OMAE THERMO-MECHANICAL DESIGN OF CANAPU PIP SYSTEM

Proceedings of the 8th International Conference on Ocean, Offshore and Arctic Engineering OMAE 009 May 31 - June 5, 009, Honolulu, Hawaii, USA OMAE 79713 THERMO-MECHANICAL DESIGN OF CANAPU PIP SYSTEM Rafael Familiar Solano PETROBRAS Rio de Janeiro, Brazil rsolano@petrobras.com.br Malcolm Carr ATKINS BOREAS Newcastle, UK malcolm.carr@atkinsglobal.com Anderson Dolinski TECHNIP Rio de Janeiro, Brazil ADolinski@technip.com ABSTRACT This paper discusses the thermo-mechanical design of the pipe-in-pipe (PIP) flowline installed in the Canapu field, located in Espírito Santo State, offshore Brazil. The pipeline is approximately 0km in length and connects the gas producing well -ESS-138 positioned in a depth of 1608m to Cidade de Vitória FPSO, located in Golfinho field. The Canapu PIP will operate under high pressure and temperature (HP/HT) conditions and is laid on the seabed. Due to the operational conditions, the thermo-mechanical design evaluated the susceptibility of the pipeline to the phenomenon of lateral buckling and pipeline walking in addition to free spanning and on-bottom stability. The lateral buckling behavior of the PIP is the major challenge for the design. It can be a safe and effective way to accommodate the thermal expansion of a hot pipeline, however high stress and strains can be developed in the buckles and a conventional stress based approach is not suited to design a pipeline that buckles laterally. The conventional stress limits are therefore relaxed and replaced by a strain limit. For this the methodology and recommendations of the SAFEBUCK JIP were adopted. The thermo-mechanical analysis selected a buckle initiation strategy based on distributed buoyancy. The strategy combines three distributed buoyancy triggers along the route together with the beneficial effect of the bathymetric out-ofstraightness. The analysis shows that this initiation strategy is robust and highly reliable. Fábio Braga de Azevedo PETROBRAS Rio de Janeiro, Brazil fabio.azevedo@petrobras.com.br Leanne Tindall ATKINS BOREAS Newcastle, UK leanne.tindall@atkinsglobal.com Carlos Eduardo Ingar Valer TECHNIP Rio de Janeiro, Brazil CEValer@technip.com From the start, this project represented a great challenge for Petrobras; it is the first PIP in Petrobras; has a low value specified for OHTC; and the pipeline is susceptible to lateral buckling. Besides all that, since the Canapu project was included among the priorities of Petrobras Plangas, it was executed as a fast track project. Keywords: Lateral buckling, thermo-mechanical buckling and pipe-in-pipe, HP/HT. INTRODUCTION To meet government requirements for gas in Brazil, Petrobras created the Plangas program. This program seeks to increase the gas production in Brazil by fast tracking the production of key gas fields. The Canapu gas field was included in Plangas program. The Canapu gas field is located in Espírito Santo State approximately 75km off the east coast of Brazil in 1600m of water, Figure 1. The development option selected by Petrobras was to tie the Canapu well (-ESS-138) back to the existing Cidade de Vitória FPSO, positioned in the Golfinho field. The gas will be processed on the Cidade de Vitória FPSO and exported to shore through the existing 1in Golfinho gas pipeline. The flow assurance studies highlighted the threat of hydrate formation during planned and unplanned production shutdowns. To mitigate this risk the flow assurance requirements defined a value for the pipeline OHTC of 0.8W/m K. This was the first major challenge for the pipeline 1 Copyright 009 by ASME

construction and operation, since it necessitated the use of a high performance PIP system. OOS - Out of Straightness PDEG-B - Gas Flow Master Plan PDET - Oil Flow Master Plan PIP - Pipe-in-Pipe PLET - Pipe Line End Termination UB - Upper Bound UOE - Pipe Fabrication Process for Welded Pipes UTS - Ultimate Tensile Strength VAS - Virtual Anchor Spacing 3LPP - 3 Layers Polypropylene Figure 1 - Location of the Canapu Field. The second major challenge results from the high temperature and pressure operation conditions. As a result the flowline may be susceptible to lateral buckling and pipeline walking. In addition, the uneven seabed presents a significant pipeline free spanning challenge. These mechanical design challenges are the main focus of this paper. Nowadays, Petrobras is involved with a large number of hot pipeline projects which will be laid in deep water. Recently, the PDET and PDEG-B projects were laid with buckle initiators (triggers) along the routes. These triggers (sleepers and distributed buoyancies) were specified to control thermomechanical buckling. The design for lateral buckling was based on the controlled lateral buckling principle, as presented by Solano et al.[1] and Sriskandarajah et al. []. This principle aims at achieving benign buckles by defining the locations of buckles through the provision of some form of engineered buckle triggering feature, at pre-defined intervals along sections of pipeline susceptible to buckling. The Canapu PIP lateral buckling design builds upon the lessons learnt during these and other projects. NOMENCLATURE API - American Petroleum Institute BE - Best Estimate CNP - Canapu DNV - Det Norske Veritas FBE - Fusion Bounded Epoxy FEA - Finite Element Analysis FJC - Field Joint Coating HP/HT - High Pressure and High Temperature JIP - Joint Industry Project KP - Kilometer Post LB - Lower Bound OHTC - Overall Heat Transfer Coefficient THE CANAPU PIP FLOWLINE SYSTEM The subsea production system for well -ESS-138 is composed of an electro-hydraulic control umbilical, a 6.65in service pipeline and an 8.65in production pipeline. The production flowline is approximately 0km in length with Pipeline End Terminations (PLET s) at each end. The flowline employs a thermally insulated flexible jumper (approximately 50m in length) upstream of the PLET-CNP-001 at the -ESS-138 well head and a flexible riser (approximately km in length) downstream of the PLET-CNP-00. Figure presents the outline for the production system and highlights the Canapu PIP. Figure - Production System of Well -ESS-138. The Canapu PIP design was developed to allow installation by the reel-lay method. This method was shown to be economically attractive and viable technically during the selection of the installation contractor (Technip). A sliding PIP system, which involves the insertion of the inner pipe inside the carrier pipe during stalk assembly, Figure 3, was used, because this methodology is most suited to construction at an onshore spool base. Copyright 009 by ASME

Northing (m) 7,787,000 7,786,500 Canapu 7,786,000 7,785,500 7,785,000 7,78,500 Golfinho 7,78,000 7,783,500 Easting (m) 7,783,000 5,000 50,000 55,000 60,000 65,000 70,000 Golfinho 0 KP (km) 15 10 5-100 Canapu -150-1500 -1550-1600 Water Depth (m) Figure 3 - Typical PIP Configuration. The PIP is composed of two concentric carbon steel pipes of 8.65 x 13.37in (19.1 x 339.7mm). The inner pipe will transport the produced gas and was specified to meet the design internal pressure, sour operational conditions and structural requirements due to the thermal loads. Centralizers are mounted onto the inner pipe at regular intervals of m in order to protect the insulation and ensure the concentricity of the two pipes. The carrier pipe was designed in accordance with the local buckling criteria of DNV-OS-F101 [3] and also in order to fulfill to the structural requirements due to the thermal loads. A three layer polypropylene coating (3LPP) was specified for the carrier pipe, together with aluminum anodes, to provide the corrosion protection. A FBE anticorrosion coating was used on the inner pipe. To meet the stringent OHTC target, the thermal insulation material Aerogel, with a thermal conductivity @ 37.5ºC of 0.0135W/mK and density of 130kg/m 3 was used in the annulus. Table 1 presents the operational data for the thermomechanical design. Table 1 - Operational Data. Parameter Value Design Life 15 years Condition Service Sour Service Water Depth 100 to 161m Design Pressure 15bar Minimum Design Temperature ºC Maximum Design Temperature 87ºC Geophysics and geotechnical surveys were carried out along the preliminary route of the Canapu PIP. The design route and bathymetry data for the pipeline are presented in Figure. -1650 Figure - Layout and Bathymetry along the PIP Route. The seabed is highly uneven and includes numerous large vertical features. In addition, the Canapu PIP crosses three existing pipelines near to the Golfinho FPSO. The pipe-soil interaction design data was based on interpretation of the geotechnical survey data along the pipeline preliminary route. The soils are variable across the route and two zones were defined. Zone 1 covered the east half of the pipeline route up to approximately 10.5km from PLET-CNP- 001 (Canapu). Zone covered the west half of the pipeline route from 10.5km to PLET-CNP-00 (Golfinho). The calculation of suitable pipe-soil interaction parameters was carried out considering the SAFEBUCK JIP methodology, as presented by Bruton et al. []. The axial and lateral friction factors calculated using this methodology are presented in Table. Table - Pipe-Soil Friction Coefficients. Friction Lateral Lateral Axial Factors Zone 1 Zone LB 0.7 0. 0.31 BE 0.58 0.66 0.50 UB 0.85 1.00 0.80 THERMO-MECHANICAL DESIGN In HP/HT pipeline design, extreme conditions can be developed within a lateral buckle; on first load the stresses can exceed yield and may involve significant plasticity; in addition, shutdowns will lead to very high stress cyclic fatigue. Lateral buckling if left uncontrolled can compromise the integrity of the pipeline. If the buckling response of the pipeline is not understood and controlled, rogue buckling could occur and hence lead to integrity issues with the pipeline. The end expansion of the pipeline will also be influenced significantly by the buckling behavior of the pipeline. 3 Copyright 009 by ASME

In addition, the operational start-up/shut-down cycles involving significant thermal gradients can lead to axial ratcheting of the pipeline. Start-up/shut-down cycles of a pipeline on a sloping seabed can also lead to axial ratcheting. Over a number of cycles, this movement can lead to very large global axial displacement, with associated overload of the spool piece or jumper. This cumulative axial displacement is described as pipe-walking [5], [6]. In the Canapu PIP design, the methodology and recommendations of the SAFEBUCK JIP were adopted [7]. Initially, a preliminary design analysis was carried out in order to identify a suitable, technically feasible pipeline design. This assessment employed the analytic models of lateral buckling and walking developed for the SAFEBUCK JIP [8], [9]. The preliminary assessment demonstrated that uncontrolled buckling was not acceptable and a buckle initiation strategy was required. An initiation strategy involving three triggers based on distributed buoyancy was identified as the most robust solution. The role of the detailed design was to verify and optimize this basic strategy. To do this, five distinct analyses were undertaken: A virtual anchor spacing (VAS) FE analysis to identify the tolerable buckle spacing for both engineered (buoyancy assisted) and rogue buckles; A probabilistic buckle formation analysis to quantify the robustness of the buckle initiation strategy; A buckle interaction and pipeline walking analysis in order to understand the global cyclic behavior of the pipeline; A pipeline free span assessment; A pipeline crossing design and analysis. Thus, the thermo-mechanical behavior of the Canapu PIP was evaluated and buckle initiators were specified along the route, to provide a robust solution for this pipeline during its life. Lateral Buckling: VAS Analysis In the VAS analysis relatively short FEA models (typically up to 6km) with fixed ends are employed to evaluate the lateral buckling response of the pipeline. Since the models are short it is possible to accurately incorporate the complex behaviors anticipated. For Canapu, this included non-linear pipe-soil interaction with soil berm response under cyclic loading; nonlinear material behavior; the sliding interaction between the inner and carrier pipe; and the effect of reeling. In the PIP design the inner flowline and carrier pipe are connected only at the end bulkheads. Elsewhere interaction between the two pipes is through contact and friction at the spacers. This sliding interaction was incorporated into the PIP FE model. The reeling process has a significant influence on the subsequent response. It is not possible to fully straighten the inner pipe following reeling; therefore the as-laid inner pipe incorporates residual stresses and strains. The flowline adopts a periodic deformation, which is governed by the pitch of the spacers. The amplitude of the deformation is low, Figure 5. Offset from Centreline (m) 0 0 5 10 15 0 5 30 35 0 5 50-0.005-0.01-0.015-0.0-0.05-0.03 Distance Along Pipeline (m) Figure 5 - Inner Pipe Post Reel Shape. The carrier pipe is fully straightened following reeling and therefore carries residual stresses but not any residual strain. Both the inner and carrier pipes do however contain plastic strain that occurs during the reeling process. There is no evidence to suggest that reeling is detrimental to the mechanical behavior of a pipeline, however as the mechanical behavior will be modified, the influence of reeling was fully assessed. The VAS analysis was performed at a number of locations along the pipeline length, and sensitivity analyses were undertaken to ensure that the influence of PIP design and reeling were fully evaluated. Typical post buckle displaced shapes from the analyses are presented in Figure 6. Figure 6a shows the response of an engineered buckle. The figure shows the displaced shape that results from three different analysis assumptions. The first analysis does not incorporate the reeling history. In this case a symmetric buckle forms. When the reeling history is incorporated a very similar displaced shape is observed. This is essentially a mode I buckle (the secondary lobes are small as a result of the increased resistance associated with the sections of pipe without buoyancy). A third analysis is also performed in which the pipe is assumed to twist through 90 between the reel barge and the seabed, so that the reeling plane is aligned with the buckling plane. In this case the FEA develops a mode II buckle; this is observed in all analyses undertaken. The change in shape is due to the residual reeling curvature in the flowline (Figure 5), which in the twisted case is in the same plane as the lateral buckling. Copyright 009 by ASME

No Reeling 10 Reeled Reeled and Twisted 8 6 0-100 -50 0 50 100 - Distance (m) - -6 (a) At Buoyancy Lateral Disp(m) Engineered Buoyancy No Reeling Reeled Reeled & twist 3 1 Inner Pipe 0-100 -50 0 50 100-1 - -3 (b) On-bottom Lateral Disp(m) Rogue On-bottom Distance (m) Figure 6 - Influence of Load history on Buckle Shape. In fact beyond KP 11 the pipeline meets all relevant unity checks at the maximum feasible spacing. Consequently, no initiation strategy is required in zone. The figure also plots the selected locations of the buoyancy triggers. The first two triggers are to control the spacing associated with rogue buckling. Trigger 3 is to reduce the risk of buckling at the pipeline crossings. To meet the design requirement, each trigger was 130m long and achieved an 85% reduction in pipeline submerged weight. The triggers were constructed from individual buoyancy modules, Figure 8. Figure 6b shows the response of a rogue buckle for the same three cases. The analysis predicts a mode 3 buckle in the reeled and unreeled cases. Once again the reel and twist analysis develops a more asymmetric response. Of course the influence of the reeling residual shape is no different to the influence of inherent OOS, and in general, it is not possible to predict the actual mode shape that a buckle will adopt. However, the different mode shapes do produce different design loads. Consequently, the VAS analysis ensured that the PIP was robust to the different mode shapes that could occur. For the buoyancy triggers the VAS analysis showed that a lateral buckle was acceptable at the maximum feasible buckle spacing (i.e. at the spacing controlled by the fully constrained force profile). This is because the buoyancy significantly reduces the lateral resistance, and hence the bending strain, imposed on the buckle. For rogue buckling, the results of the calculations are presented in terms of the Tolerable VAS; this is the VAS at which the most onerous limit state (in this case local buckling of the carrier pipe) becomes critical, Figure 7. VAS (km) 10 9 8 7 6 5 3 1 Trigger #1 Trigger # Trigger #3 On Bottom Buckles 0 5 10 15 KP 0 Figure 7 - Rogue Buckling: Tolerable VAS. The Tolerable VAS defines the maximum spacing between buckles that the pipe can tolerate. In geotechnical zone 1 the Tolerable VAS is approximately.5km. In geotechnical zone the Tolerable VAS is much greater as a result of the falling operating conditions and the reduced lateral resistance (Table ). Figure 8 - Canapu Buoyancy Modules. Buckle Formation Assessment In order to confirm that the buckle initiation strategy was able to keep the spacing below the Tolerable VAS a Monte Carlo simulation of the pipeline expansion process was undertaken. Since buckling at the triggers was acceptable at the maximum feasible VAS, the buckle formation assessment was only concerned with rogue buckles. In the analysis the pipeline was divided into a series of discrete elements (typically 100m long). The behavior of each element was described through a critical buckling force distribution. The distribution of critical buckling force was based on a model of inherent pipeline OOS and lateral pipe-soil resistance. By assigning different distributions of critical buckling force to different pipeline features (straight sections, route bends triggers etc.) a completely general behavior at each point along the pipeline was modeled. Within the Monte Carlo analysis the expansion of the pipeline system is undertaken 10 7 times. For each simulation the axial friction coefficient is selected from the defined axial friction distribution. At each element in the pipeline a critical buckling force is randomly selected from the defined distribution. The pipeline operating conditions are defined and a load uncertainty distribution is applied to these conditions. Pressure, temperature, depth and lay tension all varied along the length of the pipeline and the operational analysis considered a transient temperature profile to represent the heatup process. 5 Copyright 009 by ASME

The analysis then simulated the start up process to identify the buckle locations for the randomly selected set of design parameters. Once the buckled configuration was identified the results for that simulation were stored. The key results are: Whether the simulation has produced a buckle or not (to give the overall probability of at least one buckle); The total number of buckles in the pipeline (to give the distribution of number of buckles in the pipeline); Each element that produced a buckle is recorded (to give the probability of buckling at any location); For each element that produced a buckle the VAS is recorded (to give the VAS probability distribution at any location); VAS (km) 7 6 5 3 1 Tolerable VAS Probable: 3 Triggers Probable:No Triggers Probable: 3 Triggers & Bathymetry 0 5 10 15 KP 0 Figure 9 - Comparison of Tolerable and Probable VAS The key result from the analysis is a distribution of Probable VAS along the pipeline. This is defined as the VAS that has an acceptably low probability of being exceeded. In this case a probability of exceedance of 10-3 was taken to be a suitable design value. The result of the buckle formation simulation is presented in Figure 9. The figure shows the Tolerable VAS obtained in the VAS analysis and compares this to the Probable VAS calculated in the Monte-Carlo simulation. If the Tolerable VAS exceeds the Probable VAS then the design is robust. The red line in Figure 9 shows the Probable VAS in the absence of a buckle initiation strategy (uncontrolled buckling). This shows that uncontrolled buckling is not acceptable between KP 3 and KP 10. This confirms the requirement for an initiation strategy to ensure the loads in the rogue buckles are acceptable. The blue line in Figure 9 shows the Probable VAS for the three buoyancy trigger design strategy. The initiation strategy significantly reduces the Probable VAS for rogue buckling. This is less than the Tolerable VAS over the whole pipeline route and is therefore acceptable. This reduction is Probable VAS is due to the high formation reliability of the triggers. The figure also shows the results of the simulation when the influence of seabed bathymetry is considered. In this analysis the most significant bathymetric features were explicitly modeled. These were assigned a distribution of critical buckling force based on FEA of the measured bathymetry. Since there are a number of significant features, this results in a further reduction in Probable VAS. Overall this produces a very robust design position. Buckle Interaction And Walking The VAS analysis and buckle formation analysis are the main parts of the lateral buckling design. As a final stage of the design process a FE model of the full pipeline between Canapu and Golfinho was constructed. The 0km model incorporated the seabed bathymetry and all of the key pipeline features (route curves, buoyancy, PLET s etc.). The aim of the analysis was to investigate the global expansion and walking behavior of the pipeline and to identify any severe bathymetric features which should be incorporated into the buckle formation assessment (either beneficial or detrimental). The model was subjected to ten operational cycles, with the sensitivity of the results to axial friction coefficient and seabed stiffness considered. The effective force profile developed in the pipeline during operation is presented in Figure 10. Effective Axial Force (kn) Peff - 1st Load Peff - 10th Load Bending Moment (Buckling Plane) 00 600 00 00 0 KP (m) 00-00 0 5000 10000 15000 00000-00 -00-00 -600-600 -800-800 -1000-1000 -100-100 -100-100 Figure 10 - Effective Force and Bending Moment. The figure shows the effective force profile at maximum operating conditions during the 1 st and 10 th cycles. The bending moment profile in the horizontal plane is superimposed on the plot. Buckles can be observed from the effective force and bending moment profiles at the three trigger locations; KP5.75, P11.15 and KP18.65. In addition to these planned buckles, a number of rogue buckles have formed. These can be identified by the location of significant horizontal bending moment in Figure 10. The location and VAS associated with the rogue buckles are illustrated in Figure 11, which compares the FE results to the probabilistic simulations. The VAS identified in the FE analysis is smaller than the Probable VAS for all rogues. The FE analysis provides detailed information on the influence of bathymetry on buckle formation and buckle interaction. However, each FE analysis provides only a single prediction of possible behavior, which depends upon a large number of highly uncertain parameters, and is thus not a demonstration of pipeline integrity. The FE predictions in Figure 11 are for the best estimate conditions Bending Moment (knm) 6 Copyright 009 by ASME

whereas the probabilistic results relate to a 10-3 probability of exceedance. The FE analysis is extremely informative, but it is important to quantify the reliability of the buckle formation process; this is achieved by the probabilistic simulation. VAS (km) 7 6 5 3 1 Probable:No Triggers FEA: Best Estimate Probable: 3 Triggers & Bathymetry 0 5 10 15 KP 0 Figure 11 - Probable VAS and FE predictions. The full length FE model was used to evaluate the potential for pipeline walking [5], [6]. The end expansion predicted in the model is illustrated in Figure 1. Axial Displacement (m) 0.1 0.1 0.10 0.08 0.06 0.0 0.0 PLET-CNP-001 PLET-CNP-00 Pipeline Free Spanning The effective force profile in Figure 10 shows a number of locations where the force is found to reduce but there is no significant horizontal bending moment. These areas are associated with free span features; the drop in force is caused by feed-in to the spans. The full length FE model identified the potential for a large number of free spans along the pipeline route. The severity of these pipeline free-spans were assessed in accordance with DNV-RP-F105 [10]. The assessment considered local buckling, fatigue and span interaction. Due to the deep water and modest environmental loading only local buckling was a significant concern. Local buckling was evaluated in accordance with the load controlled checks specified in DNV-RP-F105. There is an inevitable inconsistency between the strain based design for lateral buckling and the load based design for spanning. The highest strains were due to the seabed bathymetry and it may have been possible to apply a less conservative strain based criteria. However, this was not implemented and the load based limits were applied. The vast majority of the spans were small and complied with the requirement of DNV-RP-F105. However, a small number were found to exceed the allowable bending moment and span repairs were implemented. These employed mechanical supports, capable of repairing spans up to 5m in height, Figure 13. 0.00 0 6 8 10 Load Cycle Figure 1 - Pipeline Walking Response. The figure shows the end expansion over 10 start-up shutdown cycles. On first load the maximum expansion occurs at the pipeline outlet (Golfinho, PLET-CNP-00) not the hotter inlet (Canapu, PLET-CNP-001). This is due to the proximity of bucking to the pipeline inlet, which reduces the end expansion. Over the 10 cycles the expansion at the outlet is quite steady (between 15mm on load and 60mm on unload), with little indication of walking. At the inlet the end expansion does exhibit a low level walking response, growing from about 100mm to 10mm over the simulation. Away from the pipeline ends similar levels of axial ratcheting were observed at the lateral buckles. However, no significant walking was observed at the steep slope between KP 10 and KP 15. Given the modest end expansion and the low walking rate, no specific mitigation measures were adopted. Figure 13 - Mechanical Span Support. The full length FE model was modified to incorporate the span repairs and the analysis was repeated to confirm that the post-repair configuration complied with the requirements of DNV-RP-F105. A typical output from the analysis is presented in Figure 1. 7 Copyright 009 by ASME

Z (m) KP -1600-1601 3. 3.1 3. 3.3 3. 3.5-160 -1603-160 -1605-1606 -1607-1608 Bathymetry Laydown Hydrotest 1st Load Span Repairs Depth (m) -16-18 -130-13 -13-136 -138-10 18.5 18.6 18.7 18.8 18.9 19. 19.1 KP Trigger #3 Oil Prod Bathymetry Laydown Hydrotest Operation EHU Gas Lift Z (m) -1600 KP -160 6. 6.1 6. 6.3 6. 6.5-160 -1606-1608 -1610-161 -161-1616 -1618-160 Bathymetry Hydrotest Span Repairs Figure 1 - Post Span Repair Analysis. Laydown 1st Load The figure shows results from the full length model for two 500m sections of the pipeline. No repairs were required between KP 3 and 3.5. However, this section illustrates the behavior at a significant vertical feature, in which feed-in during operation allows the pipe to conform to the seabed. Between KP 6 and 6.5 three significant spans occur, which were repaired using mechanical supports, Figure 13. Pipeline Crossings The Canapu PIP has to cross three service lines in the approach to Golfinho (an oil production line, a gas lift line and an electro-hydraulic umbilical). The design must maintain a 300mm clearance between the Canapu PIP and the crossed services. The crossing construction is based on stacks of concrete mattresses (each mat is 6096mm x 39mm x 9mm). The height of each mattress stack is limited to four mattresses high in order to remain within the capability of the existing mattress installation equipment. There is significant bathymetric variation in the vicinity of the crossings, so six mattress stacks of varying height were employed, Figure 15. Figure 15 - Behavior at Crossing. The vertical upset introduced by the crossing design could lead to either lateral or upheaval buckling. It is undesirable for buckling to occur at the crossing. Therefore to prevent this a buoyancy trigger is located close to the crossing. The buckle formation and FE analysis both showed that this trigger buckles in preference to the crossing. On installation the PIP spans between the crossing of the gas lift line and the EHU. However, on operation, the expansion causes the PIP to touch down between these crossings. All spans associated with the crossing were assessed and found to be in accordance with DNV-RP-F105. CONCLUSIONS The design of the Canapu PIP system presented a number of complex challenges. The pipeline was the first PIP in Brazil, was installed by the reel lay method, was susceptible to lateral buckling and experienced significant spanning. To meet these design challenges the PIP was designed in accordance with the SAFEBUCK design methodology. This process involved a new, comprehensive probabilistic approach to buckle formation, which is suitable for general application. The thermo-mechanical behavior of the Canapu PIP was evaluated and buckle initiators were specified along the route, providing a robust solution for this pipeline during its life. ACKNOWLEDGMENTS The authors would like to thank Petrobras for sponsoring this work and giving permission to publish this paper. The authors also acknowledge the Technip and Atkins Boreas project team for their comprehensive work being provided throughout the design and installation phases of the project development. REFERENCES [1] R. F. Solano, V. R. Braga, Dr T. Sriskandarajah, P. Tanscheit and A. Bedrossian, Lateral Buckling Design Philosophy and Implemented Solutions, Rio Pipeline Conference, IBP1193_07, October, 007. [] Dr T. Sriskandarajah, R. F. Solano, V. R. Braga, P. Tanscheit and A. Bedrossian,, Free Span Rectification of PDEG-B Oil Lines Subject to Thermal Expansion and 8 Copyright 009 by ASME

Lateral Buckling, Rio Pipeline Conference,IBP1176_07, October, 007. [3] Det Norske Veritas, Offshore Standard F101, Submarine Pipeline Systems, January 000. [] D. Bruton, D. White, M. Carr, and J. Cheuk, Pipe-Soil Interaction During Lateral Buckling and Pipeline Walking The SAFEBUCK JIP, Paper 19589, OTC 008. [5] Axial Creeping of High Temperature Flowlines Caused By Soil Ratcheting. Tornes, K., Jury, J., Ose, B., Thompson. OMAE 000. [6] Lateral Buckling and Pipeline Walking, a Challenge for Hot Pipelines. Carr M., Bruton, D. and Leslie, D. Offshore Pipeline Technology Conference 003,Amsterdam. [7] D. Bruton, M. Carr, M. Crawford, and E. Poiate, The Safe Design of Hot On-Bottom Pipelines with Lateral Buckling using the Design Guideline Developed by the SAFEBUCK Joint Industry Project, DOT 005. [8] M. Carr, F. Sinclair, and D. Bruton. Pipeline Walking- Understanding the Field Layout Challenges and Analytical Solutions Developed for the Safebuck JIP. SPE Projects, Facilities & Construction Journal. September 008. [9] M.Carr, and D.Bruton. Design Guideline. SAFEBUCK JIP. August 00. [10] Det Norske Veritas, Recommended Practice F105 Free Spanning Pipelines, February, 006. 9 Copyright 009 by ASME