[oc] end fittings for composite risers
TRANSCRIPT
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Development and Qualification of End Fittings for Composite Riser PipeStephen Hatton, Luke Rumsey, Praveen Biragoni, Damon Roberts, Magma Global Limited
Copyright 2013, Offshore Technology Conference
This paper was prepared for presentation at t he Offshore Technology Conference held in Houston, Texas, USA, 69 May 2013.
This paper was selected for presentation by an OTC program committee following review of information contained in an abstract submitted by the author(s). Contents of the paper have not beenreviewed by the Offshore Technology Conference and are subject to correction by the author(s). The material does not necessarily reflect any position of the Offshore Technology Conference, itsofficers, or members. Electronic reproduction, distribution, or storage of any part of this paper without the written consent of the Offshore Technology Conference is prohibited. Permission toreproduce in print is restricted to an abstract of not more than 300 words; illustrations may not be copied. The abstract must contain conspicuous acknowledgment of OTC copyright.
AbstractWhilst much interest is often focussed on the composite pipe body it is often forgotten that a reliable end fitting is a
prerequisite of a successful composite pipe application and further that the design challenge of the end fitting is more
challenging than the pipe itself.
The purpose of this paper is to present the end fitting arrangement for a composite pipe manufactured from carbon fibre and
PEEK polymer. The paper describes the design approach and testing/qualification program employed for the end fitting to
demonstrate reliable application in a critical and structurally demanding application such as for a deep water dynamic riser.
The paper summarises historical design approaches and design alternatives and explains the reason for selecting the proposed
arrangement. It presents the design process used to develop the design and to predict its structural response. It discusses
manufacturing issues and describes the test program conducted to prove the end fitting performance.
Industry focus is often on the pipe rather than the end fitting. However the latter often presents a more difficult design
challenge. Without a reliable design solution and methodology composite pipe application cannot be considered. The
development work presented in the current paper therefore presents an important step towards application of compositetechnology on demanding riser applications.
An end fitting with reliable structural and sealing performance is a prerequisite for the successful application of compositepipe. Historically, this has proven to be a difficult challenge and end fitting performance limitations have been cited for the
slow introduction of composite pipe technology. The current paper presents a new design approach to the problem, made
possible by a unique manufacturing process and a better understanding of the composite metallic interface, which together
allows the end fitting problems to be resolved. The paper describes fundamentals of the design approach, the development
work conducted, manufacturing and qualification testing. The paper discusses the function specification, key design features,FEA approach and results, codes and standards, testing results under combined load conditions and fatigue testing results.
Introduction
Design, fabrication and installation of riser systems for floating production is a complex challenge. Constraints such asmaximum payload capacity and design issues such as internal and external corrosion, fatigue capacity, thermal insulation,
weldability and susceptibility to sour service conditions conspire to make riser design one of the most demanding challenges in
the offshore industry today. The magnitude of this challenge has grown, perhaps non linearly, as water depths have continued
to increase over the last 20 years. Also operating pressures and temperatures have increased, particularly in the last 10 years,
with the focus on presalt and subsalt reservoirs.
Composite pipe technology has been proposed many times in the last 20 years as a potential technology to resolve some of thedesign challenges [1,2]. Composite pipe can bring important performance advantages over steel pipe risers and in some
instances unbonded flexible risers with metallic armor layers. These advantages include:
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Lower weight
Improved fatigue capacity
Corrosion resistance
Higher strain limits
High strength
Ease of deployment
The design advantages offered by composite materials can simplify the problems that riser designers typically face when
working with steel pipe and designing for demanding applications. In such cases it is typical that many of the following design
strategies need to be employed inorder to achieve the levels of performance required:
More sophisticated analysis methodologies
Provision of higher quality design basis assumptions (eg environmental and fluid data)
Increased materials and component testing
Application of higher strength steels
Weld quality improvement (Welding and inspection)
Internal cladding such as inconel or polymer liners
Application of VIV strakes
More compliant riser configurations eg the use of buoyant configurations
Chemical injection for corrosion control and process flow management
Higher capacity installation vessels
Higher capacity host production vessels Lower motion host production vessels
More demanding inspection strategies
More regular replacement strategies
The above design strategies for steel pipe risers have cost and schedule impacts meaning that riser costs can be a significant
percentage of the overall development cost. Unfortunately, the complexity and knock on effect of some of the above issues canbe underestimated or even missed at the early project stages leading to a difference in the as installed riser cost compared to
the predicted cost, leading to project cost and schedule overruns.
Therefore the proposed use of composite pipe brings the potential for some simplification of the riser design process and the
potential for improved performance, reliability and overall cost benefits. Whilst there appears to be a belief across the industry
that composites can deliver these benefits there is also an uncertainty as to whether the technology is sufficiently mature forproject application. One of the key design areas that needs to be better defined is that of the composite pipe end fitting, which
despite previous initiatives remains an area of some uncertainty. This is because the problem of terminating a composite pipewith a metallic end fitting, poses some seriously complex design and manufacturing challenges.
Whilst throughout this paper we use the term Composite, it should be remembered that the term Composite is allencompassing and as such covers technologies from e-glass /polyester structures at one end of the performance spectrum
through to carbon fibre/ PEEK at the other. The final end termination design and reliability are ultimately dictated by the
inherent performance on the Composite choosen for the application in question.
Background and Historical Approach
Composite pipe must be terminated with an end fitting that allows connection of the composite pipe to standard metallic oil
industry interfaces such as API or ANSI flanges or hub connections. The design of these composite end fitting is potentiallymore complex than the pipe itself.
Composite materials are technically more complex to use in design when compared to metallic materials due to their an-
isotropic properties. When such materials are interfaced with metallic components the problem is even more complex. The
difference in structural properties between composite materials and metallic materials, typically steel and titanium, makes thedesign of the interface highly problematic. The main challenge is that steel and composite materials have very different
coefficients of thermal expansion and thermal conductivity and also different stiffnesses and Poissons ratios. Therefore as the
end fitting or pipe is heated or externally loaded the metallic and composite materials respond in a different structural mannerand this can lead to failure of the interface by cracking or disbondment of one material away from the other. As a minimum,
this may cause a failure of the end fittings ability to maintain leak tight integrity. In the extreme, this may reduce the end
fittings structural capacity or fatigue performance. Failure of such interfaces can be progressive, such that on repeated loading
the failure mechanism progresses along the interface until it fails catastrophically. The magnitude and complexity of this
problem is evident from the complexity of the solutions proposed spanning many years.
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TrapLock
The most popular approach has been the Trap-Lock solution patented by Baldwin/Reigle/Drey in 1997 [3]. [1]In this design
the composite material is wound over a metallic mandrel that has been machined with a profile as shown in Figure 1. Each
element of the profile provides the ability to transfer load from the composite into the metallic component via a combination of
tensile and shear mechanisms.
The load path in such a design is dependent on many factors including fiber direction, stiffness, pretension etc. Often a hoop
wound material is laid over the top of the trap lock or an outer steel collar is applied. High levels of preload are required at the
interface to maintain contact between the inner metallic mandrel and outer composite materials under all load and temperatureconditions.
The detail design of the Traplock must prevent the majority of the load from being taken at the first thread. This is achieved
by a combination of preloading of the composite material and optimising the structural stiffness of the mandrel. This preventshigh local stresses and potential for local failure which can progress along the interface on subsequent loading. To avoid this
failure process it is necessary to preload the traplock interface by winding the fibres under tension and/or auto frettage where
the steel mandrel is plastically deformed by internal pressure after winding. This can give good results but remembering thatcomposite materials can creep and initial preloading can be reduced over time, an effect that is accelerated under elevated
temperatures. This effect must be taken into account.
In the Traplock design the profile uses one or more "traplock" grooves in the exterior surface of the mandrel into which thefilament or reinforcements of the composite tube are wound and/or compacted.The axial load is transferred between the
composite tube and the end fitting through bearing on the load-carrying face of the traplock grooves. The surface area of the
load-carrying face is one of the parameters determining the strength of the joint or interface. The bearing area can be increasedby increasing the height of the load-carrying face but the bearing stress which the composite material can support is limited
and therefore the diametrical envelope required by a single traplock groove can become quite large as the height of the load-
carrying face is increased.
The diametrical requirements of the joint can be reduced by the use of multiple traplock grooves, but as discussed above it isimportant to make sure that the grooves carry equal load. This response is optimized by varying the thickness of the root of
the mandrel so that it becomes progressively stiffer from its tip to its other end. This approach further reduces the difference in
stiffness between the two materials and thus its ability to better share the load.
An additional design issue is that of establishing and maintaining a pressure-tight seal between the composite tube and the end
fitting. This is particularly true if the composite tube has an internal pressure sealing liner, which is terminated at the tip of the
traplock mandrel. Designing a reliable elastomeric seal at this location can be a challenge due to practical manufacturing issuesaround the need to lay-up and cure resins, which can easily contact seal surfaces and contaminate them. The reliance on an
adhesive bond between the end fitting and the tube liner is also not that reliable due to the differential movement in the axial
direction that is inherent in traplock system operation. As the fitting moves outboard under load, the liner material andadhesive typically cannot accommodate this differential movement without the possibility of high streses and ultimately
cracking, tearing or disbonding. The solution proposed is to use an elastomeric seal at the interface between the liner andtraplock mandrel in a location where it cannot be easily damaged or contaminated during pipe construction. This is typically
located at the tip of the Traplock mandrel and needs to incorporate profiles that are pressure energised when the pipe is
internally pressurised.
A commercial consideration of this design approach is that it is a one off shot and in the event that a reliable seal is not
achieved it is not possible to replace it, the joint is effectively irreparable.
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Figure 1 - Traplock Design Example
Swaged End Fitting
An alternative approach shown in Figure 2 is to use a metallic inner and outer sleeve that effectively sandwiches the composite
material and its internal/external liners.
This is a relatively simple design approach that has the benefit of assembly onto plain ended composite pipe. The end fitting
can be assembled in a number of ways but typically the end fitting is fitted over the pipe end as an assembled unit. The inner
spigot has an interference fit with the bore whilst the outer sleeve is a close sliding fit. After assembly of the end fitting overthe plain pipe end the outer sleeve is swaged to reduce its internal diameter and generate a high interference with the
composite pipe. The high interference generates a combination of friction and mechanical interference that is capable of
transferring structural loads.
The inner mandrel is machined with a serrated or saw tooth outer profile that is designed to engage with the composite pipeliner. Additionally, it also incorporates a nose seal, typically an elastomeric O ring that seals against the bore of the liner pipe.
The issues associated with this arrangement are:
Unlike with the trap-lock design, where the axial stiffness of the mandrel can be readily optimised, load transmissionfrom the steel to the composite elements is focussed on the first thread and this can lead to higher local stresses
The inner mandrel design results in a bore restriction
Sealing mechanisms are typically O ring technology and it is not easy to incorporate preloaded seal solutions
There is a high stress concentration factor at the transition from the stiff steel outer sleeve to the nominal compositepipe through both the high difference in bending stiffness and fact that the composite pipe within the fitting ismaintained in hoop compression whilst at the exit from the end fitting the pipe transitions to hoop tension causinghigh local bending.
The steel outer sleeve and inner mandrel interface with the weaker liner and external coating materials and not withthe structural core composite material.
There is potential for damage of the inner liner surface resulting in uncertain mechanical interface as the innermandrel is inserted
Thermal and structural loads can result in differential movements between the steel and composite elements that needto be considered in the design process
High local stresses at the interface between the composite elements and steel elements result in high potential forcreep and relaxation affecting the long term response.
To overcome some of these issues it is typically necessary to thicken the wall of the composite pipe locally to achieve an
acceptable strength consistent with the pipe strength performance. The construction of this end thickening may need to bedifferent to that of the pipe construction to accommodate the nature of the loads at this location.
Such a design is well proven in smaller diameter hose products but is not considered as attractive for larger diameter, highpressure and dynamic applications. This is because the loads required for the swaging process get prohibitively large, the
swaging process is complex metalurgically, especially where there may be H2S and the end fitting must accommodate
dynamic loading
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Figure 2 - Swaged End Fitting Design Example
Metallic Liner End Fitting
Recent composite riser pipe development has considered a Hybrid composite tube [4,5], where a thin walled steel or titanium
liner is used along the entire pipe length and which is structurally continuous, via a weld, with the end fittings. The externallyapplied composite material is then used to enhance the steel pipe performance and is wound over the top of the steel liner. This
reinforcement is achieved by a combination of hoop and axially placed fibers.
This arrangement significantly reduces the complexity of the end fitting design and has three important benefits:
The metallic liner provides an impermeable membrane that precludes the problem of gas migration across the liner,which is a common problem for many polymer materials
It simplifies the end fitting design by avoiding the need for a seal assembly between the pipe and end fitting as thepipe and end fitting are welded.
It allows axial and hoop loads to be accommodated separately ie it allows axial loads to be primarily accommodatedby the steel pipe and the hoop (burst) loads accommodated by the composite material
In the extreme design case, the composite fibers can be used to provide only hoop or burst reinforcement and all axial loads are
taken by the steel liner. In this case there is no need for a trap-lock interface as all loads are hoop. Where the compositematerial is configured to take both hoop and axial loads the trap-lock style interface may still be required.
This configuration does not offer such a large weight saving benefit as an all composite pipe structure, due to the higherweight of the liner, its minimum bend radius and fatigue performance are dictated by the steel pipe and by selected weld
details rather than the composite material and there remains a design challenge regarding differential material strains when
subjected to temperature gradients and/or external loads.
In the hybrid tube approach corrosion is a remaining issue and often the steel liner needs consideration to both internal and
external corrosion allowances. An external corrosion allowance may be required if there is potential for water to ingress
between the steel pipe and composite material where such an interface can present a challenging corrosion environment. Theneed to provide such corrosion allowances can further degrade the weight benefits of such a structure and when this is coupled
with fatigue and corrosion issues of the steel pipe it is not evident that such an approach offers a significant advantage over an
all steel approach.
Where this approach does offer an important benefit is for ultra-high pressure applications where without such an approach
extreme metallic wall thicknesses may be required, potentially beyond the practical limits of manufacturability or weldability.In such an application, the hybrid tube approach may be considered an important enabling technology.
During manufacture the composite is cured at an elevated temperature but when the pipe cools down the steel pipe contracts
more than the composite material and a gap occurs between the two. To prevent the two materials adhering to each other and
causing unwanted stresses this process must be managed by using a release agent applied to the pipe prior to coating.Additionally, a rubber layer is applied as the first layer before the composite material is wound this is also applied to the
traplock area so that the composite materials can be completely encased with an outer rubber layer.
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Winding the composite material into the 3D traplock profile must be conducted accurately and care must be taken to manage
winding turnarounds to make sure that fiber angles are correct and material placement achieved to avoid potential for gaps and
defects during the turnaround process.
After the composite has been fully laid and the traplocks filled with 90degree fiber the whole joint undergoes a curing and then
autofrettage process that produces an interference fit between the steel pipe and the composite material. During the
autofrettage process it is difficult to control the interference between the traplock and the composite material due to its muchgreater hoop stiffness than the pipe body itself. Consequently, there is a possibility that at the critical end locations optimum
preloading of the traplock profile is difficult to achieve.
Figure 3 - Metallic Liner Design Example
Magma End Fitting Design
The end fitting design, developed by Magma and discussed in the following sections builds on the experience and knowledge
developed from previous end fitting arrangements. The end fitting is designed to allow termination of fully composite pipes
that have no metallic liner. Fundamentally, the design separates structural and sealing functions and thereby simplifies the
problem of having materials with dissimilar structural and thermal properties. The end fitting manufacture is also somewhat
simplified by virtue of the composite material selected, which is a thermoplastic rather than a thermoset as typically used in theabove designs. This is an important difference as it allows materials to be easily added at multiple manufacturing stages which
cannot be easily achieved with the thermoset process.
A schematic of the arrangement is shown in Figure 4. The composite material under consideration is a high performance
composite comprising high strength carbon fibre and Victrex PEEK.
PEEK is a thermoplastic semi-crystalline polymer material with a glass transition temperature of 143degC but with goodstructural properties well above this temperature up to 200deg C.
The carbon fibre used is supplied by Toray, a high performance fibre with a tensile strength of 4.9GPa.
Both materials offer good general performance characteristics and together produce a composite material with some of the
highest possible performance characteristics with respect to structural strength but also corrosion, fatigue, chemical resistance,
aging and permeation. The combination of PEEK and carbon fiber is well proven and regularly used in critical oil & gas welldrilling and completion and aerospace applications.
An important point, with respect to the proposed end fitting design, is the ability of the manufacturing process to allow thewall thickness to be built-up locally at the pipe end whilst maintaining the structural performance through the full section,
consistent with the main pipe body and with no strength knock-down.
This build-up can be conducted during the initial pipe manufacture or as a secondary thickening process onto the end of plainended pipe as discussed above. The properties of this build-up can also be optimised in terms of fiber orientations to best suitthe structural needs of the end fitting. Also, importantly from an operational point of view, an end fitting can be cut off a pipe
and a new build-up and end fitting applied. This could be conducted many years after the initial pipe manufacture.
Once the pipe end has been thickened in this manner the material can be machined to accurate dimensions to suit the end
fitting design.
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In the proposed design the pipe end is gradually thickened in two stages. The first thickening provides a transition from the
nominal pipe dimensions to the much stiffer end fitting and therefore allows the stress concentrations that occur at this point to
be appropriately managed.
The second thickening is more significant than the first and allows provision of tapered section to interface with a steel outer
collar. The benefit of increasing the wall thickness in this manner is that it significantly increases the local strength and
stiffness of the pipe at this section. This allows higher local loads to be applied with small strains.
The taper angle selected is a compromise between generating radial preload, minimising axial movement, allowableinterlaminar shear properties and spreading peak loads. The angle was selected only after completing a number of tests to
understand the effects of different angles and also surface finishes and friction factors.
The outer steel collar has a similar mating tapered profile to the composite pipe and is preloaded onto the taper using a
hydraulic tool. The preload is then locked into the end fitting using a locking nut that is wound in to take up the gap that occursduring preloading. This approach ensures that the preload is accurately applied and takes account of the deformations that
occur in both the steel and composite materials. The preload is selected to be greater than the highest service load, ensuring
positive load is maintained.
Sealing is provided by the use of a conventional pressure energised bore seal with an AX profile. This arrangement provides asmooth internal bore and as a percentage of the end fitting preload is taken by the seal it is considered to be a high integrity
sealing arrangement. The bore seal material can be either PEEK or stainless steel. On the composite pipe side the seal sits in aPEEK lined pocket and on the steel side a conventional inconel inlaid pocket is used.
Figure 4 - Magma End Fitting Design Example
Magma End Fitting Design Issues
The design features that were identified as being important at the start of the development process are summarised as follows:
Be structurally as strong as the rated pipe
Have a fatigue performance as good as the pipe
Have smooth unobstructed bore
All seals can be back pressure tested and are replaceable
The end fitting can be disassembled, inspected and reassembled with replacement parts as necessary
The design must use a preloaded seal to ensure high integrity
Metallic sealing surfaces are stainless steel inconel inlaid to prevent seal surface degradation
The design is preloaded to ensure fatigue performance of steel components and ensure seal integrity at high load
The overall dimensions are minimised to reduce weight and cost
Critical interface stresses are minimised to avoid the problem of creep which in all composite design problems is akey design issue encountered in the development of the end fitting are discussed below.
Creep is time, temperature and load dependent and generally it can degrade the long term strength of a composite structure. Inthe case of the end fitting, creep has the potential to reduce initially applied preloads such that over time the sealing integrity
or load capacity is compromised. The consequence of creep in the current end fitting design is to allow preload at the end face
of the build up section to be relieved such that under extreme loading a gap may appear and this can influence sealperformance and ultimately fatigue response of the laminate and particularly the steel components.
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In this respect the combination of PEEK and carbon fiber is a key advantage. PEEK has a high creep resistance, even at
elevated temperatures, in comparison with other polymers. Also carbon fiber, which has a negligible creep response, further
stabilises the polymer.
In the current design the interface between the steel and composite materials is formed by relatively large accurately machined
and well toleranced surfaces allowing high local stresses to be avoided. This combined with the high section stiffnesses helps
to reduce the potential for creep and end fitting relaxation.
However, the design is also attractive from the point of view of thermal and external loading as it does not rely on bondedsurfaces and can accommodate some relative movement between the steel collar and the composite pipe. The highly preloaded
connection ensures that even when the end fitting is heated the resulting differential movement can be accommodated.
FEA APPROACH and RESULTS
The end fitting design has been subjected to a full FEA evaluation using ANSYS. Rather than using a complex composite
modelling process a simplified approach is used that quickly and efficiently provides results for both the composite andmetallic components. The analysis process uses material data generated by testing, primarily on full scale pipes but also some
coupon materials. This test data is also supplemented where necessary by specialist material design software to produce
characteristic orthotropic material properties.
The starting point for the analysis work was a series of small scale pull tests to investigate the response of taperedcomposite/metallic interface and understand the issues related to slip, friction factors, surface finish, creep and necking. This
also allowed benchmarking of ANSYS results against test data.
This preliminary testing was conducted using 2in 10,000psi rated bore pipe specimens, shown in Figure 5 which were axially
loaded and accurately monitored to understand their response. This provided confidence in the proposed interface and
provided confidence in selected friction factors, which are surface finish dependent.
Figure 5 End Fitting Pull Test Configuration
Subsequently, a full FE model of the end fitting was prepared that correctly models the interfaces between all the steel andcomposite components including the gasket. The end fitting was modelled axisymetrically using PLANE182 elements and
using a range of orthotropic properties to model the different sections and layers of the pipe and build-up.
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Figure 6 End Fitting Assembly
The geometry used for the analysis is a 3.25in 5,000psi end fitting. A 2D axisymmetric model is used, created by importing a
3D Parasolid file, generated in Solidworks CAD software, and refined within the ANSYS Workbench software ensuring thatthe geometry accurately reflects the detailed design and enables better mesh and analysis. The model is quarter sectioned and
small chamfers and bolt holes are removed.
The entire model is meshed using 4-noded plane182 elements, with linear elastic behaviour. The pipe and collar sections of the
model are divided into smaller areas such that a regular pattern of mesh is achieved. Mapped mesh is applied on these sections.
Other steel parts such as the Locking Ring, Steel Hub, and threaded part of the collar are meshed using the free meshingoption. The collar is also divided into smaller areas and free meshed such that the contact surfaces on the inner layer and the
steel have the same mesh pattern.
In the contact regions, the mesh is refined to achieve virtually matching mesh. The contact regions are modelled, using
TARGET171 and CONTA169 elements, as asymmetric frictional contacts.
The model contains 5 contact regions to simulate the locking taper friction, the pipe end abutment with the hub, the bore seal
behaviour and influence of elastomer (rubber) which is used as a compliant filler between the internal diameter of the steel
collar end transition to the composite pipe. This manages local contact stresses. The five areas are as follows:
1. Outer surface of the locking taper of the pipe and the corresponding inner surface of the collar, the friction factor isassumed 0.3.
2. Forward surface of the pipe and the aft face of the hub, friction factor 0.2.3. Top left surface of the bore seal and forward and bottom surfaces of the inner pocket layer, friction factor 0.2.4. Top right surface of the bore seal and bottom taper face of the hub, friction factor 0.4.5. Elastomer, between the thin end of the collar and composite build up, friction factor 0.6.
The FEA model is loaded using a range of load conditions including internal pressure, bending and axial load. The internalpressures used in the current study are relatively low but other studies have considered operating pressures of 15,000psi and
test pressures in excess of 45,000psi.
Model pressures are as follows:
Operational Pressure 3,103 psi (214 bar)
Test Pressure 5,555 psi (383 bar)
A PEEK bore seal is used. This has a small interference fit with the pocket and results in an initial contact pressure when the
end fitting is assembled. The seal is then pressure actuated as the bore pressure increases. In accordance with the findings ofprevious analyses and tests, the end fitting taper needs to be pre-loaded to minimise movement at the taper face when high
COLLAR
LOCKING
TAPER
HUB
LOCKING
RING
m-pipe
BORE SEAL
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external loads are applied to the end fitting. An axial pre-load of at least 1.3 times the maximum axial load (pressure end cap
plus applied load) is applied. Practically, this is achieved on the end fitting by using a hydraulic puller on the collar and then
holding this with a threaded locking ring. Once the locking ring has taken up the slack and has been tightened up against the
hub, the hydraulic puller is released. This process, using a hydraulic tool rather than a applying torque, ensures that an accuratepreload is applied without the need to make assumptions for friction factors or component deformations. In the FEA the
preload is achieved by thermally shrinking a ring of elements until the appropriate preload is achieved.
The material properties used, the contact definitions and loading applied are explained in the following sections.
Structural Loading and Boundary Conditions
The structural loading applied in the analysis is summarized in Table 1. The analysis is carried out in 2 load steps as below:
Load steps Load type Application type Quantity
1 Pre loadShrink a row of collar elements to achieve stresses of ~66
MPa in the collar by applying -100C
(34 tonnes)
2
Axial load
Operational Pressure Load23.9 MPa
(10.87 tonnes)
Test Pressure Load42.75 MPa
(19. 5 tonnes)
Internal
pressurePressure Load 38.3 MPa
Table 1 Design Load Cases
Figure 7 - Loading and Constraint
The results from the analyses of the model under loading, as described in the previous section, are summarised and illustrated
in the following tables and images.
The positions of the nodes 2356 & 2394 are within the pipe laminate build-up and are at the thick and thin ends of the taper,
Figure 8 Nodes 148 & 1015 are within the steel (collar) and at the thin and thick ends of the taper, local to the peak hoop
stresses indicated.
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Figure 8 Key Node Positions
Contact pressures at the locking taper are shown in Figure 9 for preload and preload plus pressure load. It can be seen that thenominal contact pressure load and distribution is relatively low along the majority of the contact face and that it does not
increase significantly as the internal pressure is applied. This is a result of the the preloading process.
Stresses in the steel collar and locking rings are relatively low even at the critical threaded collar sections.
a) b)
c)
Figure 9 - Contact pressures (MPa) at the taper. a)Preload b) Preload+Operational Pressure c) Preload+Test
Pressure
Node 2394Node 2356
Node 1015
Node 148
Node 277
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a)
b)
Figure 10 - Equivalent (Von Mises) stresses (MPa) after a) Preload+Operational Pressure b) Preload+Test Pressure
NODE Location Sz Sy Sxy Von Mises
2356 Pipe Thick End of taper 24% 29% 5% N/A
2394 Pipe Thin End of taper 25% 17% 0.2% N/A
148 Collar Thin End of taper 22%
1015 Collar Thick End of taper 19%
277 Collar in region of the locking taper 22%
Collar to Lock Ring thread roots. Secondary stress. 50%
Collar adjacent to Lock Ring take-up. Primary stress. 27%
Bore Seal Nib Root 17% 27% N/A
Bore Seal ID centreline 45% 30% N/A
Table 2 - Stresses after Load Step 1 (Preload) % Utilisation of Ultimate StrengthSz=radial, Sy=hoop, and Sx=axial.
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NODE LocationSz Sy Sxy VM
2356 Pipe Thick End of taper 24% 30% 6% N/A
2394 Pipe Thin End of taper 27% 17% 1% N/A
Pipe Laminate ID at start of build-up 28% N/A
Pipe laminate ID 18% 21% N/A
Pipe laminate OD 10% 21% N/A
148 Collar Thin End of taper 39%
1015 Collar Thick End of taper 26%
277 Collar in region of the locking taper 33%
Collar to Lock Ring thread roots. Secondary stress. 17%
Collar adjacent to Lock Ring take-up. Primary stress. 38%
Bore Seal Taper Root 9% 26% N/A
Bore Seal ID centreline 56% 68% N/ATable 3- Stresses after Load step 2 (Preload + Internal Pressure) % Utilisation of Ultimate Strength
Sz=radial, Sy=hoop, and Sx=axial.
LocationMax Contact Stress
(MPa)
Pipe seal pocket 5
Hub seal pocket 6
Table 4 Bore Seal ContactStresses after Load Step 1 (Preload)
LocationMax Contact Stress
(MPa)
Pipe seal pocket 50
Hub seal pocket 60 (Mean) 80 (Peak)
Table 5 Bore Seal Contact Stresses after Load Step 2 (Preload + Pressure)
Location Contact Stress (MPa)
Average over the contact face 30
Peak value at contact face edges 142
Table 6 Locking Taper ContactStresses after Load Step 1 (Preload)
Location Contact Stress (MPa)
Average over the contact face 40
Peak value at contact face edges 186
Table 7 Locking Taper Contact Stresses after Load Step 2 (Preload + Pressure)
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Testing Approach
A series of mechanical tests have been performed on the proposed end fitting. The objective of each test is to prove the
structural performance and seal integrity for a range of applied loads.
Pressure
The internal pressure test comprises a closed end pipe filled with water. The end of the pipe is terminated with a Magma endfitting and assembled with a PEEK bore seal. The pump allows accurate control of the pressure so that any form of test profilecan be conducted.
There are four typical test profiles:
1. Step test: pressure increased in steps to a target maximum and then released in steps2. Dwell test: valve locked at a target pressure and left for a period of days3. Cyclic test: pressure repeatedly increased and decreased in quick succession4. Failure: pressure increased in steps until failure
After each test the fittings are inspected for any evidence of leaks. The pressure data is analysed to check for changes in the
pressure profile. A concrete bunker ensures the safety of personel during all tests that operate near predicted failure pressures.
Individual loadsIn addition to the tensile tests performed to evaluate the metal-m-pipe interface (see FEA Approach and Results), two
individual load cases have been assessed: axial tensile load and cantilever bending load. The test approach includes multiplefibre optic strain gauges on the surface of the pipe in the axial and hoop direction. In addition, a video extensometer is used for
the tensile test to monitor the strains at any location along the pipe or end fitting. Calibrated load cells are linked to the
hydraulic cylinders that apply the loads.
Combined loads
The combined load test has been assembled to enable any combination of internal pressure, axial tension, and cantilever
bending loads to be applied. The following combinations were chosen:
1. Internal pressure and tensile load2. Internal pressure and bending load
3. Internal pressure, tensile and bending loads4. Failure: pressure and bending to failure
Figure 11 - Combined Load Rig GA
Fatigue
Fatigue has been evaluated using a resonant cantilever beam arrangement. A vibration motor with offset weights was secured
to one end of an m-pipe test sample with the other end terminated in a Magma end fitting bolted to a vertical bracket. Thepipe was instrumented with fibre optic strain gauges so that the oscillation frequency and strain response could be monitored at
all times. The pipe was filled with water and pressurised to a nominal level throughout the resonant fatigue tests.
m-pipe
End Fitting
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Figure 12 - Resonant Fatigue Rig GA
Test Results
Pressure
A number of pipes were prepared over a range of two sizes: 2in and 3.25in. A summary of the key results is presented in Table
8.
m-pipe Test Target Pressure Test Description Result
2in 5000psi Step 11,000psi 2900psi steps, hold for 2 minutes
at each step
No sign of leak
Step 14,500psi 2900psi steps, hold for 2 minutes
at each step
No sign of leak
Dwell 14,500psi 2900psi steps, hold for 2 minutes
at each step. Dwell for 150 hours(~6 days)
No sign of leak
2in 10,000psi Failure - 2900psi steps, hold for 2 minutes
at each step continue untilleak/failue
Data from step dwells
indicated no sign of leaks.Pipe burst at 26,527psi
3.25in 3100psi Step 5555psi 1450psi steps, hold for 5 minutes
at each step
No sign of leak
Step 6600psi 1450psi steps, hold for 5 minutes
at each step
No sign of leak
Dwell 6600psi 1450psi steps, hold for 5 minutes
at each step. Dwell for 65 hours(~2.5 days)
No sign of leak
Cycle 5555psi Increase to 5555psi in 1 minute,
dwell for 1 minute, releasepressure in 1 minute. Repeat 10
times
No sign of leak
Table 8 - Pressure Test Results Summary
Cantilever End Fitting
Vibratory Motor
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Each test sample was assembled and dismantled several times during the sequence of pressure tests showing that the seal and
the seal pocket remained undamaged.
Figure 13 - Photo of a 2in 5000psi m-pipe in the pressure chamber
Individual loads
The test results for the individual loads are shown in Table 9.
m-pipe Test Load Test Description Result
2in 5000psi Tensile 120kN 20kN steps, dwell for 1 minute at eachstep
No sign of axialslippage or damage
Cantilever Bend 0.9kNm 0.15kNm steps, dwell for 1 minute at
each step
No sign of damage
Table 9 - Individual Load Test Results
The tensile only test could have presented difficulties due to the hoop shrinkage as a result of the Poissons effect. The benefits
of the selected manufacturing method include the ability to vary the fibre angles in the design of the additional thickeningaround the fitting and so the Poissons effect did not cause any issues during this test.
Combined Loads
The combined load tests followed on from the individual load tests using the same equipment. The test plan consisted of a
matrix of target loads and pressures and in all cases the pipe axial and hoop strains were monitored. The key results are
summarised in Table 10.
TestTarget
Load/PressureTest Description Result
Tensile &Pressure
112kN,10,000psi
20kN steps at 4 different levels of internalpressure, dwell for 1 minute at each step
No sign of axial slippageor leaks
Cantilever Bend& Pressure
0.9kNm,10,000psi
0.2kNm steps at 4 different levels of internalpressure, dwell for 1 minute at each step
No sign of damage orleaks
Tensile,Cantilever Bend
& Pressure
40kN,0.6kNm,
10,000psi
2 tensile load steps, 2 levels of internalpressure, 4 steps of increasing bending
moment
No sign of damage orleaks
Cantilever Bend
& Pressure to
failure
10,000psi Increase bending moment until failure 150mm stroke limit
reached on the hydraulic
cylinder no failure
Table 10 - Combined Load Test Results (2in 5000psi Rated)
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These tests present themselves as a fairly rigorous series of events with axial pipe strains in excess of 1.8%. The combined
bending and pressure test highlighted a beneficial aspect of m-pipe in that the axial strain due to pressure had to be over-
come by the compressive strain due to bending before failure in the pipe could occur. In this case the limit of the vertical
hydraulic cylinder was reached and so the test was repeated without internal pressure so that a failure could be achieved. Thefailure occurred on the top surface of the pipe (compressive load) near the exit of the end fitting.
Figure 14 - Combined Load Test Rig
Figure 15 - Bend to Failure
Fatigue
A series of ring-down tests were performed to establish the resonant frequency of the system. This was quite simply achieved
by knocking the end of the pipe with a rubber mallet and recording the axial strains in the fibre optic sensors. The time-strain
plot gives the damping characteristics and an indication of the natural frequency.
The motor weights were initially balanced to give 75% of their maximum vibrational force. An internal pressure of 1000psi
was applied using a water filled pump. The motor control was increased until a frequency close to the natural frequency wasachieved (~10Hz). The motor speed was adjusted to give a target peak-peak strain value in the fibre optic strain sensors that
were positioned in the middle of the pipe. Assuming the bending moment increases linearly towards the fixed end of the pipe
the strain was calculated at the point the pipe exits the end fitting and this extrapolated value is shown in Table 11.
For this particular test the end fitting setup represented an early design of the bore seal and as such does not contain some of
the design improvements from subsequent pressure testing.
The results are summarised in Table 11.
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Strain level at End Fitting/Pipe Interface Number of cycles Result
~0.8% strain peak-peak 2.4millionNo leak,
No structural damage
1.1% strain peak-peak 10.0million
Small leak observed
from seal at 7.8million
cycles,
No structural damage
~1.45% strain peak-peak 0.5million
No structural damage
unable to reach higher
strains due to motorlimit.
Table 11 - Resonant Fatigue Test Results
Further resonant fatigue testing is planned with 2in 15,000psi m-pipe samples and 3in 10,000psi samples to increase the
sample data.
Conclusion
It is believed that carbon fibre/PEEK pipe has the potential to be an enabler for future oil and gas developments in applications
where metallic based pipe technologies struggle to meet project requirements. However, composite pipe end fitting design
issues have often been highlighted as a concern, of sufficient magnitude, to preclude serious consideration of the technology.These concerns have related to structural performance, sealability and general long term reliability of composite materials and
specifically composite/metallic interfaces.
The work reported evidences carbon fibre/PEEK materials and manufacturing methods that together allow a different endfitting approach to be adopted. This approach resolves previous issues facilitating a solution with increased performance and
reliability. Importantly, this removes a serious obstacle to the future broarder application of composite pipe technology.
The key design features of the end fitting design are enabled by the manufacturing method and the selection of carbon fibre
and PEEK as the composite materials of choice. This process allows the end of the pipe to be readily thickened as a secondary
manufacturing operation to provide a reliable structural interface with the steel end fitting. Additionally, the end fitting design
specifically manages the issues related to differential thermal expansion and Poisson ratio effects between metallic and
composite elements. The combination of these features is that the end fitting can be designed to stronger than the base pipe inall respects.
References
[1] Melve, B. First Offshore Composite Riser Joint Proven on Heidrun Offshore Magazine
[2] Blanc, L. Composites Cut the Riser Weight by 30-40%, Mass 20-30% Offshore Magazine
[3] Baldwin, D.Interface System Bewteen Composite Tubing and End Fittings United States Patent 6,042,152
[4] Guesnon, J. Caillard C. Hybrid Tubes for Choke and Kill Lines, Offshore Technology Conference OTC 14021
[5] Cederberg, C.Design and Verification Testing Compsoite-Reinforced Steel Drilling Riser Final Report RPSEA
07121-1401