otc-24229-ms.pdf

OTC 24229

Riser Concepts for High Motion Vessels in Ultradeep Water Brian S. Royer, Thomas L. Power, Stress Engineering Services, Inc., William Head, RPSEA

Copyright 2013, Offshore Technology Conference This paper was prepared for presentation at the 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 been reviewed by the Offshore Technology Conference and are subject to correction by the author(s). The material does not necessarily reflect any position of the Offshore Technology Conference, its officers, or members. Electronic reproduction, distribution, or storage of any part of this paper without the written consent of the Offshore Technology Conference is prohibited. Permission to reproduce in print is restricted to an abstract of not more than 300 words; illustrations may not be copied. The abstract must contain conspicuous acknowledgment of OTC copyright.

Abstract The paper addresses early work executed during the research study for RPSEA project 10121-4401-01 Ultra-Deepwater Riser Concepts for High Motion Vessels. The steel catenary riser is the simplest riser configuration but has limitations when attached to a high-motion vessel, such as a conventional semi-submersible or a ship-shape floating, production, storage, and offloading vessel (FPSO), in water depths approaching 10,000 feet. Further challenges are posed by the high-pressure, high-temperature reservoirs that can be encountered at these water depths. Alternate riser arrangements to meet the demands of both ultra-deepwater and high motion vessels are investigated. Production is assumed to be high pressure (20,000 psi SITP), high temperature (350F), and sour. Configurations include wave shapes and other methods to reduce the coupling between vessel motion and the fatigue-critical riser touchdown zone. The work described in this paper is primarily concerned with static configuration and vessel payload considerations of the various riser material/configuration options. Introduction This paper and an anticipated follow-on paper will present the results of a currently active RPSEA study whose objective is to evaluate potential riser concepts for high-motion vessels operating in waters up to 10,000 feet deep in the Gulf of Mexico (GoM). Among the goals of this study are:

Identify technology gapsin particular, those that represent concepts/aspects that appear analytically promising

but have yet to be field-proven. Determine feasibility for particular concepts. Identify preferred riser options.

The study scope to be discussed in this first paper includes:

Problem framing

Literature search Identification of candidate systems Compilation of design basis

Preliminary sizing Static analysis/design Intermediate conclusions Refinement of options/goals for remaining work

The follow-on paper is anticipated to cover dynamic system analysis (coupled analysis of floater, risers, and buoys), fatigue considerations, final comparison of options, and recommendations for further study. As is typical for RPSEA studies, a project working group was formed, comprising riser specialists from five oil companies with ultra-deepwater interests. The project working group is responsible for providing guidance to the study in order to insure that its results are valid and useful to the industry at large.

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The approach taken for this study is to use information gathered from published work and to extend the understanding with particular focus on high-motion vessels. Six riser types are considered (see Figs. 1-6):

Steel (Rigid) Catenary Riser (SCR) Steel (Rigid) Lazy Wave Riser (SLWR) Steep Wave Riser (SWR) Compliant Vertical Access Riser (CVAR) Tension-leg Riser (TLR) Hybrid Riser Tower (HRT)

The design basis for this study includes a shut-in tubing pressure of 20,000 psi at the subsea wellhead, determined with the guidance of the project working group. This effectively restricts the study to rigid pipe construction (steel, titanium, or fiber-reinforced composite) and eliminates currently available flexible pipe products. The potential benefit of using a high-integrity pressure protection system (HIPPS) to reduce the ultimate pressure demand on the system is evaluated.

Extensions to Current State of the Art.

Over the past 10 years, projects have demonstrated that conventional SCRs are feasible in water depths greater than 5,000 feet from a variety of host platforms that exhibit limited motions of the riser attachment point (e.g., Na Kika, Independence Hub, Thunder Horse, Perdido, and others). Banon et al. (2007) demonstrated analytically that even in 10,000-ft water depths in the GoM, SCRs can be made feasible with careful design of the host platform. When attached to hosts with significant heave dynamics like the ones considered in this study, however, conventional SCRs are perceived to have shortcomings with respect to fatigue life and compression in the lower portion of the riser.

Regarding the other riser types, important work for 10,000-ft GoM waters was performed approximately ten years ago

for the SLWR, TLR, and HRT concepts serving an FPSO host (see DeepStar, 2005). Since that time, ultra-deepwater projects for both the HRT (Song and Streit, 2011) and the TLR (INTECSEA, 2012) have been executed or are now under construction. The industrys first steel lazy wave risers were installed in 2009 in approximately 5900-ft waters offshore Brazil, servicing a turret-moored FPSO (Hoffman, 2010).

The current work represents an extension to what has been considered or built previously in ultra-deep water by studying

risers operating in scenarios characterized by the combination of:

High-motion host vessels (conventional Semisubmersible or ship-shaped FPSO) Shut-in pressure of 20,000 psi Water depth of 10,000 ft.

Ultra-deepwater Challenges.

There is currently very little pipeline infrastructure in the GoM that reaches fields with water depths approaching 10,000 ft. Consequently, new projects at these depths require either the construction of costly export pipelines or the utilization of FPSO/FSO facilities to host the production and the use of tankers to transport the production to market.

Ultra-deepwater prospects often have very high reservoir pressures and temperatures. The mechanical requirement to

contain the high pressure at elevated temperature drives up riser wall thickness, which, when combined with unavoidably long riser length, can result in very large riser payloads and resultant increase in host vessel size and cost. Furthermore, a very-high-pressure environment results in a very low sour threshold concentration of H2Sperhaps imperceptibly small, in which case, the production must be considered sour. For example, for the present 20,000-psi case, the NACE MR0175 threshold partial pressure of H2S of 0.05 psia is reached at a concentration of only 2.5 ppm, which will be very difficult to detect. The selection of appropriate materials for sour service can have a significant cost impact.

This study will address each of these challenges.

Design Basis and Environments Table 1 summarizes the design basis for this study, compiled in consultation with the project working group. This combination of functional requirements represents a significant step-out compared to currently deployed systems.

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Table 1: Design Basis

GENERALDesignLife 25 yearsWaterDepth 10,000 feetShutInTubingPressureatSubseaWellhead 20,000 psiMax.OperatingPressureatSubseaWellhead 12,000 psiMax.OperatingTemperatureatSubseaWellhead 350 degFMax.TemperatureforShutin

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Host Vessel Options and Number of Risers. Options for high-motion host platforms that are also suitable for the GoM are fairly limited. Two basic classes of

platform were determined appropriate for this application: weathervaning (i.e., turret-moored) FPSOs and conventional semis.

Turret-moored FPSO. Turret-moored FPSOs come in two basic configurations: disconnectable-turret and captive-turret. The disconnectable

turret option allows the FPSO to depart the site when extreme environmental conditions are expected, while captive-turret (permanently moored) FPSOs are designed to remain on-station in all environments. It is not yet clear whether or not FPSOs operating in the GoM have to be able to flee hurricanes. The one FPSO that has been installed in the GoM (Song and Streit, 2011) and known FPSOs being planned for GoM deployment are all of the disconnectable-turret variety. Internationally, however, permanently moored FPSOs have been installed in typhoon-prone regions. (See, for instance Franz et al. (1996), and CNOOK, Ltd. (2001).)

The use of a disconnectable turret limits the riser payload capacity because the disconnected turret must be able to buoy

the full riser/umbilical payload as well as the weight of all the mooring lines in absence of the host FPSO. Turret buoyancy (and therefore riser payload capacity) is size-limited to dimensions that can be accommodated by the host vessel. For this study, the riser payload capacity for a disconnected turret is assumed to be 3000 kips, which should leave some capacity for dynamic control umbilicals (not included here). Aspects of disconnecting the turret from the FPSO are not trivial but are beyond the scope of the current study.

Riser payload for the captive turret option is typically limited by the capacity of the bearings between the turret and the

ship-shaped FPSO. This limit is significantly higher than the limit imposed by disconnected turret size/buoyancy. For this study, the payload capacity for a captive-turret FPSO is assumed to be three times that of the disconnectable turret FPSO, or 9000 kips.

The choice made between the disconnectable- and captive-turret options will determine the environments to be considered

with the FPSO on-station. Maximum on-station offset for the captive-turret FPSO will be assumed to be 7.5 percent of water depth, or 750 feet.

Conventional Semi. The riser payload for a conventional semi is assumed to be limited only by the size of the semi and the other payload

items that need to be accommodated. No particular payload limit is being assigned to the conventional semi for this study. Number of Risers. The prescribed number of risers (6 initial, 2 future) is driven by the assumed production capacity. Feasibility of a given

riser concept will be determined in part by its ability to be supported in the necessary numbers without exceeding the payload capacity of the potential host vessels being considered.

Wet Insulation.

All concepts considered in this study utilize wet (subsea) trees, so unless an insulating annulus is explicitly provided by the use of pipe-in-pipe riser construction, no annulus will be present (as would automatically be present with a dry-tree completion). Pipe-in-pipe construction at this water depth is considered infeasible with currently available installation vessels and is therefore not being considered. Consequently, the risers in this study carry a 3.5-inch thick blanket of wet insulation with a density of 50 lb/ft3, except as noted for the composite-reinforced pipe option. These values correspond to typical values of density and U value for appropriate insulation materials. Note that a technology gap exists here for temperatures at or approaching 350 deg F; the highest temperature to date known to the authors for which wet insulation has been used is 298 deg F. All riser pipe, including jumpers, for each of the concepts being considered is assumed to be fully covered with wet insulation.

VIV Suppression.

Strakes will be assumed to cover the risers and jumpers in the top 6,000 feet of the water column. The strakes will be considered neutrally buoyant for the purposes of this study and will be assumed to effectively eliminate fatigue damage due to VIV.

Corrosion Allowance, Cladding Thickness, and Sour Service.

For this study, it is assumed that carbon steel pipe wall-loss corrosion is addressed either by adding a corrosion allowance to the ID of the pipe wall or by cladding the pipe ID with a corrosion-resistant alloy (CRA). The corrosion allowance of 3 mm and cladding thickness of 3 mm shown in Table 1 were chosen as both expedient from a modeling perspective and

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representative of typical practice. The expediency comes from using the same values for corrosion allowance and cladding thickness, resulting in a non-corroded pipe that has the same dimensions whether clad or not.

The requirement for sour service means that the materials coming into contact with the well fluids must be NACE-

MR0175-compliant. It also means that carbon steel in contact with well fluids will have de-rated fatigue S-N curves.

Design Environments. Tables 2 and 3 contain descriptions of the design environments considered, which are appropriate for the central GoM. A

condensed 32-bin, 8-heading wave scatter diagram (not shown), also appropriate for the central GoM, was supplied by the project working group and will be used for fatigue life assessment.

Table 2: Design Wave Environments

Table 3: Design Current Environments

Riser Options Considered Configuration Geometries.

Figures 1 - 6 illustrate the various configuration geometries being considered in this study. The shut-in tubing pressure of 20,000 psi rules out currently available flexible pipe, so these concepts as considered in this study (including jumpers) are constructed entirely out of rigid pipe and components. The concepts are presented in order of increasing complexity/cost. Table 4 presents the concepts and outlines the complexities of each. It may be noted that, while the TLR and HRT concepts are somewhat mature according to the definition in the table, neither has yet been built (or contemplated) with rigid pipe jumpers.

100yearWinterStorm

100yearLoopCurrentAssociatedWave

100yearHurricane

1000yearHurricane

Hs(ft) 29.0 7.5 52.0 66.0Tp(s) 13.0 7.5 15.4 17.2

JONSWAP 2.0 1.0 2.4 2.41hrVwind(ft/sec) 82 35 150 187

SurfaceCurrentSpeedVc(ft/sec) 2.0 7.5 7.0 8.5

100yearHurricaneAssociatedCurrent

1000yearHurricaneAssociatedCurrent

100yearWinterStormAssociatedCurrent 100yearLoopCurrent

Depth Speed Depth Speed Depth Speed Depth Speed(feet) (ft/sec) (feet) (ft/sec) (feet) (ft/sec) (feet) (ft/sec)0 7.0 0 8.5 0 2.0 0 7.5125 5.2 150 6.3 150 2.0 164 7.5250 0.0 300 0.0 150 0.0 984 2.1500 0.0 500 0.0 500 0.0 2,953 1.0

6,561 0.310,000 0.3

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Figure 1: Simple Catenary Riser Configuration (SCR)

Figure 2: Lazy Wave Riser Configuration (SLWR)

Figure 3: Steep Wave Riser Configuration (SWR)

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Figure 4: Compliant Vertical Access Riser Configuration (CVAR) (Note: No vertical access is actually contemplated for this application)

Figure 5: Tension Leg Riser Configuration (TLR)

Figure 6: Hybrid Riser Tower Configuration (HRT)

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Table 4: Riser Concepts

Materials of Construction. The following materials are considered for construction of the various riser concepts:

Carbon Steel Line Pipe (API 5L X70), welded Carbon Steel Line Pipe (API 5L X70) with CRA-clad ID, welded High-strength Steel OCTG (5CT Q125) with CRA-clad ID, threaded and coupled Super Duplex Stainless Steel Pipe, welded Titanium Pipe, welded Carbon Fiber Composite-Reinforced Carbon Steel Line Pipe, welded or mechanically connected

This list is thought to be in order of increasing system cost, except that the cost of the composite option is unknown at this

time. Two notable omissions deserve explanation. Aluminum construction was ruled out because of unresolved issues with aluminum welding and a fracture toughness too low for a non-inspectable application. Nickel-based alloy pipe (solid CRA pipe wall) was ruled out due to cost thought to be approaching that of titanium, but without titaniums low weight and low stiffness benefits. Wall Thickness Sizing To size the riser pipe wall thickness, two cases were checked: (1) Gas-to-surface Shut-in, with SITP at the subsea wellhead and temperature at or below 250 deg F, and (2) Operating, with maximum flowing pressure at the subsea wellhead and temperature at 350 deg F, which requires the use of a temperature knock-down factor as shown in the following tables. The shut-in case was found to control wall thickness in all cases. The maximum differential pressure occurs at the top of the riser where the internal pressure is reduced from the subsea wellhead pressure by the hydrostatic head of the pressurized gas column.

For consistency, except as noted, wall thickness sizing was performed using Annex B of API RP 1111. In practice, application of RP 1111 to anything other than API 5L line pipe requires particular qualification of the pipe according to Annex A. It is also noted that use of the performance extensions represented by Annex B requires additional specification and verification for any pipe material, including API 5L line pipe. There is a significant perceived economic benefit to making a carbon steel option viable for this application, and if this is within reach, the cost of the necessary specification and verification to allow use of Annex B criteria would likely be negligible in comparison to the benefit.

Carbon Steel Line Pipe.

Wall thickness required for pressure containment for the API 5L X70 carbon steel line pipe option appears in Table 5. For simplicity, the tabulated value of 1.468 in. is rounded up to 1.500 in. for use in modeling. For the option without cladding, an additional 3 mm (0.118 in.) corrosion allowance is added to the pipe wall to arrive at a final required wall

Concept AbbreviationMaturityinUltradeepWater*

Complexity/CostAdders

SimpleCatenary SCR 3 Basecase;simplest

Lazywave SLWR 3 BuoyancyModules

Steepwave SWR 1 BuoyancyModules+RiserBase

CompliantVerticalAccess CVAR 1 BuoyancyModules+RiserBase

TensionLeg TLR 2 SubmergedBuoyancyStructure+Jumper+BuoyancyTethers+TetherBases

HybridRiserTower HRT 3 SubmergedBuoyancyCan+Jumper+Riser/CanTether+Gooseneck+RiserBase

*1=studied,notbuilt;2=underconstruction;3=currentlyinservice

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thickness for manufacture of 1.618 in. For the clad option, 3 mm of CRA cladding is added to the ID of the pipe to protect it from corrosion. In both cases, the carbon steel wall thickness available to contain pressure (1.500 in.) is the same. Neither the corrosion allowance steel nor the cladding is allowed to participate in resisting loads, though both the corrosion allowance steel and the cladding are accounted for in payload and riser weight calculations.

Table 5: 8.625-inch OD X70 Wall Thickness Calculations

High-strength Steel with Cladding. Table 6 contains the results of the wall thickness sizing exercise for the high-strength steel (Q125) option. CRA cladding

for this option is required in this sour service application because the material is not NACE MR0175-compliant. The benefit of using high-strength steel is reduction of the required wall thickness of the riser pipe and the resulting reduction in riser payload.

For simplicity, the tabulated value of 0.931 in. is rounded up to 0.95 in. for use in modeling. The 3 mm of CRA cladding

added to the ID of the pipe to protect it from corrosion is not allowed to participate in resisting loads, but it is accounted for in payload and riser weight calculations.

Table 6: 8.625-inch OD Q125 Wall Thickness Calculations

Super Duplex Stainless Steel. The super duplex (SD) stainless steel pipe is assumed for the purposes of this study to have the same yield and ultimate

strengths as X70 line pipe. This results in the same required wall thickness for pressure containment as for the X70 option. The payload benefit of using SD is that no corrosion or H2S mitigation (either corrosion allowance or cladding) is required, resulting in a somewhat reduced riser payload compared to carbon steel options. A further benefit will be realized in the dynamic portion of this study, and that is that the SD will not be subject to the same S-N curve penalties for sour service as the un-clad carbon steel options.

Welded Titanium.

Titanium pipe delivers several benefits relative to steel pipe as well as a significantly higher cost per pound. It is not subject to corrosion or embrittlement, its yield strength (110 ksi) is over 50 percent greater than that for the X70 pipe, its elastic modulus (16.5 106 psi) is about half that of steel, and its unit weight (276 lb/ft3) is about half that of steel. Use of titanium pipe therefore results in significantly reduced riser payload. Depending on market demand, it may also be more quickly obtainable than clad pipe, which could have a favorable impact on project schedule. It should be noted that titanium

fd fe ft

20,000psi 18,958 psi 18,960 psi 250Fmax. 0.75 1.0 1.0 70ksi 82ksi 1.468 in.12,000psi 10,958 psi 10,960 psi 350F 0.75 1.0 0.933 70ksi 82ksi 0.980 in.Notes:

(a)AnnexBimposesadditionalrequirementstodemonstratemanufacturercanachieverequiredburstpressure.(b)Assumes15pcf(gas)contentsand10,000ftverticalrise.(c)Doesnotincludeanycorrosionallowance

TensileStrength

CalculatedWallThicknessperAnnexB(a)(c)

Pressureat

Wellhead

PressureatTopofRiser(b)

WellFluidTemp.

APIRP1111FactorsYield

StrengthDesignPressure

fd fe ft

20,000psi 18,958 psi 18,960 psi 250Fmax. 0.75 1.0 1.0 125ksi 135ksi 0.931 in.12,000psi 10,958 psi 10,960 psi 350F 0.75 1.0 0.933 125ksi 135ksi 0.603 in.Notes:


TensileStrength


PressureatWellhead


WellFluidTemp.

APIRP1111FactorsYield

StrengthDesignPressure

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cannot be used in applications coming into contact with hydrofluoric acid (HF) or pure (dry, neat) methanol, though dilution of methanol with 10 wt. percent water removes the methanol concern (Baxter, Schutz, and Caldwell, 2007).

Table 7 contains the calculated required wall thickness for titanium pipe; for simplicity, the tabulated value of 1.188 in.

was rounded up to 1.200 in. for use in modeling.

Table 7: 8.625-inch OD Titanium Pipe Wall Thickness Calculations

Composite-Reinforced Steel Pipe (CRP). The carbon fiber reinforced pipe considered in this study consists of a carbon steel liner wrapped on its outer surface with

carbon fiber/epoxy composite. The steel liner was sized to withstand collapse (pressure only, no bending) at 10,000 ft water depth, and the carbon fiber composite was sized to carry all tensile hoop stress and 60 percent of the axial load, including the substantial pressure end load created by the large internal pressure. Preliminary sizing indicated the need for the composite to carry a significant amount of axial load.

This concept follows in the footsteps of a recent study performed for RPSEA that investigated a composite-reinforced

steel drilling riser. Riser joints have all-steel ends and composite-wrapped transition and uniform pipe regions. The uniform cross section described in the previous paragraph describes about 90 percent of the riser joint length.

No detailed study or design methodology past what is described above was performed. The goal was to develop with as

little effort as possible a reasonable cross section for the purpose of this comparative study. The resulting cross section consists of an 8.625 in. OD 0.500 in. wall thickness steel liner surrounded by a 1-inch-thick layer of carbon fiber epoxy composite, surrounded by thermal insulation out to an overall OD of 15.625 in., similar to the overall OD of the other options in the study. This results in one inch less insulation thickness than the other cases, but it is assumed that this is partially compensated for by the presence of the carbon fiber composite layer.

Since the liner is carbon steel, it is subject to the same need for either corrosion allowance or cladding as the other carbon

steel options considered. Payload and riser weights for the composite option therefore include the weight of 3 mm of corrosion allowance or cladding.

Carbon Steel Line Pipe with HIPPS.

Another option for reducing riser payload is to reduce the pressure demand on the riser system (and thus wall thickness) by use of a high-integrity pressure protection system (HIPPS). In the event of a HIPPS failure, the system must still contain the system shut-in pressure, but this pressure need not be treated as the design pressure, since HIPPS failure would represent an accidental occurrence.

Two approaches were taken to sizing an X70 riser wall for the HIPPS option:

1. Treat the top-of-riser pressure (18,960 psi) corresponding to SITP as an incidental overpressure in the sense of RP 1111 Figure 2 rather than a design pressure and size wall accordingly using RP 1111 Annex B.

2. Using the top-of-riser pressure (18,960 psi) corresponding to SITP, size the wall to keep the API RP 2RD membrane von Mises value e (see Eq. (9) of APR RP 2RD) below 90 percent of yield, and check the RP 1111 combined tension and pressure equation (Eq. (8) of RP 1111).

fd fe ft(d)

20,000psi 18,958 psi 18,960 psi 250F 0.75 1.0 1.0 89.9ksi 106.3ksi 1.188 in.12,000psi 10,958 psi 10,960 psi 350F 0.75 1.0 1.0 82.1ksi 99.6ksi 0.786 in.Notes:

(a)AnnexBimposesadditionalrequirementstodemonstratemanufacturercanachieverequiredburstpressure.(b)Assumes15pcf(gas)contentsand10,000ftverticalrise.(c)PerBaxter,Schutz,andCaldwell(2007OTC18624)(d)Doesnotincludeanycorrosionallowance

TensileStrength(c)

CalculatedWallThicknessperAnnexB(a)(d)

PressureatWellhead


WellFluidTemp.

APIRP1111Factors YieldStrength(c)

DesignPressure

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The first approach is more conservative than the second approach, but does not provide much advantage (see Table 8). For the purpose of this study, the thinner wall based on the second approach (see Table 9) was used going forward, without any rounding. For this particular case, sizing the wall to achieve e = 90 percent of yield strength put the RP 1111 combined tension and pressure load right at 0.96, which happens to be the limit appropriate for both hydrotest and extreme loads.

A third approach in which a reduced (say, 5,000-psi) design pressure corresponding to HIPPS operation was dicsussed,

but this methodology does not result in sufficient wall thickness to insure survival of exposure to full SITP in the event of a HIPPS failure. The project working group was reluctant to embrace such an approach.

In an actual design situation, the subject of design criteria for a scenario involving HIPPS should receive particular

attention. This attention is outside the scope of the present study. The sizing exercise represented here is not proposed as a valid design methodology, but rather is being used as a means to bound the problem while maintaining a defensible safety rationale.

The carbon steel pipe with HIPPS will still require the same corrosion allowance or cladding as the carbon steel pipe

(without HIPPS) case. The benefit of the HIPPS is the reduced wall thickness required to contain pressure.

Table 8: 8.625-inch OD X70 Pipe Wall Thickness Calculations with HIPPS Approach 1

Table 9: 8.625-inch OD X70 Pipe Wall Thickness Calculations with HIPPS Approach 2

Design Criteria The primary design criterion for this study will be the combined load Method 1 approach being embodied in the new API standard being developed to replace RP 2RD (Stanton, et al. 2010). The fundamental equation for this criterion is:

2

2

b

eiD

yy pppF

MM

TT

(1)

where

T is the effective tension in the pipe

Ty is the yield tension in the pipe = SA

fd fe ft

20,000psi 18,958 psi 18,960 psi 250Fmax. 0.75 1.0 1.0 70ksi 82 ksi 1.333 in.Notes:


YieldStrength

TensileStrength


PressureatWellhead


WellFluidTemp.

APIRP1111FactorsIncidental

Overpressure

APIRP2RDBurst

Pressure(d)CombinedLoad(e)

MembraneStress(c)

CalculatedWall

Thickness(f)

20,000psi 18,958 psi 18,960 psi 250Fmax. 70ksi 82 ksi 531 kips 20,834 psi 0.96 63.0 ksi 1.034 in.Notes:

(a)Assumes15pcf(gas)contentsand10,000ftverticalrise.Thisvaluenotusedforpayloadcalculation.(b)APIRP1111Figure2PressureLevels(c)Prescribedvalue;90%ofyieldstrength.(d)APIRP11114.3.1andAnnexB,whichrequirestestingtoverifyburstpressure(e)APIRP11114.3.1.2,Equation(8)(f)PerAPIRP2RD1998Edition,Equation9.Onlytensionandpressureloadsareconsidered.Doesnotincludeanycorrosionallowance

TensileStrength

CatenaryRiserTopTension(a)

APIRP1111CheckPressure

atWellhead

PressureatTopofRiser(a)

IncidentalOverpressure(b)

WellFluidTemp.

YieldStrength

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S is the specified minimum yield strength of the pipe

A is the cross-sectional area of the pipe wall

M is the bending moment in the pipe

My is the membrane yield moment = 2SI / (D t)

I is the area moment of inertia of the pipe wall

D is the pipe outside diameter

t is the nominal thickness of the pipe, reduced for corrosion, wear and/or erosion

pi is the internal pressure

po is the external pressure

pb is the API RP 1111 burst pressure

FD is a design factor that depends on the occurrence likelihood of the loads being considered.

For calculated values of loads M, T, pi, and pe, Equation (1) can be solved for the calculated value of combined load FDcalc, and the utilization ratio FDcalc/ FD can be determined. When this ratio is less than unity, the design criterion is satisfied by the calculated load, and when it is greater than unity, the design criterion is violated. Evaluation Approach The basic objective of this work is to identify and evaluate candidate riser systems for servicing high-motion vessels in 10,000-ft water depths in the GoM for 20,000-psi wells.

All-Carbon-Steel SCR Not Likely Feasible.

The requirement of 20,000 psi drives the requirement for sour service. The combination of high-motion host and sour service is expected to rule out SCRs constructed entirely of carbon steel from a fatigue perspective, and that fact sets the stage for this study. What alternatives to carbon steel SCRs exist for these conditions? The answers come along two lines of inquiry: (1) What configuration alternatives are there to simple catenaries? (2) What material alternatives to carbon steel exist?

Configuration Alternatives.

The configuration alternatives shown in Figures 1-6 can be categorized as follows:

1) Conventional catenary a) SCR

2) Catenary-type with distributed buoyancy a) SLWR b) SWR c) CVAR

3) Discrete Buoy with jumper a) TLR b) HRT

Alternatives to the conventional catenary all employ buoyancy to isolate the otherwise fatigue-critical lower portion of the

riser from the motions of the host vessel. A secondary benefit of the buoyancy is to reduce riser payload. For the discrete-buoy options (TLR and HRT), buoy size is primarily governed by simple vertical equilibrium, and the buoy supports most of the weight of the riser system. From a payload perspective, these systems are attractive when limiting riser payload is a primary consideration. For the distributed-buoyancy options (SLWR, SWR, CVAR), the buoyancy requirement is more complex; the buoyancy has a stronger influence on the character of the response of the riser system, the distribution of buoyancy along the length of the riser has to be determined, and the proportion of system weight carried by the buoyancy is less than for the discrete-buoy systems.

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For the discrete-buoy systems, the following early design tasks will be addressed in this paper: Size buoy Determine appropriate length L and nominal offset H for the jumper(s) Size wall thickness and stress joints for rigid pipe jumper Assess payload feasibility for disconnectable- and captive-turret FPSOs

For the distributed buoyancy systems, the following early design tasks will be addressed in this paper:

Size riser wall thickness Determine rough extent and location of distributed buoyancy Assess payload feasibility for disconnectable- and captive-turret FPSOs

Material Alternatives.

The wall thickness sizing discussed previously surveyed the material options under review for this study. The wall thickness sizing results are considered configuration-independent, though not every combination of materials will be considered for every configuration. Table 10 shows the material/configuration combinations that will be studied.

Table 10: Material Configuration Combinations

As noted previously, there is economic incentive to show feasibility for carbon steel solutions wherever possible;

therefore, all of the configuration options will be considered using carbon steel construction. The benefit provided by the freedom from corrosion allowance or cladding provided by super duplex stainless steel (SD)

is worth considering across all configurations; since the wall thickness for SD differs from that of carbon steel only by the 3 mm corrosion allowance, a separate model will not be created for the SD case. The primary difference between carbon steel and SD will be the use of more forgiving S-N curves for the SD material.

Titanium is seen as a potential enabler for the simple catenary because of its high strength-to-weight ratio and its

invulnerability to corrosion and embrittlement, and it is potentially attractive as a rigid jumper material because of its high strength and low modulus. Assuming the viability of a titanium simple catenary, no compelling driver for accepting titaniums high cost is expected in the other catenary-type (distributed buoyancy) options.

High-strength steel is seen primarily as a potential payload-reducer. Since it is not at all clear that a high-strength steel

SCR is viable, the high-strength steel option will be considered for all distributed-buoyancy options. Since jumpers by their nature are payload-friendly, no real benefit for jumpers is expected from high strength steel construction.

CarbonSteel

SuperDuplex* Titanium

HighStrength

Steel

Composite Reinforced

Pipe

Simple Catenary Riser Yes Yes Yes Yes Yes

Lazy Wave Riser Yes Yes No Yes Yes

Steep Wave Riser Yes Yes No Yes Yes

Compliant Vertical Access Riser Yes Yes No Yes No

Jumper forTLR or HRT Yes Yes Yes No No

*Separate riser models for Super Duplex will not be created.

ConfigurationGeometry

Pipe Material

14 OTC 24229

The composite-reinforced pipe (CRP) is expected to provide a good deal of payload reduction, but its viability in a simple catenary is not clear enough to rule it out for the other catenary-like options. There is no driver to consider the CRP option for jumpers, because payload is not expected to be a major driver for jumpers. The omission here of a CRP CVAR is not by any particular design other than reduction of study effort; it can be reinstated if work on other options suggests that a CRP CVAR would be worth studying.

Host Vessel Alternatives.

As noted, the design basis for this study includes the following three host vessel alternatives:

Disconnectable-turret FPSO with a riser payload capacity of 3000 kips Captive-turret (permanently moored) FPSO with a riser payload capacity of 9000 kips Conventional Semi-submersible with no prescribed cap on riser payload

Discrete-Buoy Options Two discrete-buoy riser concepts are presented. These are the tension leg riser (TLR) and hybrid riser tower (HRT). The similarity between the two is that each consists of a long riser supported by a single submerged buoy structure and a jumper that connects the host vessel to the riser, spanning from the host to the buoy. The submerged buoy structure isolates the riser from the vessel motions, and it supports all of the riser weight and some of the jumper weight. This aspect of discrete buoyancy riser concepts is well-suited to applications with limited riser payload capacity such as the disconnectable-turret FPSO considered in this study. The large motions of the host vessel are largely accommodated by flexure of the jumper; the riser below the buoy receives minimal loading from host motions. The primary differences between the two concepts are:

The TLR buoy is a large structure tethered to the seafloor servicing multiple risers and jumpers, while the HRT buoy is attached only to the top of the riser or riser bundle.

The TLR system risers are catenaries, while the HRT riser or riser bundle is composed of near-vertical pipe(s) The HRT considered for this study consists of a single riser pipe rather than a riser bundle. This arrangement is sometimes referred to as a single-line hybrid riser (SLHR) or single-line offset riser (SLOR). The riser pipe materials, insulation coating, and VIV suppression approach outlined previously are applicable to both the jumper and the risers for the discrete-buoy options. The long riser portions are not thought to pose a challenge to feasibility for these concepts. Riser weight is carried by the buoy, and the riser is largely isolated from dynamic loads. The jumpers on the other hand have to accommodate the host vessel motions. The remaining discussions will therefore concentrate on the jumpers. As noted in Table 10 and the accompanying text, only carbon steel and titanium construction for the jumpers will be evaluated in this phase. Super duplex stainless steel will eventually be considered as well, but will not be looked at separately until the fatigue evaluation in the next phase. Catenary Jumper Sizing.

The jumper sizing consists of the wall thickness sizing already discussed (see Tables 5 and 7) as well as a quasi-static exercise to determine an appropriate jumper length L and horizontal span distance H between the ends of the jumper in the nominal configuration. The primary considerations are pipe bending in the sag-bend region of the jumper in the near offset condition and jumper end rotations in the far offset condition, which must be accommodated by the end termination components. Conditions after lowering the host end of the jumper 500 feet to simulate a turret disconnect scenario were checked as well but found not to control.

The high pressure and high operating temperature reflected in the design basis preclude the use of flexible pipe

construction as well as the use of elastomeric flexible joints at the ends of the jumper. Thus, rigid pipe with tapered stress joint end terminations is the only available option. This represents a departure from previously installed systems, which have relied on flexible pipe construction for the jumpers.

It was required in the quasi-static sizing exercise that the jumper and its stress joints satisfy the design criterion (1) with a

design factor FD of 0.8 at all points along their lengths when subjected to the maximum vessel offset of 7.5 percent of water depth and gas-to-surface shut-in conditions. The buoy was assumed to remain stationary during this exercise. Additionally, the host end riser support is assumed to rotate 10 in the plane of the jumper to represent pitch/roll of the host vessel. There are multiple conservatisms built into this feasibility exercise (particularly for the disconnectable turret case) that should be

OTC 24229 15

carefully evaluated in an actual design situation. Note that stress joints are required at both jumper ends, but the stress joint at the host end will generally be the more challenging to design.

Steel. A steel catenary jumper length L of 4400 feet and nominal horizontal span H of 2500 feet was found to be adequate, but a

steel stress joint sized for this application was found too long to be feasible. A titanium stress joint was sized for this steel catenary jumper and found reasonable (taper length of 35 ft).

Titanium. The titanium catenary jumper requires less length L and nominal horizontal span H when compared to steel pipe

construction. Values of L = 3800 ft and H = 2000 ft were found to meet the design criteria. The accompanying stress joint sized for the host end of the jumper is longer but thinner than the titanium stress joint designed for steel catenary jumper because of the significantly lower end tension in the much lighter all-titanium jumper. The stress joint taper length of 45 feet sized for this case is still considered reasonable.

Buoy Sizing.

TLR. As part of DeepStar work, a TLR for 10,000-ft water depth has been described (DeepStar, 2005). The sizing of the

submerged TLR buoy for the current paper took a similar approach. One submerged buoy structure is sized to support the full initial and future production riser count given in the design basis in addition to load from a gas export riser. Assuming carbon steel construction throughout, the total vertical load from risers is estimated to be 8300 kips. The estimated vertical load from the jumpers is 2000 kips. The buoy is assumed to provide a net uplift of 1.5 times the riser load plus mooring load.

The total required displacement of the submerged buoy, which takes into account the weight of the buoy materials, is projected to be approximately 316,000 cubic feet (20,200 kips of displaced seawater). Various shapes have been proposed for TLR buoys. It is anticipated that for the current study, the buoy outside dimensions would be on the order of 120 ft 120 ft 30 ft and that the shape of the buoy would either be a square ring or an H shape in plan. Several tethers would be used to attach the buoy to the seafloor. Ballasting and installation of the TLR submerged buoy are beyond the scope of this paper.

HRT. The following assumptions were made when approximating the size of the HRT buoyancy can:

Carbon steel riser pipe and jumper. Combined weight in seawater for gooseneck (jumper to riser connection component) plus buoy tether is 135

kips. Vertical load from jumper is 165 kips. Maximize riser pipe tension at top of near-vertical riser without exceeding the design criterion (1) with FD = 0.8

value at SITP (ignoring bending). Given the above assumptions, the required net buoyancy of the buoyancy can is estimated to be 1400 kips (635 mT).

This is close to the Cascade-Chinook value, which is reported as 670 mT (Song and Streit, 2011). The Cascade-Chinook cans are 107 feet long and 21 feet in diameter; the steel HRT of this study should be expected to have similar dimensions. Catenary and Catenary-Like Options with Distributed Buoyancy The catenary-type riser concepts with distributed buoyancy considered are the lazy wave riser (SLWR, see Figure 2), the steep wave riser (SWR, see Figure 3), and the compliant vertical access riser (CVAR, see Figure 4). The primary purpose of adding the distributed buoyancy is to induce a compliant shape that isolates the lower portion of the riser from the motions of the host vessel to an extent that results in acceptable minimum fatigue life in the riser. Among these options, the simplest step-out from the simple catenary riser is the SLWR. The remaining two distributed buoyancy concepts (SWR and CVAR) increase system complexity by introducing a riser base that anchors the riser to the seafloor and provides an interface between the riser and the flowline. One candidate riser base system that shows promise for this application is the roto-latch and rigid pipe jumper system employed by the Cascade-Chinook hybrid riser towers (Song and Streit, 2011), provided the system can be adapted for service at 20,000 psi Extent of Distributed Buoyancy; Balanced Buoyancy Approach.

Distributed buoyancy is provided by individual buoyancy modules placed at uniform intervals along the length of a portion of the riser. The portion of riser over which the buoyancy modules are distributed will be referred to as the buoyancy region. The term total net buoyancy will be used to refer to the total uplift provided by all the buoyancy modules on the

16 OTC 24229

riser (weight of seawater displaced by buoyancy modules minus air weight of buoyancy modules), which is a direct measure of the number of buoyancy modules installed.

For this study, buoyancy is applied using a balanced approach such that the buoyancy region has a negative submerged

weight per unit length that is equal in magnitude to the positive submerged weight per unit length of the un-buoyed portions of the riser. This results in similar maximum curvatures in the buoyed and un-buoyed portions of a given riser, and makes the resulting geometries for a given buoyancy section length and location independent of the riser pipe material employed. Exact balance of this type between the buoyed and un-buoyed portions can only hold for a single contents weight; once the buoyancy distribution is balanced for a particular contents weight, the balance will be only approximate when the contents weight changes. For this study, the buoyancy distribution was balanced for each pipe material option using a contents weight of 30 lb/ft3, which is the average of the gas and oil contents weights in the design basis. This balanced approach has to be modified somewhat for the CVAR, as explained later.

Lazy Wave and Steep Wave. For the lazy wave riser, a buoyancy region length of 2500 ft was selected by extrapolation from an existing lazy wave

riser system. This buoyancy region length is designated Option 1. Applying the Option 1 buoyancy region length of 2500 ft to the steep wave configuration, however, results in a nominal inclination of the bottom of the riser that is over 20 off of vertical; this is believed to be excessive from an installation standpoint, so a second, larger buoyancy length (4000 ft - Option 2) was defined. With this increased amount of buoyancy, the nominal inclination at the riser base decreased significantly to 11.5, giving a more installable nominal configuration. Note that the roto-latch connection was stated by Song and Streit (2011) to have been designed with an angular deflection of 20.

CVAR. The CVAR (Ishida, et al., 2001), (Mungall, et al., 2004) can be thought of in some ways as a special case of the SWR

because both riser configurations terminate near the seafloor at a riser base. The significant differences between the two are listed below:

For a given water depth, the CVAR length is shorter overall and does not include a true sagbend or overbend in

the nominal configuration as does the SWR. The CVARs riser base is set closer to the host floater so as to create a nominal riser geometry that is more

vertical at both ends. In order to facilitate its particular shape, the CVAR has two adjacent portions of its length that carry buoyancy

modules. For this study, the upper is neutrally buoyant (zero or near zero submerged weight) while the lower has the balanced buoyancy described earlier.

For the CVAR, the Option 1 and Option 2 designations are still used, but they do not refer to the same lengths of balanced

buoyancy region as they do for the lazy wave and SWR. Rather, they refer to the same total net buoyancy (for a given riser material) as they do for the lazy wave and SWR configuration. Thus an Option 2 CVAR and an Option 2 lazy wave riser constructed from the same materials carry the same number of buoyancy modules, if the same size modules are in use on the two risers. Since, however, some of the buoyancy modules on a CVAR go to creating the neutrally buoyant section, fewer are available to create the balanced buoyancy region; the Option 2 CVAR will therefore have a shorter length of balanced buoyancy than the Option 2 lazy wave (or SWR).

For this study, the neutrally buoyant region for the CVAR was selected to be 1250 feet long, and this value remains

constant for the different buoyancy options. Table 11 provides geometry descriptions of the catenary and the catenary-type distributed buoyancy riser concepts. The

first (i.e., un-highlighted) buoyancy option listed for each riser configuration represents a first-trial reasonable equilibrium configuration that is thought to be installable. The second (highlighted) buoyancy option listed for each riser configuration represents second-trial increased buoyancy configurations created in an attempt to reduce riser payload to a level acceptable for a disconnectable turret FPSO. Table 12 summarizes the total net buoyancy requirements for the various material options and coverage options.

OTC 24229 17

Table 11: Catenary and Catenary-like Distributed Buoyancy Configurations

In an actual design scenario, the effect of contents weight variations on the performance of the the various distributed-

buoyancy configurations would need to be taken into account. The present feasibility study will not attempt to optimize performance by fine-tuning buoyancy across multiple contents scenarios.

Table 12: Buoyancy Requirements

Payload Results. The chart in Figure 7 shows riser payload demand on a host supporting the entire initial and future production riser count

given in the design basis. The represented configurations are the first trial buoyancy options (non-highlighted lines in Table 11). The following observations may be made:

All of the steel options have payload requirements that are above the 3000-kip assumed maximum allowable

payload for a disconnectable turret. Among the catenary riser material options, only titanium has a small enough total payload to work with a

disconnectable turret. This could indicate an enabling technology, depending on the fatigue life estimates to be made in the next phase of work.

The composite reinforced pipe shows promise from a payload perspective as either a lazy wave or steep wave riser. A lazy wave riser with 2500 feet of buoyancy (Option 1) has a payload requirement just above the 3000-kip threshold and the steep wave riser with buoyancy Option 2 is approximately 15 percent under the assumed disconnectable turret payload threshold.

(degrees) (feet) (feet) (feet)

Simple Catenary Riser N/A 5.0 N/A 10,890 3,000

1 2,500 14,330 6,4602 4,000 15,140 6,3602 4,000 12,500 4,5753 5,000 12,250 4,0331 1250/1875(a) 10,500 2,2502 1250/3375(a) 10,500 2,250

Notes:(a) Neutrally Buoyant Region Length / Balanced Buoyancy Region Length

Compliant Vertical Access Riser N/A

SuspendedSpan

Length

HorizontalSpan

Length

Lazy Wave Riser 5.0

Steep Wave Riser 5.0

Configuration Geometry BuoyancyOption

Top Angleoff Vertical

Buoyancy Region Length

w/o HIPPS 1.618 inches 435 kips 700 kips 870 kips

w/ HIPPS 1.152 inches 298 kips 477 kips 596 kips

1.068 inches 310 kips 496 kips 610 kips

0.618(a) inches 184 kips 295 kips N/A

Notes:(a) Thickness of steel pipe; carbon fiber composite is 1 inch thick

Carbon Steel

High Strength Steel

Composite Reinforced Pipe

Riser Pipe Material

Wall ThicknessIncluding 3mm

Corrosion Allowanceor CRA Cladding

Total Net Buoyancy

CoverageOption 1(2500 ft)



18 OTC 24229

Figure 7: Total Riser Payload, Buoyancy Trial 1

A second trial set of riser configuration/buoyancy option configurations was compiled and evaluated in an attempt to

drive down the payload requirements, particularly of steel risers. These are the highlighted lines in Table 11. The chart showing the results for this second trial is shown in Figure 8. The carbon steel with HIPPS case and the high strength steel case look feasible from a payload standpoint as a CVAR with buoyancy Option 2. The carbon steel case, however, does not show promise for being able to be used at 20,000 psi with a disconnectable turret FPSO for any of the distributed buoyancy configuration options.

The composite reinforced pipe was also evaluated in the second trial, indicating that with buoyancy Option 2, it becomes

payload-feasible as a lazy wave riser for a disconnectable turret in addition to its already established feasibility as a steep wave riser.

It should be noted that the carbon steel with HIPPS case is representative of a riser wall thickness corresponding to a

design pressure of about 12,000 psi. Therefore, this study can apply to a broader scope than just the 20,000 psi SITP defined in the design basis. If ultra-deepwater fields are discovered with reservoir pressures yielding SITP values (at subsea wellhead) in the 10,000- to 12,000-psi range, then a disconnectable turret supporting carbon steel risers with distributed buoyancy becomes a potential concept.

0

1,000

2,000

3,000

4,000

5,000

6,000

7,000

8,000

9,000

10,000

Tota

l Ves

sel P

aylo

ad fo

r 8 R

iser

s, k

ips

Carbon SteelCarbon Steel with HIPPSHigh Strength SteelComposite Reinforced PipeTitanium

Jumper(TLR or HRT)

CVARBuoy Opt. 1

Steep WaveBuoy Opt. 2

Lazy WaveBuoy Opt. 1

Catenary

Wellhead Pressure = 20,000 psiAssumed Upper Limitfor Captive Turret FPSO

Assumed Upper Limit forDisconnectible Turret FPSO

OTC 24229 19

Figure 8: Total Riser Payload, Buoyancy Trial 2

Technology Gaps The following technology gaps have been identified at this time for the design scenario presented here:

Swivel capable of withstanding internal pressure approaching 20,000 psi. Thermal insulation capable of service at 350 deg F. Riser base flow path components capable of containing 20,000 psi internal pressure. Use of composite-reinforced pipe as a flowline/production riser. Use of a captive-turret FPSO in the GoM. Proven inhibitor methodology for 20,000 psi / 350 F.

Conclusions The conclusions from this phase of the study primarily pertain to what options look promising for dynamic evaluation in the next phase. A conventional semi-submersible and a captive-turret FPSO have riser payload capacities that allow them to support heavy riser systems. A carbon steel riser sized for 20,000 psi SITP will be the base-case as a lazy wave riser for these two high-motion vessels. For completeness, a simple SCR modeled with the same carbon steel wall thickness is to be analyzed as well even though it is not expected to meet fatigue criteria. The fatigue benefit of using titanium pipe locally in the touchdown region in an otherwise steel-constructed SCR may also be investigated. For a disconnectable turret FPSO, the discrete buoy concepts, TLR and HRT, are the most promising from the standpoint of riser payload because the host vessel supports only the jumpers spanning from the turret to the submerged buoy. For these systems, carbon steel jumpers with titanium stress joints and titanium jumpers with titanium stress joints are feasible from a strength perspective. A few catenary-type distributed buoyancy riser concepts also have riser payload estimates that do not exceed the assumed limit of 3000 kips for a disconnectable turret. None of these, however, are configurations constructed of carbon steel pipe sized for 20,000 psi SITP at the subsea wellhead. The riser concepts that meet the assumed 3000 kip riser payload limit of a disconnectable turret FPSO and which will be studied further are:

Titanium Catenary Riser Composite Reinforced Pipe shaped as both a lazy wave riser and a steep wave riser CVAR (buoyancy option 2) constructed of

Q125 high strength steel

0

1,000

2,000

3,000

4,000

5,000

6,000

7,000

8,000

9,000

10,000

Tota

l Ves

sel P

aylo

ad fo

r 8 R

iser

s, k

ips

Carbon Steel

Carbon Steel with HIPPS

High Strength Steel

Composite Reinforced Pipe

CVARBuoy Opt. 2

Steep WaveBuoy Opt. 3

Lazy WaveBuoy Opt. 2

Wellhead Pressure = 20,000 psi

Assumed Upper Limit forDisconnectible Turret FPSO

20 OTC 24229

carbon steel sized with HIPPS. Note that the HIPPS case can also represent a riser design for 12,000 psi SITP (at subsea wellhead).

Further optimization will be explored to find a steep wave buoyancy option that allows the use of high strength steel or carbon steel with HIPPS. Addition of lightweight buoyancy coating to the upper-middle portion of a steel SCR to reduce payload burden may be considered as well. Acknowledgements The authors would like to acknowledge the efforts and support of the RPSEA Project 10121-4401-01 Project Working Group, its project champion, Pierre Beynet (BP), and the operating companies they represent: Shell, Petrobras, ExxonMobil, ConocoPhillips, and BP. Funding for the project is provided by the Research Partnership to Secure Energy for America (RPSEA) through the Ultra-Deepwater and Unconventional Natural Gas and Other Petroleum Resources program, authorized by the U. S. Energy Policy Act of 2005. RPSEA (www.rpsea.org) is a non-profit corporation whose mission is to provide a stewardship role in ensuring the focused research, development and deployment of safe and environmentally responsible technology that can effectively deliver hydrocarbons from domestic resources to the citizens of the United States. RPSEA, operating as a consortium of premier U. S. energy research universities, industry, and independent research organizations, manages the program under a contract with the U. S. Department of Energys National Energy Technology Laboratory. References

1. Frantz, J. P., Koster, T. E., Liang Yu, Haire, W. M., and Shimamura, Y. 1996. Liuhua 11-1 Development-Design and Fabrication Considerations for Nanhai Sheng Li for the FPSO. Paper OTC 8188 presented at the Offshore Technology Conference, Houston, Texas, 6-9 May.

2. CNOOK, Ltd. 2001. FPSO "Nanhai Endeavor" Launched Ahead of Schedule. Press Release dated 19 June, 2001, Beijing. 3. Ishida, K., Otomo, K., Hirayama, H., Okamoto, N., Nishigaki, M., and Ozaki, M. 2001. An FPSO with Surface Wells and

Workover System in Deepwater. Paper OTC 12990 presented at the Offshore Technology Conference, Houston, Texas, 30 April-3 May.

4. Mungall, C., Haverty, K., Bhat, S., Andersen, D., Sarkar, I., Wu, J., and Martensson, N. 2004. Semisubmersible Based Dry Tree Platform with Compliant Vertical Access Risers. Paper OTC 16199 presented at the Offshore Technology Conference, Houston, Texas, 3-6 May.

5. DeepStar. 2005. DeepStar VII Project Semisubmersible in Ultra-deepwater (8,000 10,000). DSVII CTR-7404-G. Referenced by permission.

6. Banon, H., Nygard, M., Loken, R., Lu, R., and Garrett, D. 2007. HP/HT Semisubmersible for 10,000 ft in Gulf of Mexico. Paper OTC 18467 presented at the Offshore Technology Conference, Houston, Texas, 30 April-3 May.

7. Baxter, C. F., Schutz, R. W., Caldwell, C. S. 2007. Experience & Guidance in the Use of Titanium Components in Steel Catenary Riser Systems. Paper OTC 18624 presented at the Offshore Technology Conference, Houston, Texas, 30 April-3 May.

8. ANSI/NACE. 2009. Petroleum and natural gas industriesMaterials for use in H2S-containing environments in oil and gas production MR0175/ISO 15156.

9. American Petroleum Institute. 2009. Design of Risers for Floating Production Systems (FPSs) and Tension-Leg Platforms (TLPs) Recommended Practice 2RD 1st Ed. 1998, Reaffirmed 2006, Errata 2009.

10. Stanton, P., Kavanagh, K., Garrett, D., and Mork, K. 2010. New Code for the Design of Dynamic Risers for Floating Production Installations. Paper OTC 20816 presented at the Offshore Technology Conference, Houston, Texas, 3-6 May.

11. Hoffman, J. 2010. Parque das Conchas Pipeline, Flowline, and Riser System Design, Installation and Challenges. Paper OTC 20650-MS presented at the Offshore Technology Conference, Houston, Texas, 3-6 May.

12. Song, R., and Streit, P. 2011. Design of the Worlds Deepest Hybrid Riser System for the Cascade & Chinook Development. Paper OTC 21338 presented at the Offshore Technology Conference, Houston, Texas, 2-3 May.

13. American Petroleum Institute. 2011. Design, Construction, Operation, and Maintenance of Offshore Hydrocarbon Pipelines (Limit State Design) Recommended Practice 1111 4th Ed 2009, Errata 2011.

14. INTECSEA Worley Parsons Group. 2012. Global Frontiers Quarterly Journal Q2-2012.

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