Proceedings of OMAE2005 24th International Conference on Offshore Mechanics and Arctic Engineering (OMAE 2005) June 12-17, 2005, Halkidiki, Greece OMAE2005-67216 REELING EFFECT ON THE ULTIMATE STRENGTH OF SANDWICH PIPES Xavier Castello Ocean Engineering Department, COPPE/UFRJ Rio de Janeiro, Brazil xavier@lts.coppe.ufrj.br ABSTRACT Sandwich pipes composed of two steel layers separated by a polypropylene annulus can be used for the transport of oil&gas in deepwaters, combining high structural resistance with thermal insulation in order to prevent blockage by paraffin and hydrates. In this work, sandwich pipes with typical inner diameters of those employed in the offshore production are analyzed numerically to evaluate the ultimate strength under external pressure and longitudinal bending as well as the effect of the reeling installation method on the collapse pressure. Numerical models were developed using the commercial finite element software ABAQUS. The validation was based on experimental results. The analyses for combined loading were performed using symmetry conditions and the pipe was reduced to a ring with unitary length. The analysis of bending under a rigid surface was simulated numerically according to the experiments performed using a bending apparatus especially built for full scale tests. Symmetry conditions were employed in order to reduce the analysis to a quarter of a pipe. Mesh sensitivity studies were performed to obtain an adequate mesh refinement in both analyses. The collapse pressure was simulated numerically either for the pre or post reeling process. Bauschinger effect was included by using kinematic hardening plasticity models. The influences of plasticity and out-ofroundness on the collapse pressure have been confirmed. NOTATION D Dmin Dmax Dn K Kco L M diameter minimum diameter maximum diameter nominal diameter curvature collapse curvature length moment Segen F. Estefen Ocean Engineering Department, COPPE/UFRJ Rio de Janeiro, Brazil segen@lts.coppe.ufrj.br P Pd Pco Ri Re t x, y, z ∆ ε ε eqp θ σ σ eq pressure design point collapse pressure internal radius external radius thickness coordinate axes ovalization strain equivalent plastic strain rotation angle stress Von Mises equivalent stress INTRODUCTION New concepts of especial submarine pipes have been proposed mainly to allow an adequate oil&gas transportation flow. It is the case of both Pipe-In-Pipe (PIP) and Sandwich Pipe (SP). PIP are composed of two concentrically mounted steel pipes with the annular space filled with circulating hot water or materials with pre-defined thermal insulation properties. The objective of this type of pipe is to increase the thermal insulation capacity to prevent blockage of the line caused by dropping the fluid temperature to that required for paraffin or hydrate formation. One of the advantages of PIP system is the possibility of using materials with excellent thermal properties, considering that the structural integrity is provided independently by the outer and inner steel layers [1]. In the case of SP, object of this study, the annular layer characteristics differ from PIP by satisfying simultaneously mechanical and thermal requirements. Consequently, it is an interesting option for ultra deepwaters under conditions which require thermal insulation for flow assurance. Numerical and experimental studies carried out by Estefen et al. [2] for the limit strength, under combined external 1 Copyright © 2005 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use pressure and bending, indicate that SP are viable for application in water depths up to 3000m. Inter-layer contact behavior, i.e. the degree of adhesion, was observed to have significant influence on the collapse pressure. Among the advantages compared with single wall pipe, it was observed a substantial higher bending capacity for equivalent external pressure, with similar steel weight and less submerged weight. Structural resistance is affected by the chosen material of the annular layer. Polypropylene was adopted in the present work for reasonably low cost, good mechanical properties and relatively low thermal conductivity. Active heating system using electrical wires, associated with passive insulation due to the annular material was evaluated by Su Jian et al. [3, 4]. The study of the heat transfer in transient flow for a deepwater scenario indicates the necessity to input a certain amount of heating to avoid line blockage in case of production interruption or for long distance tieback. In this work numerical modeling of the ultimate strength was evaluated for combined external pressure and longitudinal bending of the sandwich pipe with geometric properties (Table 1) similar to the prototypes to be fabricated by TENARISCONFAB as part of a collaboration project with COPPE/UFRJ to perform experimental tests at the Submarine Technology Laboratory (LTS). Mesh refinement studies were conducted to obtain the numerical model to be used to evaluate the prototypes ultimate behavior. Symmetry conditions for halftransversal section and half-length were considered to minimize the computational effort. Later, another model was developed to study the residual effect of the reeling installation method through numerical simulation of the process of bending and rectification of pipes used at LTS. Symmetry of half-transversal section and half-length in relation to the actual pipe length used in the apparatus was adopted. Kinematic hardening plasticity models were introduced to simulate the Bauschinger effect for the metal layers. After the bending and rectification of the sandwich pipe the collapse pressure was determined to verify the influence of this process on the ultimate strength. Moreover, the curves for moment-curvature and stress-strain are determined to provide a better understanding of the pipe stress and strain distributions. SANDWICH PIPE MODEL FOR PRESSURE AND BENDING ANALYSIS The finite element program ABAQUS [5] was employed for the structural analyzes. Another program written in FORTRAN language was used initially to generate the mesh with nodes and elements of the sandwich pipe with geometric imperfection in the form of initial ovalization. For all models symmetry conditions for half-transversal section and halflength were assumed. Figure 1 shows an example of the model used for the ultimate strength analysis. One solid element is considered in length and for each metal layer and two solid elements for the polymer layer. The symmetry conditions applied to the X-Y and Y-Z planes are the displacement node constraints in the respective normal directions. Table 1 presents the geometric properties of the analyzed sandwich pipe. The steel pipes were selected according to API 5L standard. Geometry was determined based on TENARISCONFAB pipe availability in stock and in accordance with the LTS hyperbaric chamber capacity (50MPa), considering that experimental tests are planned. Figure 1: Finite element mesh for the sandwich pipe ring Table 1: Geometrical properties of the analyzed sandwich pipe Re (mm) t (mm) Dn (in) Ri (mm) 6 5/8 77.75 84.15 6.4 Annular 84.15 103.15 19 8 5/8 103.15 109.55 6.4 Initial ovalization of 1% was adopted for the mesh sensivity analyzes, which is calculated by the expression: D − Dmin . (1) ∆ 0 = max Dmax + Dmin Smaller diameter along the pipe length is coincident with the bending plane to generate lower bound values representative of the ultimate strength. For both external pressure and bending simulation acting separately Riks method and automatic increment control have been employed, respectively. Combined loading was initially implemented by fixed increments of external pressure followed by incremental rotation until buckling failure has been achieved. The external pressure is applied through surface load on the external pipe. The bending was induced by the use of a reference node located at x=L (model length), y=0 and z=0, where the rotational displacement in z axis is applied. Automatically generated kinematic coupling equations are used to couple the degrees of freedom of the nodes of the transverse plane to the reference node. It is assumed that this plane remains flat and normal to the neutral line during the loading. Additionally, the x degree of freedom coupling induces a plane strain state for the section, which is needed for simulating a long pipe. The curvature applied to the pipe is calculated by the following equation: K =θ L , (2) where the rotation angle θ (in radians) is input for each load step. The use of displacement control for the bending loads allows extending the analysis even after the ultimate load, thus generating the unloading moment-curvature relationship. 2 Copyright © 2005 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use However, in some cases the analysis ends just after reaching the maximum moment because of excessive element distortion leading to difficulties in obtaining the pipe ultimate strength. The collapse curvature corresponds to the maximum moment reaction at the reference node. In all models the mesh was generated using threedimensional quadratic solid elements C3D27, with twentyseven nodes and three degrees of freedom per node. As the polymer has volumetric incompressible characteristics, elements with mixed formulation C3D27H were used in the annular layers. Mixed formulation consists of calculating element pressure stress independently of the displacement stress to avoid that a small change in displacement produces extremely large changes in pressure. This happens when Poisson ratio reaches a value close to 0.5 or the bulk modulus is much larger than the shear modulus. MATERIAL MODELING API X-60 steel was used for inner and outer pipes. The proportional limit stress and yield stress are 310.38MPa and 481.46MPa, respectively. True stress versus logarithmic plastic strain curve, shown in Figure 2, was obtained by tensile tests as well as the derived elastic modulus (E) of 206,863MPa and Poisson ratio (ν) of 0.3. 800 True Stress (MPa) 700 600 dε p = where: H = 9 1 Sdσ S, 2 4 H σ eq dσ 11 dε 11p or H = (4) dσ 11 −E. dε 11 The notation indicates x direction for the uniaxial case. The dependency of the stress on the strain is simulated for the entire elastic-plastic regime and the material non-linear effects are considered. For the simulation of the reeling process a kinematic hardening plasticity model is added to the isotropic model, so that the Bauschinger effect can be simulated during the cyclic loading for more accurate results. The Bauschinger effect is characterized by a reduced yield stress under reverse loading after plastic deformations have occurred in the initial loading. This phenomenon tends to reduce its influence as the cycles progress. However, in the present simulation only one cycle and a half is performed, therefore it is not necessary to analyze the effects for stabilized cycles. Similar to the previous plasticity model, the subsequent yielding surface is represented by: 1 ⎛3 ⎞ 2 (5) f (σ − α ) = ⎜ (S − a )(S − a )⎟ = σ eq , ⎝2 ⎠ where α is the actual position of the yielding surface and a is the center of the surface in the deviatoric stress space. Now, the center of the surface translates in the direction of the plastic deformation increment. Figure 4 shows the representation of the movement of the yielding surface. 500 σ1 400 300 Initial yielding surface 200 Subsequent y. s. 100 0 0 0.025 0.05 0.075 0.1 0.125 0.15 0.175 0.2 σ2 Logarithmic Plastic Strain Figure 2: True stress-logarithmic plastic strain curve API X-60 steel Figure 3: Representation of the initial and subsequent yielding surfaces of the J2 flow rule theory for isotropic hardening in the stress space The steel layers are modeled by plasticity theory of potential flow rule J2 associated with isotropic hardening and Von Mises yielding criteria for the combined loading models. Summarizing, the initial yielding surface represented in the stress space in Figure 3, changes only in size, which is defined by the yielding surface expression with radius equal to Von Mises equivalent stress calculated at the last increment as a function of the stress state. The yielding surface expression is given by: ⎛3 ⎞ f (J 2 ) = ⎜ S ij S ij ⎟ ⎝2 ⎠ where 1 2 = σ eq , σ1 Initial yielding surface Subsequent y. s. σ2 (3) S ij is the deviatoric stress tensor. Figure 4: Representation of the initial and subsequent yielding surfaces of the J2 flow rule theory for kinematic hardening in the stress space The incremental plastic strains are obtained solving the flow rule as a function of the local hardening module (H) and the stress state σ , resulting in the Prandtl-Reuss expression: 3 Copyright © 2005 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use In the non-linear kinematic hardening model, the evolution law of the yielding surface consists of a component of kinematic hardening, which describes the translation of the surface in the stress space by the variable α . This component is defined by the combination of a pure kinematic term (Ziegler’s linear hardening law) and a relaxation term, which introduces the non-linearity. The law is expressed by: 1 (6) dα = C (σ − α )dε eqp − γ α dε eqp , σ eq where C and γ are material parameters adjusted by cyclic tests, C is the initial kinematic hardening modulus and γ determines C variation rate. The finite element program provides an option to include the kinematic hardening and Bauschinger effect consequently, through a uniaxial tensile test, being only recommended when a few cycles will be performed. Solid polypropylene was modeled as non-linear elastic material (hyperelastic) and volumetrically incompressible. A stress-strain curve is used by the program to obtain an Ogden approximation adjust, where the strain energy is given by the expression: N U = ∑ i =1 2 µ ⎛ αi αi αi ⎜ λ1 + λ2 + λ3 − 3 ⎞⎟ + ⎠ α i2 ⎝ ∑ D (J N 1 i =1 el ) −1 2i , (7) i where N, µi, αi and Di are material parameters. −1 The deviatoric strains are given by λi = J 3 λi , where λi are the components of the principal strain. J parameter is the ratio of total volume variation, J = dV dV0 , where V and V0 are the final and initial volumes, respectively. Jel is the ratio of elastic volume variation given by J el = J J th . For the case where no thermal strains occur ( J th = 1 ), we get J el = J . Shear stress and compressibility modules (bulk modulus) are given by µ 0 = N ∑µ i =1 i and K = 2 , respectively. D1 The parameters are defined by uniaxial tensile and compressive tests. Approximate results are obtained only for uniaxial tensile test. The Di parameter is obtained by volumetric compression test. Utilizing N equal to 2 and informing the uniaxial stressstrain curve of polypropylene for ambient temperature, Figure 5, the parameters are obtained by the automatic adjust of the finite element program, as below: µ1 = -769.66 µ2 = 944.98 α1 = 17.68 α 2 = 15.42 D1 = 0 D2 = 0 MESH SENSIVITY The same numerical model employed in previous works [2] has been validated with experimental tests of small-scale specimens and a parametric study indicated possible geometries for offshore application which are similar to those here analyzed. Mesh sensivity has also been considered. The intention now is to prepare a numerical model for further experimental correlation with full-scale prototypes (COPPE/UFRJ-TENARISCONFAB collaboration project). The present work includes a mesh sensivity analysis for a pipe with length equal to that adopted for tests at the LTS bending apparatus. The half ring presented earlier is first modeled with 10, 16, 20 and 24 elements in the circumferential direction for 5, 6.4, 10 and 20mm longitudinal length. It is well known that during the collapse of pipes by external pressure shell bending in the circunferencial direction occurs caused by the ovalization increase and, therefore, the mesh should be refined circumferentially. The results for collapse pressure (Pco) with 1% of initial ovalization are presented in Table 2, where Nc is the number of elements in the circunferencial direction and the deviation is relative to the Pco obtained for the long pipe (Table 3), taken as reference after Nc is selected. Table 2: Sandwich pipe ring mesh refinement results for external pressure collapse L (mm) Nc Pco (MPa) Deviation (%) 10 45.05 -1.12 16 44.4 0.34 10 20 44.3 0.56 24 44.25 0.67 10 45.15 -1.35 16 44.5 0.11 6.4 20 44.4 0.34 24 44.35 0.45 10 45.25 -1.57 16 44.5 0.11 5 20 44.4 0.34 24 44.4 0.34 Figure 5: Adjusted uniaxial stress-strain curve for polypropylene 4 Copyright © 2005 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Colapse Pressure, Pco (MPa) 46 45.5 45 44.5 44 43.5 5 10 15 20 25 30 Nc L = 5mm L = 6.4mm L = 10mm Figure 6: Collapse pressure results for sandwich pipe ring mesh refinement Figure 6 indicates that values of Pco tend to converge for Nc greater than 20. It is also observed that the ring lengths (L) 5 and 6.4mm converged with a deviation of 0.34%. In order to analyze a numerical model to evaluate a more precise Pco (used as reference in Table 2) a sandwich pipe is modeled with the same length as that of reeling simulation apparatus. Half-length (2600mm) and half-transversal section are considered and the pipe end is radially constrained, Figure 7. The ovalization is uniform for all pipe length and equal to 1% as before. Nc is assumed as 20 and the element length L is refined according to Table 3, where Nl is the number of elements in the longitudinal direction. The collapse pressure was practically not influenced by the length variation. Computer time increases too much for Nl greater than 65, without benefit for Pco accuracy. The 40mm element length was adopted to obtain Pco reference in previous sandwich pipe ring mesh refinement (Table 2). The 65mm element length was not adopted because the next simulation (reeling) uses contact elements and the analysis does not converge with coarse longitudinal mesh. Table 3: Collapse pressure results for long sandwich pipe using longitudinal mesh refinement L (mm) Nl Pco (MPa) 65 40 44.45 40 65 44.55 20 130 44.55 10 260 44.55 6.4 406 44.55 5.2 499 44.6 Figure 8: Collapse curvature for the long sandwich pipe Ring meshes for collapse curvature results were evaluated for the same three cases (5, 6.4 and 10mm length) and compared with the result for the long pipe, Figure 8. Table 4 shows the results, where the first line is related to the long pipe (with 65 elements in longitudinal direction and 2600mm length), with collapse curvature value taken as reference for deviation calculation. Table 4: Collapse curvature results for different sandwich pipe ring lengths compared to the long pipe L (mm) Kco (1/m) Deviation (%) 1.119 40 1.168 4.38 10 1.121 -0.18 6.4 1.103 1.43 5 It can be noted that the 6.4mm element length is adequate for further analyzes. Moreover, the 6.4mm ring length generates a deviation of only 0.34% for collapse pressure (Table 2) against 0.56% for the 10mm ring length. LIMIT STRENGTH CURVES FOR SANDWICH PIPE The numerical ring model with 20 circumferential elements and 6.4mm length has been employed to generate the ultimate strength curve of the sandwich pipe under combined external pressure and longitudinal bending. Sandwich pipe geometries are indicated in Table 1. Initial ovalization of 0.2% was adopted. Figure 9 shows the ultimate strength curve under combined external pressure and bending. The design point Pd indicates approximately the values of pressure and curvature acting on the pipe during the installation process for 3,500m water depth. Figure 10 presents the moment-curvature curves for different initial applied pressure given as a percentage of the collapse pressure. Figure 7: Long sandwich pipe with longitudinal mesh refinement 5 Copyright © 2005 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 51 Collapse Pressure, Pco (MPa) 60 Pressure, P (MPa) 50 40 Pd 30 20 10 0 0 0.2 0.4 0.6 0.8 1 50 49 48 47 46 45 1.2 44 Curvature, K (1/m) 0 Figure 9: Ultimate strength under pressure and bending 0.2 0.4 0.6 0.8 1 1.2 Ovalization (%) Figure 11: Influence of the initial ovalization on the collapse pressure of the sandwich pipe 300000 Moment, M (N.m) 250000 These results show the importance of having a good fabrication control and also that the pipe must be stored and installed appropriately to avoid excessive out-of-roundness. 200000 150000 APPARATUS FOR EXPERIMENTAL SIMULATION OF THE REELING INSTALATION PROCESS 100000 50000 0 0 0.2 0.4 0.6 0.8 1 1.2 Curvature, K (1/m ) 0%Pco 20%Pco 40%Pco 60%Pco 80%Pco The apparatus for the simulation of the reeling process allows the pipe to undergo different longitudinal deformations when reeled on a cylindrical surface, reeled-off and rectified using reverse bending. Figures 12 and 13 shows a schematic view of the reel-on and the subsequent reel-off process on a typical reel-lay ship, as indicated in Figure 14. Figure 10: Moment-curvature curves for different initial pressure acting on the sandwich pipe Each point representing the combined load for ultimate strength in Figure 9 is related to the maximum moment for each initially prescribed pressure (0, 20, 40, 60, 80% of collapse pressure) in Figure 10. The design point (Pd) is calculated adopting a 1.33 safety factor over the collapse pressure and a maximum longitudinal strain according to API of 0.5%. Expression (8) relates curvature K and longitudinal deformation ε: ε = KRe (8) INFLUENCE OF THE COLLAPSE PRESSURE OVALIZATION ON Figure 12: Reel-on process, bending THE The same mesh from the previous analyzes has been also employed to evaluate the influence of different magnitudes of initial ovalization (0.1, 0.2, 0.5 and 1%) on the collapse pressure. The results presented in Figure 11 indicate a reduction of 12.1% in collapse pressure for an ovalization increase from 0.1 to 1%. Figure 13: Reel-off process, unbending/rectification 6 Copyright © 2005 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Figure 16: Extended actuators (bending) Figure 14: Typical reel-lay ship [6] The load history (moment-curvature curve) experienced by a given section during the whole installation process is shown in Figure 15, according to stages 1 to 5 indicated in Figures 12 and 13. Figure 17: Retracted actuators (rectification) REELING PROCESS SIMULATION MODEL Figure 15: Moment-curvature loading history during the reeling process [6] As the maximum curvature imposed on a pipe occurs in the first stage (1-2), and no plastic deformation is induced in the intermediary cycle 2-3-4, the laboratory test is carried out for stages 1-2-3-5-1. Thus, in order to simplify the test without affecting the final result, the load stage 3-4 (bending on the aligner) is not reproduced by the apparatus. The pipe is placed between the bending formers, which are connected by hydraulically operated rods, forming a single piece that involves the pipe. Pipe ends react against rollers, allowing it to rotate and slide longitudinally, but preventing it from moving transversally. In these experiments, the curvature radii of bending and rectification formers are 6 and 30 meters, respectively. Figures 16 and 17 show the final movement positions of the formers when the hydraulic actuator is extended and retracted, respectively. The reeling installation process was originally simulated for single wall pipe and the results compared with those from full scale tests [7]. In this paper the model was adapted to three layer sandwich pipe. It was necessary to model the sandwich pipe with the same length as used in the bending apparatus to make it possible to analyze the complete stress-strain and moment-curvature curves for the whole process. Half-length and half-transverse section symmetry conditions were considered, Figure 18. Table 1 indicates the pipe geometric properties used in the analysis. Initial ovalization of 0.2% is adopted. Half-lengths of analyzed pipe and formers are 2600mm and 2000mm, respectively. Figure 18: Step 0 – Undeformed configuration The model was implemented with C3D27 elements, the same employed in previous limit strength analyzes. Kinematic hardening plasticity model was used in addition to the isotropic 7 Copyright © 2005 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use model, as explained earlier, to simulate the Bauschinger effect for the steel layers. The bending and rectification formers were modeled through analytical surfaces. On the external pipe contact surfaces were generated to simulate the interaction forces. The curvature was induced by applying displacement at the pipe end and keeping the bending formers fixed. Reference nodes coupled with nodes on the section plane were used both at the pipe end and on the half-length symmetry section. Displacement in y direction was applied at the pipe end to induce pipe curvature on the formers. On pipe mid-section (x=0), restrictions associated with the symmetry condition were applied to the reference node, which were then transferred to nodes on the transverse section. Reaction moment around axis z could then be monitored and the moment-curvature curve plotted for the pipe mid-section. Curvatures were obtained by the rotation angle of the reference node at the pipe end. Complete bending, rectification and elastic spring-back process (stages 1-2-3-5-1) was modeled through four load steps, with external pressure applied after the rectification, as shown below: 1) Displacement of the pipe end from the initial position (step 0) to the bending former; 2) Reverse displacement to the rectification former; 3) Displacement of the rectification former away from the pipe; 4) Increments of external pressure applied up to the collapse failure. All the mentioned displacements are applied along the y axis. Figures 19, 20 and 21 show the deformed configurations according to the load steps. Single line pipe indicated in the figures represents the pipe undeformed configuration. Figure 21: Step 3 - Displacement of the rectification former away from the pipe Results of the collapse pressure of the pipe before and after the reeling process were obtained. The results of the maximum equivalent plastic strain and stress for each load step are presented in Table 5. Table 5: Results for reeling installation process, load steps 1 to 3 Variable Step 1 Step 2 Step 3 ε eqp (10-2) 2.149 4.076 4.076 σ eq (MPa) 487.6 449.8 311.5 ∆ (%) 0.85 0.62 0.56 It can be seen that in the steps 1 and 2 the stresses remained high and only decreased after the pipe was released (step 3), when a residual stress of 311.5MPa was recorded. Equivalent plastic strain reached its highest value during the second load step. At step 4 the collapse due to external pressure occurred under 47.58 MPa compared with 48.24 MPa before the reeling simulation. It represents a collapse pressure reduction of only 1.4%, despite an ovalization increase from 0.2% to 0.56%. Symmetrical distribution of the equivalent plastic deformations accumulated after the reeling process is shown in Figure 22. y x Figure 19: Step 1 – Displacement of the pipe end over the bending former Figure 22: Distribution of the equivalent plastic strains at the end of step 4 Figure 20: Step 2 – Reverse displacement of the pipe end over the rectification former The pipe deformed configuration under external pressure with collapse failure initiation at the pipe mid-section is shown in Figure 23. 8 Copyright © 2005 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Residual longitudinal stress of 89.55MPa and strain of 0.13% were observed in the outer pipe (larger cycle) in Figure 25. The inner pipe has only a small strain value, with residual stress close to zero. CONCLUSIONS Figure 23: Pipe deformed configuration at collapse pressure Figure 24: Moment-curvature for steps 0-1-2-3 of the bending and rectification process for the sandwich pipe Figure 25: Longitudinal stress-strain curves for the radially-outermost upper points of the steel pipes The reverse moment attains its maximum during step 2, when the pipe is back to its initial position (reel-off) and starts to be rectified, Figure 24. At the end of the process the pipe is released (step 3). In addition, a final curvature close to zero is obtained, which suggests that rectification of the residual curvature was successfully performed. It can be seen that the moment values necessary to induce and reverse pipe are different, due to the use of kinematic hardening plasticity models to include the Bauschinger effect. Sandwich pipe suitable for 3,500 meters water depth application was analyzed for limit strength under external pressure and bending. In addition, the effects of the reeling installation method on the collapse pressure were also evaluated. Two different numerical models were employed in the computer simulations, full length pipe and ring section. For ultimate strength analyzes of external pressure combined with longitudinal bending the ring section was employed. The reeling installation process, i.e. bending over a rigid surface and rectification through reverse bending were simulated using longitudinal symmetry for the pipe full length. Afterwards, the sandwich pipe was submitted to increments of external pressure up to the collapse failure. Some of the main conclusions are outlined below: ¾ Based on the limit strength curve it was possible to estimate a 3,500m water depth application considering a realistic initial ovalization of 0.2%, 1.33 safety factor for the external pressure and maximum longitudinal strain during installation limited by API yield strain; ¾ The ovalization increase from 0.1 to 1% implied in collapse pressure reduction by 12%; ¾ The collapse pressure obtained for the sandwich pipe after the reeling process has not been significantly reduced (1.4%) if compared with the pipe collapse capacity before reeling, despite ovalization increase from 0.2% to 0.56%; ¾ Reeling simulation indicates significant residual equivalent strains of 4.08%, although the longitudinal strain of 0.13% is still in the elastic range; ¾ The reverse moment for the pipe curvature rectification is, approximately, 87.5% of the moment to bend initially the pipe over the reel, confirming the influence of the kinematic hardening plasticity model which incorporates the Bauschinger effect; ¾ The maximum equivalent stress of 487.6MPa measured during the reeling simulation is slightly higher than the yield stress for the X-60 steel assumed for the metal layer. The results presented in this paper are associated with a research work in progress about the offshore installation and operation of sandwich pipes. Although the results so far achieved have attended the initial expectations, further experimental and numerical studies for full scale prototypes have been planned to build up the necessary confidence for the use of sandwich pipes in ultra deepwaters. ACKNOWLEDGMENTS The authors would like to acknowledge the Brazilian Ministry of Education / CAPES for the financial support to the first author and TENARISCONFAB for the help during the definition of the prototype geometric properties. 9 Copyright © 2005 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use REFERENCES [1] Grealish, F., Roddy, I. (2002), “State-Of-The-Art on Deep Water Thermal Insulation Systems”. 21st International Conference on Offshore Mechanics and Arctic Engineering, Proceedings of OMAE´02. Oslo, Norway. [6] Pasqualino, I.P., Silva, S.L., Estefen, S.F. (2004), “The Effect of the Reeling Laying Method on the Collapse Pressure of Steel Pipes for Deepwater”. 23rd International Conference on Offshore Mechanics and Arctic Engineering, Proceedings of OMAE´04. Vancouver, Canada. [2] Netto, T.A., Santos, J.M.C., Estefen, S.F. (2002), “Sandwich Pipes for Ultra-Deep Waters”. 4th International Pipeline Conference, Proceedings of IPC´02, Aberta, Canadá. [7] Daly , R., Bell, M. (2001), “Reeled Pipe In Pipe Steel Catenary Riser”. 20th International Conference on Offshore Mechanics and Arctic Engineering, Proceedings of OMAE´01. Rio de Janeiro, Brasil. [3] Jian, S., Cerqueira, D.R., Estefen, S.F. (2003), “Thermal Analysis of Sandwich Pipes with Active Electrical Heating”. 22nd International Conference on Offshore Mechanics and Arctic Engineering, Proceedings of OMAE´03. Cancun, México. [8] Odegard, J., Christian, T., Kjell, O. H., Bàrde, M. (1998), “Installation of Titanium Pipelines by Reeling Strain Analysis and Material Properties”. 17th International Conference on Offshore Mechanics and Arctic Engineering, Proceedings of OMAE´98, Lisbon, Portugal. [4] Jian, S., Cerqueira, D.R., Estefen, S.F. (2004), “Simulation of Transient Heat Transfer of Sandwich Pipes with Active Electrical Heating”. 23rd International Conference on Offshore Mechanics and Arctic Engineering, Proceedings of OMAE´04. Vancouver, Canada. [9] Pasqualino, I.P., Pinheiro, B.C., Estefen, S.F. (2002), “Comparative Structural Analyses Between Sandwich and Steel Pipelines for Ultra-Deep Water”. 21st International Conference on Offshore Mechanics and Arctic Engineering, Proceedings of OMAE´02. Oslo, Norway. [5] ABAQUS User´s and Theory Manuals (1998). Version 6.4, Hibbitt, Karlsson, Sorensen, Inc. 10 Copyright © 2005 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use

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