On-Bottom Stability Analysis of Submarine Pipelines, Umbilicals and Cables Using 3D Dynamic Modelling

Transcription

1 OTC MS On-Bottom Stability Analysis of Submarine Pipelines, Umbilicals and Cables Using 3D Dynamic Modelling Bassem Youssef and Dermot O'Brien, Atteris Pty Ltd Copyright 2017, Offshore Technology Conference This paper was prepared for presentation at the Offshore Technology Conference held in Houston, Texas, USA, 1 4 May This paper was selected for presentation by an OTC program committee following review of information contained in an abstract submitted by the author(s). Contents of the paper have not been reviewed by the Offshore Technology Conference and are subject to correction by the author(s). The material does not necessarily reflect any position of the Offshore Technology Conference, its officers, or members. Electronic reproduction, distribution, or storage of any part of this paper without the written consent of the Offshore Technology Conference is prohibited. Permission to reproduce in print is restricted to an abstract of not more than 300 words; illustrations may not be copied. The abstract must contain conspicuous acknowledgment of OTC copyright. Abstract Submarine pipelines and umbilicals are essential elements of many offshore hydrocarbon developments. In general, the most economical approach is to lay the pipelines and umbilicals directly on the seafloor with adequate self-weight to avoid any requirement to perform secondary stabilisation work such as trenching. A fundamental design requirement for this scenario is to ensure adequate pipeline stability under extreme environmental loading conditions, as excessive lateral displacements may result in pipeline or umbilical damage. The commonly used and widely accepted recommended practice DNV-RP-F109 (Ref. 3) provides three design approaches with increasing levels of complexity to perform on-bottom stability design. These three design approaches are: Absolute Lateral Static Stability, Generalised Lateral Stability and Dynamic Lateral Stability Analysis. The first method is based on two dimensional (2D) force balance equilibrium equations while the second method is based on extensive dynamic modelling results of 2D pipeline models. These first two methods are intended to provide conservative on-bottom stability designs for rigid pipelines. The third method involves detailed finite element modelling of the pipeline under a time domain hydrodynamic loading. This paper highlights the DNV-RP-F109 recommended practice limitations with respect to the on-bottom stability design of flexible pipelines, umbilicals and cables. The paper presents a dynamic simulation methodology, in accordance with the recommended practice DNV-RP-F109, for on-bottom stability design of rigid pipelines, flexible pipelines, umbilicals and cables, together with specific sensitivity analyses and comparisons with the Absolute Stability method. The dynamic modelling results and comparisons presented in this paper highlight some limitations of the Absolute and Generalised methods for flexible pipelines, umbilicals and cables and illustrate the benefits of using the three dimensional (3D) dynamic stability analyses. The results of the sensitivity analyses clearly identify important parameters for the dynamic simulation, such as the hydrodynamic load correction due to the pipeline movements, the pipeline axial stiffness and the seabed passive soil resistance, that significantly affect the on-bottom stability behaviour. Refining the identified parameters input values will lead to more accurate and cost effective on-bottom stability design.

2 2 OTC MS Introduction Pipeline on-bottom stability is a complicated interaction between the hydrodynamic loads, the pipeline structure and the supporting soil. The pipeline length can vary from a few hundred meters to hundreds of kilometers. For a long pipeline route it is more likely that one or more of the water depth, the pipeline heading, the metocean conditions or the soil conditions will vary. A 2D model will not take into account the fact that the hydrodynamic loads and soil resistance are shared along the pipeline route due to the effect of the pipeline stiffness. For the Soliton current case, which is usually high in velocity, short in period and localized in distance, the 2D on-bottom stability methods of the recommended practice DNV-RP-F109 would result in a too conservative pipeline weight requirement. In contrast with a 3D model, high hydrodynamic load events at any particular location along the pipeline route will be distributed and shared over a length of the pipeline, leading to a more realistic prediction of pipeline lateral movement. The Generalised Lateral Stability method of the recommended practice DNV-RP-F109 is calibrated for rigid pipeline regardless of the pipeline axial stiffness and bending stiffness. Therefore, the effect of changing the bending stiffness and axial stiffness values, as applicable for flexible pipe or umbilical, is not accounted for. Other important parameters such as the operational internal pressure and temperature are not considered in the 2D on-bottom stability methods calibration. As a result, the 2D stability methods may potentially under estimate the pipeline displacements. An important factor in the on-bottom stability analysis is that the hydrodynamic loads are generated for a pipeline sitting on the seabed in stationary position. The generated hydrodynamic loads are much higher than the hydrodynamic loads for a pipeline partially buried in the seabed or for a pipeline that experiences a lateral velocity. For accurate and economical on-bottom stability design, in the 3D dynamic modelling the hydrodynamic loads are corrected during the simulation to account for the pipeline lateral and vertical movements. For most of the pipeline cases, the pipeline route includes multiple changes in the heading direction and sometimes structure crossings. These changes in the pipeline route will affect the hydrodynamic loading angles and may change the behaviour of the pipeline-hydrodynamic loads interaction. For some cases different pipeline route headings could provide more stability for the pipeline if the hydrodynamic loads act parallel to the pipeline heading. An additional benefit of the 3D dynamic stability analysis is that accurate calculation of the pipeline tension load, bending moments and tie-in loads at the tie-in connection can be achieved. Considering 3D modelling, the pipeline designer can include the tie-in connection details, the actual pipe route heading, the pipe bending stiffness and axial stiffness and any bending restrictors in the model. The stability design in the vicinity of tie-in points can then be optimized compared to what can be achieved if 2D stability methods are relied upon. For small diameter pipelines, less than 10 cm for example, or for cables; a significant limitation in the current recommend practice is that the wave velocity correction due to the seabed boundary layer effect is not accounted for. The specific gravity calculated by the Absolute stability method of the recommended practice DNV-RP-F109 to achieve the on-bottom stability requirements is un-realistic. For most of the seabed materials, the boundary layer height would be in a range of few centimeters. The effect of a few centimeters boundary layer height on the correction of the wave velocity on a small diameter pipeline or cable can be significant. Using the 3D dynamic modelling the pipeline designer can overcome the limitations of the 2D stability methods of the recommended practice DNV-RP-F109 and achieve more reliable and economic on-bottom stability designs (see Refs. 11 and 12 for 3D dynamic modelling discussions and examples).

3 OTC MS 3 DNV-RP-F109 On-bottom stability design methods Absolute stability method The Absolute Stability method is calibrated to provide the minimum pipe submerged weight such that no lateral displacement occurs during the design storm return period. The method uses the maximum wave velocity expected during the entire storm period. Therefore the pipeline submerged weight estimated tends to be very conservative. Remarks and limitations of the method: The method is based on the extreme wave velocity which may occur for a few seconds during the three hours storm period at one particular location along the entire pipeline route. The method does not account for the pipe operational temperature and pressure. Generalised stability method The generalized stability method is based on PONDUS dynamic stability simulations of 2D pipeline models. PONDUS is a computer program that computes the dynamic lateral response of offshore pipelines subjected to wave and current action on a horizontal seabed (Ref. 3). The method is calibrated considering the significant wave velocity during the storm to provide the pipeline submerged weights corresponding to pipeline lateral displacement between 0.5 outer diameter (D) and 10 D under the design storm return period loading conditions. Remarks and limitations of the method: This method is based on 2D section of the pipeline, ignoring the benefits of the 3D effect on the pipeline bending and axial stiffness and soil resistance on the pipeline stability. Unlike the Absolute stability method, this method uses the significant wave velocity to estimate the pipe submerged weight values corresponding to pipeline lateral displacement of 0.5 OD and 10 OD. Similar to the Absolute stability method, the method does not account for the pipe operational temperature and pressure. The method is applicable for silica sand soil and clay soil only and cannot be used for other soil types, such as calcareous or carbonate sand soil. Dynamic stability method The dynamic stability method is considered the most sophisticated stability method as it requires a numerical modelling tool and a level of expertise to perform the modelling and interpret the results. Using the numerical modelling, the pipeline designer can overcome the limitations of the 2D methods mentioned above. Benefits of using the 3D dynamic stability modelling have been discussed extensively in the literature (see for example, Youssef et al., Ref. 10). As recommend by DNV-RP-F109, the following considerations are required for the dynamic stability modelling: Full sea-state time series using the wave spectrum should be utilized in the model; if no information is available regarding the storm duration, a typical storm period of three hours should be used. Storm hydrodynamic loads of irregular wave on the pipeline should be calculated using advanced hydrodynamic force model that accounts for the wave wake effect. During the simulation, the hydrodynamic loads should be corrected to account for the pipeline displacement and penetration experienced during the simulation time history. Soil resistance should include two parts, a pure friction term and a passive resistance term, accounting for the pipe penetration in the soil and build-up of the soil berm.

4 4 OTC MS Full pipeline length should be modelled including the actual route and the actual pipeline ends boundary conditions. For long pipeline routes, the middle section of the pipeline could be modelled with special considerations for the model boundary conditions. The operating temperature and pressure should be included in the model. Remarks and limitations of the method: The actual pipeline properties conditions can be modelled including axial stiffness and bending stiffness, nonlinear cross section properties, any bending stiffeners, any change in the pipe cross section or in the pipe weight along the pipeline route. For small diameter pipelines or cables, DNV-RP-F109 does not provide any considerations regarding the wave velocity and hydrodynamic load corrections. Therefore, the results of the 2D stability methods of the recommended practice DNV-RP-F109 and the dynamic stability modelling are usually unrealistic for small diameter pipelines or cables if the wave velocity and the hydrodynamic loads are not corrected. More discussion regarding this point will be presented in numerical modelling section. Figure 1 presents a diagrammatic sketch of the hydrodynamic-pipeline-soil model elements. Figure 1 a- Diagrammatic Sketch of Hydrodynamic-Pipeline-Soil Model, b- Force Balance Diagram on the Pipeline Section Dynamic stability modelling consideration The on-bottom stability simulations presented in this paper are performed using the CORUS-3D software. The software code is written in FORTRAN programming language and introduced to the finite element software ABAQUS through the DLOAD subroutine to perform the hydrodynamic-pipe-soil interaction calculations. Hydrodynamic loads The Morison equation is an industry recognized and simply established method used to estimate the hydrodynamic loads around a cylinder. When applied to a pipeline on the seabed, lift (F L ), drag (F D ) and inertia (F M ) loads under combined wave and steady current can be calculated as: (1) (2) (3)

5 OTC MS 5 where ρ is the density of the fluid, U w the wave velocity, U c the current velocity, the wave acceleration, D the pipe diameter and C L, C D and C M the lift, drag and inertia coefficients, respectively. The value of the Morison equation coefficients are reported in many references (for example, Refs. 7, 8 and 9) as a function of steady current to oscillatory wave velocity ratio, pipe roughness and Keulegan- Carpenter number (KC) which takes the form (KC=U w T/D). However, the Morison equation yields poor load prediction especially for the lift load component and for the irregular wave conditions (DNV-RP-F109, Ref. 3). More advanced hydrodynamic load model, Fourier load model, was presented by Sorenson et al. (Ref. 8) and based on intensive full-scale laboratory tests. This model takes into account the wave wake effect and the model is applicable to both regular and irregular waves, a wave only or wave and current. The Fourier transformation method has been used to fit the non-dimensional hydrodynamic loads and to estimate the Fourier coefficients. This hydrodynamic load model is recommended by DNV-RP-F109 (Ref. 3) to estimate the hydrodynamic loads on the pipeline. Moreover, the Fourier load model is implemented in one of the widely accepted and used commercial pipeline simulation package of AGA Level 3 software (Ref. 1). The basis for the Fourier method is that any quantity that has a periodic variation with a certain period T can be reproduced by superposition of a number of sine waves with periods equal to T and smaller, so that the i th sinusoidal wave or harmonic has a period T i = T/i, where i = 1, 2, 3,, N. The general expression of the periodic quantity F(t) is: (4) where a o,a i and b i are the Fourier coefficients derived from experimental measured loads and the wave angular frequency. In order to obtain the non-dimensional load coefficient from the Fourier analysis, the measured physical experimental loads have been normalized by. Before that, for the horizontal load case, the drag load was calculated by subtracting the inertia component from the measured horizontal load: where C a is the added mass coefficient with a value of 2.29 as determined from potential flow theory (Refs. 1 and 8). The decomposition of the non-dimensional horizontal (C H ) and vertical (C V ) loads have the expressions: is (5) (6) The total in-line load is found by adding the inertia term. where C M = C a + 1 and takes the value of 3.29 (Refs. 5 and 8). For full details of the Fourier hydrodynamic load model and the hydrodynamic loads coefficients, reference should be made to Sorenson et al. (Ref. 8) and AGA (Ref. 1). (7) (8)

6 6 OTC MS Hydrodynamic load corrections. The hydrodynamic load calculations presented in the previous section assume a pipe fixed in position with its apex just touching the soil surface. In real field scenario, the pipeline on the seabed may move both horizontally and vertically under the effect of the applied hydrodynamic loads and pipe self-weight load. Movements of the pipeline will change the relative velocity of the pipeline to the flowing water. Pipeline vertical penetration also reduces the exposure of the pipeline and therefore the hydrodynamic loads acting on it. Moreover, if the pipeline is elevated from the seabed the hydrodynamic loads will also change. These effects must be taken into account to accurately predict the hydrodynamic loads on the pipeline. Pipeline horizontal movement The hydrodynamic loads acting on the pipeline are calculated as a function of the water particle velocity and acceleration. For the case of a horizontally moving pipeline, the relative velocity and acceleration between the pipeline and flowing water will change and consequently affect the hydrodynamic load. As explained by Sorenson et al. (Ref. 8), the hydrodynamic loads on moving pipeline can be estimated by adding correction terms to the pre-estimated hydrodynamic loads on a pipeline fixed in position. (9) (10) where is the pipe acceleration, U e the effective near pipe water velocity and C D,Corr and C L,Corr the drag and lift load correction coefficient, respectively. Values C D,Corr and C L,Corr are provided in AGA (Ref. 1). The U e value can be calculated using the form: (11) (12) Pipeline Penetration A pipeline may become partially buried during the pipe laying process and can further penetrate into the seabed due to pipeline dynamic cyclic movements under small oscillatory wave action. With penetration, the pipeline becomes less exposed to the flowing water, which leads to a reduction in the hydrodynamic loads. DNV-RP-F109 (Ref. 3) recommended reducing the lateral and vertical loads as per the following reduction factors equations: (13) where R lateral is the reduction factor for lateral loads, R vertical the reduction factor for vertical load and Z P is the pipe vertical penetration. Wave boundary layer effect The seabed material roughness influences the wave velocity at the vicinity of the seabed level. The wave boundary layer height defined as the minimum distance from the seabed to a point where the wave velocity in the x-direction equals the free-stream wave velocity amplitude (Ref. 6). In general, for pipeline on-bottom stability applications the wave boundary layer will only be a few centimeters in height. The recommended (14)

7 OTC MS 7 practice DNV-RP-F109 recommends ignoring the effect of wave boundary layer on the wave velocity. This is justified for large diameter pipelines where the influence of the boundary layer on the wave velocity over the pipeline diameter is minor. However, for small pipelines or cables of 10 centimeter diameter or less, the effect of the wave boundary layer can be significant. Detailed modelling of the wave boundary layer including the eddy viscosity models, one-equation models (k-model) and two equations models is discussed and presented in Fredsoe and Deigaard (Ref. 4). A schematic sketch of the wave boundary layer is presented in Figure 2-a. The wave velocity profiles in the boundary layer at different phases predicted by eddy viscosity models as discussed by Fredsoe and Deigaard (Ref. 4) is presented in Figure 2-b with the full line presents a steady eddy viscosity and the dashed line presents a time-varying eddy viscosity. Figure 2 a- Wave Boundary Layer, b- Wave Velocity Profile within the Boundary Layer at Different Phase (Ref. 4) The wave boundary layer thickness for a pipeline on smooth or rough seabed soils can be estimated based on the methods introduced by Fredsoe and Deigaard (Ref. 4). For smooth seabed soil, the height of wave boundary layer (δ) can be approximated as: where Re = U w a/v is the amplitude Reynolds number defined by water particle excursion distance a, wave orbital velocity U w and the kinematic viscosity of water ν. The water particle excursion distance a is defined as a = U w T/2π where T is the wave period. On a rough seabed soil, δ is related to surface roughness and can be estimated as: (15) (16) where k N is the Nikuradse roughness height and is dependent on seabed conditions. Under hydrodynamically rough flow conditions, the Nikuradse roughness height can be estimated as k N = 30z 0 where z 0 is the seabed roughness height. The average wave velocity profile in a wave boundary layer considering a logarithmic form is presented in Cheng et al. (Ref. 2). Considering a simplified linear profile of the wave velocity within the boundary layer height, the average wave velocity over the pipe diameter can be calculated as follows: (17)

8 8 OTC MS (18) Soil Models One of the most important parts of the on-bottom stability dynamic model is to model the full soil resistance which consists of a pure Coulomb friction part and a passive resistance part. The Coulomb friction value provides the lateral soil resistance capacity as a ratio of the pipe vertical weight. DNV-RP- F109 recommended for a concrete coated pipe a Coulomb friction value of 0.6 for sand or rock soils and a Coulomb friction value of 0.2 for clay soil. The passive soil resistance part accounts for the soil resistance capacity due to pipeline penetration in the soil. Considering a pure Coulomb friction only and ignoring the passive soil resistance term in the on-bottom stability modelling may lead to larger predicted pipeline horizontal displacements. The modified Coulomb friction model or the tri-linear Coulomb friction model (see Figure 3) accounts for the passive soil resistance and requires much less efforts to be implemented in the finite element model. The model simply provides modified Coulomb friction values equivalent to combined pure Coulomb resistance and passive soil resistance at different pipe lateral displacement levels. However, the tri-linear model requires geotechnical laboratory testing of the soil sample to provide the tri-linear Coulomb friction model parameters. Figure 3 Diagrammatic Sketch of Pipe-Soil Interaction Models Used in this Paper The Verley and Sotberg (Ref. 14) soil resistance model is calibrated to model the passive resistance term of silica sand soil. DNV-RP-F109 recommended the use of Verley and Sotberg soil model to predict the passive soil resistance of the silica sand soil. Full details of the model can be found in Verley and Sotberg (Ref 14) and DNV-RP-F109 (Ref. 3). Similar passive soil resistance models for clay soil and calcareous soil are presented in Verley and Lund (Ref. 15) and Youssef and Cassidy (Ref. 13), respectively.

9 OTC MS 9 Flowchart of the CORUS-3D code and the main procedure of the software are presented in Figure 4. Figure 4 Flowchart and Procedure of the Dynamic Stability Software Dynamic stability simulations The effects of the various parameters discussed above have been evaluated through performing a series of 3D dynamic stability simulations with various sensitivity cases to gauge the effect of each parameter on the predicted behaviour of a sample pipeline. Details of the performed analyses inputs including the sensitivity cases details are presented in Table 1. Table 1 Dynamic Simulations Input Data and Details Item Base Case Sensitivity Cases* Case 1 Case 2 Case 3 Case 4 Case 5 Case 6 Case 7 Outer Diameter 220 mm 25.4 mm Wall thickness 11.5 mm 12.7 mm Submerged weight 200 kn/m 11 kn/m Specific gravity Significant Wave height Peak period Steady current Water depth 7.5 m 10 sec 0.5 m/s 50 m

10 10 OTC MS Item Wave spreading factor 2 Base Case Sensitivity Cases* Case 1 Case 2 Case 3 Case 4 Case 5 Case 6 Case 7 Hydrodynamic load correction Yes No Soil Model Pure Coulomb friction Various Pressure/ Temperature No Yes Ends Boundary Conditions Axially fixed Laterally free Axial and lateral fixity Routing Straight Dogleg Bending and axial rigidity (EI and EA) 8.2E6 N.m 2 1.5E9 N 1E4 to 1E8 1E7 to 1E10 1E2 2.2E5 Wave Boundary layer No Yes * The blank cells values are as per the Base Case values Base Case The base case of the dynamic stability simulation examples is a 2,500 m long rigid pipeline in a 50 m water depth. The pipeline is modelled using 500 hybrid linear pipe element (PIPE31H) each 5 m length. The simulation is performed for a 30 minute storm loading and considered the hydrodynamic load correction due to the pipeline movements. In this example, the soil type considered is silica soil and the soil resistance considered is a pure Coulomb friction with a value of 0.6. The contribution of the passive soil resistance has not been included in this example and will be presented in the following examples. The pipeline ends boundary conditions are set to prevent the axial pipeline displacement and pipeline rolling. This example assumes the 2,500 m section of the pipeline is part of a longer pipeline route. The wave velocity and hydrodynamic loads along the pipeline route are generated using the AGA (Ref. 1) software assuming the wave heading direction is perpendicular to the pipeline route and considering the pipeline initial location, i.e. before any pipeline lateral or vertical displacements. The sea surface was generated using the JONSWAP wave spectrum. It should be noted that the wave velocity and hydrodynamic loads at each pipeline node are different due to the wave spreading and the random numbers used during generating the irregular wave. As recommend by DNV-RP-F109 to account for the random seed effect on the pipeline lateral displacement, at least seven analyses with randomly chosen seeds should be performed. For this example, ten pipeline simulation cases are performed using ten full generated sea-states with randomly selected seeds. The final pipeline lateral displacements of the ten simulation cases are presented in Figure 5. As shown in the Figure, the average, minimum and maximum displacement values varies between the ten simulation cases because of the change of the random seeds and random numbers used in generating the irregular sea-states. The average lateral displacement for the ten simulation cases varies between m and m with an average value across all ten simulation cases of m. Figure 5 Pipeline Final Displacement Ten Simulation Cases

11 OTC MS 11 Detailed time history of the lateral displacements at three points along the pipeline route, at 500 m, 1250 m and 2000 m, for the Seed 3 simulation case is presented in Figure 6. The detailed displacement time history presented in Figure 6 illustrates how the lateral displacement of different points along the pipeline route varies with time as the hydrodynamic load changes. It is also shown in the graph that the lateral displacement of pipe node 251, at 1250 m, increased from about 8.20 m to about m within a few seconds around the 1550 s time period. To investigate this displacement further, the applied wave velocity and hydrodynamic loads are presented in Figure 7 to Figure 10 in blue lines. The actual pipe velocity and the calculated hydrodynamic load corrections are also shown in Figure 7 to Figure 10 in red lines. Figure 6 Pipeline Displacement Time History Seed 3 Figure 7 Wave Velocity Time History Figure 8 Drag Load Time History Figure 9 Inertia Load Time History

12 12 OTC MS Figure 10 Lift Load Time History As shown in the Figure 7, a peak wave velocity of about 1.1 m/s is formed around the 1550 s time period, which is almost double the value of the second highest wave velocity. Corresponding to this high wave velocity value, high drag, inertia and lift loads are generated. During the dynamic simulation at the buildup of the wave velocity, the pipeline displaces laterally causing the pipe to gain velocity as shown by the red line in Figure 7 and as a result of this pipeline velocity, hydrodynamic load corrections are calculated and applied during the simulation (see the red lines in Figures 8 to 10). Correcting the pre-generated hydrodynamic loads by applying the hydrodynamic load corrections prevents the pipeline from experiencing larger unrealistic lateral displacement. It should be noted that every single pipe node during the dynamic simulation will experience different time history of wave velocity, pipe velocity and hydrodynamic load corrections. Sensitivity Simulation Cases The base case simulation considering Seed 3 hydrodynamic loads is selected for the comparison purpose in this section and to measure the sensitivity of the 3D simulation to the model input parameters. Case 1 Hydrodynamic load correction. To highlight the effect of the hydrodynamic load correction on the final pipeline displacement, the base case example considering Seed 3 hydrodynamic loads is repeated, deactivating the hydrodynamic load correction due to pipeline displacement. Figure 11 shows the final lateral displacement results with and without the hydrodynamic load correction. Figure 11 Pipeline Final Displacement W/WO Load Correction As shown in Figure 11 the final pipeline displacement for the Seed 3 simulation without hydrodynamic load correction due to the pipeline displacement is almost four times the final pipeline displacement for the same case with hydrodynamic load correction considered. The time history of node number 251 lateral displacement is shown in Figure 12. It is clear from Figure 12 that the points in time where the effect of not considering hydrodynamic load correction are most significant coincide with the points in time where the pipe velocity is significant in the base case, i.e. around the times 160, 420, 650, 1080 and 1550 s (see Figures 7 to 10).

13 OTC MS 13 Figure 12 Time History of Displacement W/WO Load Correction Case 2 Soil resistance modelling. The soil resistance in the base case simulations is modelled considering pure Coulomb friction soil resistance term only appropriate for silica sand soil. To highlight the effect of the soil resistance modelling on the on-bottom stability simulation results, the base case example considering Seed 3 hydrodynamic loads is repeated three times to simulate the following soil resistance conditions: silica sand soil using Coulomb friction term and passive resistance term, calcareous sand soil using equivalent trilinear Coulomb friction model and calcareous sand soil using Coulomb friction term and passive resistance term. Table 2 presents the soil models input data used in each of the simulation cases. It should be noted that the soil model data for the last two cases are for the same calcareous soil sample. Table 2 Soil Models Input Data Item Silica Sand Coulomb friction Silica Sand Coulomb friction + Passive term Calcareous Sand Trilinear Coulomb friction Calcareous Sand Coulomb friction + Passive term Initial distance, passive term 0.02D 0.01D Breakout friction 0.75 Breakout distance 0.50D 0.25D 0.25D Residual friction Residual distance 1.00D 1.00D 0.75D Results of the four simulations are presented in Figure 13. The results highlight the effect of the soil resistance model on the final pipeline displacement. For the silica sand soil cases, addition of the passive term reduces the maximum lateral displacement by approximately 30% relative to the base case. While for the calcareous soil cases, the final lateral displacement is higher than the cases for silica sand soil. This result highlights the fact that the calcareous soil has less lateral resistance capacity than the silica sand soil. The final lateral displacement of the tri-linear model is almost 10% higher than the final lateral displacement of the case with Coulomb friction term and passive resistance term. The vertical pipeline penetration at the end of the simulations is presented in Figure 14. As shown in the figure, the vertical penetration varies along the pipeline route due to the change in the hydrodynamic lift load distribution along the pipeline. Moreover, the vertical penetration for the calcareous sand soil cases is higher than the vertical penetration of the silica sand soil. While there is no record of the vertical penetration for the silica sand case with soil modelled using pure Coulomb friction term.

14 14 OTC MS Figure 13 Pipeline Final Displacement Soil Models Figure 14 Pipeline Vertical Penetration Soil Models Case 3 Operational pressure and temperature. Pipelines are more susceptible to displace laterally when the operational pressure and temperature are accounted for in the model. The base case example considering Seed 3 hydrodynamic loads is repeated including an operational pressure of 50 bar and operational temperature of 70 C at start of the pipeline, length 0.0 m, and linearly reduced to 25 C at the end of the pipeline, length 2500 m. The ambient temperature used is 20 C. The final displacement of the base case Seed 3 and the case considering the pressure and temperature is shown in Figure 15. It is clear from Figure 15 that the final lateral displacement for the case with operational pressure and temperature is in general higher by approximately 25% than the case without operational pressure and temperature. The maximum tension force recorded during the simulation along the pipeline for both cases is presented in Figure 16. As expected the maximum tension values of the case with the operational pressure and temperature is higher than the case without the operational pressure and temperature. Figure 15 Pipeline Final Displacement W/WO Operational P&T Figure 16 Pipeline Maximum Tension W/WO Operational P&T

15 OTC MS 15 It should be noted that the axial forces, bending moments and section stresses can be extracted from the dynamic simulation program; however for the paper space limitation, the results of the axial forces only will be presented. Case 4 Pipeline end boundary conditions. To highlight the effect of the pipeline end boundary condition effect on the pipeline displacement and tension loads, the base case example considering Seed 3 hydrodynamic loads is repeated considering the pipeline is fixed at both ends. The final pipeline displacement and maximum tension loads are presented in Figure 17 and Figure 18. As shown in the figures the lateral displacement at the middle section of the pipeline has not been changed much. While the tension load along the pipeline has increased by about 40% for the case with fixed pipeline end. Therefore, it is important to account for the pipeline's actual end constraints in the dynamic simulation to correctly estimate the tension load and the pipeline stresses. Figure 17 Pipeline Final Displacement Figure 18 Pipeline Maximum Tension Loads Case 5 Pipeline route. One of the important aspects of modelling the pipeline in a 3D domain is to consider the actual pipeline route rather than a simple straight route. To highlight the effect of considering the pipeline route on the on-bottom stability analysis, the base case example considering Seed 3 hydrodynamic loads is repeated considering a lateral offset of 350 m at the middle point of Seed 3 alignment and considering a bend radius of 750 m as presented in Figure 19. For comparison purposes, the hydrodynamic loads estimated for Seed 3 is used for the route 1 and route 2 simulations ignoring any hydrodynamic loads correction due to the change in the wave heading angle. The final pipeline displacements of the three cases are presented in Figure 20. Figure 19 Pipeline Routes

16 16 OTC MS Figure 20 Pipeline Final Displacement As shown from the results, the final pipeline lateral displacement changed significantly at the middle section of the pipeline route. The effect at the pipeline's end location, 100 m or 200 m from each end, could be influenced by the pipeline end boundary conditions. Case 6 Flexible and rigid pipeline. To demonstrate the effect of pipeline axial stiffness and bending stiffness on the pipeline on-bottom stability, the base case example considering Seed 3 hydrodynamic loads is repeated with replacing the rigid pipe axial stiffness of (1.51E9 N) and bending stiffness of (8.21E6 N.m 2 ) with a flexible pipeline axial stiffness value of (3.30E8 N) and bending stiffness value of (5.89E4 N.m 2 ). The pipe self-weight, outer diameter and hydrodynamic loads are kept unchanged as per the base case data. The final lateral displacement results of the rigid and flexible pipeline cases are presented in Figure 21. The maximum tension loads recorded during the simulation time is presented in Figure 22. Figure 21 Pipeline Final Displacement Figure 22 Pipeline Maximum Tension Loads As shown in Figure 21, the predicted final displacement of the flexible pipeline case is higher than the final displacement of the rigid pipeline case by about 17% while the maximum tension force recorded for the flexible pipeline is about 50% of the tension loads recorded for the rigid pipeline case. Therefore, disregarding the flexible pipeline properties in the on-bottom stability simulation may lead to conservative tension loads prediction and unconservative lateral displacements. To further explore the effect of changing the pipeline axial stiffness and bending stiffness on the pipeline final displacement and tension load prediction, a test matrix of 20 simulation cases is performed considering the axial stiffness and bending stiffness combinations presented in Figure 23. The actual axial and bending stiffness of the rigid and flexible pipeline cases are also shown in Figure 23. It should be noted that not all the test matrix cases represent real pipeline conditions; however, the test matrix cases are considered

17 OTC MS 17 acceptable for the purpose of this study to investigate the effect of the axial stiffness and bending stiffness on the dynamic simulations. An example of the pipeline final lateral displacement for the cases with axial stiffness of 1.0E9 N is presented in Figure 24 while varying the pipeline bending stiffness. Figure 23 Test Matrix Pipeline Stiffness Figure 24 Pipeline Final Displacement Varying EI Values As shown in Figure 24 the change in the pipeline bending stiffness has a minor effect on the final pipeline displacement for all the cases except for the case with the highest bending stiffness of 1.0E8 N.m 2. This behaviour is expected to be a result of the bending stiffness being sufficiently high to redistribute the hydrodynamic loads along a sufficient axial length of the pipeline leading to a more stable pipeline. As shown in the figure, the maximum lateral displacement of the case with bending stiffness of 1.0E8 N.m 2 is m which is about 84% of the maximum displacement value of the case with the bending stiffness of 1.0E4 N.m 2. Figure 25 presents the maximum lateral displacements of all the test matrix cases. While Figure 26 presents the maximum tension forces recorded along the pipeline route during the simulation. Figure 25 Pipeline Maximum Displacement

18 18 OTC MS Figure 26 Pipeline Maximum Tension Loads As shown in Figure 25 the maximum lateral displacement along the pipeline at the end of the simulation decreases with the increase of the pipeline bending stiffness and axial stiffness. However, the effect of the increasing the axial stiffness is much higher than the effect of increasing the pipeline bending stiffness on the final pipeline displacement. For example the cases with pipeline bending stiffness of 1.0e8 N.m 2, the maximum lateral displacement for the case with axial stiffness 1.0E7 N is about 21 m which is almost double the maximum lateral displacement of the case with axial stiffness 1.0E10 N. As shown in Figure 26, the maximum tension loads along the pipeline are much higher for the cases with high axial stiffness values. Case 7 Wave boundary layer. For small diameter pipelines, 10 cm diameter or less, or cables the 2D stability method of DNV-RP-F109 (Ref. 3) provides unrealistic pipe specific gravity requirement. As discussed in Cheng et al. (Ref. 2) it has been speculated that the exclusion of the wave boundary layer effect is attributing to the unrealistic prediction of the required pipeline specific gravity. Considering the wave boundary layer Equations 15 to 18, the wave velocity acting on the pipeline can be estimated accounting for the seabed material roughness height. To highlight the effect of the wave boundary layer, the reduced wave velocity is calculated for a wave velocity of 0.5 m/s and wave period of 10 s. The example considered four different seabed materials and considered pipe outer diameter values of 2.54, 5, 10, 20 and 40 cm. As shown from the reduced wave velocity results presented in Figure 27, the boundary layer has significant effect for the cases with small diameter pipelines and for the cases of rough seabed material. Figure 27 Wave Velocity Considering Boundary Layer Figure 28 presented an example of the required pipeline specific gravity to achieve the DNV-RP-F109 Absolute stability requirement for different pipeline outer diameters under peak wave velocity of 0.5 m/ s and wave period of 10 sec without considering the wave boundary effect. The figure also presents the required specific gravity requirement for the same pipeline case considering the wave boundary effect for the four different seabed materials.

19 OTC MS 19 Figure 28 Pipeline Specific Gravity for Absolute Stability To highlight the effect of the wave boundary layer on the dynamic stability simulation, the hydrodynamic loads are generated for a 2.54 cm outer diameter cable in 50 m water depth. A submerged weight of 11 N/ m and tension capacity 220 kn are considered. The specific gravity of the cable is The metocean data are as per the base case data (Table 1). Two cases have been simulated with and without considering the wave boundary layer effect. The final cable displacement and maximum tension load results from the two simulation cases are presented in Figure 29 and Figure 30, respectively. Figure 29 Cable Displacement W/WO Boundary Layer Figure 30 Cable Maximum Tension Loads W/WO Boundary Layer As shown in the figure, the cable is predicted to experience significant lateral displacement of about 61.0 m for the case ignoring the wave boundary layer effect. When the wave boundary effect is taken into account, the maximum predicted cable displacement along the cable route is about 1.2 m. As DNV-RP-F109 (Ref. 3) states "It is not recommended to consider any boundary layer effect on the wave induced velocity", consideration of this effect should be used with caution. Conclusion This paper discusses the on-bottom stability analysis of submarine rigid and flexible pipelines, umbilicals and cables in the light of recommended practice DNV-RP-F109. The limitations of the current recommended practice 2D design methods are first addressed before presenting the 3D dynamic analysis modelling tool. Using 3D modelling, the pipeline designer can perform more reliable and cost effective on-bottom stability design. The pipeline designer can define the important analysis parameters such as the pipeline axial stiffness, bending stiffness, material nonlinearity, the actual pipeline route, the pipeline end conditions,

20 20 OTC MS the operational pressure and temperature and the soil resistance models etc. The limitations of the 2D onbottom stability design methods of DNV-RP-F109 and the advantages of performing the 3D dynamic onbottom stability modelling are highlighted with the results of the dynamic stability simulation examples. Based on the results of the simulation analyses presented in the paper, the following conclusions can be drawn: The dynamic stability simulation should account for the effect of the random seed and random numbers used in generating the sea-states. As presented in the simulation section, the final pipeline displacement will vary along the pipeline route due to the change in the random seed. As recommended by DNV-RP-F109, at least seven analyses with randomly chosen seeds should be performed. The hydrodynamic loads acting on the pipeline should be corrected during the simulation to account for the experienced pipe velocity and or vertical penetration. Uncorrected hydrodynamic loads will lead to overly conservative predictions of pipeline lateral displacement. The soil resistance including the pure Coulomb friction term and passive resistance term should be modelled in the on-bottom stability analysis. Moreover, the soil resistance model used in the analysis should correctly represent the soil conditions in the field. The pipeline is shown to have different lateral displacement behaviour under the same metocean loads with changing soil resistance model. The pipeline operational pressure and temperature should be included in the on-bottom stability model as the pipeline is more susceptible to displace laterally when the operational pressure and temperature are accounted for. Moreover, the predicted pipeline maximum tension forces and stresses are higher when the operational pressure and temperature are considered. Special consideration should be given to modelling the pipeline end conditions and the pipeline actual route. The pipeline is shown to experience higher tension loads when the pipeline ends are restrained. Furthermore, the full pipeline route including the correct end restraints and any change in the route heading should be included in the model to correctly estimate the pipeline lateral displacement and the maximum tension forces. The pipeline axial stiffness and bending stiffness have significant effect on the pipeline displacement and the maximum tension loads. Therefore, for flexible pipelines, umbilicals and cables the Generalised stability method of the recommended practice DNV-RP-F109 may lead to unconservative on-bottom stability designs as the displacement of a flexible pipeline or cable (with low axial stiffness) could be higher than the displacement of a similar rigid pipeline. References The wave boundary layer affects the wave velocity in the vicinity of the seabed. The recommended practice DNV-RP-F109 does not provide any guidelines on the wave velocity calculations for small diameter pipelines or cables. Ignoring the effect of the wave boundary layer in on-bottom stability analysis of small diameter pipelines or cables can lead to overly conservative design results. However, caution needs to be exercised in implementing the effect of the wave boundary layer effect in the design as it is not listed in the recommended practice. 1. AGA (2008), Submarine pipeline on-bottom stability, Report no. PR , Analysis and Design Guidelines, American Gas Association, Houston, Texas. 2. Cheng, L., An, H., Draper, S. and White, D. (2016), Effect of wave boundary layer on hydrodynamic forces on small diameter pipelines journal of Ocean Engineering, Volume DNV (2011), On-bottom stability design of submarine pipelines, Recommended practice, DNV- RP-F109, Det Norske Veritas (DNV).

21 OTC MS Fredsoe, J. and Deigaard, R. (1992), Mechanics of Coastal Sediment Transport. World Scientific, Singapore. 5. Hughes, S.A. (1993), Physical Models and Laboratory Techniques in Coastal Engineering, Advanced Series on Ocean Engineering. 6. Sana, A. and Tanaka, H. (2007), Full-range equation for wave boundary layer thickness, Journal of Coastal Engineering, 54, Sarpkaya, T. and Rajabi, F. (1979), Hydrodynamic drag on bottom-mounted smooth and rough cylinders in periodic flow, Offshore Technology Conference, Houston, Texas. 8. Sorenson, T., Bryndum, M. and Jacobsen, V. (1986), Hydrodynamic forces on pipelines- model tests, Report no. PR , Danish hydraulic Institute (DHI), Pipeline Research Council International (PRCI). 9. Sumer, B.M. and Fredsøe, J. (1997), Hydrodynamics around cylindrical structures, vol. 12, World Scientific Publishing, London. 10. Youssef, B.S., Cassidy, M.J. and Tian, Y. (2010), Balanced three-dimensional modelling of the fluid-structure-soil interaction of an untrenched pipeline, International Offshore (Ocean) and Polar Engineering Conference, Beijing. 11. Youssef, B.S., Cassidy, M.J. and Tian, Y. (2011), Probabilistic model application in the integrated stability analysis of offshore on-bottom pipeline, International Conference on Ocean, Offshore and Arctic Engineering, Rotterdam. 12. Youssef, B. S., Cassidy, M. J., and Tian, Y (2013) Application of Statistical Analysis Techniques to Pipeline On-Bottom Stability Analysis. Journal of Offshore Mechanics and Arctic Engineering 135(3). 13. Youssef, B.S. and Cassidy, M.J. (2014), Calibration of Verley and Sotberg Soil Resistance Model for Pipelines Placed on Calcareous Soils, International Ocean and Polar Engineering Conference, Busan, Korea. 14. Verley, R. and Sotberg, T. (1992), A Soil Resistance Model for Pipelines Placed on Sandy Soils, OMAE Volume 5-A. 15. Verley, R. and Lund, K. (1995), A Soil Resistance Model for Pipelines Placed on Clay Soils, OMAE Volume 5.