes and standards have been developed sincethe start of the installation of subsea pipelines. The main design codes are discussed inChapter 3 and their advantages and shortcomings are discussed. The DNV-RP-F109 ischosen to be the most appropriate and relevant design code.

One of the main objectives of this Thesis, as discussed in section 1.3, is to use the state-of-the-art design code and show the design methodologies and procedures in the sensitivityanalysis of on-bottom stability of subsea pipelines. The analysis is done in Chapter 6 andthe sensitivity analysis is performed accordingly using the SESAM: StableLines software,which is developed according to DNV-RP-F109. A case study data from a North Seaproject of Subsea 7 is used in the analysis.

According to the case study, concrete and steel wall thicknesses are the most criticalparameters. Achieving an absolutely stable pipeline needs addition of a large concretethickness and/or steel wall thickness. Since it is not cost efficient to increase the steelwall thickness in order to increase the submerged weight, the primary stabilization methodhas been to apply sufficient amount of Concrete Weight Coating (CWC) to achieve theon-bottom stability.

Therefore, allowing a pipeline lateral movement of up to 10 pipe diameters can minimizethe additional CWC and steel wall thickness requirements.

7.2 Recommendations

Applying CWC has been the primary stabilization method to achieve the on-bottomstability. But there is a practical limit to how much CWC can be applied due to the lim-itation of the tension capacity of a pipe-laying vessel. A secondary stabilisation methodmay have to be adopted such as through lowering the pipeline into the seabed by pre-laydredging or post-lay trenching or by using on-seabed restraints. On-seabed restraints can

69

7.2. RECOMMENDATIONS

involve covering the pipeline by crushed rock. However, where trenching and backfillingor rock dumping are not technically feasible or cost efficient, the designer may have toresort to more costly and technically challenging solutions such as anchoring the pipelineto the seabed.[27]

Even if in this Thesis, the Absolute and Generalized stability methods are used andseem sufficient, the use of Dynamic Lateral Stability method is very important and rec-ommended for pipelines for which stability if a major design challenge. DNV-RP-F109allows the use of advanced Dynamic FEA for on-bottom stability even though it is notpopularly used by pipeline engineers. Its use can prove advantageous in reducing theadditional cost of different types of mitigation methods, which can be used for improvingthe stability.

As the analysis made in this Thesis is true only for the part of the pipeline which is notclose the lateral restraints, care must be taken to avoid the risk of accumulating excessiveloads at these lateral restraint joints. The lateral restraints can be, subsea manifolds,platforms, risers, etc.


REFERENCES

[1] ABS. Guide for Building and Classing Subsea Pipeline Systems, 2006, updated 2014.

[2] API-RP-1111. Design, Construction, Operation and Maintenance of Offshore Hydro-carbon Pipelines (Limit State Design), 1999, updated 2005.

[3] Yong Bai and Qiang Bai. Subsea Pipelines and Risers. Elsevier Ltd., Kidliington,Oxford, 2005.

[4] Yong Bai, Gerhard Knauf, and Hans-Georg Hillenbrand. Materials and Design ofHigh Strength Pipelines. The International Society of Offshore and Polar Engineers,2000.

[5] M. W. Bræstrup, J. B. Andersen, L. W. Andersen, M. B. Bryndum, C. J. Christensen,and Niels Rishøy. Design and Installation of Marine Pipelines. Blackwell ScienceLtd., Oxford, UK, 2005.

[6] BS-PD-8010:2. Code of Practice for Pipelines - Part 2: Subsea Pipelines, 2004,confirmed 2010.

[7] Subrata K. Chakrabarti. Handbook of Offshore Engineering. Elsevier Ltd, Plainfield,Illinois, 2005.

[8] DNV. SESAM User Manual, StableLines, 2013.

[9] DNV-OS-F101. Submarine Pipeline Systems, 2012.

[10] DNV-RP-C205. Environmental Conditions and Environmental Loads, 2010.

[11] DNV-RP-E305. On-Bottom Stability Design of Submarine Pipelines, 1988.

[12] DNV-RP-F109. On-Bottom Stability Design of Submarine Pipelines, 2010.

[13] Boyun Guo, Shanhong Song, Jacob Chacko, and Ali Ghalambor. Offshore Pipelines.Gulf Professional Publishing, Burlington, USA, 2005.

[14] J. R. Hale, W. F. Lammert, and D. W. Allen. Pipeline On-Bottom Stability Calcu-lations: Comparisons of Two State-of-the-Art Methods and Pipe-Soil Model Verifi-cation. Offshore Technology Conference, 1991.

[15] J. R. Hale, W. F. Lammert, and V. Jacobsen. Improved Basis for Static StabilityAnalysis and Design of Marine Pipelines. Offshore Technology Conference, 1989.

[16] Xiting Han and Fuping Gao. A Pipe-Soil Interaction Model for Anti-Rolling PipelineOn-Bottom Stability on a Sloping Sandy Seabed. International Society of Offshoreand Polar Engineers, 2013.

71

REFERENCES

[17] International Standardization Organization ISO. About ISO, 2014.

[18] ISO-13623. Petroleum and Natural Gas Industries - Pipeline Transport System, 2009.

[19] Eric Jas, Dermot O’Brien, Roland Fricke, Alan Gillen, Liang Prof. Cheng, DavidProf. White, and Andrew Prof. Palmer. Pipeline Stability Revisited. The Journal ofPipeline Engineering, (4th Quarter):11, 2012.

[20] D. S. Jeng, A. Ismail, L. L. Zhang, and J. S. Zhang. Empirical model for the predictionof lateral stability coefficient for un-trenched submarine pipes based on selv-evolvingneural network (SEANN). Ocean Engineering, 72:167–175, 2013.

[21] Prof. Daniel Karunakaran. Pipelines and Risers, University of Stavanger, 2013.

[22] Andrew C. Palmer and Roger A. King. Subsea Pipeline Engineering. PennWellCorporation, Oklahoma, 2008.

[23] Dr. Qiang. On-bottom Stability Design of Subsea Pipelines, Lecture Notes, Univer-sity of Stavanger, 2013.

[24] Jay Ryan, Dean Campbell, David White, and Eric Jas. A Fluid-Pipe-Soil Approachto Stability Desing of Submarine Pipelines, 2011.

[25] Subsea7. Environmental, Pipe and Soil Data, publisher =.

[26] Yinghui Tian, Mark J. Cassidy, and Bassem S. Youssef. Consideration for On-BottomStability of Unburied Pipelines using a Dynamic Fluid-Structure-Soil Simulation Pro-gram. The International Journal of Offshore and Polar Engineering, 21(4):8, 2011.

[27] Knut Tørnes, Gary Cumming, Hammam Zeitoun, and John Willcocks. A StabilityDesign Rationale - A Review of Present Design Approaches. page 13. Proceedings ofthe ASME 28th International Conference on Ocean, Offshore and Arctic Engineering,2009.

[28] Su Young Yu, Han Suk Choi, Seung Keon Lee, Chang Ho Do, and Do Kyun Kim.An Optimum Design of On-Bottom Stability of Offshore Pipelines on Soft Clay.International Journal of Naval Architecture and Ocean Engineering, 5:598–613, 2013.


APPENDICES

Appendix A: Manual Calculation

Stability of pipeline should be checked both for the installation and operational phases.The main analysis is done separately using the SESAM:StableLines software. Here, asample manual calculation is shown. It should be noted that this sample calculation isprepared to show the analysis procedures and illustrate on how the software works. Theresults calculated in this sample do not exactly match the results from the software. Itshould also be noted that, this sample calculation is done for a pipeline in an operationalphase and resting a sandy seabed. The data used is a real data for a North Sea projectfrom Subsea 7.

In the design process the Absolute and Generalized Stability Analyses methods arechecked for the operational phase of the pipeline.

Operational Phase

Lateral Stability

Absolute Lateral Stability

The environmental, seabed, pipeline and other relevant data are taken from Tables: 6.1,6.2 and 6.3 in Section 6.2.

The overall hydrodynamic diameter is given by:

D = 610+(2*5)+(2*55)=730 mm

From Figure 4.6, the wave parameter, Tn can be found as:

Tn =

√d

g=

√104

9.81= 3.256 sec. &

TnTp

=3.256

15.9= 0.205

From equation 4.4, φ can be calculated from which the peakedness is found:

φ =Tp√Hs

=15.9√14.8

= 4.133

Hence,

73

APPENDIX A

γ = exp(5.75− 1.15φ) = exp(5.75− 1.15 ∗ 4.133) = 2.710

Then, from Figure 4.6, by interpolating the peakedness factor, Us can be found as:

Us ∗ TnHs

= 0.18, Us = 0.818 m/s

And the ratioTuTp

can be found from Figure 4.7, and Tu is found accordingly.

TuTp

= 1.04 , Tu = 15.9 * 1.03 = 16.377 sec.

AsTnTu

=3.256

16.377= 0.199 ≤ 0.2, then equation 4.15 is used to calculate T ∗;

T ∗ =

((kt − 5.(kt − 1).(

TnTu

)

)Tu

The value of kt can also be interpolated from equation 4.16.

kt = 1.22

T ∗ = ((1.22− 5.(1.22− 1).(3.256

16.377) ∗ 16.377,

T ∗ = 16.398 secThe number of waves in the storm is calculated from equation 4.12 ;

τ =TstormTu

τ =TstormTu

=3 ∗ 3600

16.536= 659

The oscillatory velocity amplitude for single design oscillation, U∗ is then calculated fromequation 4.14 ;

U∗ = 0.5.

(√2 ln τ +

0.5772√2 ln τ

).Us.RD

where, spectral spreading exponent, s = 8 (6-8 in the North Sea) and taking the waveand current directions to be perpendicular (900), then RD = 0.95 in Figure 4.9.

U∗ = 1.462 m/s

The mean perpendicular current over the pipe diameter is found from equation 4-19

Uc(z0) = Ur(zr).

(1 +z0D

). ln (D

z0+ 1)− 1

ln (zrz0

+ 1)

. sin θc

Uc = V ∗ = 0.574 m/s


APPENDIX A

The peak load coefficients are now calculated by combining equations 5.3 and 5.4 andTables 5.1 and 5.2.

K∗ = 32.84 and M∗ = 0.393

Hence from the graph in Figure 5.1, the peak load coefficients are;

C∗Y = 2.529 and C∗Z = 2.377

Load and Weight Calculations

The load reduction factors due to permeable seabed and penetration can be calculated asare explained in section 4.3.3 as:Load reduction in lift due to permeable seabed:

rperm,z = 0.7

Load reduction due to trench profile from equations 4.26 and 4.27rtrench,y =1.0rtrench,z= 1.0

Load reduction due to penetration from equations 4.24 and 4.25 , assuming, zp = 0mm:

rpen,y = 1.0 ≥ 0.3

rpen,z = 1.13 ≥ 0.3

Hence, the total reduction due to penetration, trenching and permeability as given byequations 4.28 and 4.29 is;

rtot,y = rpen,y ∗ rtrench,yrtot,y = 1.0 * 1.0 = 1.0

rtot,z = rperm,z * rpen,z ∗ rtrench,z

rtot,z = 0.7 * 1.13 * 1 = 0.791

Now that all the peak load coefficients and reduction factors are found, the peak loadscan be calculated from equations 5.1 and 5.2 as:

The Peak Horizontal Load is;

F ∗Y = rtot,y.1

2.ρw.D.C

∗Y (U∗ + V ∗)2

F ∗Y = 1.0 ∗ 1

2∗ (1025) ∗ 0.73 ∗ 2.529 ∗ (1.462 + 0.574)2


APPENDIX A

F ∗Y = 3922.12 N/m (total inertia and drag forces)

The Peak Vertical Load is;

F ∗Z = rtot,z ∗1

2∗ ρw ∗D ∗ C∗Z(U∗ + V ∗)2

F ∗Z = 0.791 ∗ 1

2∗ (1025) ∗ (0.73) ∗ 2.377 ∗ (1.543 + 0.574)2

F ∗Z = 2915.94 N/m (total uplift force)

Passive Soil Resistance

For operational pipe, Ws = 2.426 kN/m

FC = Ws - F ∗Z

FC = 2.426 kN/m - 2.916kN/m

FC = -0.49 kN/m = -490 N/m

FC is less than zero, this means that the lift force is larger than the submerged weight.

The passive soil resistance can be calculated from equation 4.31 and 4.32 :

κs =(γ′ ∗D2)

FC=

(7000 ∗ 0.732)

905= 4.12 ≤ 26.7

Hence,

FRFC

= (5.0 ∗ κ2s) ∗ (zpD

)1.25 = (5.0 ∗ 4.122) ∗ (0

0.73)1.25 = 0

The passive resistance, FR is set to zero.

Stability Criteria

Absolute Lateral StabilityAs explained in Section 5.2, the pipeline is considered to satisfy the absolute stabilityrequirement if:

γsc.F ∗Y + µ.F ∗Zµ.Ws + FR

6 1.0 and γsc.F ∗ZWs

6 1.0

γsc = 1.32 for winter storms in the North Sea from Table 5.3.

1.32 ∗ (3922.12 + 0.6 ∗ 2915.94)

(0.6 ∗ 2426 + 0)6 1.0 and 1.32 ∗ 2915.94

24266 1.0

5.14 � 1.0 and 1.21 � 1.0


APPENDIX A

Both the equations do not satisfy the Absolute Lateral Stability criteria. Hence, the hor-izontal stability criterion of the pipeline’s is violated.

Generalized Lateral Stability

The analysis should further be assessed using the Generalized Lateral Stability allowinga displacement of 0.5 times the diameter (0.5D).

The design criteria for the Generalized Stability is dependent on the specific weight of thepipeline, i.e., 1.05 <Sg ≤ 3.

First N, K and L are calculated from equation 5.9 ;

N =Us ∗RD

g ∗ Tu, K =

Us ∗RD ∗ TuD

, L =Ws

0.5 ∗ ρw ∗D ∗ U2s

N =0.818 ∗ 0.95

9.81 ∗ 16.377, K =

0.818 ∗ 0.95 ∗ 16.377

0.73, L =

2426

0.5 ∗ 1025 ∗ 0.73 ∗ 0.8182

N = 0.0048, K = 17.434, L = 9.691

Therefore,

Sg = 1 +2

π∗ (N ∗K ∗ L), 1.05<Sg ≤ 3

Sg = 1 +2

π∗ (0.0048 ∗ 17.434 ∗ 9.691)

Sg = 1.52

Since the Generalized Stability Analysis is valid for 1.05 <Sg ≤ 3, the pipeline is stable.

The analysis is stopped here. If it were not stable, it would have further been checkedusing the 10 D criteria.

Vertical Stability

The criteria in equation 7.2 should be fulfilled for the pipeline to be stable vertically.

γw ∗b

(Ws + b)=γwSg≤ 1.00

The safety factor, γw = 1.1 can be applied if a sufficiently low probability of negativebuoyancy is not documented. [12]

Therefore,


APPENDIX A

1.1

1.52= 0.72 ≤ 1.0

The pipeline fulfills the vertical stability criteria.


APPENDIX B

Appendix B: SESAM: StableLines Output (For Instal-

lation and Operation Conditions

Design Data

Table 1: Pipe and soil data. [25]

Table 2: Material densities. [25]

Table 3: Environmental data. [25]


APPENDIX B

Lateral Stability

Absolute Lateral Stability


AP

PE

ND

IXB

Table 4: StableLines: Absolute Lateral Stability main page

On

-Bottom

Stability

An

alysisof

Su

bseaP

ipelines

Accordin

gto

DN

V-R

P-F

10981

APPENDIX B

Table 5: SESAM:StableLines Output: Necessary weight vs. Direction

Table 6: SESAM:StableLines Output: Concrete thickness vs. density


APPENDIX B

Figure 1: SESAM:StableLines Output: Necessary weight vs. direction presentation

Figure 2: SESAM:StableLines Output: Concrete thickness vs. density presentation


APPENDIX B

Table 7: SESAM:StableLines Output: Necessary weight vs. concrete thickness

Figure 3: SESAM:StableLines Output: Necessary weight vs concrete thickness presenta-tion


APPENDIX B

Table 8: SESAM:StableLines Output: Necessary weight vs. pipe diameter

Figure 4: SESAM:StableLines Output: Necessary weight vs pipe diameter presentation


APPENDIX B

Table 9: SESAM:StableLines Output: Necessary weight vs. water depth

Figure 5: SESAM:StableLines Output: Necessary weight vs water depth presentation


APPENDIX B

Table 10: SESAM:StableLines Output: Necessary weight vs. current velocity


APPENDIX B

Figure 6: SESAM:StableLines Output: Necessary weight vs current velocity presentation

Figure 7: SESAM:StableLines Output: L vs. K for three values of M presentation


APPENDIX B

Table 11: SESAM:StableLines Output: L vs. K for three values of M


APPENDIX B

Table 12: SESAM:StableLines Output: Iteration (Empty condition)


APPENDIX B

Table 13: SESAM:StableLines Output: Iteration (Operational condition)

Table 14: SESAM:StableLines Output: Results: All RPV combinaions


APPENDIX B

Table 15: SESAM:StableLines Output: General output


APPENDIX B

Generalized Lateral Stability (0.5D)


AP

PE

ND

IXB

Table 16: StableLines: Generalized Stability, Displacement=0.5D main page

94O

n-B

ottomS

tabilityA

nalysis

ofS

ubsea

Pipelin

esA

ccording

toD

NV

-RP

-F109

APPENDIX B




APPENDIX B





APPENDIX B

Figure 10: SESAM:StableLines Output: Necessary weight vs concrete thickness presen-tation




APPENDIX B





APPENDIX B




APPENDIX B




APPENDIX B



APPENDIX B

Generalized Lateral Stability (10D)


AP

PE

ND

IXB

Table 26: StableLines: Generalized Stability, Displacement=10D main page

On

-Bottom

Stability

An

alysisof

Su

bseaP

ipelines

Accordin

gto

DN

V-R

P-F

109103

APPENDIX B




APPENDIX B





APPENDIX B

Figure 17: SESAM:StableLines Output: Necessary weight vs concrete thickness presen-tation




APPENDIX B




APPENDIX B




APPENDIX B




APPENDIX B



APPENDIX B



APPENDIX B

Table 36: SESAM:StableLines Output: Concrete thickness vs. water depth, AbsoluteStability

Figure 22: SESAM:StableLines Output: Concrete thickness vs. water depth presentation,Absolute Stability


APPENDIX B

Table 37: SESAM: StableLines Output: Concrete thickness vs. water depth, GeneralizedStability, 0.5D

Figure 23: SESAM:StableLines Output: Concrete thickness vs. water depth presentation,Generalized Stability, 0.5D


APPENDIX B

Table 38: SESAM: StableLines Output: Concrete thickness vs. water depth, GeneralizedStability, 10D

Figure 24: SESAM:StableLines Output: Concrete thickness vs. water depth presentation,Generalized Stability, 10D


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FACULTY OF SCIENCE AND TECHNOLOGY MASTER'S THESIS