Feasibility of IGW technology in offloading hoses. MSc Thesis Report Civil Engineering

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1 Feasibility of IGW technology in offloading hoses MSc Thesis Report Civil Engineering Siebe Nooij June 2006

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3 The optimist proclaims that we live in the best of all possible worlds; and the pessimist fears this is true James Branch Cabell iii

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5 PREFACE This is the final report of my thesis research on the use of a new composite technology in offloading hoses. During my research I have gained basic knowledge in modelling losses in pipe flow by experiment and evaluating a new technology for implementation in an existing product. By submitting this thesis report, I have finished my final course to be granted a MSc degree in Civil Engineering. Although the thesis is an individual exercise, I have relied on the assistance of coaches and cooperation with others. First of all I would like to thank the representatives of the section Fluid Mechanics at the Delft University of Technology, prof. dr. ir. G.S. Stelling and drs. R. Booij, for their guidance and for making it possible to do the experiment. Furthermore I would like to thank prof. ir. J. Meek for his advice and for giving the project a sharp focus in the initiation phase. I have performed this thesis research at Gusto Engineering. I am grateful for their help and for giving me the opportunity to visit SBM in Monaco to get more feeling with the offloading hose by interviewing people that use these hoses frequently. Special thanks go out to dr. ir. J.H. Westhuis, my coach at Gusto, who always thought along and came up with ideas which had a large contribution to the final result. Furthermore I would like to thank Thana Mousawi for her kind cooperation in the development process of the new offloading hose design. Finally I am grateful for the help of the inventor of the technology, dr. ir. S. Koussios, without whom the subject of this thesis would not have existed. After having spent over eight months on this thesis research, I can say that I have very much enjoyed it. In the first place because of the chance that is given to let creativity and enthusiasm evolve to achieve a result, but most of all because of the people that were involved. v

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7 ABSTRACT Since the late 1950 s offloading hoses are used for the offloading of crude oil in the offshore industry. The main problem with the current offloading hoses is the limited flexibility and the low resistance against fatigue which results in a short lifetime (1 to 12 years). A composite technology called Integral Geodesic Winding (IGW) has been developed at Delft University of Technology which concerns a new type of composing artificial fibres and rubbers. This results in a tubular structure which resists high pressures and is flexible and lightweight. The main topic of this research is whether or not the use of IGW in offloading hoses is feasible, because the technology shows potential for implementation due to the expectation that IGW leads to higher flexibility and a longer lifetime. An important part of the research is to find the influence of the varying diameter that is inherent to the technology on the loss of pressure in the hose. Finally the feasibility is determined by the advantages and disadvantages it offers for two different types of players in the industry: the hose manufacturer and the system engineer / end user who respectively design and use the offloading facility. First an analysis of the technology was made. Comparison of IGW with existing hose technology shows that IGW puts less strain on its materials, because the structure obtains its flexibility through geometrical change of the composition instead of elastic change of the materials. Implementation of IGW showed the largest potential in three product groups within the offloading hose products. These are the floating hoses, in air catenary hoses and the submarine hoses for depths between 0 and 100 meters. To design the experiment an analysis of the relevant parameters was conducted. This resulted in 4 dimensionless parameters. The first two are the Reynolds number and the relative roughness. In the experiment these could be kept the same as under prototype conditions. Hence a relation between the pressure drop and the IGW model could be described with two dimensionless parameters, ε and α, which describe the geometry of the model. Three models, A, B and C, were constructed ranging from the structurally most beneficial (large diameter variation: A) to the hydraulically most beneficial (small diameter variation: C). A relation has been established between the pressure drop and ε and α. The results showed an unacceptable pressure drop for model A. Therefore the relation that was found was used to estimate the pressure loss for a model Z which has a geometry in between B and C. These three models have been used for a preliminary design. In offloading systems the design velocity often is 7 m/s, because this velocity generates the maximum allowable pressure surge in the case of an accidental closure of a valve. For this velocity the pressure drop was calculated for existing offloading hoses for an internal diameter of 0.6 meters. Then the internal diameters of the IGW models were calculated that produce the same pressure drop at the same flow rate. This gave the dimensions for the design. With this design method the varying diameter does not cause larger pressures for the piping and pumps in the offloading facility to achieve the same offloading rate. Then the quantity of materials, the weight and the costs were calculated. This showed that model Z is interesting to use for the design assuming the flexibility and resistance against fatigue is sufficient. This will have to be verified by experiments on IGW prototypes. Analysis of the impact that this design has on the use and exploitation of the offloading system showed benefits for the customer. The lower weight offers advantages for the transportation and safety. The increased flexibility shows advantages for the design of the offloading facility. The material costs, development costs and production switching costs served as input for the financial analysis of the feasibility for an existing hose manufacturer to implement IGW in its products. Because the manufacturer can offer his customers the mentioned advantages plus a cost reduction through a decrease of hose failures, it is believed that he can increase his market share from 15 to 25 %. However analysis of the market shows that cooperation with a large oil company is indispensable, because they give the product sufficient credibility and have the negotiation power to adapt the standard specifications. The general conclusion is that this implementation is feasible for small ratio (1,14) of the maximum and minimum diameter with an inclination of the pipe wall of 10.3 degrees. The manufacturer of offloading hoses can increase the revenues and obtain a profit margin of 30% by using the technology. His customers profit from a cost reduction of euros over a period of 20 years, the average lifetime of an offloading project, through a decrease of hose failures. Besides these cost savings, there are Health, Safety and Environmental benefits for the end user. vii

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9 TABLE OF CONTENTS List of symbols..xi 1. INTRODUCTION SCOPE OF RESEARCH BACKGROUND PROBLEM DEFINITION SPECIFICATIONS AND OBJECTIVES IGW TECHNOLOGY ESSENTIALS MATHEMATICAL DESCRIPTION ADVANTAGES IN HOSES OFFLOADING HOSES INVENTORY OFFLOADING SYSTEMS HOSES PER MOORING SYSTEM HOSE CATEGORIES FRICTION EXPERIMENT DESIGN PARAMETER ANALYSIS REQUISITES EXPERIMENT PRACTICAL EXECUTION CALIBRATION EXPERIMENT RESULTS MODEL A MODEL B MODEL C COMPARISON IGW MODELS HOSE DESIGN STRUCTURAL & HYDRAULIC INPUT FINAL DESIGN & FABRICATION COSTS IMPACT IGW DESIGN IN USE FEASIBILITY ANALYSIS MANUFACTURER COST ANALYSIS CUSTOMER BUSINESS CASE MARKET ENTRANCE BARRIERS CONCLUSIONS & RECOMMENDATIONS CONCLUSIONS RECOMMENDATIONS...54 References...57 Appendix A: Experiment configuration Appendix B: Model design & construction Appendix C: Data straight pipe & flange calibration 65 Appendix D: Measurements IGW pipe. 71 Appendix E: Weight & cost calculation. 81 ix

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11 LIST OF SYMBOLS Latin A (N) Cf D (m) D 0 (m) D eff (m) D max (m) F (N) g (m/s 2 ) H (m) h (m) H (m) H v (m) k (m) k a L (m) L r (m) Nf Pr (N/m 2 ) p (N/m 2 ) q Q (m 3 /s) r Re U (m/s) Y Z (m) z (m) external load wall friction loss coefficient pipe diameter minimum diameter IGW pipe effective pipe diameter maximum diameter IGW pipe tension force in fibre gravitational acceleration total head pressure head pressure drop pressure drop per bump roughness pipe wall dimensionless axial load length piping repetition length bump number of fibres internal pressure static pressure aspect ratio of a vessel flow rate dimensionless axial force Reynolds number average flow velocity dimensionless radius meridian profile distance to reference level Greek α β ε ζ ν (m 2 /s) ξ ρ (m) ρ 0 (m) ρ eq (m) inclination pipe wall diameter ratio flange diameter ratio inclination coefficient dynamic viscosity local loss coefficient radius minimum shell radius equatorial radius xi

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13 INTRODUCTION 1. INTRODUCTION Since the late 1950 s offloading hoses are used for the offloading of crude oil in the offshore industry. The design has always consisted of a vulcanised rubber hose with additional steel wiring or fabric plies. Innovation always relied on the use of new materials instead of a new composition of the hose. The main problem with the current offloading hoses is the limited flexibility and the low resistance against fatigue which results in a short lifetime (1 to 12 years). A new composite technology called Integral Geodesic Winding (IGW) has been developed at Delft University of Technology which concerns a new type of composing artificial fibres and rubbers. This results in a tubular structure which resists high pressures and is flexible and lightweight. This innovation does not rely on small improvements in the choice of materials, but offers a new perspective in the design and construction of the hose. Because this technology is also expected to increase the lifetime of the hose it became relevant to know whether or not the use of IGW offers advantages for the offloading hose. The main topic of this research is the feasibility of the use of IGW technology in offloading hoses. This feasibility is determined by the advantages and disadvantages it offers for two different types of players in the industry: the hose manufacturer and the system engineer / end user who respectively design and use the offloading facility. Hence not only must the design and costs be investigated, but also the impact that the use of the technology has on the design and exploitation of the offloading facility. The most important part of the research was to find the influence of the varying diameter that is inherent to the technology on the loss of pressure in the hose. This report has the following structure. In Chapter 2 the background of how the idea rose to do this investigation is described more extensively and the specifications and objectives are given. Then in Chapter 3 the technology is described in detail leading to a number of advantages of its implementation in offloading hoses. With these advantages the most promising products for IGW implementation within the offloading hose industry are selected in Chapter 4. Then in Chapter 5 a description is given of how the friction experiment was designed to obtain the desired data. In Chapter 6 the results of this experiment are given and relations between the parameters are established. This serves as input for the preliminary design of the IGW hose and her impact on the use and exploitation of the offloading facility in Chapter 7. Finally sufficient information is gathered to do a feasibility analysis with the help of a manufacturer cost analysis and a customer business case in Chapter 8. The conclusions and recommendations of the report are presented in Chapter 9. 1

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15 IGW TECHNOLOGY 2. SCOPE OF RESEARCH This chapter provides the background of how the development of a new composite technology with extraordinary characteristics gave rise to an investigation in its commercial opportunities. One of the outcomes is that the technology offers great potential in offloading hoses used for the offloading of crude oil in the offshore industry. The main research topic is the feasibility of the technology in this product. 2.1 BACKGROUND Currently a lot of fibre reinforced materials (composites) are being used in stead of steel. This results in lightweight solutions. With current composite technologies a trade-off (Figure 2-1) exists between pressure resistance, flexibility and a low weight. This causes the problem that the desired combination of these characteristics often can not be obtained or leads to expensive solutions with fabric plies. Figure 2-1: trade off composites In recent years the use of composites has gained a great deal in popularity. More often steel and other metals are replaced by these composites. Several advantages exist when comparing composites to metals. These include their low density, high resistance against fatigue and noncorrosiveness. In a course taught in December 2004 at the Delft University of Technology groups of students get to choose a patent that has not been commercialized yet and subject of the course is to find a commercial application. One of those patents was Pressurisable structures comprising different surface sections developed by Prof. Beukers, Dr. Koussios and Dr. Bergsma of the faculty of Aerospace Engineering [1]. The patented technology concerns a new type of fibre wound objects. This is graphically explained in Figure 2-2. The invention is a transformation of a former technology that concerns the geodesic winding of elliptical spheres (left hand side of the figure). Integration of these spheres leads to the Integrally Geodesic Wound (IGW) body on the right hand side of the figure. Figure 2-2: Integrally Geodesic Wound bodies 3

16 IGW TECHNOLOGY This technology enables high pressurisable tubular structures to be flexible and lightweight. A part of this invention is the use of geodesic paths on tubular structures. Fibres are wound around this structure on the shortest distance, the so called geodesic paths. The advantage of fibres on geodesic paths is that internal pressure in a structure will result in an evenly distributed tensile force in the fibres. Fibres can resist a very high tensile force. The key of this invention is the ability to wind a tubular structure with fibres running on geodesic paths. Before this invention this was only possible with helical paths. Helical winding results in insufficient flexibility and low resistance against fatigue. The invention is called Integral Geodesic Winding (IGW), a technology that does not trade off because it combines the benefits of composites (Figure 2-3). Figure 2-3: IGW does not trade-off In the commercialization research different applications in the offshore industry and gas storage were found. A promising alternative is the use of the technology in the flexible hoses that are used for loading and offloading of crude oil in mooring systems. An example of such a system is illustrated in Figure 2-4. The concerned hose is seen floating on the water surface. The IGW technology offers many advantages which will come to light further on. Figure 2-4: Single point mooring TANIQ has been founded to investigate the possibilities of the technology and has already been awarded a valorisation grant of euros which is used at this moment for the construction of a scaled prototype of the IGW technology. This prototype will be tested in August Besides the construction of this proof of technology it is desirable to investigate the feasibility of implementation of IGW technology in offloading hoses. 2.2 PROBLEM DEFINITION It is unknown if the use of the IGW technology flexible offloading hoses is feasible and for which type of offloading hoses. Besides its structural design, the consequences of the use of IGW in offloading hoses for the offloading system design are an important issue. The last problem, the feasibility of IGW in offloading hoses is the main topic of research. An important part of this is the analysis of the loss of pressure head caused by the varying diameter. The other investigation, i.e. the structural design of the hose, is being conducted in another graduation project at the faculty of Aerospace Engineering at the time of writing this report [2]. The preliminary outcome of structural design and the data on pressure head loss serve as input for a preliminary design of the IGW hose which is subjected to a feasibility analysis. 4

17 IGW TECHNOLOGY The design is done according to the industry specifications. It has to be noted that these specifications are minimal and to make the technology feasible in this case it either has to perform better or cost less than the conventional offloading hose (see Figure 2-5). Finally it is unmistakable that the market should be ready for innovation for a successful implementation. Figure 2-5: Conventional offloading hose build up 2.3 SPECIFICATIONS AND OBJECTIVES First the specifications that the new design with IGW has to meet are presented. These result in the research objectives that have to be investigated to analyse if the implementation of IGW in offloading hoses is feasible. Specifications The new hose has to be designed in such a way that the capacity of the pumps and the piping of the offloading facility is not affected. This specification makes sure that the market is ready for the new product. It has to be possible to execute the conventional testing and maintenance procedures without a problem. The customer, which designs and exploits the offloading facility, must benefit from the new design. The manufacturer of hoses who can adapt the technology has to benefit from the new IGW hose. The hose design must meet the OCIMF industry specifications that are presented in paragraph 4.3. This is part of the other investigation conducted at the faculty of Aerospace Engineering. Main objective The main objective is to establish a relationship by experiment with which the pressure drop in the IGW hose can be predicted for various shapes. Additional objectives The additional objectives are established to gather the required information to finally make a feasibility analysis. These objectives are: To analyze the IGW technology for offloading hoses. This leads to the specific advantages of IGW for the use in offloading hoses. To choose appropriate offloading hose products for IGW. These are the products that have the potential to benefit from the specific advantages of IGW. To make a preliminary design of the IGW offloading hose. This evolves from the pressure drop experiment and from structural design investigation. An optimal shape that meets industry specifications and is a good compromise between strength, flexibility, durability and streamlining, is chosen. To compare the system life time cycle costs. Impact on different aspects such as the design of the flexible, installation procedures, maintenance procedures and ancillary equipment is quantified in money. 5

18 IGW TECHNOLOGY To analyze HSE issues and market entrance barriers. Together with the cost comparison, analysis of the market and health, safety and environmental issues indicate whether or not the use of IGW in offloading hoses is feasible. 6

19 IGW TECHNOLOGY 3. IGW TECHNOLOGY In this chapter the technology which is to be implemented in the hoses is explained. This serves as input for the design of the experiment in Chapter 5 and the final design of the offloading hose in Chapter 7. First the essentials are explained. Then the mathematical background [3,4] is dealt with and examples are given of how variation of the parameters can result in different configurations with different characteristics. Finally the advantages and especially those advantages that it offers for flexible offloading hoses are introduced. 3.1 ESSENTIALS The technology is a novel class of filament wound articulated pressurizable structures (APS) that contain alternating convex and concave areas. Different geometries for such structures can be found in Figure 3-1. Filament winding is towing fibres around structure to obtain a composite vessel. Composite vessels generally are made in rigid form, because the majority is created by means of wet winding or the use of pre-impregnated fibre bundles. However in this case the tows are applied dry and will remain on place due to geodesic winding or by friction. Figure 3-1: APS with different geometries Because of this dry winding, a flexible matrix (a matrix is a layer, i.e. rubber, which holds the fibres together) can be applied during or after fibre placement resulting in a flexible structure. Another interesting characteristic of the APS is that the convex and concave regions both are practically isotensoid over the complete range of the applied tow length. An example of a single isotensoid fibre path is drawn on the structure in Figure 3-2. Figure 3-2: Geodesic fibre path 7

20 IGW TECHNOLOGY 3.2 MATHEMATICAL DESCRIPTION For discussion of the technology the netting theory is used. Two important and limiting assumptions in this theory are that the wall thickness remains small and that the contribution of the stiffness of the matrix is neglected. The composite vessel can thus be considered as a thin shell of revolution, subjected to internal pressure (Pr) and external loads (A) as depicted in Figure 3-3. Figure 3-3: Loads on a membrane element The pressure vessel can be considered optimal in terms of utilising the maximum strength of the fibre bundle when the in-plane shear stress of the laminate is zero. Assuming that a number Nf of fibres is loaded with a constant tensional force, F, the orientation of the fibre for a given geometry can be found through the following formula: 2 FNf cosα cos β = Prπρ + A (1) ρ is the radius and the definition of the angles can be found in Figure 3-4. The contour of the vessel is called the meridian profile. Because the radius, ρ, changes with the vertical position, z, the fibre path constantly is reoriented (change of the angles α and β) maintaining an isotensoidal path. Figure 3-4: Pressure vessel loads and basic geometry 8

21 IGW TECHNOLOGY Now the following dimensionless coefficients are introduced in which ρ_0 denotes the minimum shell radius: ρ A Y = (2) ka = (3) 2 ρ 0 π Pr ρ 0 With these coefficients the following parameters, q and r can be defined: Y q = Y eq min 2 (4) ka r = (5) 2 Yeq The subscript eq stands for equatorial, being the largest diameter of the body. q is referred to as the aspect ratio of the vessel, a shape parameter, and r stands for the dimensionless axial force, a load parameter. The shape of the meridian profile depends entirely on q and r. However not every combination of values results in the right kind of meridian profile due to manufacturing issues and desired use. For the design of an APS the external forces, F and Pr, and the minimum shell radius, ρ_0, can be used as the input parameters. The only parameter left for variation is ρ_eq and by choosing its value q and r are fixed. When q and r are fixed, only one meridian profile is possible. Thus by varying the equatorial radius different shapes with different qualities can be created. Manufacturing issues A basic cell belonging to the APS is entirely optimal only in the case where r = -1/q. Entirely optimal means here that the fibres run over geodesic paths over the entire trajectory. However for this value a manufacturing issue becomes important, because the maximum height of the profile is equal to the height it achieves at the polar opening (see the vessel depicted in the lower right corner of Figure 3-1). In this case the resulting cells cannot be stacked onto each other, because the lower surface of the upper vessel contacts the upper surface of the lower vessel. Thus no space remains for the fibres. If r is chosen close to zero and q relatively small, the concavity of the structure becomes quite small (see the vessel depicted in the upper left corner of Figure 3-1). This means that the fibre bundle is able to contact the concave surface at every point, because the fibre bundle is able to make the twist gradually. Figure 3-5: Influence of the r factor on the resulting meridian profile It is also possible to choose the radii and then vary the external loads. The resulting meridian profiles are given in Figure 3-5. As discussed before the loads are included in the r factor. If the vessel is loaded under tension and positive internal pressure, which is the case under normal practice with hoses, a higher pressure results in a decrease of Ka and of r. A decrease of the r factor, results in a longitudinal compression of the vessel. This can be interpreted from the meridian profiles which are given for different values of r in Figure 3-5. A large negative r factor, this is the case with a large negative axial force and a positive internal pressure, results in a donut shaped meridian profile. With r smaller than -1/q the height at the pole opening is not the largest anymore and it becomes impossible keep the fibres on the surface and still continue their course to the next surface. 9

22 IGW TECHNOLOGY The following parameterisation in elliptical coordinates of Y is introduced: Finally the description of the meridian profile, which depends only on q and r can be derived, resulting in: (6) (7) 3.3 ADVANTAGES IN HOSES When looking at the technology and its practical application, there appear to be a number of advantages for her implementation in flexible hoses. To obtain insight in the working principle, the IGW structure is compared to the existing design that shows the greatest similarity. This is a helical wound structure. There are two important differences between IGW and helical winding. The first one is the behaviour of the fibres when subjected to a load displacement (see Figure 3-6). In the helical wound structure the angle between the fibres changes radically. The same displacement in the IGW structure causes a slight change of the angle between the fibres. Another important difference is that for the same translation or flexure, the elongation of the fibre materials and the matrix materials is less with IGW than with her helical competitor. These two aspects make that IGW is more flexible and puts less stresses on its materials than helical winding. Figure 3-6: Fibre movement under load displacement A graphical summary of the advantages of IGW in hoses, which are explained underneath, is given in Figure 3-7: High pressure: composite structures have proven to be advantageous in high pressure hose applications. Flexibility: the fact that the structure comprises of convex and concave areas makes it more flexible than the helically wound alternative. Because the materials elongate less, the structure also becomes more flexible. Resistance against fatigue: if the IGW technology is compared with the technology that shows the greatest similarity, helical winding, it appears that the fibres in the IGW structure displace less and elongate less under a load displacement of the structure. This results in lower stresses in the matrix and fibres and thus in a higher resistance against fatigue. Wind ability: because the fibres are wound geodesically, the matrix can be applied afterwards. This means that the fibres need not be impregnated, which halves the production time and thus reduces production costs. 10

23 IGW TECHNOLOGY Simplified composition: Because the fibre winding can take up the stresses, no combination of different materials, i.e. fabric ply s and steel helix wires, are needed to achieve the desired characteristics. This results in an easier to construct composition involving fewer layers than conventional hoses. Light structure: Since fewer layers are applied and no steel is necessary in the construction of the hose, a lighter hose can be constructed. This means that less buoyancy has to be applied to obtain the desired buoyancy and ancillary equipment has to withstand lighter loads. Resistance against impact: As already has been proven in tests with an LPG tank constructed with the former technology (see Figure 2-2), the product has a very high resistance against impact. The structure can absorb the impact by temporary deformation. Figure 3-7: Advantages of IGW in hoses 11

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25 OFFLOADING HOSES 4. OFFLOADING HOSES There is a wide variety of mooring systems that use offloading hoses. The conditions under which they must function vary as well. Even within one system there is a difference in the specifications of the different segments of which it is made up. In this chapter all the mooring systems are first described. Secondly a short description of the different types of hoses that can be found within one system explains the need for a categorization beyond the type of mooring system. Finally all the hoses are divided according to their characteristics into 4 main product lines and then the product lines are selected that fit within the boundary conditions of the investigation. 4.1 INVENTORY OFFLOADING SYSTEMS Mainly there are four different set ups for a loading/offloading system of crude oil and related products [5]. Though other configurations are possible, these four cover the larger part of the possibilities CALM buoys A Catenary Anchor Leg Mooring (CALM) buoy is depicted in Figure 4-1. The mooring buoy has a turntable that can rotate around its vertical axis. The tanker moors to this turntable and connects to the floating hose strings that are attached to the turntable as well. The whole system can rotate freely because of the forces exerted by the currents and waves. This is called wheathervaning. The buoy itself is kept in place by the anchor chains. Through a swivel the fluid is conveyed to the submarine hose strings that connect via the pipeline end manifold to the submarine pipeline. Figure 4-1: CALM buoy system Submarine hoses The submarine hoses consist of one or more, but generally not exceeding three, separate strings. The configuration needs to be such that the strings do not become overstressed by becoming over extended. Also the string can not be made too long, because then the hose strings can kink or become damaged by abrasion on the sea floor. In general there are three different configurations for the submarine hose. All three configurations must be dimensioned in such a manner that they maintain the proper configuration in oil filled as well as sea water filled conditions. Choice depends on environmental and operational conditions. 13

26 OFFLOADING HOSES Chinese lantern hose configuration (see Figure 4-2) Underwater floats are attached to the hose string and the string should maintain an acceptable configuration in the most extreme conditions. The hoses that connect to the pipe line end manifold should be reinforced at the attached end to avoid concentrated bending stresses. Figure 4-2: Chinese lantern configuration Lazy-S hose configuration (see Figure 4-3) Floatation in the form of bead floats or adjustable buoyancy tanks is provided to achieve the proper hose configuration. Again this is done to avoid damage by kinking or wear on the hoses. Figure 4-3: Lazy-S configuration 14

27 OFFLOADING HOSES Steep-S hose configuration (see Figure 4-4) This configuration is well suited for CALM buoys in water depths exceeding 45 meters. The capacity of the buoyancy tanks must be such that positive buoyancy is maintained for hoses both in oil filled as well as water filled conditions. Figure 4-4: Steep-S configuration Floating hose system The hose string needs floatation media through integral floatation. The string usually consists of varying degrees of stiffness to help transition stress between the string and the ship s manifold and between the string and the buoy. Single or twin hose main strings are used with a diameter up to 750 mm. Because of the ship s derrick s lifting capacity the rail hose s (hose nearest to the ship) diameter is usually limited to 400 mm. Tapered hoses are used to gradually reduce the diameter when needed. The rail hose also needs an extra floatation medium to compensate the extra weight of the installation devices attached to the hose SALM buoys A SALM (Single Anchor Leg Mooring) system consists of a mooring buoy attached to a gravity or piled base with an anchor chain with a swivel to allow rotation (see Figure 4-5). Like in the CALM system the tanker moors with one or more hawsers to the buoy, but in the SALM system the hose strings go directly to the pipeline end manifold. A fluid swivel is mounted concentrically about the anchor leg to permit the passage of fluid while rotating in response to the moored vessel s movements. 15

28 OFFLOADING HOSES Figure 4-5: SALM buoy system The submarine hose string connects to the fluid swivel. The hose arm may be pivoted about the horizontal axis, thus permitting the angle at which the hose angle leaves the fluid swivel to vary, or the hose arm may be set at a fixed angle. The hose string may be configured in an s-shape or can start vertical and curve slowly towards the horizontal water plane. Buoyancy must be applied to obtain the desired shape. This can be done by using integral floatation or with hose floats. Integral floatation however does not function well under a large hydrostatic head. Sufficient buoyancy must be provided between the submarine and floating hoses to prevent the floating hose from being pulled underwater. Hose sections with reinforced ends must be used to distribute the bending stresses created by the transition between the submarine and floating hoses Floating (tandem mooring) In certain cases a Floating Production Storage and Offloading (FPSO) system is used for extracting the oil from the well. These ships are moored with a turret with anchor lines as is depicted in Figure 4-6. The turret works by the same principle as a single point mooring system. Thus the position of the FPSO is determined by the forces exerted by the wind and the sea, because it can weathervane. Figure 4-6: Floating Production Storage and Offloading system 16

29 OFFLOADING HOSES The FPSO frequently needs to offload its oil for which a shuttle tanker moors to its stern by connecting with a hawser. Then FPSO has a flexible hose which floats on the water. Buoyancy must be applied by using integral floatation or bead floats. This hose string is now connected to the shuttle tanker so that the FPSO can offload its products. The two ships move as one system and the shuttle tanker stays in line with the FPSO. When the shuttle tanker is full, the hose string is disconnected and either left floating on the water or stored on the FPSO Stored hose (tandem mooring) In this situation the mooring takes place in a tandem configuration again. But now the hose is used floating on the water hanging as an in-air catenary (see Figure 4-6). When the offloading hose is not used, it is stored on the FPSO on a hose reel or in a chute (Figure 4-7). The purpose of this design is that the hose can be retrieved and stored on the FPSO during heavy weather conditions. This type of hoses needs to have a very high reserve of flexibility, bend radius tolerance and axial strength. The catenary design also exists with integral buoyancy so that the hose string is not lost, even when filled with water, in the case of an emergency release. Figure 4-7: Storage on hose reel & in a chute Storage by reeling is more popular than storage in a chute, because it occupies less space, but it does put higher demands on the hose design Deepwater export line In project areas with relatively light environmental conditions often the less expensive spread mooring is used instead of a turret mooring system. The problem that this generates is that it is not possible anymore to do tandem mooring with the shuttle tanker, because of the fixed position of the FPSO. A possible solution is a deepwater export line that runs to a CALM buoy to which the shuttle tanker moors (see Figure 4-8). Here the tankers can weathervane without risking collision with the FPSO. The deepest point of the hose usually is between 200 and 300 meters. Figure 4-8:deepwater export line 17

30 OFFLOADING HOSES 4.2 HOSES PER MOORING SYSTEM The different types of existing offloading systems have been introduced. Most systems require more than one type of hose and some of the hoses occur in different systems. In Figure 4-9 and Figure 4-10 a CALM buoy and a tandem mooring configuration are depicted and it becomes clear that quite some hoses are the same. Figure 4-9: CALM buoy hose types Figure 4-10: Tandem mooring hose types Within every system different functions apply to different segments of the hose string. Without going into detail about the exact properties of every segment of the hose, those three main properties are introduced here which are varied to reach the desired design. Quantity and placement of floatation material: Floating hose strings need sufficient buoyancy to stay afloat. Certain segments that have heavy ancillary equipment attached need an even higher buoyancy reserve. Hose segments that make the transition between hanging and floating or floating and submerged need floatation material on certain spots. End reinforcement bending stiffeners: Those hose strings that make the connection between the floating or hanging hose string and rigid piping on FPSO s, tankers, buoys or PLEM s need gradual reinforcement to prevent a concentration of the bending stresses near the concerned flange. Internal diameter: Because of the limited lifting capacity of the used derrick often the hose segments that run from the water surface to the manifold on the tanker or FPSO have a smaller diameter. Thus tapered hoses which reduce the diameter in the longitudinal direction are necessary. 18

31 OFFLOADING HOSES 4.3 HOSE CATEGORIES Despite differences between separate segments of the hose string all hoses will be divided into five categories which differ on important aspects. It should not be forgotten that within every product line variations such as those mentioned in 4.2 exist. First the main specifications defined by Oil Companies International Marine Forum (OCIMF) which apply to all conventional types of hoses are given here [6]. Then the aspects on which the product lines deviate from each other are mentioned. General specifications: Diameter: 150mm<D<600mm Pressure rating: 15 bar (19bar or 21 bar available on request) Operating pressure: bar gauge to designated pressure rating Fluid temperature: -20 C <T<82 C Ambient temperature: -29 C<T<52 C Temporary elongation: < 2.5% (material related) Permanent elongation: < 0.7% (material related) Resistance to: 25% aromatic content in petroleum products vflow for D<400mm: 21 m/s vflow for D>400mm: 15 m/s Axial strength: 37 tons for D=600mm The loading and offloading hoses can be divided into the following five groups: the floating hose string, the catenary hose string, the submarine hose string (<100m) and the deepwater submarine hose string (<300m). For every product line the typical properties and the mooring systems of 4.1 in which it can appear are given. Floating hose string The floating hose string occurs in CALM, SALM and floating tandem mooring systems: A buoyancy reserve of 20% must be obtained. Under all conditions it may be bent to a bending radius of 6 times the nominal diameter. However a smaller bending radius is desirable. Catenary hose string This product line only occurs in tandem mooring systems where the hose is deployed from the FPSO and left hanging between the FPSO and shuttle tanker: An axial strength of 37 tons must be obtained, because of the fact that the string has to carry its own weight. A minimum bend radius of 4 times the nominal diameter is required, but for reasons of space occupancy of the hose reel often a more flexible design is desired. Submarine hose string (<100m) This product line often is used for the submarine parts of the hose strings in CALM and SALM systems: Because of the hydrostatic head of the sea water a differential pressure of 3 to 5 bars can occur (the string always remains filled with sea water, oil or condensate). A minimum bend radius of 4 times the nominal diameter is required. Deepwater submarine hose string (>300m) This product line generally is used in deepwater export lines with spread anchor moored FPSO s. However sometimes it is used in SALM and CALM buoys if depths of a 100 meters are exceeded. The hose has to withstand a differential pressure of 15 bars. Because of the length of the string it has to withstand a larger internal pressure, which often results in a pressure rating of 24 bars. Selected categories The deepwater submarine hose string does not qualify for the investigation since it has to withstand large external pressures, i.e. a differential pressure of 15 bars, which is not the key strength of the technology. To conclude the product lines A, B and C will be used in the investigation. 19

32 OFFLOADING HOSES 20

33 FRICTION EXPERIMENT DESIGN 5. FRICTION EXPERIMENT DESIGN As has been defined in the scope of research an important issue that needs to be investigated is the loss of pressure head due to the varying diameter. This is investigated through an experiment which is designed in this chapter. First the parameters of the losses in pipe flow are analysed and made dimensionless. Then a transformation into dimensionless form shows that 4 dimensionless parameters are necessary to describe the friction in the IGW model. Analysis shows that an important parameter, the Reynolds number, has to be 12,000 < Re < 100,000 during the experiment. Then three IGW geometries are chosen which can generate data that can be used for the final hose design. Based on these models, losses are estimated resulting in requirements for the experiment. Finally the experiment configuration and measurement equipment which meet these requirements are presented. 5.1 PARAMETER ANALYSIS Analysis of the parameters shows two dimensionless parameters that relate to the geometry of the IGW model. The other dimensionless parameters are the relative roughness and the Reynolds number Losses in pipe flow In this paragraph a review is given of the losses that occur in pipe flow as a preparation for the design of the experiment. In Figure 5-1 different types of losses are indicated for a pipe configuration [7]. The Bernoulli equation is used to describe the total H (indicated as e.n. in the figure). In pipe flow it is common to indicate pressures in the height of the column of liquid that can be associated with this pressure. In this case it is meters of water column. entry loss H h outflow loss Figure 5-1: losses in pipe flow (CT 2100, Fluid Mechanics) H 2 p U = + (8) ρ g 2g H (m) = total head 2 p( N / m ) = static pressure 3 ρ( kg / m ) = density fluid 21

34 FRICTION EXPERIMENT DESIGN 2 g( m / s ) = gravitational acceleration U ( m / s) = fluid velocity The first two terms of the equation form the pressure head, h (indicated as p.n. in Figure 5-1). p h = (9) ρg The difference between the total head and the pressure head is called the velocity head. Along the pipe losses occur. These can be found in the drop of H. There are two different types of losses, i.e. local losses and wall friction. Local losses are caused by local discontinuities such as an inlet, sudden change of diameter, constriction, outlet etc. These losses cause a sudden drop of the energy line. Wall friction is energy dissipation between the fluid and the pipe wall. This loss occurs all along the pipe. The total pressure drop is the sum of these two types of losses. H (m) = pressure drop ξ( ) = local loss coefficient Cf ( ) = wall friction loss coefficient L(m) = length D(m) = diameter 2 2 U U L H = ( ξ ) + ( Cf ) (10) 2g 2g D The first member of the equation is the sum of the local losses. The second member of the equation is the sum of the losses due to the wall friction in the different pipe segments of which the setup consists Dimensionless parameters The goal of the experiment is to obtain more knowledge about the pressure drop in the new hose constructed with the IGW technology. Therefore certain relationships between this loss and the other parameters have to be found. The pressure drop in the IGW hose is described by the following formula: H 2 U L = C + Σ 2g D f Hv 0 (11) D0( m) = minimum diameter IGW pipe Hv(m) = pressure drop per bump The total pressure loss thus is generated by the wall friction and the varying diameter. The pressure drop caused by the varying diameter can not be measured separately. The reason for this is that both losses occur along the pipe. However their contributions are split for better understanding of the flow phenomena. For the loss due to wall friction the classic formula is used with the smallest diameter of the IGW pipe, D 0. Now two new parameters are introduced that describe the shape of the hose (see Figure 5-2): D max = the maximum diameter (in the convex part) Lr = repetition length 22

35 FRICTION EXPERIMENT DESIGN Figure 5-2: hose shape parameters First the independent parameters on which H depends are presented. Then with the help of the pi-theorem the number of required dimensionless parameters is calculated. Finally the parameters are made dimensionless and their meaning is discussed. H depends on the following parameters: 1 ν ( m 2 s ) = dynamic viscosity k (m) = roughness pipe wall H = f ( D0, D max, Lr, U, ν, k) The pi-theorem states: # independent parameters - # base dimensions = # required dimensionless parameters 6 2 (length and time) = 4 Thus 4 dimensionless parameters have to be formed, resulting in: C fd D = ( D max 0, tan 1 D ( D Lr max 0 UD ), ν 0 k, D 0 ) D = D max ε (12) 0 This is a parameter which describes the geometry of the IGW model. It is the ratio of the minimum and maximum diameter. Different combinations of q and r (these parameters are described in chapter 3) give different values for the geometry. This shape parameter is important for the local losses generated by the bumps. 1 D max D0 α = tan ( ) (13) This is the second parameter which describes the geometry of the test model. Its value indicates the inclination of the imaginary line that runs from the highest point to the lowest point of the pipe wall. This shape parameter is important for the local losses generated by the bumps. Re UD ν Lr 0 = (14) This is the Reynolds number. Re has to indicate a similar flow regime in the model as in the prototype. If Re is to be kept constant a different velocity must be used in the experiment, because D and ν are different in the test model. This parameter is important for the loss generated by the wall friction. 23

36 FRICTION EXPERIMENT DESIGN k D0 This is the relative roughness coefficient. This coefficient does not have to be exactly the same in the model as in the prototype, but it should indicate the same hydraulic roughness conditions in both cases. This is explained further on. This parameter is important for the loss generated by the wall friction. The parameters ε and α are related to q, the aspect ratio of the vessel, and r, the dimensionless axial force. q and r were introduced in Chapter 3. As can be deducted from Graph 5-1 and Graph 5-2, which depict the relations between the parameters, a combination of ε and α leads to a unique combination of q and r. q and r determine the meridian profile (the shape of the IGW model) with equation 7. Because ε and α lead to a unique solution for the shape, they are appropriate to use as parameters for the experiment. (15) Graph 5-1: relation between q, ε and r Graph 5-2: relation between ε, α and r For the analysis of Hv Carnot s equation is used. This equation is designed to calculate losses in an expansion. The losses in the contractions are marginal compared to the losses in the expansions and therefore are not taken into account. The loss caused by the expansion depends on the velocity difference and a coefficient which describes the influence of the inclination of the expansion on the loss. Data of the relation for a single expansion between this coefficient (notated with ζ), the surface ratio and angle of the expansion is given in Graph 5-3. The angle used in the graph is two times larger than the angle used in the IGW model, α. 24

37 FRICTION EXPERIMENT DESIGN ε 2 ζ 2α ζ 2α Graph 5-3: Local loss coefficient for a single expansion The data in this graph is valid in the case of a single expansion. However when expansions and contractions follow each other up rapidly, which is the case in the IGW model, these losses calculated with these figures probably result in conservative values. This is because the flow does not reattach to the wall at the widest point. The streamlining of the IGW model also causes lower losses than the experimental data for the single expansion. Analysis of the local loss with Carnot s equation with the IGW model dimensionless parameters results in: Hv ( U = ζ U 2) 2g 0 2 = ζ (1 A0 ) A2 2 U 2 0 2g D = ζ (1 ( D 0 max ) 2 ) 2 U g This leads to: U 0 Hv = ζ (1 ε ) (16) 2g In this loss the coefficient, ζ, is a function of α and ε. Without this coefficient the formula calculates the loss caused by a sudden expansion for an angle of 90 and straight piping after the expansion. Since in the IGW model the angle of the expansion is different and the expansion is not followed by straight piping, which is true in the case of a single expansion, the coefficient ζ is introduced: ζ ( ) = inclination coefficient ζ = f ( α, ε ) Looking at the inclination coefficient lines for different α s and ε s for a single expansion it becomes clear that the influence of α is larger. This probably also is the case for the IGW model. Therefore the influence of ε is neglected in the data processing of the IGW model. Next the repetition length, Lr, is expressed in the dimensionless parameters: Lr D max D tanα 0 εd0 D = tanα 0 ( ε 1) D = tanα = (17) Using formulas 11, 16 and 17 the total pressure drop for a certain length of IGW pipe then can be described with: 0 25

38 FRICTION EXPERIMENT DESIGN 2 2 U L L tanα L U tanα 2 2 H = Cf + Hv = ( Cf + ζ (1 ε ) ) (18) 2g D0 D0( ε 1) D0 2g ( ε 1) 5.2 REQUISITES EXPERIMENT With the dimensionless parameters that were introduced in the previous paragraph an analysis is made of the requisites for the experiment. First the prototype conditions under industry practice are dealt with. Then the dimensions of 3 models are determined to provide the desired data. Through the dimensionless parameters the magnitude of the parameters in the model are found which form the requisites of the experiment such as the necessary pressure head Prototype conditions First the flow conditions of a conventional offloading hose in normal practice are determined. To get an impression of the friction coefficients the pressure drops calculated with the formula used by ITR Spa, a hose manufacturer, are used. Then the Reynolds numbers are calculated which indicate the range in which the experiment should be conducted. Here new parameters are introduced and some values for normal practice are given: *10 < ν ( m s ) < 1*10 = typical range of kinematic viscosity for processed oil 1 Q( m 3 h ) = flow rate 6 k (m) = 15*10 = roughness (value extracted from a piping design calculation) Cf ( ) = friction coefficient (only wall friction) 3 dp dx( Nm ) = pressure drop per meter An example has been worked out for diameters of respectively 0.15m and 0.6m (see Table 5-1 and Table 5-2). Re can be calculated from the data. If Re > 12,000 it is sure the flow is turbulent. This is the case for both diameters for the given range of flow velocities and viscosities for which Re is calculated. According to Blasius [8] if the following statement is true, then the turbulent flow can be considered hydraulically smooth: Re 65D / k (19) This condition is met for all velocities of the table. The friction coefficients then can be calculated with the Nikuradse formula. This relative roughness is a dimensionless parameter that in the experiment must be scaled in such a way that the flow conditions are hydraulically smooth as well C f = ,221 (Re) (20) Do=0.15m, ν1=5*10^-6 & v2=1*10^-5, k= Q (m^3/h) U (m/s) Re, v1 Re, v2 Cf1 Nikuradse Cf2 Nikuradse Cf1 ITR Cf2 ITR , , , , , , , , , , Table 5-1: prototype conditions for D=0.15m 26

39 FRICTION EXPERIMENT DESIGN Do=0.6m, ν1=5*10^-6 & v2=1*10^-5, k= Q (m^3/h) U (m/s) Re, v1 Re, v2 Cf1 Nikurads e Cf2 Nikurads e Cf1 ITR Cf2 ITR 6, , , , ,179, , , ,651, , , ,123,142 1,061, , ,476,999 1,238, Table 5-2: prototype conditions for D=0.6m The first four columns in the tables contain data that are valid for the existing hoses and for the new IGW hose. The 5 th and the 6 th column are the friction coefficients calculated with Nikuradse. Since the flow regime is turbulent and hydraulically smooth, there is no need to experiment the whole region up to an Re of 2,500,000. Thus a relation can already be found with an experiment with 12,000 < Re < 100,000 setting the first boundary condition for the experiment. Another issue that makes testing for large Reynold numbers unnecessary is that the velocities above 7 m /s, which are specified by the hose manufacturers, under normal practice never occur. The last two columns contain friction coefficients calculated with the formula used by ITR Spa [9], a hose manufacturer (see Figure 5-3). These friction coefficients are larger than the coefficients calculated with Nikuradse. Since the formulas of Nikuradse are more reliable, these will be used for the data processing of the experiment measurements in Chapter 6. As can be seen the friction coefficient becomes smaller as the velocity becomes larger. In this formula a larger velocity makes the friction coefficient smaller through the Reynolds number. -4 Figure 5-3: Formula for pressure drop by hose manufacturer ITR Spa The influence of the hydraulic regime can be demonstrated very well with the formula of ITR. As the relative roughness (K/D) becomes larger, the 1 st member of the equation increases and the friction coefficient increases as well. If the velocity becomes larger, the 2 nd member of the equation decreases because Re increases and thus the friction coefficient decreases IGW model choice In this paragraph the dimensions of the models which are used for the experiment are determined. In order to achieve this first the desired format of the outcome of the experiment is defined. Then together with the preliminary results of the research of the design of the hose the quantity and dimensions of the models are generated. 27

40 FRICTION EXPERIMENT DESIGN Two different ways of analysing the data are used in this research. First for a practical comparison a comparison is made of the IGW pipe to a straight pipe. The diameter of the straight pipe that causes the same pressure drop per meter as the IGW pipe with the same flow rate is called the effective diameter. Thus for a known α and ε the ratio of the effective diameter to the minimum diameter of the IGW pipe can be found: 1) for a combination of D max 1 D max D0 ε = and = tan ( ) D0 Lr α => Deff D0 D eff = effective diameter The second approach for data analysis is calculating the pressure drop per bump by subtracting the loss caused by the wall friction and dividing the pressure drop by the number of bumps. Therefore equation 11 is used again. Then Hv can be plotted against U^2/2g for given values of α and ε: 2) 2 U L Σ Hv = H Cf (11) 2g D For D max 1 D max D0 ε = and α = tan ( ) => Hv D0 Lr plotted against 2 U 2g The preliminary results of the structural hose design showed an optimal fibre configuration for the geometry which has an ε and α of respectively 1.6 and From a hydraulic point of view it is clear that a larger difference of the diameters causes a larger pressure drop. To establish relationships between α, ε and Hv a minimum of 9 models is required. However the budget and time span of the research allowed the construction of 3 models. Thus the choice is made to construct 3 models that are realistic alternatives for the IGW hose that range from the structural most beneficial alternative to a hydraulically beneficial alternative being almost straight (Table 5-3). By neglecting the influence of the diameter difference, ε, on the inclination coefficient, ζ, a relation still can be established. IGW model geometries D_0 (mm) Dmax (mm) Lr (mm) ε α (degrees) Model A Model B Model C Table 5-3: IGW geometries chosen for the experiment Requirements Because the conditions in the prototype that must be simulated and the geometry of the model are known the velocities that must be used in the test model can be calculated through the Reynolds number. With these velocities and a conservative estimation of the pressure drop in the IGW models the limits of the outcome of the experiment are known. In the experiment the smallest diameter in the varying cross section is D_0, the nominal diameter. The velocity, U, is determined in the smallest cross section. This is m because this is a standard diameter used for piping in the laboratory. v = 1 E-6, the viscosity of water, is used in the experiment. The roughness, k, must be smaller than 1 E-5 m. This is no problem since PETG, a 28

41 FRICTION EXPERIMENT DESIGN smooth plastic, will be used for the construction of the test model. The length of the test model, L, is chosen to be 5 meters. Now equation 18 is used again. 2 L U tanα 2 2 H = ( Cf + ζ (1 ε ) ) (18) D0 2g ( ε 1) The first member between the apprentices is the friction caused by the pipe wall. This is calculated with the formulas of Nikuradse and Blasius. The second member between the apprentices represents the pressure drop caused by the bumps. For this a conservative value is calculated which determines the upper boundary of the outcome of the experiment. Model C, which is expected to generate the highest loss, is used for this with a cross section ratio of ε^2 = 1.6^2 = 2.56 an α of The value for the inclination coefficient, ζ, of a single expansion then is 0.96 (see Graph 5-1). This value is used to estimate the loss per bump. Then: 2 ζ (1 ε ) 2 = 0.96 Calculations are made using different Reynolds numbers as input. The results can be found in Table 5-4. Each calculation is done without and with a high estimate of the influence of the concavity. The expected value of H, the pressure drop, will be in between the calculated values with the influence of the bumps (Hmax) and without the influence of the bumps (Hmin). Thus as long as the calculated range of H and the flow rate is feasible with the available equipment and the flow regime is similar, that is hydraulically smooth, the experiment can be conducted. Underneath the table the different steps are given which lead to the pressure drop caused by the wall friction. D=0.0825m, k= , v= , L=5, α=26.1, ε=1.6 Re U (m/s) Q (l/s) Cf bl Cf nik Hwall (m) Hbumps (m) Hmax (m) 10, , , , , , , , , Table 5-4: Limits pressure loss in IGW model The Reynolds parameter is chosen larger than which indicates turbulent flow. Because the hydraulic conditions of Blasius equation (19) are met, this configuration is hydraulically smooth. The velocity is calculated out of the Reynolds parameter with the diameter of the model and the viscosity of water. Re 65* D / k (=3,466,667) (19) The flow rate is then calculated which is important for the flow rate measurement equipment that must be chosen further on. 2 Q 0.25* pi * D0 * U = (21) 29

42 FRICTION EXPERIMENT DESIGN The friction factor for wall friction, Cf, is calculated with that formula for straight pipe flow by Blasius or Nikuradse that corresponds with the Reynolds number: < Re < , Blasius -> C f = 0,316 (Re) (22) < Re < , Nikuradse -> C f = ,221 (Re) (23) Now that the wall friction is known the minimal pressure head can be calculated. This is the second part of the equation. The first part is the additional loss caused by de varying diameter. Adding this loss leads to the maximum pressure head loss that can be expected. For the range of Reynolds numbers that was determined to make the experiment worthwhile, 12,000 < Re < 100,000, a minimum pressure head of 1.83 meters must be available for the IGW model. To be able to conduct the experiment at higher velocities as well, flow measurement equipment which can measure up to 30 l/s must be used. The loss caused by the bumps is estimated very conservative. Thus it is probable that the velocities in reality are higher resulting in higher Reynolds numbers and thus a larger simulated range. 5.3 PRACTICAL EXECUTION With the requisites of the experiment an appropriate set up for the experiment has been designed. First an overview of this set-up and its components is given. Next there is a brief description of the measurement method and the equipment used for this. Finally a result of the construction of the IGW test models which influences the experiment execution is presented Set-up Based on advantages and disadvantages (see appendix B) of possible configurations, measurement methods and available equipment a set-up has been chosen which is described in this section. The set up is illustrated in Figure 5-4. Figure 5-4: Overview set-up The pressure is delivered by a reservoir that has a level of approximately 5 meters above the experiment floor. The flow rate is adjusted with a valve until the desired flow rate is obtained. This flow rate can be read with the flow rate measurement flange. For accurate measurement an area of 20 times the diameter before and 10 times the diameter after the flange is minimally necessary. The diameter that is used in the whole experiment set up is a standard diameter of 82.5mm. The first part of the IGW model serves as a region in which the flow stabilizes until a linear pressure drop is obtained. This region can maximally be 30 times the diameter. The next section is the linear region in which the pressure head measurements are done. This section has a minimum length of 1.6 meters, but it probably is larger. The last section of the IGW model serves to exclude the disturbances of the outflow. 30

43 FRICTION EXPERIMENT DESIGN The setup ends with steel piping with a slightly higher level for the outflow. This serves as a minimal pressure so that the model remains full of water and under pressure under all circumstances. In the experiment however this higher outflow level was attained by a gradual inclination of the outflow pip over a distance of approximately 15 meters, where the water flowed back into the laboratory system. The real set-up can be found in Figure 5-5. Figure 5-5: Experiment set-up in practice Measurement method In the experiment two parameters, the flow rate and the inclination of the total head level, need to be measured. The flow rate is measured with a flange with a design according to the ISO norm. The contraction in the flow generated by the plate (see Figure 5-6) causes a difference in pressure on both sides of the plate. This pressure differential is measured with the attached manometer. With the formula provided by ISO the flow rate can be calculated from this pressure differential. Figure 5-6: Measurement flange The other parameter, inclination of the total head line, can be found by measuring pressures on different points in the linear region. By taking another look at the Bernoulli equation (8), it can be seen that if U is constant the last term does not have to be included in the measurements to find the difference in H. 31

44 FRICTION EXPERIMENT DESIGN H = z + p ρg 2 U + 2g (8) U can only be the same in the different measuring points if the surface of the cross-cut is the same. By placing the points in the smallest diameters (82.5mm) of the IGW models an attempt is made to make the last term, the velocity head, constant. Now the pressure can be measured without the velocity head. This pressure is called the static pressure: p h = z + (9) ρg To measure the static pressure, holes are drilled in the pipe wall perpendicular to the flow direction as depicted in Figure of these points are used for each of the 3 set-ups. The points are connected with tubes to a tap block allowing measuring the pressure of the first 6 points relative to the last point by opening and closing the right taps. The two points between which the pressure differential is measured are connected to a manometer as used with the measurement flange. Figure 5-7: Measuring points for static pressure & tap block Test models The dimensions of the test models required for satisfactory experiment execution were given in Table 5-3. In this paragraph the requisites for the test model are discussed and the final results of the test model construction are compared to these requisites. For the construction the following requisites apply: The lining of the model must be smooth The pressure holes must not disturb the flow The pressure holes must have an arrangement for injecting dye The transition between the different segments must be smooth The surface area of the cross sections must be constant in the smallest parts The construction method used for the models is elaborated in Appendix B. The final result is 8 segments for each geometry. In Figure 5-8 each of the 3 geometries a segment is depicted. 32

45 FRICTION EXPERIMENT DESIGN Figure 5-8: Segments for model A (left), model B (middle) and model C (right) Looking at the construction results all requisites are met except the requisite of the constant surface area of the cross sections at the smallest diameters for model C (see Figure 5-9). Reasons for this are described in the appendix. The consequence is that the static measurements do not lead to the energy level since the velocity head is not the same for all pressure points due to variation in the cross sections. Thus a different measurement method is chosen for this model which is explained in 6.3. Figure 5-9: Unsatisfactory cross section dimension results for model C 5.4 CALIBRATION For proper measurement of the pressure drop, the experiment set-up first is calibrated. This is done by measuring the pressure drop with a straight pipe (see Figure 5-10). The large difference of approximately 30% between the measured values and the values found in literature indicated an error in the measurement. This error was found in the outcome of the measurement flange leading to the decision to calibrate the flange. All supporting data for this paragraph can be found in Appendix C. 33

46 FRICTION EXPERIMENT DESIGN Figure 5-10: Straight pipe set up Straight pipe check The experiment set up is checked with measurements on four points in a straight pipe. R squared, the relative predictive power of the model, was between 0,98 and 0,995 showing a strong correlation between the measured pressures (see Graph 5-4). calculation Cf static pressure relative to point 4 (cm water column) y = -8.94x y = -6x y = -4.44x y = -3.6x y = -2.54x position pressure points (m) Graph 5-4: 1st measurement straight pipe U = 2.48 U = 2.08 U = 1.76 U = 1.57 U = 1.26 Linear (U = 2.48) Linear (U = 2.08) Linear (U = 1.76) Linear (U = 1.57) Linear (U = 1.26) Linear (U = 2.48) However comparing the friction factor, Cf, which can be calculated from the pressure drop, to the data that can be generated with the formulas by Nikuradse a difference of around 40% is found which is unacceptable. Nikuradse: C f = ,221 (Re) (23) 34

47 FRICTION EXPERIMENT DESIGN Comparison Cf meaured data with Cf from literature U flange (m/s) Cf_Nikuradse Cf_data Deviation (%) Table 5-5: Deviation friction factor measurement with literature Therefore measurements were done to evaluate whether the flow rate measurements conducted with the measurement flange were correct. Two methods were used for this. First water velocities were measured in the axis with a mill and corrected to obtain the average velocity over the cross section. Secondly a volumetric measurement in time was done to obtain the flow rate. Both measurements showed a velocity that was 25 to 32 % larger than the velocity measured with the flange giving reason to believe the flange does not work properly. Now the velocities in the 1 st measurements in Graph 5-4 are corrected with 28 %, the average of the deviation. With these new velocities and the measured pressures the Cf s are calculated again from the measured pressures and with Nikuradse. The results are given in Table 5-6. Now the average difference is 9,3 %, an acceptable value for this type of experiment. Comparison Cf meaured data - Cf from literature for velocity corrected with 28% U flange * 1.28 (m/s) Cf_Nikuradse Cf_data Deviation (%) Table 5-6: Results Cf comparison for corrected velocities Flange calibration From the extra measurements of the flow rate it can be concluded that the measurement flange is not working properly. A rather remarkable conclusion, since according to the ISO norm the flange has the right dimensions and the condition of straight piping length upstream and downstream was met. Since especially the volumetric measurements were accurate and the error that the flange measured was reproducible, the decision was made to calibrate the flange with series of volumetric measurements. Since the flow rate did not exceed 30 litres per second, a flange with smaller opening (β = 0.7) was used for this calibration to obtain more accurate measurements. 35

48 FRICTION EXPERIMENT DESIGN new formula measurement flange Q (l/s) dh (cm water column) Q = (dH) R 2 = volumetric measurement Power (volumetric measurement) Graph 5-5: Calibration flange The new formula also contains the squared root of dh as the formula of the ISO norm. The difference is mainly in the multiplier which now is 2.41 instead of The R squared shows that the line is a very good fit to the data. The formula that is used in the experiment then becomes: Q = * dh (24) 36

49 EXPERIMENT RESULTS 6. EXPERIMENT RESULTS In this chapter the results of the experiment are presented for all 3 IGW models. For every model the pressure drop is plotted and the quality of the measurement is described. For model C an alternative measurement method is introduced due to an error in the construction. Then the turbulence is discussed, which can be quit strong, based on the visualisations done with dye and wires. Finally the results for every model conclude with a practical comparison to an effective diameter of a straight pipe and a relationship between Cfd and Re. All supporting data for this chapter can be found in Appendix D. 6.1 MODEL A This is the model that has the strongest wavy geometry. The measurements showed a high friction, high turbulence and rapid mixture of the fluids Measurements The measurements had a high linear correlation between the pressures with an R squared higher than 0,99 for all velocities (see Graph 6-1). Therefore there is no significant variation in the velocity head, making these values useful for interpretation of the energy level drop. pressure head relative to point 7 (cm water) pressure drop model A y = x y = x y = x y = x y = x U = 2.13 U = 1.94 U = 1.66 U = 1.40 U = 1.05 Lineair (U = 2.13) Lineair (U = 1.66) Lineair (U = 1.40) 0 Lineair (U = 1.05) Lineair (U = 1.94) position (cm) Graph 6-1: Static pressures for Model A Visualisation of the flow shows a very turbulent character. Pieces of nylon string with a length of 1 centimetre were placed on a ¼, ½, and ¾ of the bellow. On all spots the strings flipped in all directions in which they could (360 degrees). Injection of dye shows that the flow constantly hits back against the main flow direction over the entire length of the bellow in which it is injected. This is even the case when injection takes place on ¾ of the bellow length Conclusions Comparing the results of the dh/dx with a straight pipe for the same range of flow rates, a straight pipe with a diameter of 49.5 mm roughly produces the same pressure drops. This leads to a Deff / D_0 of 0.6. Next the Carnot losses per bump are calculated by subtracting the loss due to wall friction of the total loss and dividing this by the number of bumps. The results can be found in Graph 6-2. A linear relation fits the data very well and the line almost goes through the origin, which is a very good result since zero velocity should not produce a pressure drop. 37

50 EXPERIMENT RESULTS Carnot loss model A delta_hv (cm) U^2/2g (cm) y = x R 2 = model A Lineair (model A) This results in a velocity head multiplier of: Graph 6-2: Carnot loss model A 2 ζ (1 ε ) 2 = Then ζ, which mainly depends on α, becomes: ζ = (1 1.6 ) = For the geometry presented here it can be concluded that the friction is very high, because of the high turbulence. This turbulence also causes a very rapid mixture of the fluid over the entire cross section. 6.2 MODEL B This is the model that has the medium wavy geometry. The measurements showed substantial friction, substantial turbulence and rapid mixture of the fluids Measurements The measurements had a high linear correlation between the pressures with an R squared higher than 0,98 for all velocities (see Graph 6-3). Therefore there is no significant variation in the velocity head, making these values useful for interpretation of the energy level drop. pressure head relative to point 7 (cm water) pressure drop model B y = x y = x y = x y = x y = x y = x Position (x) Graph 6-3: Static pressures model B U = 2.94 U = 2.64 U = 2.35 U = 2.06 U = 1.69 U =

51 EXPERIMENT RESULTS Visualisation of the flow shows a quite turbulent character. Pieces of nylon string with a length of 1 centimetre were placed on a ¼, ½, and ¾ of the bellow. On all spots the strings flipped in all directions of the 180 degree window of the main flow direction. Injection of dye shows that the flow frequently hits back against the main flow direction over the entire length of the bellow in which it is injected. This is even the case when injection takes place on ¾ of the bellow length Conclusions Comparing the results of the dh/dx with a straight pipe for the same range of flow rates, a straight pipe with a diameter of 64 mm roughly produces the same pressure drops. This leads to a Deff / D_0 of Next the Carnot losses per bump are calculated by subtracting the loss due to wall friction from the total loss and dividing this by the number of bumps. The results can be found in Graph 6-4. A linear relation fits the data very well, but the line misses the origin, which can not be correct since zero velocity should not produce a pressure drop. delta_hv (cm) Carnot loss model B Hv = (U^2/2g) R 2 = model B 0.80 Linear (model B) u^2/2g (cm) Graph 6-4: Carnot loss model B However the inclination of the line is used for calculation of the velocity head multiplier: 2 ζ (1 ε ) 2 = Then ζ, which mainly depends on α, becomes: ζ = (1 1.3 ) = For the geometry presented here it can be concluded that the friction is rather high, because of the turbulence. This turbulence also causes a pretty rapid mixture of the fluid over the entire cross section. 6.3 MODEL C This is the model that has the small wavy geometry. The measurements showed friction almost similar to a straight pipe, no substantial turbulence and normal mixture of the fluids. 39

52 EXPERIMENT RESULTS Measurements The measurements had a weak correlation between the pressures with an R squared around 0.74 (see Graph 6-5). Therefore there is a significant variation in the velocity head, making these values unacceptable for interpretation of the energy level drop. static pressures model C pressure head relative to point 7 (cm water) y = x R 2 = y = x R 2 = position (cm) Graph 6-5: Static pressures for model C meas 2 (U=3.49) meas 3 (U=1.74) Lineair (meas 2 (U=3.49)) Lineair (meas 3 (U=1.74)) The reason for this low quality data was the variation in the surface areas of the cross sections at the pressure points as was indicated in paragraph Therefore instead of measuring static pressures, the total head is measured for model C: H 2 p U = + (8) ρ g 2g These measurements (Graph 6-6) include the velocity head (the last member of the equation) and thus lead directly to the total head line of which we want to know the inclination (dh/dx). The practical execution of these measurements can also be found in Appendix D. The results of the measurements show a strong correlation with an R squared of over 0.98 for all velocities. dynamic pressure model C 50 pressure head relative to point 4 (cm water) y = x y = x y = x y = x y = x y = x U = 3,47 U = 3,16 U = 2,78 U =2,41 U = 1,86 U = 1,24 Lineair (U = 3,47) Lineair (U = 3,16) Lineair (U = 2,78) Lineair (U =2,41) Lineair (U = 1,86) Lineair (U = 1,24) -10 position (cm) Graph 6-6: Dynamic pressures for q =

53 EXPERIMENT RESULTS Visualisation of the flow shows low turbulence. Pieces of nylon string with a length of 1 centimetre were placed on a ¼, ½, and ¾ of the bellow. On all spots the strings flipped in a 40 degree window of the main flow direction. Injection of dye shows that the flow does not hit back against the main flow direction. This is also the case when injection takes place on ½ of the bellow length proving the absence of vortices in the bellows Conclusions Comparing the results of the dh/dx from Graph 6-6 with a straight pipe for the same range of flow rates, a straight pipe with a diameter of 80.2 mm roughly produces the same pressure drops. Due to an error in the construction D_0 varies between 82 and 85 mm giving an average D_0 of 83.5mm. This leads to a Deff / D_0 of Next the Carnot losses per bump are calculated by subtracting the loss due to wall friction of the total loss and dividing this by the number of bumps. The results can be found in Graph 6-7. A linear relation fits the data very well, but the line misses the origin, which can not be correct since zero velocity should not produce a pressure drop. delta_hv (cm) Carnot loss model C Hv = (U^2/2g) R 2 = model C Linear (model C) U^2/2g (cm) Graph 6-7: Carnot loss model C However the inclination of the line is used for calculation of the velocity head multiplier: 2 ζ (1 ε ) 2 = Then ζ, which mainly depends on α, becomes: ζ = (1 1.1 ) = For the geometry presented here it can be concluded that the friction is almost the same as a straight pipe. There also is a normal mixture of the fluid over the entire cross section. It has to be noted that the measurements for model C are less reliable, because due to the construction problems the test model did not have the exact dimensions it should have had. Instead of circular shaped cross sections, the cross sections were elliptical because of internal stresses that were introduced in the materials during the vacuum moulding process. 6.4 COMPARISON IGW MODELS With all the results of the 3 models, an overview is given of the data. The comparison of the IGW models to straight piping gives practical input for the next chapters in which the IGW offloading hose is compared to substitute products (see Table 6-1). It is clear that model A leads to large diameters to obtain an acceptable effective diameter. 41

54 EXPERIMENT RESULTS Deff / D_0 Model A 0.59 Model B 0.77 Model C 0.97 Table 6-1: Comparison to straight piping For each model the corresponding inclination coefficient, ζ, has been calculated. Comparison of these coefficients to those of a single expansion makes clear that the loss in one bump is much smaller in the IGW pipe (see Table 6-2). This difference is roughly 90% for the models A and B. However for model C the difference is 70%. This indicates that for smaller angles the gap between the inclination coefficients for the single expansion and the repeated expansion of IGW becomes smaller. IGW model Single expansion model α ε ζ (1- ε^2)^2 ζ(1- ε^2)^2 ζ A B C Table 6-2: Comparison inclination coefficients A plot of the ζ s for the IGW model and the single expansion leads to Graph 6-8. As was explained in paragraph 5.1.2, ζ in the case of a single expansion principally depends on α. Assuming that the ζ in the IGW model fully depends on α leads to the relation drawn in the graph. This assumption and the scarceness of data points mean that care must be taken with the interpretation of this data. zeta IGW bump vs single expansion zeta_igw = alpha alpha (degrees) R 2 = IGW single expansion Macht (IGW) Graph 6-8: Relation between inclination coefficients ζ = α (25) The scarceness of data also leads to a high unreliability of extrapolation of the results. However a rough estimate is illustrated with the sole purpose of indicating the expectation that ζ will become smaller for high values of α (see Graph 6-9). The reason for this is that with a very high angle of the pipe wall the flow eddies will not penetrate completely into the bellows, because the contractions follow each other up rapidly. 42

55 EXPERIMENT RESULTS Thus from a certain value an increase of α leads to a decrease of the loss per bump and thus to a decrease of ζ. However the relation can not be drawn that the loss per meter pipe is decreased by this phenomenon since a higher α leads to an increase of the number of bumps per meter. Extrapolation results IGW zeta alpha (degrees) IGW results Power (IGW results) Graph 6-9: Expected behaviour for higher alpha For further experimenting it would be interesting to test if variation of ε while keeping α constant has an influence on ζ. This will indicate whether the method used here with a ζ based on α is accurate. Therefore it is recommended to test 7 models between the models A and B as listed in Table 6-3. This will provide the minimum data to establish relationships for the independent parameters α and ε. The choice for the extra geometries is explained in the next chapter. Necessary IGW model geometries = already experimented α = 5.71 α = 10.3 α = 15.6 ε = 1.1 ε = 1.14 ε = 1.3 Table 6-3: Recommended models for further testing Finally it must be noted that dynamic interaction between the flexible pipe wall and the fluid is not taken into account since the models were constructed rigidly. For accurate measurements of the friction loss it is desirable to use models with the correct elasticity since it is plausible that the pipe wall takes on a configuration that is more beneficial in terms of friction losses. Another important aspect that can be experimented with these elastic models is possible flow induced vibrations. 43

56 EXPERIMENT RESULTS 44

57 HOSE DESIGN 7. HOSE DESIGN The geometry chosen for the design has important structural and hydraulic implications. The IGW hose is compared to the straight hose. The effective diameter has to equal the internal diameter of the straight hose. This comparison shows a large increment of the weight and costs of the hose for an increment of ε and α. Hence ε=1.14 and α=10.3 were chosen for the final design, because this is a cost competitive design which is expected to be sufficiently flexible. The labour and material costs for an IGW hose amount to euros and the development and new equipment acquisition amounts to 10,5 million euros. Finally the impact of IGW on the use of offloading hoses shows great benefits. Supporting data for this chapter can be found in Appendix E. 7.1 STRUCTURAL & HYDRAULIC INPUT The design consists of five layers and floatation material. These are discussed here starting with the 1 st layer on the inside (see Figure 7-1). An inner gas and oil tight layer of NBR (rubber) is used similar to the liner applied in conventional offloading hoses. The 2 nd layer consists of Aramide fibres applied according to the IGW principle and serves as the 1 st carcass. These fibres take up the internal pressure and axial loadings. The 3 rd layer is a second oil and gas resistant layer of rubber which contains the gasses and fluids in the event of rupture of the inner liner. The 4 th layer consists of Dyneema fibres applied according to the IGW principle and serves as a 2 nd carcass which takes up the internal pressures and axial loadings in the event of rupture of the first carcass. The Dyneema fibres also function as an impact absorbing layer for accidental collisions. The 5 th and outer layer is made out of EPDM, a rubber, and protects the construction from intrusion of the sea water and UV rays. It also is abrasion resistant to protect the inner layers. The position of the floatation material depends on a design fabrication consideration presented in 7.2. Aramide NBR Oil and gas resistant layer Dyneema (second carcass) EPDM Figure 7-1: Composition IGW hose From a structural performance point of view a higher ε and α is desired because of the increased flexibility. However a larger ε and α lead to a higher weight of the hose and increased costs. Also the effective diameter is decreased. The optimization taking into account the hydraulic and cost considerations takes place in the next paragraphs. Regarding the hydraulic input an important consideration is if a flexible design can be attained without affecting the equivalent diameter of the hose too much. A possible solution is to apply a straight liner. Then the equivalent diameter equals the inside diameter of the hose, because there are no constrictions. Since the offloading hose will require buoyancy material to keep afloat this can be applied between the liner and the IGW fibre layers. However this will be a costly solution, because the required characteristics for this material are extraordinary. It must be sufficiently flexible (to maintain IGW s benefits) and sufficiently incompressible (too maintain a low density and conduct the flow). 45

58 HOSE DESIGN Figure 7-2: Application of a straight liner Because of these complications the choice is made not to use a straight liner. Then the hydraulic design consideration is whether or not an increased friction is acceptable. The design parameter for the pumps and piping on board is determined by the pressure surge caused by an accidental closure of the valves. The size of the pressure surge depends on the working pressure which depends on the friction that must be overcome in the system to attain a certain flow rate. If this increased friction is allowed, the pumps and piping on board of the tankers and FPSO s must be improved to work with the higher pressures required to maintain the same flow rate. Such adaptations would have to be initiated through new design specifications which take a long time, i.e. 10 years, to implement. Thus it takes a long time before the market really exists to which such a product can be sold making the development impossible. Hence the hose will be designed with an effective diameter equal to the straight hose that it must replace so that the IGW hose can be used in all tankers and FPSO s without modifications. 7.2 FINAL DESIGN & FABRICATION For the final design three different geometries are compared. Because of the extreme reduction of the equivalent diameter for model A, this model is left out of consideration. A new model, model Z with ε=1.14 and α=10.3, has been chosen between the two other models of which data exists. The latest outcome of the design study showed that the fibre configuration is better if the design is made with a smaller axial load. This results in a large increment of α and a small increment of ε with respect to model C. 0.6 meter is an offloading hose diameter that is often sold. Therefore a comparison is made between IGW and the conventional type of hoses for this diameter. Because 7 m/s is a standard offloading rate which is used in a design, the pressure drop is calculated for this flow rate for an offloading hose with a diameter of 0.6 meters. This results in a dh/dl of The formulas established in the previous chapters are used to determine the effective diameter. The supporting data can be found in Appendix E. 2 L U tanα 2 2 H = ( Cf + ζ (1 ε ) ) (18) & D0 2g ( ε 1) ζ = α (25) Iteration resulted in a D_0 of 0.67m for model Z to obtain an effective diameter of 0.6m. First in a program designed in Matlab the composition of the fibre and rubber structure has been calculated. With this data the weight and internal and external volume of the hose has been calculated (see results in Table 7-1). These weights together with the volumes of the hose are used as input to do a buoyancy calculation. According to the OCIMF specifications, the hose must have 20 % buoyancy reserve when it is completely filled with water to avoid accidental sinking of the hose string. The amount of floatation material using closed cell foam polyurethane is calculated from this spec. In the table the weights are given empty and filled with sea water for the IGW hose and the hose of the most innovative competitor, Trelleborg. 46

59 HOSE DESIGN Comparison weight empty and weight filled IGW hose Trelleborg hose (D=0.6m) empty, no empty with filled with material empty with filled with floatation floatation seawater costs floatation seawater (kg) (kg) (kg) ( ) (kg) (kg) model B (D_0=0.77) model Z (D_0=0.67) model C (D_0=0.62) Table 7-1: IGW geometries with an effective diameter of 600 mm From these figures it can be concluded that the IGW hose weighs less empty than the Trelleborg hose. However the internal volume of the hose is large because of the large internal diameter to minimize the local pressure losses. This large internal volume gives a large additional weight if the hose is filled with sea water. As can be seen in the table, model Z more or less has the same weight as the Trelleborg hose when filled with sea water. Another important aspect is that the material costs increase rapidly with an increase of ε or α. Finally model Z is chosen for further feasibility analysis, because it is expected to score better than model C on the flexibility and still has the potential to offer a cost competitive solution to the existing offloading hose. Another reason for this choice is that the lifting equipment does not have to be adapted to handle the filled hose. Future test results will have to verify whether or not the assumption of sufficient flexibility is true for model Z. Fabrication is a process of constructing layer by layer and with two curing sessions in an autoclave. An inflatable mandrel or a jig saw mandrel is used that has the outer dimensions of the inner liner of the IGW hose. Over this mandrel the NBR is applied in strips. Then the first fibre layer of Aramide is wound around the structure using the braiding technique with which multiple fibre bundles are wound at the same time as in Figure 7-3. Then the hose is cured for the first time to prevent the first two layers from bonding completely with the other layers. By doing this a structure with two separate carcasses is created between which oil can run in the case of a rupture of the 1 st carcass towards the leak detection system localized near the flange. Figure 7-3: Braiding examples Next the second layer of NBR is applied and the Dyneema fibres are wound around the structure. Subsequently sufficient floatation material is applied to keep the hose string afloat. Finally the EPDM layer is constructed and the hose is cured in the autoclave. Afterwards the mandrel is deflated and extracted if the inflatable mandrel is used. Otherwise the jig saw puzzle is extracted piece for piece. 7.3 COSTS The costs that are discussed here only consist of the research and development costs and the material [10] and labour costs for the manufacturing of a hose. These costs amongst others are used in the next chapter to do a cost analysis for the manufacturer and the customer. A distinction is made between constant and variable costs. 47

60 HOSE DESIGN Cost breakdown for IGW hose Fixed costs euros Variable costs (per hose) euros R&D hose and flange Labour Production design Material hose Certification flanges Braiding equipment acquisition Total Total Table 7-2: IGW hose cost breakdown The fixed costs for the development are a little over 10 million euros. Normally the costs of development of an offloading hose are around 5 million. Since part of the development here also concerns the development and construction of a braiding machine, the costs are a bit higher. The total costs for the construction of the hose (development costs, factory rent and sales costs excluded) amount up to 60 % of the sale price of the conventional hose. In the next chapter the feasibility for implementation in the production line of an existing hose manufacturer is treated. 7.4 IMPACT IGW DESIGN IN USE The impact that implementation of IGW in the offloading hose has on the complete design and exploitation of an offloading system is significant. All changes are caused by the increased flexibility, lower weight and varying diameter. The main changes in use are determined in this paragraph and these serve as input for the customer business case in the next chapter. The guidelines for offloading facility design from OCIMF are used for this paragraph [11,12,13]. Influence varying internal diameter on pumps and piping By taking a larger internal diameter in the design the varying diameter does not generate an extra pressure drop when compared to the nominal diameter for which it is designed. All piping and pumps on the tankers and FPSO s thus can remain unchanged. However a reducer is necessary to fit the hose to the tanker manifold. influence of lower weight on equipment The lower weight of the hose, 1371 kg instead of 2500 kg (Kleline by Trelleborg), creates easier handling. However this only offers benefits for the transport and installation, because the hose weighs the same under operation since then it is filled with oil or water. This is caused by the larger internal volume of the IGW hose. Influence of increased flexibility on equipment With the increased flexibility the hose can be reeled on small diameter drums (diameter drum < 4 times the diameter of the hose). Another advantage is that the hose s increased flexibility is likely to absorb the pressure surge in case of an emergency stop. This leads a higher design velocity because the pressure surge is reduced in the hose string reducing the risk of damage to piping and valves on the other side. Testing will have to point out whether this is true. Influence of weight and flexibility on safety With the handling of the heavy conventional offloading hoses, accidents occur often. IGW s lighter and more flexible design leads to safer handling of the hose when it is empty. This offers advantages for the safety during transport and installation. Influence of varying internal diameter on testing and maintenance After use the offloading hose is flushed with water. A problem that might occur with the varying diameter is that the concave areas fill up with wax, due to lower flush velocities. Whether or not this is the case should be tested in a laboratory. Another aspect is that the hose before use and sometimes during its lifetime must be vacuum tested to see if any delaminating occurs. This visual inspection is difficult due to the internally varying diameter. 48

61 FEASIBILITY ANALYSIS 8. FEASIBILITY ANALYSIS For the feasibility of implementation of a new technology in the production process two aspects are important. Adaptation of the technology must be financially attractive for the manufacturer and the new product must be interesting for the customer. In a market with annual sales volumes of 388 million adaptation of the technology by the manufacturer shows large revenues with 30% profit margins. The customer gains from the technology through a rough euro savings per offloading project. However interesting these figures look, the market can only be penetrated by cooperation with a word renowned oil company due to the conservative attitude of the offshore industry. 8.1 MANUFACTURER COST ANALYSIS The manufacturer cost analysis serves to show the financial attractiveness of production with IGW. A case in which the assumption is made that adoption of the IGW technology lets a company X increase her market share from 15 to 25 % shows production potential Market description The manufacturers are large rubber producing corporations. Many of them work together or have co developed with large oil companies setting the industry standards. Looking at the clients the situation is a bit more complex, because there actually are two parties who decide whether or not to buy a product. The system engineer that engineers the offloading system and the end user that exploits the offloading system form this decision making unit together. However the role they play is quite different. The engineering company assembles the system and searches appropriate hoses. Hose manufacturers do quite some direct marketing to these system engineers, because if their hose is chosen in the design phase there is a good chance that it will be really used. However the end user, for whom the system is designed, always can cancel the choice. Therefore hose manufacturers tend to live up to the standards that these end users, the oil companies, make up. It is a very important factor to have credibility with the end user, since the offloading hose is used in the core process of the offloading facility, i.e. the offloading of crude oil. Failure of such a key element costs the end user a fortune. Offloading market players Hose manufacturers System engineers End users Bridgestone Industrial SBM Shell Trelleborg Oil & Marine Modec BP Technip Saipem Exxon Mobil Goodyear Brown & Root Total Dunlop Oil & Marine Pusnes Repsol Parker ITR Bluewater Statoil DANTEC Intec Engineering Aral Edelbuttel + Schneider Table 8-1: Offloading market players To estimate the size of the market for the product lines that were chosen in Chapter Fout! Verwijzingsbron niet gevonden. the supply record of Bridgestone [14] and market prospects of SBM are used. The price of a hose nowadays is between 20,000 and 50,000 euros depending on the characteristics (single carcass, double carcass, submarine, diameter, floatation). The most important parameters for the price are the diameter and the number of carcasses. Since in most cases the large diameters and double carcasses are used, the average price of a hose is estimated to be 40,000 euros. 49

62 FEASIBILITY ANALYSIS Bridgestone has a 35 % worldwide market share in these hoses in From 1972 until 1998 they sold 21,267 hoses. Assuming that their market share was 35 % over that period, the average worldwide sales amount 21,267 / (27*0.35) = 2250 hoses per year. The assumption is made that this average is valid for The market at that time can be divided into near shore offloading facilities (70%) and (ultra)deepwater exploration (30%). Based on market prospects of SBM the number of (ultra)deepwater fields that will be explored until 2012 has an annual growth of 28%. Every new field that is developed uses an offloading system. Thus this growth rate can be used for the offloading hose market for (ultra) deepwater exploration. The growth rate of the other market of near shore offloading is based on worldwide port annual expansion (5%) and on the worldwide oil consumption (4% annual growth). This leads to a growth rate of 5% of this market. The average annual growth rate for the total market then becomes 0.3*28% + 0.7*5% = 11.9%. With this growth annual sales amount 9,700 hoses in is first year that the IGW hose development can be finished. The total sales in euros then are around 388 million per year in Business development Here an analysis is made whether or not the IGW technology applied in offloading hoses creates an opportunity for a hose manufacturer to expand its business by adopting the technology. To measure the feasibility of such a business development the revenues and the profit margins are analysed. The base assumption of the analysis is that a medium sized company, company X, has a market share of 15% and can expand her market share up to 25% by using IGW technology in her products due to various advantages for her customers which are defined in the next paragraph. Another assumption is that all existing equipment which can still be used in the production and testing is already depreciated and therefore does not have to be taken into account in the cost calculation. This can be done because production methods in the last decades have changed very little. Note that depreciation of money is not taken into account and that the market from 2011 onwards has a constant size. costs per hose: euros, total market size: 9700 all other figures in millions market share # of hoses sold revenues sales hoses costs R&D + expansion 20 Fixed costs marketing fabrication hoses profit revenues - costs Table 8-2: Profit and loss implementation IGW Revenues are calculated with an average price of the hose of 40,000 euros. The R&D costs and acquisition of braiding equipment are 10.5 million. Another 9.5 million euros are reserved for the necessary expansion of existing equipment because of the production which will take place on a larger scale. The constant costs per year consist of maintenance (1M), rent of production hall (1M) and organization (4M). In the first three years an additional 3 million per year is reserved for marketing of the introduction of the product. The variable costs depend on the number of hoses manufactured and these costs include the distribution. 50

63 FEASIBILITY ANALYSIS The revenues show a large increase and the profit is only negative in the 1 st year leading to a cash flow deficit of 7 million euros. This deficit will have to be paid out of profits of the preceding years. After the first years the profit goes up to 30% of the revenue making it an attractive technology to implement. Main reasons for the high profitability are the reduction of labour due to the braiding production technique and the scale benefits of a larger production. 8.2 CUSTOMER BUSINESS CASE A hose manufacturer will not adopt the IGW technology if it is not convinced her customers will want to buy the product. Therefore the customer business case is analysed here in which safety, environmental and cost issues are addressed. This shows vast advantages for the customer and also forms the basis for the assumption used in the previous paragraph that a 10% extra market share is gained by adopting the IGW technology Safety, environment & costs In the past a lot of accidents happened in the offshore industry and nowadays still the risk of working in the offshore industry is high compared to other industries. Another issue that is very important is the environmental hazards that exist in the offshore exploitation of oil and gas. Present attitude of the oil companies is that they are willing to invest a lot of money to lower the risks of accidents which involve oil spills or injuries. Therefore it is important that the use of IGW can address these issues. The lower weight of the hose 1530 instead of 2360 kg improves the personal safety of the workers due to easier handling. The use of IGW in combination with Dyneema fibres provides a good impact resistance, which reduces the risk of oil spillage in the event of an accidental impact with for example a tug boat. IGW thus provides safety and environmental benefits without expenses. In the previous paragraph an average price of the conventional hose was assumed to be 40,000 euros. The calculation of the profitability of the hose that uses IGW was done with the same price. However looking at the profit margins shows that there also is enough budget to adopt a different pricing strategy if the market expansion targets are not met. By lowering the price the expansion can then be realized. To conclude the purchasing costs are the same or can even be lower than the conventional hose Life time cycle costs With the purchasing cost comparison the life time cycle cost calculation can be made which shows the benefit for the customer which is an important criterion for the feasibility of adoption of IGW technology for the manufacturer. Interviews with superintendents from SBM learned that the average lifetime of a hose segment in the main line is between 5 and 12 years and the average lifetime of a tanker rail hose (last two in the line attached to the tanker manifold) is between 1 and 3 years. It is expected that the IGW hose has a longer lifetime, due to its ability to bend through geometrical change of the construction instead of elastic change of the materials leading to a high fatigue resistance. Therefore the average lifetime of the conventional hose is assumed to be 10 years and the average lifetime of an IGW hose 20 years for the hose segments in the mainline. The tanker rail hoses are assumed to have an average lifetime of 2 years for the conventional hose as well as the IGW hose since the failure of these hose segments is caused by mechanical damage (wear and tear over the tanker rail and the tanker). The normal lifetime for which an offloading facility is designed is 20 years and this time span will be used as input here. 51

64 FEASIBILITY ANALYSIS Life time cycle cost calculation corrected to the first year IGW costs costs base year conventional costs costs base year 10x2 tanker rail hoses 30 mainline segments x4 tanker rail hoses 30 mainline segments 30 segments after 10y total total Table 8-3: Life time cycle costs The tanker rail hoses will have to be replaced with the same frequency. The first acquisition of the mainline is the same in both cases, but the mainline of the conventional system needs to be replaced after 10 years. An effective annual rent of 5% is used. The effective rent correction for savings and expenses (tanker rail hose purchases) that run through the whole system lifetime is: effectiver ent = n= 1to n = 73% The effective rent correction for expenses done after 10 years (mainline purchase for replacement) is: effectiver ent = = 63% After correction of all costs and savings to the base year it becomes clear that IGW offers the customer a cost reduction of euros in the system design. It can thus be concluded that the manufacturer financially has a strong business case with which it can convince its clients to buy the product. 8.3 MARKET ENTRANCE BARRIERS Having a financially fit business case and a feasible adaptation in a production line in combination with some safety and environmental benefits often is enough launch a technology. However in this specific market two important factors may block the entrance to the market. The first factor is a psychological one. The attitude in the offshore industry towards the products they use is if it works it works. If it doesn t work, but I know how to replace it in time, it still works for me. This is a very important aspect, because a new technology that offers cost reduction often cannot replace an existing technology even if it is a bad one. As long as they know how to work with it, the incentive to change the methods is low. This is especially the case for a radical new technology with no track record such as Integral Geodesic Winding. The second factor is that it concerns a new technology which might work better, but might not comply with the current specification of OCIMF which are designed for conventional production techniques. The solution with which both market entrance barriers can be overcome is to attract a large oil company in the development and certification process. They can guarantee a certain part of the market, they have the negotiation power to change the standards and they give the product the credibility it needs to convince the rest of the market to buy it. 52

65 CONCLUSIONS & RECOMMENDATIONS 9. CONCLUSIONS & RECOMMENDATIONS This report is intended to evaluate the feasibility of the implementation of IGW technology in offloading hoses. In this chapter the conclusions and recommendations of the research are presented. 9.1 CONCLUSIONS The conclusions are split up into the general conclusion which answers the main research question and the partial conclusions on which the general conclusion relies. General conclusion The general conclusion is that this implementation is feasible for small ratio (1,14) of the maximum and minimum diameter with an inclination of the pipe wall of 10.3 degrees. The manufacturer of offloading hoses can increase the revenues and obtain a profit margin of 30% by using the technology. His customers profit from a cost reduction of euros over a period of 20 years, the average lifetime of an offloading project, through a decrease of hose failures. Besides these cost savings, there are Health, Safety and Environmental benefits for the end user. The assumption was made that this geometry offers sufficient flexibility to be advantageous in use and to have a high resistance against fatigue. Experiments on IGW prototypes in August 2006 will have to verify this. Partial conclusions The design of the IGW hose is a trade off between hydraulic and structural properties. The larger the variations in diameter become the more beneficial the flexibility is expected to be. However this also results in a larger friction for the flow. Three models (A, B and C) were designed for laboratory testing to obtain data on friction. The geometries for these models range from the structurally most beneficial (A) to the hydraulically most beneficial (C). The measured pressure drop was split into two different contributions, the drop caused by wall friction and the drop caused by the varying diameter (a Carnot type of loss), to establish a formula for the total drop. In the literature experimental data exists for the Carnot losses of single expansions. The geometry of the pipe can be considered as a repetition of single expansions. However the experiment showed that this repetition induces a strong reduction of the pressure drop. The values of the pressure drop per bump in the IGW pipe are roughly 90% smaller than the values for a single expansion. The total pressure drop, H, depends on the diameter ratio of the pipe, ε, the inclination of the pipe wall, α, the minimum diameter, D 0, The length of the model, L, the average velocity in the minimum cross section, U, and the gravity acceleration, g. This formula is valid for 1.1< ε<1.6 and 5.71< α<26.1: 2 L U tanα 2 2 H = ( Cf + ζ (1 ε ) ) and D0 2g ( ε 1) ζ = α Also a more practical comparison was made that served as input for the final design of the hose. The design parameter for piping and pumps on the ship are determined by the pressure drop and the pressure surge caused by an accidental closure of a valve. Usually a design velocity of 7 meters per second is used and therefore the pressure drop was calculated with this velocity in the IGW pipe. Then the diameter of the straight pipe, the effective diameter, that would have produced the same results was calculated. The ratio of the effective diameter, D eff, divided by the minimum diameter in the IGW pipe, D 0, produced the following results: Model A: D eff / D 0 = 0.59 Model B: D eff / D 0 = 0.77 Model C: D eff / D 0 =

66 CONCLUSIONS & RECOMMENDATIONS Analysis of the market psychology indicated that a product that needs adaptation in the piping and pump design is hard to sell. Hence it can be concluded that the IGW hose has to have such dimensions that it produces the same pressure drop as the existing hoses. A low D eff / D 0 ratio results in a relatively large design of the hose. The offloading hose has to have a 20% buoyancy reserve when filled with water. The larger the internal volume, the more floatation material has to be applied and the larger the costs become. Therefore a geometry in between the models B and C was chosen that could still be cost competitive. The assumption was made that this geometry offers sufficient flexibility to be advantageous in use and to have a high resistance against fatigue. Experiments on IGW prototypes will have to verify this. This design with 1400 kilograms weighed 1100 kilograms less than the existing hoses which results in increased safety because of easier handling. The increased flexibility leads to a higher resistance against fatigue which leads to euro cost reduction per offloading project for the end user if the product is priced equal to the existing products. Because of the good value proposition towards the end user, it is assumed that a manufacturer can increase his market share from 15% to 25%. This increase in market share in combination with the low switching costs and low labour costs with the use of IGW creates profit margins of 30% for the manufacturer. Because IGW offers advantages for the end user as well as for the hose manufacturer its implementation in offloading hoses is found to be feasible when the data on the flexibility and resistance against fatigue for model Z is verified. 9.2 RECOMMENDATIONS The recommendations of the research are split up into two categories. First the recommendations for further testing are presented. Then it is suggested to investigate a different design where the floatation material is placed on the inside of the hose. Further testing Three aspects require further testing. These are presented here in order of their relevance starting with the most important tests. The first aspect that needs testing is the flexibility of model Z which was used for the basis of design. The prototypes, which are finished and tested in August 2006, have dimensions that come closest to the dimensions of model A. The results of these tests serve as input for the final design and construction of model Z. If the possibility is there, it is recommended to also test model C. Secondly more models need to be tested to obtain data on the friction that the varying diameter generates. Because this friction depends on two parameters, α and ε, 9 models are necessary to establish relations with a high level of accuracy. Leaving model A out of the equation, because of its extreme friction, results in the need to construct and test 7 more models in the range between B and C. Finally the bellows in the IGW hose are at risk to be filled up with wax with certain types of crude oil. Therefore testing should investigate this risk and find out whether this is harmful or provides advantages due to better streamlining. Different design It is recommended to investigate if a suitable material, which resists high pressures and remains flexible, can be found to serve as floatation material. Placement of this floatation material in the bellows of the IGW hose leads to a lighter design due to two reasons. First of all the flow is streamlined due to this construction which leads to a better ratio of the effective diameter and the minimum diameter of the hose. Hence the hose can be designed with a smaller diameter while the pressure drop remains the same. Secondly the internal volume is decreased by placing the floatation material on the inside which results in a smaller amount of floatation material through the 20 % buoyancy reserve requirement. 54

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69 REFERENCES References [1] Pressurisable structures comprising different surface sections S. Koussios, O.K. Bergsma, A. Beukers 2002 [2] Design offloading hose with IGW technology T. Mousawi To be published in October 2006 [3] Articulated pressurisable structures: design & applications S. Koussios, O.K. Bergsma, A. Beukers 2005 [4] Filament winding: a unified approach S. Koussios 2004 [5] SPM hose system design commentary Oil Companies International Marine Forum 1993 [6] Guide to purchasing, manufacturing, and testing of loading and discharge hoses for offshore moorings Oil Companies International Marine Forum 1991 [7] Collegehandleiding vloeistofmechanica J.A. Battjes 2001 [8] Polytechnisch zakboek Koninklijke PBNA 1997 [9] Offshore and onshore installations ITR Spa 1998 [10] Material property data March 2006 [11] Guidelines for the handling, storage, inspection and testing of hoses in the field Oil Companies International Marine Forum 1995 [12] Recommendations for the equipment employed in the mooring of ships at single point moorings Oil Companies International Marine Forum 1993 [13] Single point mooring maintenance and operations guide Oil Companies International Marine Forum 1995 [14] Offshore hose manual Bridgestone

70 REFERENCES 58

71 APPENDIX A Appendix A: Experiment configuration In this appendix the comparison is made between the different experiment configurations and measurement methods. This leads to the choice for the experiment set up and measurement method presented in paragraph 5.3. Possible set ups Here possible configurations for the experiment are discussed. Three concepts are compared. The first is to use two reservoirs with different levels of the water surface. The second concept is to use a pump to add the pressure head. The third concept is to connect the model to laboratory piping. Finally all concepts are compared on advantages and disadvantages. The total pressure head loss over the entire set-up is: 2 2 U U L H = ( ξ in + ξout) * + Cfd * * (27) 2g 2g D Set up with two reservoirs In the set up with two reservoirs (see Figure A 1) a constant level of the water surface needs to be maintained if one desires to measure the loss over the whole model, because then the total pressure drop is known. After the second reservoir another reservoir can be placed to measure the flow rate. In the line that represents the energy level two steep sections can be found. These are the local losses at the inlet and the outlet. The rest of the line represents the pressure head loss due to the wall friction and the varying diameter. Figure A 1: Set up with two reservoirs Advantage: The difference in pressure level over the system is easily found. Disadvantages: Local losses at the inlet and outlet can become as large as the loss due to the varying diameter calculated if no special precautions are taken. This would result in a total pressure loss of a few meters making the construction of the model very difficult. Without these local losses the difference still is quite large. For variation of the added pressure head a reservoir with an adjustable level needs to be constructed as depicted in Figure A 2. This construction is quite complex. 59

72 APPENDIX A Figure A 2: Adjustable inlet Set up with pump In the set up with the pump the energy level at the start of the test object is the capacity of the pump minus the local inlet loss. Again the level at the outlet reservoir is maintained constant. The pressure head loss due to the wall friction and the varying diameter is the difference between the highest point of the energy level and the level of the outlet reservoir minus the local outlet loss. Advantage: The set-up is relatively easy to construct. Figure A 3: Set up with pump Disadvantages: A large inaccuracy exists in the pressure that is added to the system making it hard to find accurate values for the sought friction coefficient if one desires to measure the pressure drop over the entire set up. A relatively complex construction is needed at the inlet to be able to generate different velocities. Set up connected to piping In this set up (see Figure A 4) the pressure is delivered by the piping system of the laboratory. To regulate the flow a valve is used. This makes it necessary to apply manometers to measure pressure differences. These pressure differences can be measured inside the IGW model or over the whole model. The flow rate can be measured with a measurement flange since the local loss this causes does not form a problem, because it is outside the pressure measurement region. 60

73 APPENDIX A Figure A 4: Set up connected to piping laboratory Advantages: The method for the flow rate measurement is easy. The flow is easy to regulate. Disadvantages: A lot of local losses exist in this set up, decreasing the available pressure over the IGW model. Measurement methods The different configurations and their energy losses have been discussed. Besides choosing a setup the choice that must be made is whether the whole setup or part of the setup will be used to measure the loss of pressure head. As discussed in the previous paragraph the contributions of the local in and outlet losses can be quite large in the total loss. Because these contributions are not part of the pressure head loss that is sought, they make the measurement less accurate. However if one measures the pressures in the section of the test object in which the flow is stabilized these inaccuracies are not included. Figure A 5: Measuring inside the IGW model In the previous paragraph the pressure head loss was schematized as the red line in figure A5. However the influence of the local losses reaches further than that schematization. A gradual transition is what happens in reality. This is illustrated with the blue line. The part where the measurements need to be done to find Cfd is the linear part of the dashed blue line. It is expected that the inlet losses or flow stabilizations do not extend further than 2.55 meter (30 times the nominal diameter) and the outlet losses no further than 0.85 meter (10 times the nominal diameter). Thus in the region between these two sections the pressure loss is probably linear. In a total length of 5 meters this linear area comprises 1.6 meter. By measuring on three points within this section, the arched region, this statement can be verified. If it is found to be linear the measuring points that are further apart can also be used until the boundaries of the linear region are found. This part of the test model can be used for calculation of the pressure head drop due to wall friction and varying diameter so that Cfd can be found. 61

74 APPENDIX A An important parameter if one chooses to measure inside the test model is the setup that will be used since this influences the size of this linear part. For example the pump can be responsible for a larger non-linear part than the inlet with a reservoir. It is also possible to decrease the size of the non linear parts by using straight piping before and after the model. Final set up The choice is made to measure accurately. This can be done by measuring inside the test model perpendicular to the flow. This so called static pressure can be read from a manometer. The holes made to measure these pressures can also be used for injecting dye to visualize the turbulence of the flow. This helps to see if the measured data is valid by inspecting the flow attachment at the measurement points. Since no reservoirs are needed to measure pressure differences over the set up, the set up that is connected to the piping of the laboratory offers the most advantages and is used for the experiment. The flow can easily be regulated and the velocities are easy to measure. 62

75 APPENDIX B Appendix B: Model design & construction In this appendix the model design and construction and the final results are elaborated. For every IGW model eight sections have been constructed of a maximum length of 0.7 meters to obtain sufficient length per model. Longer sections were not possible because of the maximum that the vacuum deformer, a device necessary for the production, can manage. For each model a mould was made out of several layers of glued MDF wood in a CNC mill. The length depends on the geometry of the model and will be between 0.6 and 0.7 meters. The beginning and end are always the smallest diameter, D_0. The mould comprises half of a section plus 1.5 cm extra length at the end, beginning and bottom. Over this mould a 5 mm thick PETG plate is placed which is heated and sucked over the mould. The extra length of the mould is necessary so that the PETG follows the relevant profile of the mould well. Afterwards the extra length is removed with a horizontal saw and a sander. It was this process that causes a lot of heat and induced stresses in the materials in the sections of model C. This resulted in disappointing results for the geometry of this model. The flanges are lathed out of PVC. A mould is used to mark the centre and the holes for the bolts correctly. Finally the centre is lathed and the holes are drilled. Two of the PETG parts that were vacuum formed are glued together over the longitudinal axis to obtain one section. A glue, Loctite 495, was used that is so thin that the capillary properties of the fluid make that it sucks into the seam and glues the entire contact surface. A spring was used to measure the force that was necessary to break a glued seam with a length of 10 cm. At 40 kg it went off. This showed that the glue method was more than sufficient to hold the internal pressures caused by the flow. 63

76 APPENDIX B Arrangements were made for the manometers. These 1 mm holes were drilled perpendicular to the flow in the smallest diameters. Blocks with the same geometry as the outside of the pipe wall were constructed out of black PVC and placed over the pressure holes. These blocks have 8 mm holes with thread in which connectors are placed. From these connectors tubing runs to the manometer tap blocks. Capillary tubes used for the injection of dye for visualisation of turbulence were made which also fit into these holes. Finally the flanges were glued on the pipes using a silicon adhesive. The flanges were lathed diagonal on the inside to obtain a large contact surface between the pipe wall and the flange. The end of the pipe wall comes level with the surface of the flange to ensure good connection between the sections. The beginning and end of each section have the smallest diameter, D_0, which prevents extra local discontinuities. Finally the construction results in 8 sections for each model providing approximately 5 meters of testing length per model. 64

77 APPENDIX C Appendix C: Data straight pipe & Flange calibration This appendix provides the data on which the results of the flange calibration in paragraph 5.4 were based. The reason for doubting the velocities measured with the measurement flange was the measurement done with the straight pipe. The corresponding friction coefficients with a deviation of 40% seemed very high. Therefore the velocities were checked with a micro impeller and with volumetric measurements. This showed that the measurement flange produced incorrect data. Because the error was reproducible the flange has been calibrated with volumetric measurements. The measured data of the straight pipe, flange, micro impeller and volumetric measurement are presented in this appendix. This table gives the measured values of the first measurement of the static pressure relative to point 4 for the straight pipe. The last column shows the large deviation between the friction coefficient calculated with the measured combination of the pressure drop and flow velocity and the friction coefficient calculated from the flow velocity with the Nikuradse equation. First measurement with straight pipe (βflange = 0.8) U (m/s) h1 (cm) h2 (cm) h3 (cm) h4 (cm) Dh/dx (cm/m) Cf nikuradse Cf measured deviation (%) Graphically the measured pressures show a linear relation. calculation Cf static pressure relative to point 4 (cm water column) y = -8.94x y = -6x y = -4.44x y = -3.6x y = -2.54x position pressure points (m) U = 2.48 U = 2.08 U = 1.76 U = 1.57 U = 1.26 Linear (U = 2.48) Linear (U = 2.08) Linear (U = 1.76) Linear (U = 1.57) Linear (U = 1.26) Linear (U = 2.48) Since the measurement of pressures is fairly easy with the use of manometers, there was reason to believe that the measurement flange produced incorrect data. Therefore a check was performed with a micro impeller that measures flow velocities in a point (see figure). This micro impeller is placed inside the pipe in the axis. 65

78 APPENDIX C Because the flow velocities are not equal over the whole cross section of the pipe, a correction was used to obtain the average flow velocity. In the figure the ratio of the velocity in a point divided by the average flow velocity is depicted. Measurements with the micro impeller took place in the region of the axis at Reynolds numbers of about 3*10 5. Hence the average velocity was found by dividing the measured velocity in the axis by The results of the measurements are presented in the table. From the number of revolutions per second, n/t, Uimp can be calculated with the following formula: n Umill = (28) t The results show an average flow velocity that is approximately 28% higher than the velocities measured with the flange. Check with micro impeller (βflange = 0.8) t (s) n imp n/t (1/s) U imp (m/s) DH flange (cm) U flange (m/s) U imp average (m/s) Deviation (%)

79 APPENDIX C Since the flange had the right dimensions and was placed according to the ISO standards, the measurements of the micro impeller had a remarkable outcome. Therefore a second check was performed with a volumetric measurement. A flexible hose was attached to the end of the piping. When the flow was fully going, the flow rate was measured with the measurement flange. Then without stopping the flow the end of the flexible hose was inserted into the container maintaining a similar angle to prevent extra losses. At this moment the stopwatch was started and before the water level reaches the edge of the container the hose was pulled out and the stopwatch was stopped. Then using a balance the full container was weighed. From the weight of the water and the time used to fill the container the flow rate was calculated. The results are shown in the table underneath. Again a large deviation in the flow rates was found. This time the deviation was 25%. Check with volumetric measurement (βflange = 0.8) DH flange (cm) Q flange (m³/s) weight empty (kg) weight full (kg) t fill (s) Q vol meas. (m³/s) deviation (%) With two alternative flow measurements that produce similar results it became clear that the measurement flange formula was incorrect. Because the error was reproducible the choice was made to calibrate the flange because of the large number of flow measurements that had to be conducted in the experiment program. The error also was reproducible for β=0.7 and because of the fact that the flow rates were lower than initially expected, this flange is used for more accurate measurements. Since the micro impeller is less accurate because it disturbs the flow in the pipe, the volumetric method was used for the calibration. Because of the high importance of a good calibration, extensive measurements were performed. For each measurement the flange was read out three times: 1 visual measurement and 2 camera measurements (see table). When the camera measurements produced similar results, the visual measurement was excluded to obtain the final dh. If this was not the case the average of all three measurements was taken. 67

80 APPENDIX C Volumetric calibration (βflange = 0.7) nr. dh visual (cm) dh1 camera (cm) dh2 camera (cm) dh camera (cm) dh final (cm) note average of all 1 3 three only camera only camera average of all three excluded from measurements due to overflow average of all three average of all three only visual The measured dh of the flange together with the new flow rates are presented in the table underneath. Measurement number 5 is excluded from the final data since water was flowing over the edge of the container due to a bad configuration of the flexible pipe. Volumetric calibration (βflange = 0.7) Nr. dh (cm water) Weight empty (kg) Weight full (kg) t fill (s) Q (l/s) Finally the results are plotted in a graph and a power line is fitted, because the original ISO formula is a constant value times the approximate squared root of the pressure differential: 0.49 Q = 1.98( dh ) (29) The new formula shows a good fit with a power relation. The difference is in the multiplier which now is 2.41 instead of

81 APPENDIX C new formula measurement flange Q (l/s) dh (cm water column) Q = (dH) R 2 = volumetric measurement Power (volumetric measurement) 69

82 APPENDIX C 70

83 APPENDIX D Appendix D: Measurements IGW pipe This appendix provides the data support for the experiment results that were presented in Chapter 6. it is divided into three parts, each corresponding to one of the models. For model A and B static pressure measurements are used. For model C dynamic pressure measurements are used and the way this was executed is presented in the corresponding section. Model A In the table underneath the values of the measurements of the flange, dh, and the pressure head relative to point 7, h1-h7, are listed. Corresponding velocities are calculated with the new formula of the flange. Static pressure head measurements model A dh (cm Nr. -> h1 (cm) h2 (cm) h3 (cm) h4 (cm) h5 (cm) h6 (cm) h7 (cm) water) Pos. -> U= U= U= U= U= Graphical presentation of the data shows a strong linear correlation between the pressures which is a satisfying result. For each flow velocity a linear relation is calculated. pressure drop model A pressure head relative to point 7 (cm water) y = x y = x U = 2.13 U = 1.94 y = x y = x y = x U = 1.66 U = 1.40 U = 1.05 Lineair (U = 2.13) Lineair (U = 1.66) Lineair (U = 1.40) Lineair (U = 1.05) Lineair (U = 1.94) position (cm) dh/dx results from the linear relations that were calculated. With dh/dx the total drop over the measurement section is calculated. Then the pressure drop caused by the wall friction is calculated. Loss due to wall friction is subtracted from the total loss and leads to the pressure drop caused by the shape created by the application of IGW. Dividing this loss by the number of bumps in the measurement section leads to the drop for 1 bump. 71

84 APPENDIX D Calculation Carnot loss per bump for model A dh/dx u^2/2g (cm) total drop (cm) Cf drop cf (cm) drop igw (cm) drop 1bump (cm) The pressure drop per bump can then be plotted against the velocity head. The data shows a strong linear correlation and the results of this approach thus can be considered satisfactory. delta_hv (cm) Carnot loss model A U^2/2g (cm) Hv = (U^2/2g) R 2 = model A Linear (model A) With this relationship for the Carnot loss the effective diameter for a range of velocities is calculated. The effective diameter is the internal diameter of a straight pipe that produces the same pressure drop for a given flow rate. Note that the calculations for the velocities above 2.13 m/s are an extrapolation of the results and thus introduce an extra uncertainty. First the calculations are done to find the pressure drop in model A, dh/l (see table below). Pressure drop for IGW model A Input Q (m^3/s) Calculated u Re (m/s) Cf dh/l igw dh/l wall dh/l total , , , , , , In offloading systems the offloading hoses are specified to withstand a flow velocity of 21 m/s. however in reality the velocity is around 7 m/s to reduce the impact of the surge pressures which depends on the operating pressures. Since the effective diameter differs slightly with the velocity, Deff is chosen for U=7.23 as a design parameter. This is the green row in the tables. This leads to a Deff / D_0 =

85 APPENDIX D Pressure drop for equivalent straight pipe Input Q (m^3/s) Calculated Deff U (m) (m/s) Re Cf dh/l , , , , , , Model B In the table underneath the values of the measurements of the flange, dh, and the pressure head relative to point 7, h1-h7, are listed. Corresponding velocities are calculated with the new formula of the flange. Static pressure head measurements model B dh (cm Nr. -> h1 (cm) h2 (cm) h3 (cm) h4 (cm) h5 (cm) h6 (cm) h7 (cm) water) Pos. -> U= U= U= U= U= U= Graphical presentation of the data shows a strong linear correlation between the pressures which is a satisfying result. For each flow velocity a linear relation is calculated. pressure drop model B pressure head relative to point 7 (cm water) y = x y = x y = x y = x y = x y = x U = 2.94 U = 2.64 U = 2.35 U = 2.06 U = 1.69 U = Position (x) 73

86 APPENDIX D dh/dx results from the linear relations that were calculated. With dh/dx the total drop over the measurement section is calculated. Then the pressure drop caused by the wall friction is calculated. Loss due to wall friction is subtracted from the total loss and leads to the pressure drop caused by the shape created by the application of IGW. Dividing this loss by the number of bumps in the measurement section leads to the drop for 1 bump. Calculation Carnot loss per bump for model B dh/dx u^2/2g (cm) total drop (cm) Cf drop cf (cm) drop igw (cm) drop 1bump (cm) The pressure drop per bump can then be plotted against the velocity head. The data shows a strong linear correlation and the results of this approach thus can be considered satisfactory. However one aspect of the formula is out of the ordinary. That is the rather large constant of cm. Extrapolation towards a velocity head of zero leads to a pressure drop per bump of cm which is not possible. Therefore extrapolation cannot be used for values lower than the experimented range. delta_hv (cm) Carnot loss model B Hv = (U^2/2g) R 2 = model B Linear (model B) u^2/2g (cm) With this relationship for the Carnot loss the effective diameter for a range of velocities is calculated. The effective diameter is the internal diameter of a straight pipe that produces the same pressure drop for a given flow rate. Note that the calculations for the velocities above 2.9 m/s are an extrapolation of the results and thus introduce an extra uncertainty. First the calculations are done to find the pressure drop in model B, dh/l (see table below). In this drop the cm per bump is included. 74

87 APPENDIX D Pressure drop for IGW model B Input Q (m^3/s) Calculated u Re (m/s) Cf dh/l igw dh/l wall dh/l total , , , , , ,006, In offloading systems the offloading hoses are specified to withstand a flow velocity of 21 m/s. however in reality the velocity is around 7 m/s to reduce the impact of the surge pressures which depends on the operating pressures. Since the effective diameter differs slightly with the velocity, Deff is chosen for U=7.17 as a design parameter. This is the green row in the tables. This leads to a Deff / D_0 = Pressure drop for equivalent straight pipe Input Q (m^3/s) Calculated Deff U (m) (m/s) Re Cf dh/l , , , , , ,322, Model C In the table underneath the values of the measurements of the flange, dh, and the pressure head relative to point 7, h1-h7, are listed. Corresponding velocities are calculated with the new formula of the flange. Static pressure head measurements model C dh (cm Nr. -> h1 (cm) h2 (cm) h3 (cm) h4 (cm) h5 (cm) h6 (cm) H7 (cm) water) Pos. -> U= U= U= Graphical presentation of the data shows a weak linear correlation between the pressures which is not a satisfying result. For each flow velocity a linear relation is calculated. The relative predictive power, R 2, is 0.70 and 0.78 for these relations. The reason for this low quality data was the bad result of the construction of the segments for model C. 75

88 APPENDIX D static pressures model C pressure head relative to point 7 (cm water) y = x R 2 = y = x R 2 = meas 2 (U=3.49) meas 3 (U=1.74) Lineair (meas 2 (U=3.49)) Lineair (meas 3 (U=1.74)) position (cm) Therefore the choice was made to measure the total head instead of the pressure head. Thus the models had to be adapted to make these dynamic pressure measurements possible. Instead of measuring perpendicular to the flow, the flow now had to stream right into the pressure holes. To achieve this the construction in the picture was made. In the whole setup four of these dynamic measurement points were made. In the figure below the overview of the new setup is given. Because the inflow and outflow was considered to have a minor influence on the pressure drop line, the measurement sections were placed apart wider to have a larger range for higher accuracy. 76

89 APPENDIX D In the table underneath the values of the measurements of the flange, dh, and the total head relative to point 4, H1-H4, are listed. Corresponding velocities are calculated with the new formula of the flange. Dynamic pressure head measurements model C dh (cm Nr. -> H1 (cm) H2 (cm) H3 (cm) H4 (cm) water) Pos. -> U= U= U= U= U= U= Graphical presentation of the data shows a strong linear correlation between the pressures which is a satisfying result. For each flow velocity a linear relation is calculated. dynamic pressure model C 50 pressure head relative to point 4 (cm water) y = x y = x y = x y = x y = x y = x U = 3,47 U = 3,16 U = 2,78 U =2,41 U = 1,86 U = 1,24 Linear (U = 3,47) Linear (U = 3,16) Linear (U = 2,78) Linear (U =2,41) Linear (U = 1,86) Linear (U = 1,24) -10 position (cm) 77