bluuebunny/crossref_metadata_2025_split huggingface dataset:
Transportation, Launch, and Self-Upend Analysis of the Kilauea Jacket Using Proven Analytical Techniques
Authors: Sircar S. , Chandra T.K. , Mills T.R.J. , Roberson W.P. 
Published: October 2008

ABSTRACT
The self-upending, light-weight Kilauea jacket was designed with a minimum amount of flooding and venting equipment. The installation procedure was sensitive to changes in the jacket's weight and buoyancy properties. As a result, this jacket was subjected to extensive installation engineering and careful weight monitoring. This paper describes the methodology and results of the transportation, launch, self-upend, flotation, set-down and on bottom stability analyses of the Kilauea jacket. All hydrodynamic and friction coefficients and weight margins used are based on experience gained from past projects. Controlled ballasting was used to set the jacket down on the seabed and to obtain initial leveling and on-bottom stability.

INSTALLATION PHILOSOPHY

Installation Overview
In the final design stage, it was envisioned that the Kilauea jacket would be installed based on the following general procedures:Load out the jacket from McDermott's fabrication yard on to the launch barge Intermac 650 (I-650) with the barge freely floating at an approximate draft of 30 feet. At this draft, the barge launch ways are level with the fabrication skid ways.After load out, reduce the barge draft to the design Gulf tow draft (24 feet) and install transverse and longitudinal sea fastenings and prepare the jacket and the barge for tow.Reduce the barge draft to 14 feet for river tow and tow the jacket using three river tugs to the Eugene Island sea buoy.Based on good weather forecast, replace the three river tugs with three ocean going tugs and start the Gulf tow while continuing to ballast the barge down to arrive at the Gulf tow draft.Prior to the arrival at the launch site, and on confirmation of good weather forecast for launch, barge ballasting is continued to arrive at the pre-launch configuration.Prepare to launch the jacket against the current and facing the derrick barge. Two tugs and two head lines from the derrick barge are connected to the top four corners of the jacket (see Figure 1).Once all sea fastenings have been cut, and the I-650 is in the pre-launch draft and trim condition, the jacket is jacked until it starts sliding by itself and launch begins. After launch, the jacket self-upends itself and is then pulled in towards the derrick barge using the two headlines.Once the jacket is at the stern of the derrick barge, the derrick barge crane is attached and lifts the jacket about 10 feet to gain more positive control of the jacket. This is followed by lowering of the jacket to the seabed using a combination of hooklC1ad and controlled ballasting of the jacket's compartments. Further ballasting is then used to level the jacket to within acceptable limits and to achieve the desired on-bottom weight.Piles are then launched and self-upended from the side of the derrick barge and stabbed into the pile sleeves using the assistance of Remotely Controlled Vehicles (ROVs). This is followed by pile driving and grouting.
A Comparison of the Erosion Rate of Casing Steels by Sand/Oil Suspensions
Authors: Clark H.M. 
Published: October 2008

ABSTRACT
The erosion rates (expressed as g/mm2/min) of three casing steels (P110, N80 and K55) and hot-rolled 1020 steel in quartz sandoil suspensions, GFN 140, solids loading 0.5 - 4 wt%, at speeds from 9.35 - 18.7 m/s have been measured in a slurry pot tester. The velocity exponent was determined as 2.4 and the relationship of erosion rate to solids loading and particle size (for the particle size range 53-180, µm) determined. The erosion rates for all steels were foundto lie within a scatter band of ±m 15%, and could not be related predictively with the measured tensile properties. Results are discussed in terms of the rate of particle impact on the eroding target and the need to control flow conditions through appropriate design to minimize erosion.

INTRODUCTION
This work grew out of the observation of the erosion of oil-well components by fine sand entrained in the petroleum stream in wells off the SE coast of Trinidad (Caribbean). The effect of erosion is progressive, more sand entering the system as damage proceeds, thus engineering parts spend much of their life under conditions of steadystate erosion, component life being determined by the rate of material loss per unit time.
This investigation has focused on the rate of steady-state erosion of three casing steels under flow conditions similar to those found in oilwells and platforms. The flow rate through down-well screens is in the order of 1.5-3 m/s, while across chokes it may be at least an order of magnitude higher. The average sand concentration in a producing well in this field is about 0.5 vol %(1.3 wt%) and may rise to as much as 5 vol.% (13 wt.%).

EXPERIMENTAL

Sand
Sand used in this work was obtained from a sand separator on an off-shore Trinidad oil platform. X-ray analysis showed it to consist almost entirely of ? X-quartz, with traces of feldspar, calcite and clay minerals. The particle size distribution is shown in Fig. 1. The AFS grain fineness number was 140 (106 µm) and the density of particles was 2620 kg/m3. For most tests, the sand was used with the as-received particle size distribution, except as noted. The shape of fresh sand particles is shown in Fig. 2a. One test was conducted with silicone-coated glass beads in the size range 106-150 µm (density 2420 kg/m3). These are shown in Fig. 2c.

Oil
Suspensions were made up using light furnace oil, flash point 67°C, density 850 kg/m3 and viscosity 1.5. 10−3 Ns/m2 at 40°C.

Steels
Three grades of oilwell casing steel were examined, API P110, N80 and K55, having yield stresses in excess of 760, 550 and 380 MPa respectively. The P110 and N80 contained 0.3 °C and 1.9 Mn and were in the quenched and tempered condition. The K55 contained 0.47° C and 1.4 Mn and was hot-rolled. Coupons were cut parallel with the axis of the casing from which tension and erosion test specimens were turned. Tests were also performed on hot-rolled 1020 steel to provide a performance datum.
Diverless Flowline Tie-In Techniques for the DISPS Project
Authors: Bruton A.S. , McKeehan D.S. , Metcalf P. 
Published: October 2008

ABSTRACT
This paper describes the DISPS Project design, development and prototype testing, in the dry, of three diverless tie-in systems for application to flowlines of 4 - 12 inches diameter in water depths of 350 - 750 metres. Full scale simulation of direct pull-in, lateral deflection and vertical stab techniques using two tools for flowline pull-in and connection was undertaken and the design limits established.

INTRODUCTION
Within the DISPS development, diverless flowline tie-in concepts were recognised, as key flowline issues. A study was undertaken and this included the evaluation of possible flowline installation methods, pipeline materials and available connection hardware which would be suitable for diameters of 4 to 12 inches in water depths of 350 to 750 metres.
The main conclusion of this study was that diverless flowline tie-ins play a considerably greater role in dictating field development and installation strategies than flowline tie-ins in conventional diver depths. These diverless tie-in techniques warranted a design philosophy which favoured a highly adaptable system over a case-specific system.
The following engineering targets were established for the DISPS development of diverless flowline tie-ins.The flowline tie-in bays within the template or satellite well structure should be capable of accommodating any of four types of tie-in:Direct Pull-in for Steel Pipe (Figure 1)Lateral Deflection for Steel Pipe (Figure 2)Vertica1 Stab and Hinge-over for Flexible Pipe (Figure 3)Slack Loop Pull-in for Flexible Pipe.The tie-in system should be compatible with flowline installation by laying, reeling or towing.The alignment mechanism should be by structural interaction between a passive guide and pull head without manipulation of the flowline.All pull-ins should be achievable with a single pull wire acting through the pipe centreline.Recoverable tools should be used for pull-in and make-up of the mechanical connector. Both alignment modules should be compatible with these tools.The pull-in and make-up procedure should be performed in discrete steps and allow support vessel demobilisation after docking. This enables a smaller diving support vessel to be used for subsequent connector make-up.The alignment components should be modularised such that finalisation of the template or satellite well configuration can be deferred prior to deployment. This allows field development plans to remain flexible while the seabed structures arebeing fabricated without compromising template/flowline interface areas.The pullhead should allow pigging and hydrostatic testing of the flowline before the connection is made to inboard piping on the subsea structure. This allows the pipe-lay contractor to satisfy contractual obligations even if the final connection is deferred. In keeping with this objective, the flowline remains fixed after docking, with final stroking of the connector by movement of the inboard piping.
India's Programme on Polymetallic Nodules
Authors: Yates J. , Gupta K.R. 
Published: October 2008

Abstract
This article is a review of India's programme aimed at developing deep sea manganese nodules on a commercial basis. A brief history of the programme is followed by an outline of current work in the areas of survey and exploration, deep-sea mining technologies, extractive metallurgy and environmental studies. There then follows an analysis of the resources available to the project both fiscal and in terms of available manpower. The programme is viewed in the wider context of Indian science and technology policy in general and the ocean sciences in particular, emphasising major factors which have led to its implementation. The article concludes by detailing some current trends within the programme and likely future options for its development.

Introduction
Since Independence in 1947, successive indian governments have placed great emphasis on the development of indigenous science and technology, viewing this as a means of narrowing the technological gap between India and the more industrialised western nations, and thereby improving the average standard of living within the country. This strategy has led to the development of a range of ambitious programme; space, nuclear, electronics, Antarctic, and the subject of this article deep-sea mining. Although India has had a long sea-faring tradition, it is only after Independence that priority has been given to the R&D programme related to the tapping of living and non-living marine resources. Oceanographic research and development has been given extra impetus during the last 25 years with the establishment of the Institute of Oceanography, the marine work of the Oil and Natural Gas Commission, and the establishment of the Marine Geology Division in the Geological Survey of India. Studies by naval research groups and participation in major international programmes like IIOE, GRAP and MONEX have also played an important role 1 As a result of this activity, major advances have been made in indian oceanography in recent years.
In order to plan, coordinate and develop an indigenous capability in Oceanographic research, the Government of India set up an Ocean Science and Technology Agency (OSTA) in the Department of Science and Technology in 1976. A separate Marine Research and Development Fund was set up under OSTA to support R&D in this area. Subsequently, a Department of Ocean Development (DOD) was created in July 1981, charged with matters regarding coordination, security, regulation and development of the ocean including 2,3promotion of R&D in the marine sector.surveys to locate and map the availability of living and non-living resources of the Exclusive Economic Zone and of the continental shelf (excluding hydrocarbons).development of skills and manpower.protection of the marine environment in the high sea areas.survey, exploration and development of deep sea bed mining.promotion and coordination of Antarctic Research. The basic policy underlying this ocean development has been defined in the Ocean Policy statement adopted by the Parliament in November 1982.
Veslefrikk Field Development Concept
Authors: Solberg L. 
Published: October 2008

VESLEFRIKK FIELD DEVELOPMENT CONCEPT

ABSTRACT
This paper gives a brief description of the Veslefrikk Field development concept, from the start of the conceptual phase to the construction and start-up phases. The different main contracts are described as well as some challenges and experiences gained duringthe execution of the Project.

FIELD HISTORY AND RESERVOIR
The decision to develop the field was taken in 1986, just as the oil price had suffered a dramatic drop. As a consequence of this, the field had to be developed to meet a low oil price.
Veslefrikk is located in block 30/3 situated 145 km west of Bergen in WestNorway, and close to the Oseberg Field.
Statoil was appointed as operator in 1979, and the first oil was found in 1981. The field was declared commercial by Statoil in the autumn of 1986.
The owners of the field are:(Available In Full Paper)
The reservoir is estimated to contain 36.4 million m3 oil and 4 billion m3 gas.
The detailed engineering started in October 1987; the production start-up was 26 December 1989, that is to say 27 months later.
The daily production is estimated to be 10.000 m3 oil (65000 bbl/day) and 1.5 million m3 gas.
The field will reach peak production in 1990. After 3 to 5 years the output will decrease and the production will come to an end after 20 years. Thewater depth at the field is 174m.
The reservoirs are located at the depths of 2800 m and 3050 m. The total number of wells is 18 of which 11 are producers and 7 are injectors.

PROJECT PLANNING
Studies for development of the Ves1efrikk Field were started in the autumn of 1985.
Because of a water depth of 174m a whole range of field installations were studied. During the spring of 1986 theoil prices fell drastically presenting a new challenge in the development i.e. tomeet an oil price down to 12 USD/bb1.
The following 3 main concepts were evaluated in detail:Integrated process, drilling and quarters platform like Statfjord and Gullfaks, but downscaled.Floating production platform connected to a fixed drilling platform of steel or concrete.Subsea wells linked to a floating production platform. During technical and financial evaluations the integrated process, drilling and quarters platform turned out to be more expensive for this field compared to the selected alternative.
The floating production platform was evaluated with two options, one as a purpose-built platform, and the selected alternative with conversion of an excisting rig. The selected alternative showed a cost saving and also a schedule saving of one year.
As a result of these evaluations the optimal solution appeared to be drilling facilities supported by a steel structure in combination with a floating production and quarters unit see figure.
Design and Installation of the Veslefrikk Jacket
Authors: Baerheim M. , Fossan T.I. , Eriksen K. 
Published: October 2008

ABSTRACT
The Veslefrikk A jacket, standing in 174 meter water depth in the North Sea, was lift installed in May 1989. The jacket is the heaviest lift installed offshore structure in the world.
The jacket is relatively tall and slender, and the lift weight was close to the capacity of the available heavy lift crane vessels. Structural design and weight optimisation were thereforeimportant.
This paper gives a general description of the jacket and highlights some of the special features and lessons learnt in the design and installation of the structure.

INTRODUCTION
The Veslefrikk field is developed with a two platform concept. The drilling equipment is placed on the Veslefrikk A jacket, while processing equipment and quarters are located on the semisubmersible platform Veslefrikk.
The field development is further described in reference 1.
Being 184 metres long and with a lift weight of 9086 tonnes, the jacket was lifted and installed by the world's largest dual crane semi-submersible lift vessel. The lift took place inshore at the mouth of Sognefjorden. The jacket was then towed in a partly submerged position to the field for the upending, mating and piling operations.
Various special design methods and features were introduced in the jacket design. Most of these were introduced since the weight of the structure during lift was critical, others because thejacket is slender and dynamically sensitive.

JACKET DESCRIPTION
The Veslefrikk jacket is a four legged structure, as shown in fig 2. Both the faces and the horizontal levels are X-braced. The jacket extends from the sea bed up to elevation +9.8 metre.
The top and bottom dimensions of the jacket are 18.3 by 18.3 and 54.9 by 54.9 metres respectively. These dimensions were largely determined by lift clearance requirements. A separately installed deck support frame extends up to elevation +25.5 metres.
The legs have an outer diameter of 1.3 metres between elevation +25 and -60 metres. The legs in this area are reinforced by 1.1 metre diameter internal tubes grouted to the outer legs. Below -60 the leg diameter increases from 2.2 to 3.5 metres.
The face bracing varies from 0.8 to 1.6 metres diameter at the top and bottom of the structure respectively. The plan level bracings are 1.1 metre in diameter. The jacket is supported by four piles at each leg.
The jacket was installed over a template having six pre-drilled wells.
The following criteria were used in the design :(Available In Full Paper)

SPECIAL CHALLENGES IN DESIGN
The decision to design a liftable jacket was taken after feasibility studies of both lifted and launched structures. The lift operation determined the geometry of the jacket and put emphasis on weight optimisation.
The jacket is slender and light and founded on relatively weak soils. This also put special emphasis on design for temporary and permanent foundation strength and dynamic behaviour of the jacket.
A Unified Design Procedure for Tubular Members
Authors: Loh J.T. 
Published: October 2008

ABSTRACT
A new unified procedure for designing tubular members under combined axial load (compression or tension), bending, and hydrostatic pressure on offshore structures, fixed or compliant, is presented. The equations are applicable to members at all water depths, but are limited to members with outside diameter to wall thickness ratios less than 80. In addition to a pure hydrostatic check against hoop collapse, the procedure recommends only one check each for the axial compression and tension cases. The new design equations have been calibrated to test data from several recently completed test programs, and are thus more reliable than currently available technology. The new equations are also simpler to use. The procedure can replace several subsections in the 18th edition of API RP 2A-ASD and the draft RP 2A-LRFD (yellow copy).
The recommended procedure provides unambiguous guidance for determining the member axial, bending, and hoop stresses from a frame analysis in order to check the design adequacy or determine the unity check (member utilization) of the member. The recommended procedure includes a reduced safety factor for hoop compression, but other safety factors in RP 2A-ASD are unchanged. Further examination of the reliability question may be warranted, but it is not envisioned that such an evaluation would alter the proposed procedure so as to become more conservative.
A comparative design study using the new equations on a 2500-foot compliant pile tower resulted in about 5% tower steel-weight saving (2300 tons) compared to the approach proposed by Chen, Bubonic, and Sohal. Even larger weight savings can be expected for TLP tendons although a comparative design example is not yet available.
The new equations are currently being reviewed as part of an ongoing API PRAC 89-56 study on analysis of the cylindrical shell database and validation of design formulations.

INTRODUCTION
Both API RP 2A-ASD1 and RP 2A-LRFD2 provide incomplete guidance for the design of tubular members under combined axial compression, bending, and external hydrostatic pressure loads, as shown in Fig. 1. The guidance in both documents is limited to checking local cross-sectional buckling. (Shear forces are not shown in Fig. 1. These forces have negligible impact on the strength of beam-columns.) In addition, there is an inconsistency in both documents relating to the allowable bending capacity term. However, this inconsistency generally results in a conservative design.
In addition to hoop stress, external hydrostatic pressure will impose a capped-end axial compression force in a member if the ends are capped, as shown in Fig. 2. On the other hand, external radial pressure only imposes hoop compression in a tubular member and no capped-end compression. Since the hoop and capped-end axial stresses from external pressure are always in compression, they are assumed to have a positive sign.
In-Situ Pile Segment Model Experiments at Harvey, Louisiana
Authors: Bogard J.D. , Matlock H. 
Published: October 2008

Abstract
Three field experiments have been conducted in normally consolidated soft clay at Harvey, La. using instrumented 3inch diameter probes to simulate pile behavior under a variety of static and cyclic axial loadings. Lateral total pressure and pore pressure were observed and shear transfer versus pile displacement were recorded. The results represent the first phase of a comprehensive program to develop an improved basis for pile design.

Introduction
As a part of a comprehensive program of research into the behavior of friction piles in clay, in situ pile segment models, or probes, have been developed to measure total radial pressure, pore pressure, shear transfer, and relative pile-soil displacement during installation, consolidation, and static and cyclic axial loading.
The experiments reported herein were the first of 47 experiments performed with the probes, which have included tests at the onshore sites of two large-diameter instrumented pile load tests, the offshore site of load tests on instrumented large-diameter driven and drilled-and-grouted piles, and three onshore sites where parallel large-diameter pile load tests are not available.
Four of the six sites contain soft-to-stiff normally consolidated clays. At the four sites, a total of 42 experiments were performed, with load testing at times ranging from minutes to three months after installation. The data are believed to represent a sound basis for developing a fundamental understanding of pile-soil interaction and for developing reliable design procedures for the design of pile foundations subjected to static and cyclic axial loading.

Methods
The pile segment models were developed at The Earth Technology Corporation (Refs 1 and 2). Two sizes, 3.00 and 1.72 inches in diameter, have been built and utilized in field tests at various locations. The 3-inch size permits investigation of various cutting shoes or end conditions and is the type employed for the tests reported in this paper.
The features of the 3-inch probe and the three stages of installation are shown in Fig 1. The tool is impact-driven or jacked into the soil beyond the end of a drilled hole. The leading section is a tube which is long enough to simulate the effects of an open-end pile in the zone where the instrumented section comes to rest. Various wall thicknesses or a fully plugged condition can be utilized. At the slip joint shown at the left, the leading section is free to move one inch with respect to the instrumented section and the relative movement is indicated with an internal displacement transducer. After installation, the leading section serves as an anchor for accurate measurement of pile-sail displacement. Soil shear transfer is measured as the differential reaction between the two load cells. The pressure measurements aid in interpreting soil consolidation and set-up processes, plus pore water and structural changes during shear.