OTC-14152-MS.pdf

Copyright 2002, Offshore Technology Conference This paper was prepared for presentation at the 2002 Offshore Technology Conference held in Houston, Texas U.S.A., 69 May 2002. This paper was selected for presentation by the OTC Program Committee following review of information contained in an abstract submitted by the author(s). Contents of the paper, as presented, have not been reviewed by the Offshore Technology Conference and are subject to correction by the author(s). The material, as presented, does not necessarily reflect any position of the Offshore Technology Conference or its officers. Electronic reproduction, distribution, or storage of any part of this paper for commercial purposes 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 where and by whom the paper was presented.

Abstract This paper presents the design and implementation of buoyed steel catenary risers (SCR). Two 6" steel catenary risers, installed in November-December 2001, were designed to include 50,000-pounds of buoyancy per riser in 3,300 feet of water for Mariner Energy, Inc.'s King Kong/Yosemite field development in the Green Canyon area of the Gulf of Mexico. The risers are suspended from the Agip Petroleum "Allegheny" tension leg platform (TLP) where reduced riser loads were required for additional process equipment and safe TLP operation. Introduction The King Kong/Yosemite development consists of three subsea gas wells located in Green Canyon Blocks 472, 473 and 516 in water depths of 3,800 - 3,900 feet. The subsea wells are commingled at a manifold located in Green Canyon block 472 into two 7 production flowlines. These 7 flowlines transit approximately 15 miles to Green Canyon Block 254, in 3,300-foot water depth, where they transition to 6 SCRs to board the Agip Petroleum Allegheny mini Tension Leg Platform, (TLP). The field layout is pictured in Figure A-1. The riser locations near the platform are illustrated in Figure A-2. During the course of the riser design, and after order of riser pipe and titanium stress joints, it was discovered that the weight of additional process equipment and the planned SCRs placed the Allegheny TLP beyond its allowable load capacity. In order to maintain a safe load budget for the TLP and to account for future process equipment addition it was decided

to reduce the total TLP load by 100-kips via buoyancy on the risers. The challenge involved making a buoyed riser system work within the constraints of a non-buoyed, simple catenary SCR. The buoyed riser design involved an iterative approach to optimize the riser and buoyancy to satisfy the requirements for design codes, S-Lay installation and project cost and schedule constraints. The position and characteristics of the buoyancy were important to produce a riser catenary that met project requirements and code stress and fatigue criteria. In the final configuration, each riser utilizes 271 separate, surface installed buoyancy modules applied over a continuous length of 800-feet, starting below the VIV strakes, to provide 50-kips of net buoyancy plus a 5% safety factor. The resultant riser has an arc length of 3,532-feet with a horizontal offset of 1028-feet at a departure angle of 8. Figure A-3 illustrates the final riser configuration. The final profile of the riser is a shallow "S" exhibiting a slight double catenary. Analysis shows that, compared to a conventional SCR, the buoyed riser exhibits increased bending and dynamic stress in the sagbend. This results from the "lighter" SCR motion characteristics and the departure from the second catenary, best illustrated as a shorter horizontal offset to touchdown than conventional simple catenary SCRs. Riser stresses are managed by placement, distribution and size of buoyancy modules. Execution of this project demonstrates the viability of the buoyed SCR for deepwater developments and how this concept can be used to expand or enhance the capacity of floating production systems.

This paper presents the design and implementation of the buoyed steel catenary risers (SCR) for the Mariner Energy, Inc., King Kong/Yosemite field development. Included is an overall description of the riser key parameters; design including static, dynamic, vortex induced and wave fatigue analyses; and installation experience. In the description of the static analysis, comparing and contrasting results of the original simple catenary SCR, the final buoyed SCR configuration and one of the iterative buoyed SCR

OTC 14152

Design and Implementation of the First Buoyed Steel Catenary Risers Douglas R. Korth, Mentor Subsea Technology Services, Inc. (a J. Ray McDermott company); Bob Shian J. Chou, Mentor Subsea Technology Services, Inc. (a J. Ray McDermott company); Gary D. McCullough, Mariner Energy, Inc.

2 D. KORTH, B. CHOU, G. MCCULLOUGH OTC 14152

configurations examined before selecting the preferred concept will be presented. The conclusion will summarize the system and the lessons learned throughout the project with recommendations presented for future development and changes that would be incorporated. Design Parameters The basic parameters of the SCR, determined during the preliminary design and before the addition of buoyancy to the system area as follows: Hang-off type: Titanium Stress Joint Stress Joint Length: 27-feet Riser Material: 6.625 O.D. x 0.791 w.t. API 5L X65 Riser Departure Angle: 8-degrees Hang-off depth: 79.5-feet External Coating: 12-14 mils Roughcoat F.B.E. VIV Strakes: 5D design Strake Length/Placement: 520-feet commencing after

stress joint Touchdown Coating: 1,000-feet of Three-Layer

Polyethylene, 125 mils. Buoyancy System Selection. The buoyancy system selection was complicated by the fact that the process system could not tolerate arrival temperatures below 50 F. Given the flowing temperature of the gas, 38 F, it was essential that the buoyancy not provide any insulation, allowing the gas to warm up in the water column. This ruled out the use of a uniformly distributed coating/buoyancy system or other system that maintained direct contact with the riser pipe. Resultantly, the selection focused on a system of independent modules. The drivers for the selection of the buoyancy system were: Configuration Dimensions

Configuration. The final buoyancy system selected utilized a series of independent buoyancy modules banded to the riser. The buoyancy modules were designed with an inside diameter larger than the pipe O.D. to provide a annulus around the riser pipe. Additionally, flow channels were molded on the inside diameter of each module to allow for continuous water flow around the riser, providing the required warming of the produced gas. Figure A-4, illustrates the general arrangement of the buoyancy module.

Dimensions. The module dimensions we driven by three factors:

Allowable Local Buoyancy. During design it was determined that the maximum allowable local buoyancy was 67.5 lbs./foot. It was found that if the buoyancy provided more that 67.5 lbs./foot the riser would begin to take a true lazy-S configuration with a flat or raised profile located mid-riser. Based on the preliminary analyses of a buoyed riser, the

existing fixed parameters of the riser material, stress joint and departure angle and project time concerns it was felt that the true lazy-S configuration would not be feasible. Consequently, it was specified for the unit buoyancy to remain below this value. This was reiterated later during the dynamic analysis where the impact of the unit buoyancy on the motion response and bending stress was further magnified.

Installation Load. The modules consist of an outer polymer shell with a wax infill material, Evasyn, with macroshperes in the matrix. The design installation load, as determined by the vessel stinger characteristics and lay tensions was 11-tons per roller pair for the Allseas vessel Solitaire. From this the buoyancy supplier, CRP, Inc., determined that the infill material density required to withstand the load was 600 kg/m3. The outer shell was found to be sufficiently durable for this application. Full scale testing of the buoyancy modules was performed prior to delivery (see figure A-5).

Handling. The module could be made in any size requested. However, for installation handling ease, it was decided to limit the length of the modules to no more than four feet. The module diameter was a result of the length restriction and the maximum allowable unit buoyancy.

The modules are attached to the riser using three inconel bands per module. Every third module, a polymer stopper clamp is installed to prevent the buoyancy modules from sliding up the pipe as it enters the water. Additionally, a transition piece with a steel clamped bulkhead is affixed on either end to provide a taper for entering the lay vessel stinger and to serve as the ultimate buoyancy holdback during operation. The resultant modules have the following characteristics: Length: 35.4-inches Outside Diameter: 26.7-inches Inside Diameter: 7.2-inches Weight (dry): 142 lbs. Buoyancy per module: 193-lbs. Unit Buoyancy: 66.15 lbs./foot Figures A-6 and A-7 show the buoyancy during installation. Design Results In this section is presented the results of the SCR design. A walk through of static, dynamic and fatigue (wave and VIV) is presented. A comparison is made at the static level of three riser configurations to illustrate the uniqueness of the buoyed riser. The three risers compared are: Simple Catenary (non-buoyed) Has no buoyancy. There

are 520-feet of strakes starting immediately after the stress joint. (See Figure 1 and A-8)

OTC 14152 DESIGN AND IMPLEMENTATION OF THE FIRST BUOYED STEEL CATENARY RISERS 3

Figure 1 Simple Catenary SCR Static Configuration

Buoyed SCR 1 This configuration has 900-feet of

buoyancy, 305 modules, starting 1,500-feet below the VIV strakes. (See Figure 2)

Figure 2 Buoyed SCR 1 Low Buoyancy Static Configuration

Buoyed SCR 2 This configuration has 800-feet of

buoyancy, 271 modules, starting 30-feet below the VIV strakes. (See Figure 3 and A-2)

Figure 3 Buoyed Riser 2 High Buoyancy Static Configuration

Static Analysis Results. The static analysis was performed using the ABAQUS program, and checked using OFFPIPE, to give the total arc length and the horizontal scope (touchdown) of the SCR as well as to calculate axial, bending and total static stress. Analysis results are shown for three configurations. The differences are shown for the simple catenary (non-buoyed) and two configurations of buoyed SCR. The results of the static analysis were used as a gateway to select the final buoyed riser configuration. Figures 4 - 6 show the total stress profiles for the three configurations.

Figure 4 Simple Catenary SCR, Total Static Stress Profile

Mariner KKY SCR (No Buoyancy Modules) Total Static Stress Distribution (with 7500 psi Internal Pressure)

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Figure 5 Buoyancy Configuration 1, Total Static Stress Profile

Figure 6 Buoyancy Configuration 2, Total Static Stress Profile The static analysis results for all three cases are

summarized in Table 1. It is shown that both buoyed configurations were within the allowable stress envelope. The buoyed SCR bending stress is shown to be almost 50% higher when compared to the non-buoyed SCR. It is also seen that total stress is governed by the bending stress for both buoyed cases. Also notable is the dramatic change in the stress profile in the area of the buoyancy. This is attributed to the steep decline of the riser after the buoyancy. Figure 5, illustrates that the buoyancy placed lower in the riser serves to decrease the bending and overall stress of the riser by lifting the sagbend by approximately 10% when compared to Figure 6, the high buoyancy case.

Riser Configuration

Simple Catenary

Buoyed 1 (low buoyancy)

Buoyed 2 (high buoyancy)

Max. Static Bending Stress 15.8 Ksi 23.5 Ksi 26 Ksi

Max. Total Static Stress 27.5 Ksi 34 Ksi 36.5 Ksi Allowable

Stress 43 Ksi 43 Ksi 43 Ksi Arc Length 3720 3826 3532 Touchdown

Distance 1420 1725 1028 Departure

Angle 8 8 8 Top Tension 136 kips 86 kips 83 kips Amount of

Buoyancy for 50-kip

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N/A 59.5 kips 50 kips

# of Buoyancy Modules Reqd N/A 305 256

Total # of modules with

5% S.F. N/A 305 271

Total buoyancy with

5% overage N/A 59.5 kips 53 kips

Table 1 Static Analysis Result Comparison Buoyed Configuration 2 (high buoyancy), while still

exhibiting a higher total and bending stress than Buoyancy Configuration 1 (low buoyancy), is much more efficient, requiring 44 fewer buoyancy modules (or 9.5-kips of buoyancy) per riser to achieve the desired 50-kip reduction. This is attributed to the buoyancy being placed closer to the surface in a near vertical orientation as opposed to being placed in the sagbend area where the riser was approaching a horizontal profile. The resultant overall project cost and schedule benefit associated with fewer buoyancy modules led to the selection of Buoyed Configuration 2 (high buoyancy) (see figures 3, 6 and A-2) as the final riser configuration.

Dynamic Analysis. The dynamic analysis was performed using the finite element program ABAQUS and various Mentor in house pre and post processors. It was determined that the limiting condition for the risers is when the TLP is in the Near position, or moving closer towards the riser touchdown point. Using this limiting condition, three cases were used to determine the riser suitability, the 100-year hurricane, 10-year hurricane + 100-year eddy current and the 10-year hurricane. Table 2 illustrates the Design load matrix with the analysis results:

Case Condition Environment TLP Position

Max Stress (ksi)

Allowable (ksi)

1 Extreme 100-yr. hurricane Near 48.5 52 2 Extreme 100-yr. hurricane Far 39.5 52 3 Extreme 100-yr. hurricane Trans. 38.4 52

4 Extreme 10-yr. hurricane + 100-year eddy Near 46.2 52

5 Operating 10-yr. hurricane Near 41.6 43.33 Table 2 Load Case Matrix with Dynamic Analysis Result

Mariner KKY SCR Total Static Stress Distribution (with 7500 psi Internal Pressure)

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Mariner KKY SCR (with 900' Buoyancy modules 1500' from Strakes) Total Static Stress Distribution (with 7500 psi Internal Pressure)

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As seen in Table 2, the buoyed riser has a comparatively high dynamic stress relative to what would be expected from a simple catenary riser. The peak stresses are all located in the touchdown regions with the bending stress being the dominant factor for the total stress.

Figures 7 and 8 illustrate the total dynamic stress profile for the 100-year hurricane, near position and the 10-year hurricane, near position, respectively.

As seen in Figure 7, the SCR follows a normal stress profile from the hang-off point until the start of the buoyancy. Approximately 10-feet into the buoyancy, the stress profile exhibits a sharp peak, increasing approximately 5 Ksi. The stress declines slowly throughout the buoyancy then drops off at the end of the buoyancy. The stress peaks once again just before the touchdown point.

Figure 7 Maximum Total Stress Distribution, 100-year Hurricane, Near Position

Figure 8 illustrates the total stress profile for the governing

operating case, the 10-year hurricane.

Figure 8 Maximum Total Stress Distribution, 10-Year Hurricane, Near Position

In general, the 10-year storm dynamic stress follows a similar profile to the 100-year case. Once again, stress peaks are seen in the buoyed region and just before touchdown. One dissimilarity is the second stress peak located in the center of the buoyed region. Also, it is seen that the stress decreases and peaks again at the end of the buoyancy where the pipe is departing into the second, shallow, catenary. These peaks and valleys in the buoyed region are directly attributed to bending stress and the higher response due to the shorter period motions of the TLP. To illustrate the governance of the bending stress on the total dynamic stress, Figures 9 and 10 illustrate the maximum axial stress and maximum bending stress for the buoyed riser in the 10-year storm near position.

Figure 9 Maximum Axial Stress Distribution, 10-Year Hurricane, Near Position

The axial stress profile shown in Figure 9, with the stress

rise throughout the buoyed area, was seen for all cases run, with changes seen in magnitude only.

Figure 10 Maximum Bending Stress Distribution, 10-year

Hurricane, Near Position

A comparison of Figure 8 with Figure 10 shows that the bending stress indeed governs the riser dynamic stress profile. The analysis to determine the desired unit buoyancy of 67.5

Mariner KKY SCR Dynamic Stress Distribution, 10 Year Hurricane Near Position

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lbs./foot indicated that the stress peaks seen in Figures 5 and 7 (static profiles) in the buoyed region and consequently in the touchdown region, would be magnified commensurately with a higher unit buoyancy.

In summary, the dynamic analysis reveals that the buoyed SCR exhibits eccentric behavior when compared to a conventional SCR. The stress peak at the buoyed region is attributed to the buoyancy induced bending and resultant change in the catenary profile and is magnified, dynamically, by the increased motions of the lighter SCR. The dynamic analysis also revealed the criticality of maintaining the unit buoyancy at a value that does not create a steep double catenary for this configuration. It was concluded that higher unit buoyancy resulted in elevation in bending stress in the buoyed and touchdown regions. However, relocating the buoyancy further down the riser, closer to touchdown, results in reduced bending stress, as was found for buoyancy configuration 1, shown in Figures 2 and 5. Fatigue Analysis. The riser fatigue analysis was performed in two parts:

Wave Fatigue Vortex Induced Vibration (VIV) The results of these two analyses were combined to

produce the final fatigue life results. Wave Fatigue. The wave fatigue analysis was performed

using the finite element analysis program, ABAQUS. The wave scatter diagram was divided into 29 bins of significant wave height and peak period combinations based on probability of occurrence. The most sever sea state within each bin was used to represent that particular data set. Using this information, a time domain, random wave simulation of the dynamic response of the riser was generated. The statistical results of the analysis and the API-X curve was used for prediction of the general fatigue life of the buoyed SCR. The characteristic sea states for the fatigue analysis are shown in Table A-1.

Based on the dynamic stress distribution along the SCR, four critical regions were identified and checked for their expected fatigue life:

Top and strake region REGION 1 Strake to Buoyancy Transition REGION 2 Buoyancy section REGION 3 Touch-down region REGION 4 The root-mean-square (RMS) stress distribution is shown

in Figure 11, illustrating the critical regions for fatigue.

Figure 11 Total RMS Stress Distribution Along the SCR Length As Figure 11 illustrates, the areas of maximum RMS stress

are located in the four identified regions. A stress concentration factor (SCF) of 1.2 was used to

evaluate these critical regions. For sensitivity, this was compared with an SCF of 1.0. The results of the wave fatigue analysis are summarized in Table 3.

Region SCF=1.0

(years) SCF=1.2 (years)

1 414 209 2 9564 4836 3 2320 1173 4 772 390 Table 3 - Minimum Fatigue Life, Wave Induced

Table A-2 shows the minimum RMS stresses and Fatigue

life along the entire length of the riser. In the wave fatigue analysis the same trend is evident as in

the dynamic analysis having a peak response noticeable in the buoyed region. This is illustrated as a higher RMS stress and, consequently, lower fatigue life.

Vortex Induced Vibration. Vortex Induced Vibration

fatigue was performed by 2H offshore using the program SHEAR 7, which predicts VIV effects based on mode superposition, and a series of pre and post processors.

VIV analysis was performed using seven current profiles acting in the transverse direction. It has been assumed that the transverse current direction provides the worst VIV response.

The resultant fatigue damage due to the VIV effect is presented in Figure 12.

Case 7, Total RMS Stress Distribution

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Figure 12 Fatigue Damage Rate Due to VIV Effects

Figure 12 illustrates the effect of the strakes on the fatigue life of the riser. It is seen that the damage rate is highest in the buoyed region and then peaks a second time at the touchdown point. Figure 13 shows the curve for the resultant fatigue life due to VIV effects.

Figure 13 Fatigue Life Due to VIV Effects

In Figure 13, it is seen that, corresponding to the fatigue

damage rate in Figure 12, the fatigue life along the riser is lowest in the buoyed region. The low fatigue life is attributed to the more active response of the riser in that region and the added current effect of the modules on the riser. The low end VIV fatigue life is seen to be 318 years for the buoyed region. Table 4 summarizes the fatigue life due to VIV effect for regions 1-4, previously defined.

Region SCF=1.2 (years) 1 1513 2 2696 3 318 4 471

Table 4 - Minimum Fatigue Life, VIV Induced Fatigue Summary. The results from both the wave/current

and VIV fatigue analyses were combined to estimate the overall minimum fatigue life for the buoyed risers.

The final results of the fatigue summary for regions 1-4 are presented in Table 5, using a stress concentration factor (SCF) of 1.2.

Region Dynamic

(years) VIV

(years) Overall (years)

1 209 1513 208 2 4836 2696 2016 3 1173 318 251 4 390 471 310

Table 5 - Summary of Minimum Fatigue Life

A complete series of fatigue tests was performed after completion of the analysis to verify the design and welding procedures. It should be noted that all fatigue samples greatly exceeded the design fatigue life. Table A-3 contains the complete fatigue stress spectrum used for fatigue testing.

In conclusion of the fatigue analysis, it is seen that the addition of buoyancy has a measurable effect on the fatigue life. The wave/current dynamic fatigue life is affected by the reduced weight and drag forces induced by the buoyancy modules on the riser. A similar effect is seen with respect to VIV fatigue. It is theorized that adding strakes to the buoyancy and/or reducing the module size (by decreasing density or increasing the length of distribution) can reduce the wave and fatigue effect on the buoyed section.

Installation The installation of the buoyed risers was performed by Allseas using the DP pipelay vessel, Solitaire. The installation was a first-end layaway from the TLP, using a pull-in winch arrangement on the platform.

Two main aspects of the installation were impacted by the addition of buoyancy:

Initiation of the SCR at the TLP Laydown of the SCR

SCR Initiation. The buoyed SCRs were initiated at the TLP using a winch arrangement on the platform for pull-in.

During the initiation, the first sections to depart the stinger were the stress joint, strakes and buoyancy, in that order. The uniqueness of the buoyed system led to a vertical S bend in the suspended pipe section.

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In order to control this, strict attention was paid, both procedurally and operationally to the vessel position and lay speed to prevent the pipe from being too greatly affected by currents or allowing to much slack in the pipe which could result in a steeper S bend.

Resultantly, there were no problems during this operation for both risers and the strain was maintained below the allowable limit of 0.3%.

The decision to initiate from the platform was made late in the design and procurement process. Resultantly, the buoyancy was designed to sustain installation compressive loads for a termination at the platform. The design compressive load for this installation scenario was 11- tons per roller pair. As a result a higher density, and hence less efficient, buoyancy material was required. If the much reduced compressive loads from the platform initiation scenario were applied, a lower density, more efficient buoyancy material could have been utilized resulting in fewer or smaller modules.

SCR Laydown. Placement of the touchdown point and

monitoring of the final departure angle is critical for all SCR installations. The buoyed SCRs have a design departure angle of 8, +/- 1 to account for installation and TLP position tolerance. A target box of 10-feet x 20-feet was established with the center of the target box being the design touchdown length of 3,532-feet.

During the laydown of the buoyed SCR and monitoring of the touchdown point, two key items became apparent.

First, the buoyed SCR has a tendency to return to a more natural shape. This was encountered during the laydown of the first buoyed SCR when too low of a tension was maintained on subsequent pipelay after setting the SCR in the touchdown target box. As a result, the touchdown point, marked on the pipe with a long baseline sonar (LBL) transponder, moved approximately 100-feet closer to the TLP from its position inside the target box.

The second item became apparent immediately following this when the surface departure angle was inspected, using an electronic inclinometer mounted on the lower tip of the stress joint. It was found that the surface departure angle, with the touchdown point 100-feet off target was 7.8, well within the installation tolerance. Upon pulling the riser so the calculated pipe touchdown point fell within the target box, the surface departure angle was re-inspected to reveal that the departure angle was now exactly 8, as designed. What was determined is that the second, steeper departure angle after the buoyancy changes easier and more rapidly than the surface departure angle as a result of the spring effect provided by the buoyancy. It is possible, as evidenced by the previously described experience, that the surface departure angle can be correct within tolerances and still have the touchdown point drastically off location. The net result of this would be increased bending stress in the critical regions of the riser, the buoyed section and the touchdown region, as well as a

reduced fatigue life. This experience begs the conclusion that the exact spotting

and control of the touchdown point is critical to the successful installation of the buoyed SCR and no assumption should be made as to the correctness of the touchdown based on surface departure angle measurement. Conclusion

It has been shown that in the static analysis, the buoyancy augmented the bending stress at the touchdown point over a simple catenary by introducing a steep departure angle, smaller than the surface departure angle, after the buoyancy. This same effect was seen in both buoyancy configurations presented. It was shown that, in the static profile, placing the buoyancy lower in the water column nearer the touchdown region reduced the bending stress. It can be concluded from this that the dynamic stress would also be reduced and fatigue life increased.

The trade off between the two buoyed riser concepts, compared statically, is a reduced stress profile for the low buoyancy case with an approximate 20% lower buoyancy efficiency, requiring 44 additional modules per riser to achieve the required 50-kip weight reduction at the platform. With both buoyed configurations falling within the design code allowable for static stress, the configuration with buoyancy placed high in the water column immediately following the strakes was selected since it provides the best project value.

Carrying the high buoyancy through the dynamic analysis it was seen that the stress continue to peak in the buoyed and touchdown regions while still remaining within allowable limits in extreme and operating events.

Fatigue analysis revealed that the buoyed region is sensitive, particularly to VIV. To reduce the VIV sensitivity, it was concluded that strakes could be added to the buoyancy modules/region.

During installation it was seen that there are sensitivities specific tot he buoyed riser, most notable the relationship between the touchdown point ands surface departure angle.

In conclusion it was demonstrated that the buoyed SCR can be successfully designed and implemented within existing code limits and using the design constraints of a non-buoyed, simple catenary SCR. Also, it was shown that sensitivities can be managed and overcome allowing implementation for future SCR developments.

Other Applications In addition to being used to reduce the load on the conventional realm of floating production facilities for SCRs such as TLPs, Spars and Semi Submersibles, some of the experiences from this project indicate potential use of SCRs on FPSOs.

One of the critical design factors for SCRs in FPSO application is the sometimes large heave motions particularly in the South Atlantic, which lead to large SCR touchdown point motions. The phenomenon experienced during


installation of the first King Kong buoyed SCR, described previously, demonstrates a de-coupling or weakening of the relationship between the surface departure angle and the touchdown point position. It is theorized that this characteristic of the buoyed SCR would serve well to diminish the touchdown point variance due to heave for an FPSO application, particularly with a configuration similar to Buoyed Configuration 1 (low buoyancy) described in this paper. Acknowledgements The authors wish to acknowledge the efforts of Mariner Energy, Inc., Mentor Subsea Technology Services, Inc. (a J. Ray McDermott company), Allseas Services, U.S. and the crew of the vessel Solitaire, CRP, Inc., 2H Offshore and all the participating individuals who made the King Kong/Yosemite buoyed riser design and installation successful.. References 1. Chou, Shian J., Korth, D.R., Suschitz, L., Mariner Energy, Inc.

King Kong/Yosemite Development, Steel Catenary Riser Design Report, Doc. No. 20258-HC-230, Mentor Subsea Technology Services, Inc., Houston, Texas.

2. American Petroleum Institute, Exploration and Production

Department, Recommended Practice 1111, Design, Construction, Operation and Maintenance of Offshore Hydrocarbon Pipelines (Limit State Design), 3rd Edition, API Publishing Services, Washington D.C., 1999

3. American Petroleum Institute, Exploration and Production

Department, Recommended Practice 2RD, Design of Risers for Floating Production Systems (FPSs) and Tension-Leg Platforms (TLPs), 1st Edition, API Publishing Services, Washington D.C., 1998


-3800

-3250

-3200

-

3150

-3100

-3050

-3150

-3200

-

3250

-

3300

-

3350

Figure A-1: Mariner Energy, Inc. King Kong/Yosemite Field Layout


Figure A-2: SCR Layout Near Platform

GC 254 GC 298

-

3300


Figure A-3 Buoyed Riser Final Layout


Figure A-4: Buoyancy Module Details

Figure A-5: Buoyancy Module Arrangement on SCR


Figure A-6: Buoyancy Module Testing, Simulated S-Lay Stinger

Figure A-7: Buoyancy Module Installation During Pipelay


Figure A-8: SCR With Buoyancy Modules Departing Lay Vessel


Figure A-9: Non-Buoyed, Simple Catenary SCR Arrangement


Strake Bare pipe Buoyant Sag bendCase 1 2 2 1.42 11.587 188 82 85 35Case 2 3 4 2.84 22.066 144 82 103 69Case 3 4 5.5 3.90 22.307 196 118 163 191Case 4 4 9 6.39 0.075 345 166 247 516Case 5 4 4 2.84 1.541 189 102 120 72Case 6 6 6 4.26 14.547 351 197 277 348Case 7 8 7.5 5.32 2.946 569 287 431 602Case 8 10 8.5 6.03 2.846 815 363 550 841Case 9 12 9 6.39 1.295 976 410 602 895Case 10 16 9.5 6.74 0.166 1393 431 596 1026Case 11 19 11 7.81 0.110 1843 494 710 1275Case 12 26 11.5 8.16 0.001 2266 722 1035 1495Case 13 1 1.5 1.06 2.748 135 59 61 28Case 14 16 7.5 5.32 0.023 1154 400 616 1037Case 15 16 12.5 8.87 0.253 1611 416 588 1236Case 16 12 6 4.26 0.647 657 344 491 619Case 17 12 12 8.52 0.262 1061 410 585 1059Case 18 8 4.5 3.19 0.001 353 201 251 207Case 19 8 12 8.52 0.068 772 308 447 870Case 20 4 7.5 5.32 2.482 304 162 238 394Case 21 21 9.5 6.74 0.021 1718 615 883 1373Case 22 22 11.5 8.16 0.009 2159 560 794 1386Case 23 19 12 8.52 0.001 1974 526 798 1361Case 24 10 12 8.52 0.023 940 357 520 943Case 25 10 9.5 6.74 0.384 845 365 529 772Case 26 7 7.5 5.32 10.570 510 255 378 514Case 27 8 8.5 6.03 1.427 656 304 446 637Case 28 14 9.5 6.74 0.604 1164 450 660 979Case 29 12 7.5 5.32 0.990 806 377 564 797

412 9560 2299 561Fatigue Life-year

Maximum RMS Stresses (psi)

Bin(Case) # Hs (ft) Tp (sec) Tz (sec)

Probability of occurrence

(%)

Table A-1: Maximum RMS Stresses and Fatigue Life for Riser Regions, Wave Fatigue


Table A-2: Combined Minimum Fatigue Life for Riser Regions

Element#

ElementLength

(ft)

ArcDistancefrom SCR

top (ft)

FatigueLife

(year)Element

#

ElementLength

(ft)

ArcDistancefrom SCR

top (ft)

FatigueLife

(year)Element

#

ElementLength

(ft)

ArcDistancefrom SCR

top (ft)

FatigueLife

(year)Element

#

ElementLength

(ft)

ArcDistancefrom SCR

top (ft)Fatigue

Life (year)Min. Fatigue Life 208 2016 251 310

24 2.5 24.25 208 64 15 550.5 2016 66 20 583 1326 163 20 2523 603825 2.5 26.75 344 65 15 565.5 3645 67 20 603 265 164 20 2543 163226 2.5 29.25 522 608 251 165 20 2563 340827 2.5 31.75 731 68 20 623 484 166 20 2583 3954228 2.5 34.25 947 69 20 643 386 167 20 2603 630829 2.5 36.75 1136 70 20 663 425 168 20 2623 148130 2.5 39.25 1280 71 20 683 506 169 20 2643 284431 2.5 41.75 1374 72 20 703 352 170 20 2663 3564132 2.5 44.25 1426 73 20 723 665 171 20 2683 554333 2.5 46.75 1442 74 20 743 344 172 20 2703 128034 16.5 56.25 1471 75 20 763 764 173 20 2723 258035 16.5 72.75 1551 76 20 783 372 174 20 2743 3208536 16.5 89.25 2012 77 20 803 627 175 20 2763 408837 16.5 105.75 2348 96 20 1183 1093 176 20 2783 110638 16.5 122.25 2442 97 20 1203 665 177 20 2803 286439 16.5 138.75 2061 98 20 1223 1304 178 20 2823 3280540 16.5 155.25 1625 99 20 1243 712 179 20 2843 275741 16.5 171.75 1872 100 20 1263 1134 180 20 2863 97342 16.5 188.25 2634 101 20 1283 974 181 20 2883 375643 16.5 204.75 2946 102 20 1303 889 182 20 2903 2560144 16.5 221.25 2828 103 20 1323 1335 183 20 2923 152945 16.5 237.75 1976 104 20 1343 796 184 20 2943 94646 16.5 254.25 1523 105 20 1363 1358 185 20 2963 696147 16.5 270.75 1844 106 20 1383 4022 186 20 2983 882248 16.5 287.25 2683 107 20 1403 18802 187 20 3003 91649 16.5 303.75 2850 108 20 1423 9625 188 20 3023 120350 16.5 320.25 2532 109 20 1443 8042 189 20 3043 1201851 16.5 336.75 1612 110 20 1463 16475 190 20 3063 241152 16.5 353.25 1358 111 20 1483 52892 191 20 3083 66653 16.5 369.75 1948 112 20 1503 32596 192 20 3103 222254 16.5 386.25 2844 113 20 1523 9683 193 20 3123 728855 16.5 402.75 2999 114 20 1543 6926 194 20 3143 90856 16.5 419.25 2147 115 20 1563 13309 195 20 3163 78157 16.5 435.75 1316 116 20 1583 49543 196 20 3183 674358 16.5 452.25 1432 197 20 3203 278259 16.5 468.75 2764 198 20 3223 59060 16.5 485.25 4239 199 20 3243 188261 16.5 501.75 4133 200 20 3263 1583362 16.5 518.25 1827 201 20 3283 83963 16.5 534.75 1248 202 20 3303 692

203 20 3323 4067204 20 3343 1299205 20 3363 423206 20 3383 964207 20 3403 1259208 20 3423 436

3429.25 385209 20 3443 681210 20 3463 8619211 20 3483 1610212 20 3503 310213 20 3523 386214 20 3543 1267215 20 3563 113482

Bare pipe Buoyant Sag bendStrake


Stress RangeKsi Axial Total Axial Total Axial Total Axial Total0 0 0 0 0 0 0 0 01 23643040 15210313 21729577 14904202 20432722 9299369 20290872 123890512 65196 543989 15235 531416 681 452318 1236 3644993 2211 108104 1237 49659 0.0 14782 546 966334 416 22242 341 2651 0.0 545 488 279645 340.4 6509.1 340.4 440.7 0.0 46.3 0.3 5852.16 0.0 1496.4 0.0 215.9 0.0 2.2 0.0 380.37 0.0 1232.0 0.0 44.1 0.0 0.2 0.0 0.28 0.0 590.7 0.0 0.0 0.0 5.9 0.0 5.99 0.0 187.0 0.0 0.2 0.0 0.0 0.0 0.010 0.0 5.0 0.0 0.0 0.0 0.0 0.0 0.011 0.0 0.9 0.0 0.3 0.0 0.0 0.0 0.012 0.0 0.0 0.0 0.0 0.0 0.0 0.0 0.013 0.0 2.2 0.0 0.0 0.0 0.0 0.0 0.014 0.0 0.3 0.0 0.0 0.0 0.0 0.0 0.015 0.0 2.2 0.0 0.0 0.0 0.0 0.0 0.016 0.0 0.0 0.0 0.0 0.0 0.0 0.0 0.017 0.0 0.0 0.0 0.0 0.0 0.0 0.0 0.018 0.0 0.0 0.0 5.9 0.0 0.0 0.0 0.019 0.0 0.0 0.0 0.2 0.0 0.0 0.0 0.020 0.0 0.0 0.0 0.0 0.0 0.0 0.0 0.0

More 0.0 0.0 0.0 0.0 0.0 0.0 0.0 0.0

Maximum Total Tensile Stress (ksi) 39.72 23.6 51.05 34.71

Top Middle Bottom-Max. StressRainflow Stress Count (per Year)

Bottom

Table A-3: Fatigue Stress Spectrum for Fatigue Testing

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