Reeled Mechanically Lined Pipe: Cost Efficient Solution for Static and Dynamic Applications in Corrosive Environment

Reeled Mechanically Lined Pipe: Cost Efficient Solution for Static and Dynamic Applications in Corrosive Environment Sylvain Denniel Technip Aberdeen UK sdenniel@technip.com Tomasz Tkaczyk Technip Aberdeen UK ttkaczyk@technip.com Aurelien Pepin Technip Aberdeen UK apepin@technip.com Abstract Mechanically lined pipe is composed of carbon steel and a thin internal layer of corrosion resistant alloy (CRA). In contrast to clad pipe, where carbon steel and CRA are bonded metallurgically, the two layers are adhered by means of interference fit, in a lined pipe. Although susceptible to liner wrinkling when bent, lined pipe offers more attractive procurement cost and shorter lead time. To address an increasing demand for subsea transport of corrosive constituents, Technip proposed a design of lined pipe, which enables reeled installation under atmospheric pressure without the risk of wrinkling of corrosion resistant lining. In contrast to the method where internal pressure is used to enable reeling, standard reeling operation procedures are used for safe and reliable installation of the Technip lined pipe with an appropriately selected liner thickness. The Technip design was patented and qualified following the procedures in DNV-RP-A203. The solution, which was endorsed by DNV in 2011, will be installed for the first time in the UK sector of the North Sea in 2013. Lined pipe is a superior alternative to chromium stainless steel pipe for deep water high pressure/high temperature steel catenary risers (SCR). The alloy liner materials have an excellent fatigue performance in sour environments. In addition, lined pipe does not suffer from significant strength reduction at high temperature. To extend the field of application of reeled lined pipe, large scale bending and fatigue trials have been carried out on 10.75 pipes with alloy 625 liner. It has been shown that the Technip lined pipe is suitable for demanding SCR service post reeled installation. This paper discusses the qualification of lined pipe for reeling and the follow up qualification programme, where reeled lined pipe was qualified for SCR service. In addition, the cost advantage of line pipe compared to other corrosion resistant solutions is discussed. Introduction The development of subsea fields, characterised by the presence of highly corrosive fluids, is a worldwide growing trend, both in shallow and deeper waters. Chromium stainless steel Page 1

pipes, e.g. 13%Cr to 25%Cr, are typically specified, leading to a significant increase in the procurement cost of the pipe work. Bi-metallic pipe, where an external carbon steel pipe hosts a thinner internal Corrosion Resistant Alloy (CRA) liner, is an attractive alternative for a range of technical and economic reasons. The two main bi-metallic pipe products available are metallurgically bonded clad pipe, commonly called clad pipes, and mechanically lined pipes, known as lined pipes. Each product is strongly differentiated from the other by their process of fabrication, which also leads to differences in procurement costs and delivery schedule. Lined pipe generally emerges as the most cost attractive design solution, as will be discussed later in this paper. It is no surprise that a subsea track record of such pipe, installed with low strain installation techniques such as S-lay, does exist. On the other hand, guaranteeing integrity of the product using the faster Reel-lay method, has long been perceived as a challenge within the industry, due to concern of liner wrinkling, when the pipeline is subjected to multiple plastic deformation events. In the past years, Technip developed and qualified a design suppressing such a risk, while ensuring that the operation still occurs at atmospheric pressure, as for any other pipeline installation operations. The design has been endorsed by DNV in accordance to the RP-A203 qualification process for new technology [1], for both static flowlines and dynamic SCR applications. The reel-lay installation of the first static flowline is scheduled for summer 2013 in the UK sector of the North Sea. Introduction to Mechanically Lined Pipe A mechanically lined pipe is a bi-metallic product, which includes an external carbon steel pipe and an internal CRA liner. The essential function of the host carbon steel pipe is mechanical strength. The function of the liner is corrosion resistance. There is a wide range of liner materials that can be specified, as a function of the corrosive nature of the fluid to be transported, i.e. alloy 316L, 904L, 825 or 625. Lined pipes have no metallurgical bond but an interference stress between the carbon steel host pipe and the CRA pipe, which is induced by the manufacturing process. The process of fabrication of a mechanically lined pipe is relatively fast once both host pipes and internal liners have been procured. A number of companies are commercially proposing such product and although their procedure of fabrication can differ in the details, the general principle of manufacture remains the same: First the liner is inserted inside the host pipe. The insertion gap between the carbon steel pipe bore and the liner outer surface is a parameter of importance. Certain manufacturers will choose at this stage to seal weld the liner to host pipe interface at both ends. The liner is then subjected to a controlled expansion operation, using either mechanical tools or hydro-forming, during which the liner is plastically deformed. In a first step, the liner outer surface comes in contact with the host pipe inner surface. As the liner diameter is further increased, the host pipe diameter is also expanded. Most manufacturers control the expansion to ensure that the outer pipe only experiences elastic deformation. However, some consider a slight plastification of the outer pipe, as part of a displacement controlled process. Others have also included heating of the outer pipe prior to expansion, as part of its general process. This allows higher interference fit to be achieved. Page 2

The expansion operation is then stopped and both liner and outer pipe are relaxed. However, as the internal liner has been subjected to plastic deformation, its diameter remains in excess of the initial inner diameter of the host pipe. This results in residual hoop stresses, tensile in the host pipe and compressive in the liner. As a consequence, an interference stress between the two metallic layers develops. The magnitude of this mechanical bond is a function of the relative strength of the host pipe and the liner as well as the expansion process undertaken by the manufacturer. Figure 1 illustrates the general principle of expansion and relaxation. The end termination of the bi-metal pipe can be completed by end seal welds between the two layers. However, it is preferred to terminate the pipe ends by an overlay section long enough to enable girth weld repair and inspection, during the pipeline assembly phase. Alloy 625 is typically considered for overlay. A particular point of focus is the liner to overlay transition, for which details of design, welding procedure and inspection method can influence the performance of the product in dynamic SCR conditions. (a) (b) (c) Hoop stress Host pipe Liner A B Hoop stress A B Hoop stress A B C Diameter C Diameter C Diameter Figure 1 Comparison of Different Manufacturing Processes for Mechanically Lined Pipes The manufacturing process must ensure that there is no moisture ingress into the annulus at any stage as this may lead to an explosive phase change during pipe coating at a high temperature, and consequently, collapse of the liner. Once delivered at the spoolbase, the pipe joints are welded together into stalks and inspected. Well established procedures are used which have already been implemented on plastic lined pipe stalk tie-in welds on 14 reeling projects to date or on reeled clad pipe projects [2, 3, 4], over approximately 60 km of cumulative length to date. Development of a Reeling Friendly Design at Atmospheric Pressure When deciding to develop and qualify a lined pipe design that is compatible with reel-lay installation, the ability to pipe-lay at atmospheric pressure was identified as a key operational requirement. Initial development efforts considered reeling of the lined pipe while an internal pressure was maintained in the pipeline bore as a liner wrinkling prevention measure and a methodology was subjected to a patent application. However, added constraints on operational schedule, logistics and reliability, as well as increased vessel pay-load, were considered as significant enough to prioritise a working solution at atmospheric pressure. The qualification process for new technology, as outlined in DNV-RP-A203 [1], was implemented in order to qualify lined pipe products for reeling installation. According to this Page 3

process, a basis of design for qualification is first defined, clearly specifying the boundary of application of the technology for qualification. The exercise continues with the review of technology readiness of different components of the system. A philosophy of design for qualification is then defined and a detailed qualification programme is finally established. Once the planned analyses and testing scope are complete, results and findings are summarized in a technology qualification report. A certificate of fitness for service can be issued by DNV following successful review of that report. For the purpose of the assessment of the readiness level of the technology, the system was divided into two main zones. The weld region includes the girth weld and the lined pipe ends, terminated with clad overlay. The pipe body region includes the liner to overlay interface, as illustrated in Figure 2. Pipe body region Weld region Pipe body region Host pipe Liner Liner to clad transition Girth weld Clad overlay weld Figure 2 Longitudinal Section of Mechanically Lined Pipe Weld Regions and Pipe Body Regions It is not intended in this paper to review in detail all the aspects of the readiness levels, which are discussed in [4]. However the following major findings of this review included: The weld region of lined pipes was considered to be pre-qualified in principle for reeling. As previously mentioned, there is a strong track record of welding, inspecting and reellaying of bi-metallic pipes. In particular, the overlay profile of a plastic lined pipe at tiein points is identical to the pipe end profile of a mechanically lined pipe. Liner wrinkling is the main issue. Outwith the weld region, the lined pipe main body was not considered qualified for reeling at the time of the technology review. Although most aspects listed previously were considered as qualified, wrinkling of the liner during high strain multi cycle bending was identified as the principal concern. An improvement of the existing lined pipe design was required to allow compatibility with the reel-lay installation method. This design improvement is the essential novel element of the qualification exercise. Based on the above, a main focus of the initial qualification effort was to develop a methodology to suppress this risk of liner wrinkling. In order to develop a philosophy of design that would enable reelability of the mechanically lined pipe, it was first necessary to fully understand the wrinkling phenomenon, its driving parameters and more importantly, the parameters preventing it. In order to better comprehend the liner wrinkling phenomenon, it was decided to carry out an experimental program in parallel to a series of Finite Element (FE) based analyses. The Page 4

testing program essentially consisted of subjecting a series of standard mechanically lined pipe strings, i.e. not specifically designed for reeling, to conditions significantly harsher than a standard reeling operation. The appearance of wrinkles of non-negligible size was therefore expected. Calibration of the FE model, inclusive of all steps of manufacture of the mechanically lined pipe, was the basis for a sensitivity analysis to define a reeling friendly design. Experiment The pipe strings prepared for the initial experimental phase were 12.75 OD x 15.9 mm wt, X65 API grade host pipes, with 2.8 mm thick alloy 316L liners. It is to be noted that all test strings were subjected to a thermal cycle at 250ºC during coating application. This operation is known to reduce the interference stress generated during the lining process. Some of the test strings also included a CRA girth weld at mid length to investigate the liner behaviour in the vicinity of the weld. The testing sequence included the following steps: The test pipes were subjected to five symmetrical bending/reverse bending cycles, which is equivalent to two and a half times the full reel-lay sequence. A 7.84 m radius former was selected, despite the smallest radius on the Technip reel-lay vessel fleet being 8.23 m. As planned, this bending regime produced high magnitude wrinkles, which were closely monitored during the bending trial using a laser profiler. The results were converted into 3D liner impressions as shown in Figure 3 (a). No visible wrinkle growth was observed in the vicinity of the central CRA girth weld Pressure testing of the pipe strings was then carried out, in order to assess the effect of internal pressure on wrinkle height. This was measured post hydro-test, using the same surveying technique that was used during the bending trial phase. An internal hydro-test pressure of 39 MPa reduced the wrinkle height but was unable to remove wrinkles to a level whereby the liner returned to its original geometrical condition, see Figure 3 (b). Although the liner was subjected to severe plastic deformations during multi cycle bending and subsequent pressurisation, post-mortem examination revealed no damage to the liner (e.g. no fracture, cracking or signs of plastic collapse). Figure 3 Comparison Laser 3D Visual Profiles a) Post Reeling (Left) and b) Post Hydro-Test (Right) Page 5

Finite-Element Analysis All steps of the manufacturing process, including coating application simulation, and subsequent multi cycle bending and pressurisation of lined pipes, were simulated using ABAQUS. The model, as described in detail in [4], used first-order shell elements with reduced integration and kinematic hardening material models to provide a very accurate prediction of the lined pipe behaviour. Comparison of Full Scale Test Results and FE Analyses The FE predictions of wrinkle height at the 6 and 12 o'clock positions were compared with the test results. As shown in Figure 4, the FE predictions (indicated with a solid line) were in remarkable agreement with the measured wrinkle heights, hence validating the FE model. The 6 o'clock position is in compression, when on the reeling former, and in tension, when on the straightening former. In contrast, the 12 o'clock position experiences tension and compression on the reeling and straightening former, respectively. There are four stages in each bending cycle, i.e. bending on the reeling former, relaxing, bending on the straightening former and relaxing, totalling 20 stages in the bending trial. Stage 21 in Figure 4, corresponds to a condition after the hydro test. Up to three wrinkles along each test string were measured at a given stage using the laser profiler. The heights of wrinkles measured in the test strings with and without the central CRA girth weld are indicated in Figure 4 with open and closed symbols, respectively. (a) 24 21 2 nd cycle 3 rd cycle 4 th cycle 5 th cycle (b) 24 21 Straightening Straightening Wrinkle height (mm) 18 15 12 9 6 1 st cycle Wrinkle height (mm) 18 15 12 9 6 1 st cycle 3 0 Straightening Straightening Straightening 0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 Stage 3 0 2 nd cycle 3 rd cycle Figure 4 Wrinkle Height Profiles 6 o clock (a) & 12 o clock (b) 4 th cycle 5 th cycle 0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 Stage The wrinkle height, h, increases rapidly at the 6 o'clock position when the first compressive strain is applied during bending on the reeling former (stage 1 in Figure 4). During the subsequent bending of the pipe string on the straightening former (stage 3 in Figure 4), h is reduced at the 6 o'clock position due to tensile straining, while h grows rapidly at the 12 o'clock position, which is in compression. This behavioural trend is also observed during subsequent bending cycles. However, after the second bending cycle is completed (stage 8 onwards), only a small difference between wrinkle heights at the maximum compressive strain is observed in subsequent bending cycles, suggesting a stabilisation of the wrinkling phenomenon. The maximum wrinkle height after five bending cycles (stage 20) was measured as 7 mm at the 6 o'clock position (last in tension) and 20 mm at the 12 o'clock position (last in compression). The internal pressure, which was applied following bending, reduced h to a maximum of 8 mm and 1.5 mm at the 12 and 6 o'clock positions, respectively. It appears, therefore, that small wrinkles can be removed by a typical hydro test pressure. Page 6

Larger wrinkles, however, cannot be flattened out even by internal pressure close to the pipe burst capacity. Further parametric studies with the validated FE model allowed identification of critical parameters which may influence the onset of wrinkling and subsequent wrinkle growth. Under multi cycle plastic bending, the wrinkle size may depend on liner dimensions (D/t ratio), nominal bending strain (ε b ), number of bending cycles. The interference stress appears to play a role in the first bending cycle. However, as it reduces as a result of plastic deformation, its influence decreases sharply in subsequent bending cycles. To a certain extent, it can be argued that the interference stress only delays the onset of wrinkling rather than prevents it. Design Philosophy A design improvement of the liner was established on the basis of the following philosophy to enable reelability of the lined pipe at atmospheric pressure. Within the pipe body, wrinkle initiation is avoided by choosing an appropriate D/t ratio of the liner. An extensive parametric study was carried out using the validated FE model to examine the effect of host pipe and liner geometry, yield strength variations, wall thickness variations and manufacturing/installation parameters on liner wrinkling. This philosophy led to the definition of a formula establishing an appropriate D/t ratio of the liner to avoid the risk of wrinkling regardless of the lined pipe manufacturing method. Although high interference stress is beneficial, the novel design adopted the prudent approach not to rely on it. Firstly its magnitude may vary significantly between pipe joints and within each pipe joint. It is therefore difficult to guarantee a high enough interference stress value for design purposes. Furthermore, as previously discussed, the FE parametric studies have shown that interference stress may be negligible, as first relaxed during coating application and then further reduced during subsequent plastic straining events. The philosophy of design and associated lined pipe design improvement, as presented above, is now the object of a patent application. A comprehensive full scale qualification programme has been completed successfully on a range of pipe strings of different sizes, liner grades and produced by different suppliers. The robustness of the approach has been fully demonstrated and a statement of fitness for service was issued by DNV in April 2011. The highlights of this qualification programme are presented in the following section. An Extensive Qualification Programme for Both Static and Dynamic Applications Static Seabed Applications To validate experimentally the benefits of the optimised liner thickness, reeling simulations were undertaken on 6.625, 8.625, 10.75 and 12.75 OD lined pipes, which were manufactured to Technip s specifications by a number of lined pipe suppliers. The effect of coating application on interference stress was accounted for by systematically subjecting test pipes to a thermal cycle, representative of conditions experienced by a pipe joint during application of external coating at a temperature of up to 250 o C. Pipes with and without girth weld were tested to assess the effect of the CRA girth and clad overlay welds on the liner behaviour in its vicinity. Although reeled lined pipes are normally subjected to two bending cycles during installation, up to seven bending cycles were applied during a series of reeling simulation campaigns. Here, a bending cycle is defined as bending onto a reeling former, Page 7

followed by bending onto a straightener former and finally relaxing. No visible wrinkles or any other form of liner damage were observed visually nor by laser measurement during bending of the lined pipes detailed in Table 1. Inspection included visual inspection at different steps of reeling, laser inspection over a representative 1m long section of each lined pipe and post-mortem inspection of cut pipe sections. Pipe Diameter (mm) Host Pipe Material Liner Material Former Radius (m) ε b Number of Cycles 6.625 168.3 X65 Alloy 904L 8.23 ± 1.00% 3 to 6 8.625 219.1 X60 Alloy 316L 8.23 ± 1.31% 5 to 7 10.75 273.1 X65 Alloy 625 9.75 ± 1.38 % 4 P-SC 10.75 273.1 X65 Alloy 625 8.23 ± 1.63 % 5 12.75 323.9 X65 Alloy 316L 10.50 ± 1.52% 3 to 6 Table 1 Summary of Reeling Simulations Test Series P = perpendicular reeling simulation, see description later in this paper SC = Strain capacity limit test, see description later in this paper In one particular instance, it was decided to undertake repeated reeling cycles of a pipeline string, whereby an obstruction was applied onto the reeling former to intentionally generate lifting of the pipeline from the former and leading to a localised strain concentration, see Figure 5. The intent of such testing was to reproduce the consequence of the statistically very unlikely event of two adjacent pipe joints of significant difference in plastic moment capacity (a relevant mismatch feature, which would be outwith normal project acceptance criteria). Following each bending cycle applied, an increasingly more pronounced local obstruction was positioned onto the former. Such incremental tests were conducted until first signs of wrinkling were obtained. In the context of that test, these were only witnessed once the host pipe was also experiencing early sign of buckling. This experiment highlighted the robustness of the design methodology as it demonstrated that wrinkles could only be envisaged in a situation where the host pipe was itself about to buckle. The standard design practice of the host pipe negates risks of such a buckling event to occur. As a matter of fact, as will be discussed later in this paper, the design cases where a mechanically lined pipe is most likely to be selected are for pressure governed cases. In this instance, the buckling of the host pipe is very unlikely as it is much thicker than the minimum reelable wall thickness. The authors also decided to test the scenario of a pipeline first spooled onto a storage facility, used for transport from the spoolbase to a site closer to the offshore field, and subsequently trans-spooled onto the reel-lay vessel for installation. To conservatively simulate this approach, which was implemented on the Snohvit field in 2004, the pipeline test strings were subjected to two bending cycles in one plane, followed by two other bending cycles in the perpendicular plane. No wrinkle was found at the end of this exercise. Page 8

Figure 5 Simulation with Local Obstruction Another reeling simulation was undertaken to capture the margin of the lined pipe design in terms of nominal bending strain. The pipes were subjected to repeated reeling simulations, with reeling formers radii incrementally decreasing, from the storage drum radius of 9.75 m for which the pipe was designed. Repeated reeling cycles were carried out down to a former of 5.5 m, until first signs of onset of wrinkling were witnessed. The nominal reeling strain experienced was 1.77 times the nominal design reeling strain. It is worth highlighting that the test strings used for this experiment had previously experienced not only the perpendicular reeling simulation described in the previous paragraph but also a full scale fatigue test which was stopped, once the number of cycles subjected exceeded the targeted DNV-RP-C203 [5] C curve threshold. This fatigue test programme is discussed later in this paper. In Service Simulation It is known that the thermal expansion coefficient of the CRA liner material is typically greater than that of the carbon steel pipe. This implies that when subjected to temperature, the liner would have the tendency to expand more than the host pipe and therefore that a longitudinal compression stress may develop in such liner. It was decided to demonstrate that the reeling friendly mechanically lined pipe design was not at risk of wrinkling in the critical in-service scenario of rapid depressurisation of a pipeline experiencing a lateral buckle, for which the temperature loading is still exercised. After successful installation tests, where the 6.625 and 12.75 lined pipes were subjected to thermal cycles and reel-lay installation simulation as previously described, all specimens were subjected to a hydro-test to simulate pre-commissioning. Each test string was then tested in a four-point bending configuration as shown in Figure 6. A total of 20 bending cycles with a maximum strain of 0.5% were applied to each pipe to simulate the scenario of a pipeline subjected to several lateral buckling events. During cyclic bending, the test strings were subjected to internal heating so that the liner inner surface remained at 150 o C for the duration of the test. Page 9

Each test cycle comprised of the following steps: Apply an internal pressure of 10 bars, Subject the pipeline to a bending strain of 0.5% at the pipe intrados Unload the pipe to a bending strain of 0.3 % Release the internal pressure to atmospheric level the pipe to a straight configuration. Internal inspection of the liner was carried out every five cycles and no indication of wrinkle formation was ever found. Figure 6 In-Service Testing (Left) and Liner Profile Post Testing (Right) No Wrinkle Generated Dynamic Riser Applications Reel-lay can be an attractive rigid riser installation method because the speed of installation minimises the time of the pipe-lay vessel in the vicinity of the topside, but also because the full riser section can be assembled and integrity tested in a factory-like environment, onshore, away from the critical path. As such the reel-lay method has been selected to install a range of dynamic riser systems such as Steel Catenary Risers (SCR), Lazy Wave Steel Riser (LWSR) or Free Standing Hybrid Riser (FSHR), where a vertical rigid riser section, tensioned by a subsea buoyancy can, is connected to a dynamic flexible riser, which hangs from an FPSO. Technip began qualification of reeled SCRs in 1997 and installed the first reeled SCR in 2001. More than 30 reeled SCRs have been installed by both the Apache and Deep Blue reel-lay vessels since. Gray et al. [6] compiled the list of the first 25 reeled SCRs over 15 projects. The performance of this data-set is in line with or better than the BS7608 (1993) class D fatigue curve [7] which is equivalent to the DNV-RP-C203 [5] class D curve (up to 10 7 cycles). Internally clad pipes, typically girth welded using nickel-based alloys such as alloy 625, have been used in the industry for dynamic riser applications. Unpublished test results obtained earlier by Technip are in agreement with other industry publications [8, 9] that the fatigue strength of clad SCRs is enhanced by the use of such a nickel-based alloy consumable, which provides smooth root profile and lower crack growth rate. It was therefore a logical step to investigate the dynamic endurance of a reeled mechanically lined pipe for a dynamic riser design. Fatigue / Crack Initiation Sites SCRs are exposed to a combination of axial and alternating bending loads. Due to a radial gradient associated with the bending stress, the outer surface of a pipeline experiences larger Page 10

stress than the inner surface at the same location. In addition, the outer surface has the considerable geometric discontinuity of a weld cap acting as a stress raiser. Therefore, a weld cap toe, indicated A in Figure 7, becomes a likely location for a fatigue crack initiation to occur. Removing the weld cap by grinding and flapping to create a flush surface improves the fatigue endurance of SCR welds [6]. Girth weld Pipe (a) A (b) A Host pipe C B Girth weld Liner D B Clad overlay weld Figure 7 Probable Crack Initiation Sites - (a) Standard CMn Pipeline / (b) Mechanically Lined Pipe Due to a potential stress concentration, another likely location for crack initiation is a weld root toe, indicated B in Figure 7. Hence, the misalignment of the internal SCR surfaces (i.e. hi-lo) across a weld is preferably controlled by pipe sorting or internal machining of pipe ends. Typically, a maximum hi-lo of 0.5 mm is targeted for SCR welds [6]. Welds with low defect levels can be achieved using a readily available manual gas tungsten arc welding (GTAW) process. This process offers a consistently good weld root profile, which ensures small stress concentration and minimises the severity of inherent inclusions or crack-like discontinuities that may exist at the toes of a weld. The lower productivity GTAW process has, however, since been replaced by the high quality combined mechanised STT/GMAW welding process [6]. In the specific case of a mechanically lined pipe, although many similarities can be drawn with clad pipes, the liner to clad overlay transition locations C and D in Figure 7 have been identified as potential initiation sites for fatigue failures. As there was no published data on the fatigue performance of reeled lined SCRs at the time, the authors undertook a full scale bending and fatigue test programme, in order to determine the fatigue strength and the primary location of fatigue crack initiation in mechanically lined SCRs. The results were first presented in [10] and [12] and are summarised in the following section. Qualification Programme Dynamic Application The authors undertook the programme to qualify lined SCRs for the scenario of a pipeline experiencing reeling cycles in two perpendicular planes, as described earlier in this paper. Subsequently, the specimens were subjected to a full scale fatigue simulation in air, using the TWI resonance rig, at positive stress ratio and for three constant amplitude stress ranges. A more detailed description of the test procedure and results is documented in [10]. For the purpose of these tests, six 10.75 OD carbon steel pipe grade API 5L X65, internally lined with alloy 625 were prepared and were all inclusive of a central CRA weld. Each lined Page 11

pipe section was terminated with a 50 mm clad overlay. SCR quality manual girth welds were prepared in the 5G position using the GTAW process, where 30 weld passes were deposited per weld using Special Metal FM 625 electrodes, manufactured to AWS A5.14 ERNiCRMo-3. The weld caps were ground flush after completion of girth weld. The primary inspection method was radiography. The perpendicular bending simulation was undertaken using a reeling former radius of 8.23 m. For the purpose of the resonance fatigue test, each specimen was pre-tensioned to a constant mean stress of 130 MPa. Three constant amplitude stress ranges, 120 MPa, 170 MPa and 220 MPa respectively, were tested on two pipe strings each, to achieve fatigue lives between 10 5 and 10 7 cycles. With a total of n=6 girth welds and n=12 clad overlay welds, it was possible to set statistically based target S-N curves, in accordance with the methodology derived by Maddox and Schneider, 2000 [11] to assess whether the DNV-RP- C203 [5] class D or class C performance was met for each weld. Fatigue tests were conducted until failure occurred or a specimen reached the run-out condition, defined as the number of cycles required to achieve a target C-Class S-N curve. The fatigue test results, shown in Figure 8, are presented in terms of the number of cycles to failure versus the stress range at the pipe outer surface. Stresses at the pipe outer surface have been calculated by interpolating between strain gauges readings. No stress concentration factor (SCF) has been included. The results in Figure 8 are compared with the DNV-RP-C203 C-mean and D-mean curves. In addition, the DNV-RP-C203 target curves for classes C and D are plotted in the figure. Figure 8 Fatigue Test Results Vs DNV-RP-C203 SN-Curves a) All Failure Locations (outer wall stress range at the position of crack initiation) b) CRA Girth Welds (nominal outer stress range) Most fatigue crack initiations occurred at the liner to clad overlay transition, three of which initiated from a small lack of fusion defect at location C in Figure 7. Once initiated, the embedded cracks grew both in the carbon steel pipe towards the pipe outer surface and in the CRA liner towards the pipe inner surface. Despite a smaller ligament in the CRA liner, the examination of fracture faces revealed that the embedded cracks had first reached the pipe outer surface, followed by their final growth in the direction of inner surface and pressure loss. Therefore, the lined SCR had an internal corrosion layer until the final failure. This can be explained by a lower fatigue crack growth rate of the nickel-based alloy used for the liner [10, 12]. The results for failures initiated at the liner to clad overlay transition are indicated Page 12

by the solid triangles in Figure 8 a). The fourth failure (indicated by a solid circle) initiated from a blemish at the pipe outer surface. Finally, specimens tested at the lower stress range reached the run-out condition without failure, and were used for further reeling simulation at tighter radii, to determine the safety margin in terms of strain capacity, as previously discussed. The fatigue test results for the liner to clad overlay transition in reeled lined pipe lie above the DNV-RP-C203 D target curve and possess a lower slope, see Figure 8 b). Therefore the fatigue strength of a reeled lined SCR is at least as good as the reeled carbon steel SCRs installed by Technip to date [6]. In addition the results of lined SCRs may be significantly better at low stress ranges ( 120 MPa), which would be more representative of actual SCR ranges. The fatigue performance of the reeled alloy 625 girth welds was also found to be excellent, in excess of the DNV-RP-C203 C/2 target. Since none of these welds failed, the fatigue endurance could be considerably greater. This finding is in line with previous experience of single-sided welds in clad pipes, see Wang et al. [8] for example. Furthermore, the internal pressure set for the test, in order to obtain positive R values, was relatively high. In many actual field applications, the pressure would be lower, hence leading potentially to negative R ratios, which would further increase the fatigue endurance of the girth welds for the high stress range, as tested in [8]. The full programme of testing work, alongside a further body of work highlighting the high fracture toughness of alloy 625 girth and clad overlay welds [12] was documented in a technology qualification report submitted to DNV. A certificate of fitness for service was issued in August 2012 by DNV, endorsing the suitability of the improved mechanically lined pipe for reeling for dynamic SCR applications. Cost and Schedule Incentives to Select Mechanically Lined Pipe The high fatigue performance of bi-metallic pipelines with nickel-based alloy girth welds has been discussed at length in the previous section. This is certainly not the only advantage of mechanically lined pipe over other full body CRA pipeline products. As a matter of fact mechanically lined pipe has a number of attractive cost and schedule arguments in its favour when compared with other metallic pipe solutions for the transportation of corrosive fluids. The most obvious attractions are generally lower procurement costs and shorter delivery schedules. Because of the simplicity of the manufacturing process, mechanically lined pipe tends to be significantly cheaper than hot rolled bonded clad pipe, with a cost ratio that can approach the order of two to three. There are a number of lined pipe suppliers available, irrespective of the pipeline outer diameter, whereas the supply of smaller diameter clad pipes is limited. Among the advantages of the mechanically lined pipe over full body CRA pipe is a wider choice of liner grades to choose from, as a function of the specification fluid to which the liner will be exposed. Furthermore, there is no length restriction for lower diameter mechanically lined pipes, in the order of 6.625 OD, whereas CRA pipes may only be available in 6 m lengths in this diameter range, which would increase welding times and costs. Another attraction of mechanically lined pipe over full CRA pipes is the lower derating of the carbon steel host pipe, which provides the mechanical strength, in comparison to CRA grades such as 22%Cr or 25%Cr in particular, when exposed to higher temperature. Page 13

This is of particular relevance as a high proportion of field developments with corrosive fluids often display high temperature and often high pressure specifications, which govern the wall thickness design. Comparing pure procurement cost of a reelable mechanically lined pipe to that of a full body CRA pipe is sometimes not as straight forward as with clad pipe, as the relative cost will be dependant on the wall thickness governing factor, material grade selection and pipeline outer diameter. First a relevant comparison should consider selection of material grade (CRA and liner) on an equivalent basis, principally governed by the corrosion resistance. It is often the case where 13%Cr options are being specified against bi-metallic pipes with the most expensive 625 liner grade, whereas cheaper liner alternative such as alloy 825, 904L or even 316L may suffice. The improved design enabling the safe reeling of a mechanically lined pipe at atmospheric pressure is based on optimising the liner D/t to negate any risk of wrinkling when the pipe is subjected to plastic deformation. This implies that the greater the pipeline outer diameter becomes, the thicker the liner gets, with greater procurement cost repercussions, especially when the higher specification alloy 625 liner material is used. The D/t ratio of the liner will also be installation vessel dependant as the nominal bending strain associated with the reellay vessel is also a factor of consideration. Although there is no mechanical limit to the maximum diameter that one may consider for a reeled mechanically lined pipe, it is perceived that there would be a cost incentive threshold, somewhere between 12.75 and 14 OD, depending on the installation vessel and the liner material Finally, the pipe wall thickness requirement for the pipeline is also a major factor of influence. If the pipeline wall thickness is driven by reeling, i.e. for low pressure containment and shallow water application, then a mechanically lined pipe may not be fully attractive on a pure procurement basis alone. This especially applies for smaller outer diameter pipes, for which the minimum reelable wall thickness is not significantly superior to a minimum specified liner wall thickness. On the other hand, as soon as the specified wall thickness increases, i.e. governed by burst or potentially collapse requirement, then the relative cost of mechanically lined pipe and of a full body CRA pipe is very much at the advantage of mechanically lined pipe. This is especially relevant for higher pressure and temperature pipelines, where temperature de-rating magnifies the difference, as previously discussed. The two last points are illustrated by Figure 9, which plots the relative cost of a CRA pipe to a mechanically lined pipe with a carbon steel host pipe of matching wall thickness. For the example in consideration, an alloy 825 liner material was selected to match corrosive resistance of a 22%Cr solid CRA pipe. The pipeline is designed to be installable from the Deep Blue vessel. Costs for a range of CRA pipeline wall thicknesses, from minimum reelable to thicker HP/HT pipelines are normalised against the equivalent wall thickness mechanically lined pipes, for outer diameters of 6.625, 8,625, 10.75 and 12.75. Page 14

CRA / MLP 2.50 2.00 Min wt Mid range wt HP / HT 1.50 1.00 0.50 6.625 8.625 10.75 12.75 0.00 Figure 9 Normalised Lined Pipe Procurement Costs 25%Cr CRA Pipe Vs 825 Lined Pipe Installation from Deep Blue To further illustrate a combination of the different cost incentives of selecting mechanically lined pipe over a full body CRA pipeline, it is proposed to consider the field case study of a Pipe-in-Pipe 6.625 OD flowline, inside a 10.75 outer pipe, transporting a particularly corrosive fluid at a temperature of 110 o C and a design pressure of 600 bar, to be installed in a water depth of 100 m. On the basis of material corrosion resistance alone, a 22%Cr would be pre-selected for the design. However, closer consideration of the in-service loading resulting from the design pressure and temperature, which affects the mechanical properties of the CRA pipe, leads to the selection of a more expensive 25%Cr material. An 18.3 mm thick pipe wall is necessary to meet the design requirement. It is known that procurement of such small OD, thick walled, 25%Cr is challenging and restriction of supply to 6 m long joints occurs, with unattractive delivery schedules. The joint length limitation implies doubling the number of welds, with obvious consequences on assembly line schedule and costs. An alternative solution is a mechanically lined pipe, for which the cheaper 904L liner material suffices to meet corrosion resistance requirement. In addition, it is possible to reduce the wall thickness of the host carbon steel pipe as it is less affected by temperature derating. More importantly, it is possible to source 12 m long joints with much shorter procurement schedules. When calculating the welded cost difference between the two options, it was established that a saving in excess of 50% is realistic, not taking into account the benefit of reduced project delays. Another advantage is the easier implementation of internally cladded carbon steel bulkheads. Indeed, no dissimilar metal welding issues are encountered, as the flowline interface is purely bi-metallic and the carrier pipe interface is purely carbon steel. Page 15

Conclusions A mechanically lined pipe design, which is compatible with the fast reel-lay method, at atmospheric pressure, is now qualified for both seabed static and dynamic riser service. Following a rigorous review of qualification programmes, certificates of fitness for service have been issued by DNV for both applications in April 2011 and August 2012, respectively. Mechanically lined pipe is an extremely attractive alternative to clad pipe and full body CRA pipes in terms of lead time and procurement costs, especially for higher temperature and pressure applications. As for clad pipe, it was demonstrated that a reelable mechanically lined pipe, welded in factory like conditions at the spoolbase, in accordance with procedures meeting the demand of SCR designs, meets very high fatigue performance. A first industrial application of reeled mechanically lined pipe at atmospheric pressure is scheduled for installation in mid 2013 in the UK sector of the North Sea. There is sufficient qualification evidence to be confident that this technology will be implemented for rigid dynamic riser application in the near future. References [1] Det Norske Veritas, 2001 - Recommended Practice DNV-RP-A203 Qualification Procedures for New Technology. [2] Endal, G., Nupen, O., Sakuraba, M. and Kondo, K., 2006 Reeling Installation of 15" Clad Steel Pipeline with Direct Electrical Heating and In-Line T - Proceedings of the 25 th International Conference on Offshore Mechanics and Arctic Engineering, OMAE2006-92525. [3] Haabrekke, T., Bärs, G., Kvaale, P.E. and Rørvik, G., 2006 Experiences from Welding and AUT of the 15" Norne Clad Pipeline Installed with the Reeling Method - Proceedings of the 26 th International Conference on Offshore Mechanics and Arctic Engineering, OMAE2007-29612. [4] Tkaczyk T., Pepin A. and Denniel S., 2011 Integrity of Mechanically Lined Pipes Subjected to Multi-Cycle Plastic - Proceedings of the 30 th International Conference on Offshore Mechanics and Arctic Engineering, OMAE2011-49270. [5] Det Norske Veritas, 2011 Recommended Practice DNV-RP-C203 Fatigue Design of Steel Offshore Structures. [6] Gray J.F., Howard B., Pieton A., Gallart R., 2009 The Qualification and Continued Evolution of Reeled Steel Catenary Risers - Proceedings of the 28 th International Conference on Offshore Mechanics and Arctic Engineering, OMAE2009-79176. [7] British Standards Institution, 1993 British Standard B7608:1993 Code of Practice for Fatigue Design and Assessment of Steel Structures. [8] Wang H., Widener T., Kan W.C., Sutherland J. and Jones R., 2011 Qualification of Reeled Clad SCR Weld Fatigue Performance - Proceedings of the 30 th International Conference on Offshore Mechanics and Arctic Engineering, OMAE2011-49798. [9] Weir M.S., Kan W.C. and Hoyt D.S., 2005 Inconel Welding for Enhanced Fatigue Performance Proceedings of Deep Offshore Technology Conference. [10] Tkaczyk T., Pepin A. and Denniel S., 2012 Fatigue and Fracture of Mechanically Lined Pipes Installed by Reeling - Proceedings of the 31 st International Conference on Ocean, Offshore and Arctic Engineering, OMAE2012-83050. Page 16

[11] Maddox S.J. and Schneider C.R.A., 2000 Statistical Analysis of Fatigue Test Results fo Validate a Design S-N Curve TWI. [12] Tkaczyk T., Pepin A. and Denniel S., 2012 Fatigue and Fracture Performance of Reeled Mechanically Lined Pipes Proceedings of ISOPE Conference. Page 17