bluuebunny/crossref_metadata_2025_split huggingface dataset:
An Assessment Of Grouting Materials, Placement Methods, And Monitoring Equipment For Offshore Structures
Authors: Callis C. , Knox C. , Sutton D. , Wiley S. 
Published: October 2008

ABSTRACT
Grout placement techniques, technology of grouting materials, and methods for monitoring grout quality and grout placement are presented.
Placement methods using new generation high pressure packers and inflation-grout manifolds are discussed. High pressure packers have made single stage grouting jobs a practical reality. Reliable inflation-grout manifolds have eliminated many lines previously needed for packer inflation and placement of grouting materials.
In addition to traditional grouting materials, new materials to promote high bond strength, control self-desiccation, and minimize shrinkage are discussed. Grouts are classified according to use and typical examples are given. Methods and equipment for controlling grout density and grout placement are included.

INTRODUCTION
Platform grouting is the process of placing a cementitious material into the annulus between a jacket leg and pile. This operation has advanced to the point now that an operator has a variety of methods, materials, and tools from which to choose. For example, at least five different methods for grouting platform legs are commonly used with variations existing within each method. Terms for each of these methods, although not standard, might be (1) conventional two stage, (2) packer, (3) balanced pressure, (4) two stage delayed set for insert piles, and (5) two stage inner string for insert piles.
Recent advances in methods, materials, and tools have helped make platform grouting more trouble-free and dependable. This paper will review these advances and provide an overview of current platform grouting in general.

GROUTING METHODS

Conventional Two Stage
As shown in Fig. 1, this method usually involves a mud wiper seal at the bottom of the annulus. This serves to help reduce mud contamination of the annulus as the pile is driven through the jacket leg and it also helps support the first stage of the grout. This technique normally involves running two separate grout lines from the surface. The first stage line enters the annulus just above the mud wiper seal. Second stage line enters the annulus 5 to 30 ft above the first stage grout line depending on the length of plug required. First stage grout is normally a quickset type and fills up to the second stage grout line entry level. While the first stage is setting, water should be circulated through the second stage grout line to displace any grout that may have covered the second stage port. Once the first stage grout is set, the second stage is injected to the desired height in the annulus.
An improvement to this procedure, shown in Fig. 2, makes it possible to accomplish the grouting through only one grout line from the surface. One line is run from the surface to the lower grout inlet. A pipe is run from the upper grout inlet to a sleeve-type flow control valve installed in the grout line at the same elevation as the upper grout port.
OTEC Cold Water Pipe Program Status
Authors: McGuinness Terence , Scotti Richard S. 
Published: October 2008

ABSTRACT
In this paper, the results and current work efforts of the Ocean Thermal Energy Conversion (OTEC) Cold Water Pipe (CNP) program will be presented. The program, sponsored by the Department of Energy and developed and technically managed by the National Oceanic and Atmospheric Administration, is one of the largest civil ocean engineering activities underway at the present time. The uniqueness of the OTEC GNP is derived from its physical magnitude, namely, pipes that are being designed with a diameter of 9 meters" and a length of 1000 meters. Details of the three phases of the program: the analytical procedures used to predict pipe responses, the model test projects and the large scale at-sea tests, will be discussed. The procedures used in planning the program will be highlighted. Problems and their solutions will be specified. Projects currently underway or just completed will be synopsized.

INTRODUCTION
The cold water pipes (GNP) required in Ocean Thermal Energy Conversion (OTEC) Plants present a formidable challenge to ocean engineers and marine constructors. The GNP size alone is impressive: 1000 meters in length, 9 meters to 30 meters in diameter depending on the OTEC plant size. The magnitude of the pipe forces and the pipe structural responses, plus pipe development and deployment procedures, pose additional problems for the designer. The uniqueness of the GNP as a marine structure, however, lies not so much in its unprecedented mass as in its need to be a flexible structure which interacts dynamically with the hostile ocean environment. This factor amplifies the concern for structural fatigue to the point where only abnormally low stress levels can be tolerated and factors of safety based on stresses alone become meaningless. There is only one GNP in an OTEC plant, which must be designed to survive all environmental conditions for the duration of the plant, namely, 30 years, and not be damaged nor cause damage to the other parts of the OTEC plant.
A federal program has been formulated over the past four years, directed toward the solution of the major GNP problems. The end products of the program will be designs, data and procedures that can be utilized in the final design and development of several viable alternative GNP concepts.
In this paper the current status of the CWP program will be presented. The program formulation will be discussed and the major elements of the program will be explained.

PROGRAM OUTLINE
The original concept of an OTEC GNP was that it was a simple vertical tube appendage attached rigidly to an OTEC platform. By making the pipe large in diameter, the cold water flow could be limited to 2 meters/sec to reduce pipe head losses to a minimum. To ensure pipe survivability, pipe concepts were made rigid from such materials as concrete and steel.
On The Accuracy Of Different Finite Element Types For The Analysis Of Complex, Welded Tubular Joints
Authors: Hoffman Robert E. , Sharifi Parviz E. 
Published: October 2008

ABSTRACT
There has recently been great interest in the use of three-dimensional finite elements to study local stresses in the weld areas of complex (reinforced, overlapped) tubular joints. This paper examines displacement and stress solutions obtained for selected tubular joint configurations modeled with two such thick shell and brick elements. These solutions are compared, with the aid of an a-posteriori convergence (error) criteria, to those from a well known arid widely used thin shell element. It is concluded that the three-dimensional elements examined can behave very well in typical tubular joint problems. They, however, appear to suffer more instability in irregular mesh situations than do thin shell. The proposed error criteria seems to indicate well those areas where the thin shell approximation is most suspect.

INTRODUCTION
For many years thin shell theory has served as a basis for studying stress concentration factors of tubular joints. With various thin shell finite element types, detailed parametric studies have been performed in the hope of obtaining information consistent enough to establish design curves. It is not the purpose here to review this great volume of past work, but instead to compare in the tubular joint context the behavior of one well established thin shell element with two versions of publicly available three-dimensional elements.
The approach selected is to first discuss the three-dimensional element formulations outlining their salient features. The results of a small test problem are then examined and observations made. An overview of an element residua}C basis for comparing solutions of predominately thin shell problems is presented in the section ERROR ANALYSIS. A short discussion on its potential for application to of three problems solved with both thick (compatible and incompatible) and thin shell elements are compared from the displacement, stress and residual error standpoints.

ELEMENT DESCRIPTIONS
In this study, two types of three-dimensional isoparametric solid elements are used. One is an eight noded "brick" element. The other is a sixteen noded "thick-shell" element. As described in the following section, various "incompatible" modes are introduced into the stiffness formulations of these elements to improve their flexural behavior. Both the basic (compatible) and the incompatible versons of these elements are utilized in this study.

Element Formulations
The 8 and 16 noded isoparametric elements under consideration are shown in Figure 1. The global Cartesian coordinates are denoted as (X1, X2, X3) and the element local coordinates as (available in full paper), both right handed systems. The local coordinates range from -1 to +1 with (0,0,0) located at the centroid of the element. In the isoparametric family of elements, the local and global coordinates are related as follows (see for example Reference [1]):
(Available in full paper)
Fatigue Strength Of Tubular T- And X-Joints
Authors: Dijkstra O.D. , de Back J. 
Published: October 2008

ABSTRACT
The results of fatigue tests on 40 tubular T and X-joints are presented. The chord diameters of the joints are 168, 457 and 914 mm. The most important conclusions are that the lifetime decreases with increasing joint size and that in the high cycle rance (N ? 107) some test results of the large joints fall below the AWS-X curve.

INTRODUCTION
The weather conditions which induce a fatigue crack are more severe in the North Sea than in the Gu1f of Mexico, Because of these severe conditions, a research project, sponsored by the European Coal and Steel Community (ECSC) on the fatigue behaviour of steel in offshore structures, entitled "Fatigue and corrosion fatigue behaviour of offshore steel structures", has been carried out. The research has been carried out in the U.K., France, Germany, Italy and the Netherlands. There is a special cooperation between Norway and the participating ECSC countries. In the programme, flat plate specimens and tubular joints have been tested. In addition, theoretical work (e.g. finite element analysis on tubular joints) has also been carried out. An interim report of the program was presented at a siminar in November 1978 in Cambridge.
In this paper, the results of tests on tubular joints, carried out in the Netherlands under the sponsorship of the ECSC, are discussed. Also, the results of additional work (strain measurements, crack measurements and calculations) sponsored by MaTS (Netherlands Marine Technological Research) are discussed.

TEST PROGRAMME
Forty tests on tubular joints were carried out by the Stevin Laboratory of Delft University of Technology and the Institute TNO* for Building Materials and Building Structures (IBBC-TNO). T- and X-joints were tested with chord diameters of 168, 457 and 914 mm. Table I gives a survey of the test programme.
Three small T-joints were loaded with in plan8 bending. All other joints were axially loaded. One X- joint was tested with both chord and brace axially loaded. Some joints have one (Tl and Xl) or two (T2) additional unloaded braces. Two joints are tested with a random load and all others with a constant amplitude load (R = Fmin/Fmax = 0 or -1). Four joints have been tested in artificial seawater, one of them with cathodic protection.

TEST SPECIMENS AND EQUIPMENT
The test specimens were made by a manufacturer with experience in offshore structures. The specimens were welded in a fixed position with both chord and brace horizontal. Fig. 1 shows the weld shape of a large joint. The steel quality of the joints is in accordance with the standards mentioned in table II. Fig. 2 shows the test rig for the large joints. To avoid secondary effects, all supports in the test rigs are hinges.

STRAIN MEASUREMENTS

General
Before starting the fatigue tests, the strain distribution and the maximum (hot spot) strains in the joint were measured under a static load smaller than the fatigue load. The results are related to the nominal strain in the loaded brace (SNCF = Elocal/Enominal).
Fatigue Tests On Tubular Kand Kt Joints Under Axial Brace Loading
Authors: Wylde John G. 
Published: October 2008

ABSTRACT
The results of fatigue tests on 40 tubular T and X-joints are presented. The chord diameters of the joints are 168, 457 and 914 mm. The most important conclusions are that the lifetime decreases with increasing joint size and that in the high cycle rance (N ? 107) some test results of the large joints fall below the AWS-X curve.

INTRODUCTION
The weather conditions which induce a fatigue crack are more severe in the North Sea than in the Gu1f of Mexico, Because of these severe conditions, a research project, sponsored by the European Coal and Steel Community (ECSC) on the fatigue behaviour of steel in offshore structures, entitled "Fatigue and corrosion fatigue behaviour of offshore steel structures",has been carried out. The research has been carried out in the U.K., France, Germany, Italy and the Netherlands. There is a special cooperation between Norway and the participating ECSC countries. In the programme, flat plate specimens and tubular joints have been tested. In addition, theoretical work (e.g. finite element analysis on tubular joints) has also been carried out. An interim report of the program was presented at a siminar in November 1978 in Cambridge.
In this paper, the results of tests on tubular joints, carried out in the Netherlands under the sponsorship of the ECSC, are discussed. Also, the results of additional work (strain measurements, crack measurements and calculations) sponsored by MaTS (Netherlands Marine Technological Research) are discussed.

TEST PROGRAMME
Forty tests on tubular joints were carried out by the Stevin Laboratory of Delft University of Technology and the Institute TNO* for Building Materials and Building Structures (IBBC-TNO). T- and X-joints were tested with chord diameters of 168, 457 and 914 mm. Table I gives a survey of the test programme.
Three small T-joints were loaded with in plan8 bending. All other joints were axially loaded. One X- joint was tested with both chord and brace axially loaded. Some joints have one (Tl and Xl) or two (T2) additional unloaded braces. Two joints are tested with a random load and all others with a constant amplitude load (R = Fmin/Fmax = 0 or -1). Four joints have been tested in artificial seawater, one of them with cathodic protection.

TEST SPECIMENS AND EQUIPMENT
The test specimens were made by a manufacturer with experience in offshore structures. The specimens were welded in a fixed position with both chord and brace horizontal. Fig. 1 shows the weld shape of a large joint. The steel quality of the joints is in accordance with the standards mentioned in table II. Fig. 2 shows the test rig for the large joints. To avoid secondary effects, all supports in the test rigs are hinges.

STRAIN MEASUREMENTS

General
Before starting the fatigue tests, the strain distribution and the maximum (hot spot) strains in the joint were measured under a static load smaller than the fatigue load. The results are related to the nominal strain in the loaded brace (SNCF = Elocal/Enominal).
Experimental And Analytical Studies Of The Elastoplastic Behavior Of Offshore Pipelines During Laying
Authors: Kimura Tetsuo , Idogaki Satoshi , Takada Katsumi , Fujita Yasushi 
Published: October 2008

ABSTRACT
With more and more offshore pipelines being laid in deeper water, the need for larger laying equipment has increased larger barges, stronger tensioners and longer stingers. The primary reason is that, in deeper water, the pipeline is suspended at a steeper liftoff angle, and consequently a longer over bend portion is required between the liftoff point and the welding stations In order to help relieve stress in the pipeline.
If the over bend portion can be curved until the plastic region of the pipe is reached, it will be possible to use shorter stingers and lay pipe in deeper water without any special equipment investment.
We took note of these possibilities and conducted integrated experiments consisting of material tests, pipe bending tests and pipe well denting tests. We also developed a continuous beam elastoplastic analysis program, a pipe section analysis program and a cylindrical shell analysis program. Using both test data and programs, we established a precise elastoplastic design method for laying offshore pipelines.
Measured results show good agreement with calculated results.
This elastoplastic design method accurately predicts the capacity of pipe to bear environmental loads, thereby making it possible to set optimum limitations for laying operations.
However, it should be noted that, just above the yield point of the pipe, slight variation environmental loads cause significant changes in the strain and deflection. It would therefore be desirable in actual laying operations to use a stinger that permits control and restraint of pipe curvature and deflection.

INTRODUCTION
The elastoplastic design method for laying offshore pipelines makes it possible to use shorter stingers, which results in reduced liquid forces acting on the stinger pipeline system. This in turn results in improved operating rates and increased water depths in which pipe can be laid, as well as reductions in cost. Moreover, optimum limits set on laying operations by this design method ensure operational safety.
Three problems arise when large strains are induced in the pipe during laying. It is essential for the elastoplastic design method to solve these problems. The three problems are:NonlinearityThe bending stiffness of pipe bent beyond Its proportional limit becomes nonlinear due to the nonlinearity of the steel's stress-strain characteristics and the ovaling of the pipe section. Therefore, a nonlinear analysis must be made of the portion of the pipeline between the lay barge and the ocean floor.Local bucklingIn our experiments, all pipe specimens with a D/t ratio of 56 or over failed due to local buckling. In the case of specimens with a D/t of 56, local buckling occurred after the yield point was reached, whereas specimens with D/t ratios of 78 and 97 buckled locally before their yield points were reached. Thus, the elastoplastic design must give full consideration to local buckling.
Electrical Shock Hazard Protection System
Authors: Miller Harvey A. , Yemington C.R. 
Published: October 2008

ABSTRACT
A new electrical shock hazard protection system actually predicts impending shock hazard and provides advance warning or shutdown of electrical equipment to protect deck handlers and divers working near electrical power cables. The increasing use of electrical diver support equipment, tethered vehicles, and high power remote subsea equipment has increased the need for a failsafe shock hazard protection system. At the same time, the higher voltages, higher power, and longer cable lengths now being used make it more difficult to get good protection without a high incidence of false alarms. With the new approach, it is no longer necessary to tolerate more false alarms to get better sensitivity. This new system is immune to false alarms from changes in submerged cable length, temperature, and electrical load Characteristics including balance and transients. It also gives an alarm for almost any type of malfunction within the protective system itself. On its first application in the ocean, this system operated hundreds of hours without false alarms. It then gave an alarm when it detected damage to the jacket of a cable between canisters on the remote vehicle. It subsequently shut off power when it detected further breakdown of the insulation around the conductors inside the jacket at the point of damage. Even though this ground fault occurred in an interconnect cable on a remote vehicle, the detection system at the 'surface sensed the condition immediately. Future applications for this device include use on any electrical equipment in the marine environment. Applications to high-voltage and highpower remote vehicles are emphasized in this paper.

INTRODUCTION
Picture yourself on a steel deck with seawater washing around your ankles while you try to manhandle an umbilical cable carrying 2300V electrical power. One of your foremost concerns is bound to be the integrity of the insulation in the cable. Designers of electrical equipment for use in this environment should be equally concerned, especially since techniques are now available to continuously test the insulation and predict shock hazards, while the equipment is running, without impacting normal operation.
The use of undersea electrical instrumentation and power equipment is rapidly growing in response to economic and safety demands of the expanding offshore industry. This is true for both manned and remotely controlled systems including underwater welding equipment, sophisticated manipulator bells, submersibles which are tethered to allow greatly increased mission time, and high-powered remotely-controlled vehicles such as trenchers. In the last few years, the power level has gone from 1kW to 2,000 kW; cable lengths from a few hundred feet to several thousand feet, and the transmission voltages are increasing to accommodate the higher power levels--from 24V to 120V to 480V to 2400V to 6600V. Each such extension increases the hazard to human life.
Undersea power cables, particularly working cables used in the equipments outlined above, are indeed vulnerable.