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Thermal Insulation Material for Subsea Pipelines Benefits of Instrumented Full Scale Testing To Predict the Long Term Thermomechanical Behaviour See discussions, stats, and author profiles for this pu[.]
See discussions, stats, and author profiles for this publication at: https://www.researchgate.net/publication/254519432 Thermal Insulation Material for Subsea Pipelines: Benefits of Instrumented Full-Scale Testing To Predict the Long-Term Thermomechanical Behaviour Article · January 2007 DOI: 10.4043/18679-MS CITATIONS READS 694 authors, including: Nadège Bouchonneau Valérie Sauvant-Moynot Federal University of Pernambuco IFP Energies nouvelles 17 PUBLICATIONS 168 CITATIONS 45 PUBLICATIONS 857 CITATIONS SEE PROFILE Francois Grosjean IFP Energies nouvelles 46 PUBLICATIONS 289 CITATIONS SEE PROFILE Some of the authors of this publication are also working on these related projects: Finite Element modeling and Analysis of corroded pipelines View project PhD Thesis View project All content following this page was uploaded by Francois Grosjean on 07 September 2015 The user has requested enhancement of the downloaded file SEE PROFILE OTC 18679 Thermal Insulation Material for Subsea Pipelines: Benefits of Instrumented Full-Scale Testing To Predict the Long-Term Thermomechanical Behaviour N Bouchonneau, Ifremer, IFP , Franche-Comté U.; V Sauvant-Moynot and F Grosjean, IFP; D Choqueuse, Ifremer; and E Poncet and D Perreux, Franche-Comté U Copyright 2007, Offshore Technology Conference This paper was prepared for presentation at the 2007 Offshore Technology Conference held in Houston, Texas, U.S.A., 30 April–3 May 2007 This paper was selected for presentation by an OTC Program Committee following review of information contained in an abstract submitted by the author(s) Contents of the paper, as presented, have not been reviewed by the Offshore Technology Conference and are subject to correction by the author(s) The material, as presented, does not necessarily reflect any position of the Offshore Technology Conference, its officers, or members Papers presented at OTC are subject to publication review by Sponsor Society Committees of the Offshore Technology Conference Electronic reproduction, distribution, or storage of any part of this paper for commercial purposes without the written consent of the Offshore Technology Conference is prohibited Permission to reproduce in print is restricted to an abstract of not more than 300 words; illustrations may not be copied The abstract must contain conspicuous acknowledgment of where and by whom the paper was presented Write Librarian, OTC, P.O Box 833836, Richardson, TX 75083-3836, U.S.A., fax 01-972-952-9435 Abstract External coating systems of flowlines and risers ensure both structural and thermal insulation functions which should be efficient throughout the design life in-service, typically 25 years In that context, the long term behaviour of thermal insulation materials is difficult to predict due to the coupled effects of three factors: hydrostatic pressure up to 300 bar, thermal gradient over 120°C between internal effluents and external sea water and the water absorption of constitutive materials In addition, laboratory data collected on small size specimens of insulation materials are normally used to predict the thermo-mechanical behaviour of full scale systems, but laboratory testing simply not properly simulate the service conditions, in particular the complex loading existing through the coating thickness This paper covers the background to the development of both test facilities and models to study the thermo-mechanical behaviour of production coated steel pipe in ultra deep water conditions This original work was launched to provide both experimental and computed data to better understand and predict the thermo-mechanical behaviour of insulation materials whilst considered as a full scale system On the one hand, experimental data obtained on instrumented insulated pipes immersed in large scale facilities simulating ultra deep water are presented in both steady and transient states On the other hand, a finite element model dedicated to the abovementioned insulated pipes was developed to predict their thermo-mechanical behaviour Correlation between full scale experimental data and related model predictions are discussed to validate the predictive model taking into account the coupling between hydrostatic pressure and temperature gradient Additional modelling developments to include the water absorption are planned to reach a suitable prediction of the whole service life Introduction Optimistic estimations of oil reserves in deep water and current oil & gas prices sustain the increasing interest towards offshore deepwater field production The ultra deep water (3000m water depth) is one of the next issues Indeed, 4% of the world offshore surface with WD>1500m includes sedimentary areas with hydrocarbon potential (minimum sediment thickness of 2000m) [1] Those ultra-deepwater fields, between 100 to 500 [1], are expected to be located in the Gulf of Mexico, in the Atlantic off Brazil, Nigeria and Angola, and also near Aegypta in the Nil delta It is worth noting that the hydrocarbon reserves identified and to be identified in both onshore and conventional offshore sedimentary basins represent 19% of the world surface In comparison to onshore and conventional offshore hydrocarbons, the partial exploitation of ultra-deep reserves, about 1% of the world surface, would correspond to 30 billion to 100 billion of barrels equivalent petrol [1] As a consequence, the ultra-deep offshore production representing 10% of the offshore production in 2005 is expected to grow to 25% in 2025 [2] In that context, flow assurance continues to be a critical part of system design and operations, with lower seabed temperatures - typically in the to 4°C range at 1500m-3000m depth - and rising insulation costs in deepwater [3] Among others, the heat management in normal (steady state) and dynamic (transient) operations relies on the selection of proper insulation materials and designs for subsea flowlines and risers to meet the increasing demand for deeper waters Pipein-pipe configurations are under study to optimize their performance but their heavy weight may be a limitation Advanced insulation materials and coatings are also being developed and designed for subsea use to offer both appropriate thermal and mechanical properties in ultra-deep water applications [4, 5] Moving towards ultra-deep water applications also emphasises the need to have test methods and facilities that establish whether a given coating system is fit-for-purpose - either new insulation materials/systems or existing materials/systems to be subjected to conditions for which there are no available data Full-scale thermal testing protocoles and facilities have been developed since the late 1980's to study the behaviour of thermal insulation coating systems on lengths of pipe under simulated service conditions [6, 7] Such full-scale tests are becoming part of the qualification testing since laboratory testing on small size specimen of insulation materials simply not properly simulate the service conditions, in particular the complex loading existing through the coating thickness The development and validation of a test protocol and instrumentation which will permit the study of the thermal behavior of a coated insulation pipe when subjected to conditions (300 bar) simulating ultra-deep water (nearly 3000m water depth) is the first issue of this paper In parallel, the modelling of insulation systems at a global level is a necessary basis to address their long term behaviour in service In particular, water absorption effects have to be considered at design stage for syntactic foams which are widely used as rigid and flexible insulation materials [8, 9] Indeed the long-term response of syntactic foams while exposed to seawater at high pressure is highly non-linear, particularly at elevated temperatures, revealing the occurrence of complex degradation mechanisms leading to glass microsphere filling [10-14] The modelling of water absorption for syntactic foams has been recently performed on small size specimens and validated on a large data base (4 syntactic materials with geometries and aged under 18 conditions from 4°C, 1bar up to 130°C, 300 bar) [15] But the long term behaviour of thermal insulated structures is difficult to predict due to the coupled effects of three combined factors: hydrostatic pressure up to 300 bar, inducing a stress gradient within the coating material, thermal gradient over 120°C between internal effluents and external sea water, the water absorption of constitutive materials The development of a finite element model to satisfactorily predict the thermo-mechanical behavior of a production insulated steel pipe in ultra deep water conditions (transient and steady states) is the second issue of this paper The study work presented in this paper is an expansion of the work given in [16] and was part of a PhD study Experimental set-up Insulated pipe configuration Both experimental and computational developments achieved in this work were dedicated to 1.2m length steel pipe sections industrially coated with an externally applied fivemultilayer insulation system mainly based on syntactic polypropylen (PP) material Test pipe sections were taken from a length of pipe with internal diameter of 180mm, coated under normal production conditions The geometry and composition of the 61mm thick insulation system are summarised in Figure The outer diameter is 338mm Simulated service test Pressure vessel The pressure vessel used in this work (1m diameter and 2m height) is located in Ifremer Both water temperature and pressure (up to 1000bar) are monitored and regulated from the external data acquisition unit The pressure is monitored using a pressure transducer mounted at the top of the pressure vessel A schematic view of the coated test pipe in vertical OTC 18679 position during testing in the pressure vessel is given in the Figure The pressure vessel flange allows the connection of the inner instrumentation to the outer data acquisition unit where readings obtained from all sensors are recorded Equipment of coated test pipe The design of the prototype equipment is a crucial point in the experimental set up of full-scale tests The insulated pipe section was machined at both ends to adapt two metallic steel caps (Stainless steel APX4) equipped with connectors that resist external pressure Three 10-channels connectors were necessary to allow the electrical supply of the inner heating system and to collect inner sensor data Steel caps were covered by 100mm thick Polytetrafluoroethylene (PTFE) insulating caps to limit the axial heat flow as much as possible The equipment of the insulation prototype is shown in progress on Figure and the schematic representation of the fully instrumented pipe section is given in Figure Instrumentation is detailed hereafter Heating system An original heating system was developed in this study instead of the classical circulating oil to limit the convection effects inside of the pipe The heating system consisting of heating elements (NiCr) embedded in a thin silicone layer was placed on the internal diameter of the steel pipe and kept in place by brushes It is worth noting that such an inner 'dry' configuration with no pressure and no liquid is also very beneficial for the instrumentation used inside the pipe and will simplify the modelling of the inner heat flux boundary condition In addition, the electric power applied to the system during steady state phase will provide an indirect monitoring of the radial heating flux through the coating Temperature sensors The insulated pipe section was instrumented with six commercial temperature sensors (Pt100) specified up to 200°C minimum (precision of about 0.3 % at 100 °C), located in both inner and outer parts along the pipe length and on the caps (Figure 4): - Ti (°C): inner temperature of the steel surface in the center of the pipe (one measurement); - Te (°C): outer temperature of the coating surface in the center of the pipe (one measurement); - Tb (°C): inner temperature of the steel surface in the center of one cap (one measurement); - T100 (°C): inner temperature of the steel surface along the pipe 100mm distant from cap (one measurement); - T50 (°C): inner temperature of the steel surface along the pipe 50mm distant from cap (one measurement); - TPTFE (°C): outer temperature of the PTFE surface in the center of the cap (one measurement) Besides, the outer temperature of the water in the vessel, Twater (°C), was also measured using a platinium sensor Heat flux sensors The insulated pipe section was also instrumented with four commercial heat flux sensors located in both inner and outer parts along the pipe length and the caps: - φi (W.m-2): inner thermal flux density on the steel surface in the center of the pipe (one measurement OTC 18679 with soft circular fluxmeter of 5µV.m2.W-1 sensitivity specified up to 200°C); - φe (W.m-2): outer thermal flux density on the coating surface in the center of the pipe (three measurements with a semi-rigid fluxmeter of 50µV.m2.W-1 sensitivity specified up to 100°C and 100bar, one measure with a rigid fluxmeter of 30µV.m2.W-1 sensitivity specified up to 250°C and 150bar); - φcap (W.m-2): inner thermal flux density on the steel surface in the center of the steel cap (one measurement with rigid rectangular fluxmeter of 36µV.m2.W-1 sensitivity specified up to 200°C) It should be pointed out that heat flux sensors with soft flat form were selected for the internal and external pipe surfaces to reduce errors related to the difficulties in mounting rigid flat sensors Test programs Two cases were considered both considering a length of steel pipe coated with an externally applied multilayer insulation material as described previously: No external pressure was applied in the pressure vessel (external pressure was bar) Hydrostatic pressure simulating in-service conditions in ultra-deep water was applied to the test pipe section The first case was performed as a basis of comparison for the numerical simulation to the experimental results The second case was conducted because it was the main objective to evaluate the thermal performance of the insulation coating system in steady and transient conditions while immersed in seawater and subjected both to a temperature gradient across its full thickness and a hydrostatic pressure identical to those which it would experience when in-service in ultra-deep water Both test programs are shown in Figure and Figure Each step of the testing program of case simulating ultradeep water immersion is described hereafter A preliminary test of the instrumented insulated pipe section referred to as step was performed in fresh water under ambient temperature and pressure prior to the simulated service test This first step was necessary to check that the prototype equipment and related instrumentation had been properly installed In step (about 30 minutes duration), the outer pressure was increased to 300bar to simulate in-service on the seabed at 3000m depth Heating power values of 120W and 240W were applied in steps and to reach representative thermal gradients through the coating thickness, respectively around 50°C and 110°C, given the outer water temperature thermoregulated around 15°C, the OHTC of the insulation coating under consideration and its surface area The duration of each step 2, and was typically around days On completion of the test, the heating circuits were switched off allowing the internal temperature of the coated test pipe to decrease down to temperature of the sea water surrounding the pipe in the pressure vessel (step 5) When the response of the test pipe to the removal of the temperature gradient had stabilised, the hydrostatic pressure in the pressure vessel was rapidly reduced to atmospheric (step 6) In both cases, it should be emphasised that a test lasting approximately 10 days cannot be used to predict the long term evolution of the insulation coating systems for which water uptake and creep cannot be neglected Numerical model of simulated service test Numerical model A two-dimensional axisymetric numerical model of the insulated pipe section was developed using finite element multiphysics commercial software where mechanical and thermal aspects are coupled under the following assumptions: - Coating and pipe materials are simulated as solids with linear elastic behaviour (no creep) - The thermal conduction of coating and pipe materials is simulated using Fourier's law - The natural convection between the coating surface and the external water is simulated using equations derived from Newton's law The convection coefficients were calculated from experimental temperatures Te and Twater according to [17] Material properties and boundary conditions Thermal and mechanical properties of each constitutive material of the test section are reported, respectively, in Table and Table Values collected from experimental measurements performed at bar on small samples [15] or from the literature were used as input data in the simulation Boundary conditions used were conduction along the inner surface of the steel pipe, insulation at both inner end caps and convection along the external surfaces in contact with water For tests under hydrostatic pressure, a stress condition is applied on the external surfaces of the structure The displacements perpendicular to the symmetry conditions are also blocked The initial conditions (temperature or pressure) depend on the different sequences of the tests performed on the structure The geometry of the computational domain, the numerical mesh and the boundary conditions are shown in Figure The mesh was locally refined near interfaces and sensors to enhance the resolution Applications of the numerical model Simulation at design stage At the conception stage of the pipe equipment, the thermomechanical simulation of the insulated pipe section performed in steady and transient states on the basis of material geometries and physical properties can provide: - the insulation coating thickness shrinkage / swelling under increasing pressure / temperature along the length of the insulated pipe immersed in water under pressure; - the external surface temperatures and heat flux distributions along the insulated pipe section Numerical results help for the design of end caps for example but such simulations remain only indicative since physical properties used as input data are obtained through material testing under laboratory conditions which does not reflect full scale conditions Investigation of insulation properties During the testing stage of one insulation pipe section under bar (case 1), simulations were performed under OTC 18679 transient and steady state conditions The comparison between experimental and simulation results contributed to validation of the model and to assessing the prototype instrumentation During the testing stage of one insulation pipe section under 300bar (case 2), there were no direct measurements of the radial heat fluxes since commercial sensors are limited to a lower pressure range Hence the thermo-mechanical model was used to determine the heat transfer coefficient and the material properties from the experimental data Steady state conditions The radial heat flux was determined by simulation in order to fit the experimental temperature distribution Then the heat transfer coefficient and the apparent thermal conductivity of the insulation material (apparent meaning that the thermal conductivity was averaged on the insulation cross section under thermal gradient) were calculated using classical analytical expressions The analytical expression for the radial heat flux under steady state conditions for a one-dimensional conduction problem in a composite cylinder, evaluated on the inner surface of the structure taken as reference, is given by: (1) Q = -U.S (Text - Tint) In a case of a multilayer structure, and by assuming that the thermal contact resistance between each layer can be neglected, the heat transfer coefficient U of the structure can be expressed in terms of constitutive material thermal conductivities with the following relation: U= ⎡ ⎛ Di +1 ⎞ ⎤ (2) ⎟⎟ ⎥ ln⎜ n ⎢ ⎜ ⎢ ⎝ Di ⎠ ⎥ S× ⎢ 2πLλi ⎥ i =1 ⎢ ⎥ ⎣⎢ ⎦⎥ ∑ With these hypotheses, and according to the characteristics given by manufacturers, the overall heat transfer coefficient also defined as "U value" is 4.2W.m-2.K-1 at 20°C This coefficient is a representative thermal characteristic of the entire system: steel pipe and insulation coating Transient state conditions Both thermal conductivity and heat capacity can be determined directly by an optimisation program developed with a commercial software This routine allowed the evolution of the inner temperature of the steel pipe Ti to be adjusted during the testing time simulated by an analytical relation to the experimental data by optimisation (square root method) of the thermal parameters of the material studied, here the syntactic foam In the case of a one-dimensional radial transfer in a onelayer structure limited by radii r=rint and r=rext and of length l, the heat equations (temperature and heat flux) are: ∂ ⎛ ∂T ⎞ ∂T for rint < r < rext ⎜r ⎟= r ∂r ⎝ ∂r ⎠ a ∂t with T=T0 for t=0, and (3) Φ = −λS ∂T ∂r (4) Applying a Laplace transform to the variable t, these equations lead to: d θ dθ p + − θ dr r dr a (5) and φ = −λS ∂θ ∂r (6) The quadrupoles notation [18] is well suited to relate the Laplace transforms of the temperatures and fluxes at inner and external boundaries obtained by solving the above equations: ⎡θ int ⎤ ⎡ A B ⎤ ⎡θ ext ⎤ ⎢φ ⎥ = ⎢C D⎥ ⎢φ ⎥ ⎦ ⎣ ext ⎦ ⎣ int ⎦ ⎣ (7) θint et θext correspond to the transforms of the inner and outer surface temperatures of the cylindrical structure respectively, and A, B, C and D are analytical relations involving Bessel functions and the geometrical characteristics of the structure This development is applied to the multilayered structure under test (6 layers including steel pipe), submitted to outer convective losses Equation (7) becomes: ⎤ ⎡ ⎡θ int ⎤ ⎡ Ai Bi ⎤ ⎢1 ⎥ ⎡ θ water ⎤ (8) = h ⎥ ⎢φ ⎥ ⎢C D ⎥ ⎢ ext Sext ⎥ ⎢ i⎦ ⎣φconvective ⎦ ⎣ int ⎦ i =1 ⎣ i ⎣ ⎦ with θwater and φconvective the Laplace transforms of the water temperature and the convective heat flux respectively ∏ Temperatures and heat fluxes evolutions are obtained in the time space by means of the numerical inversion of each Laplace transform Transient model hypotheses are: - One-dimensionnal axisymmetric - Constant convective heat transfer coefficient and water temperature - Constant heat flux equal to the value of the steady state heat flux measured in the structure (experimental values at bar and simulated values for the tests under 300 bar) - Initial temperature of the entire structure supposed to be stabilized at the water temperature; - The external surface area is constant (no thermal expansion and no pressure effect) Results Main experimental and computational results are summarised in this section relative for both cases under consideration Case 1- 1bar Steady state Experimental heat fluxes measured in the steady state under 1bar and 120W, then 1bar and 240W, are reported in the Table One can notice that inner heat fluxes are smaller than the heating mat power values, due to thermal losses at both ends and possibly in the inside of the pipe But the particularly OTC 18679 low value measured under 240W (even smaller than the external heat flux) reveals that the soft heat flux sensor used at the interface between the heating mat and the steel pipe does not sustain the high temperature environment which it is subjected to Three methods are proposed to determine the heat transfer coefficient U Method A The heat transfer coefficient U is determined using equation (1) directly from the experimental external heat fluxes and temperatures measured The precision of the U value is 3.1% based on the sensor measurement errors given by manufacturers The apparent thermal conductivity values of the syntactic PP were then derived from equation (2) Results reported in the Table show that U values and apparent thermal conductivities are not significantly influenced by the thermal gradients (differences are within incertainty ranges) U values measured are smaller than the one given by the insulated pipe manufacturer and apparent thermal conductivities of syntactic PP are also slightly lower than values measured on small specimens (Table 1) Method B In the absence of any external heat flux sensor, the external heat flux could be approached to first approximation by substracting the heat losses measured at both caps from the power of the heating mat Values of 115W and 226W were obtained for, respectively, 120 W and 240W heating power Internal heat flux densities were calculated from the aforementioned heat fluxes by dividing flux values by 0.57m2 (internal heating mat surface area) The simulation of thermomechanical behaviour of the insulated pipe section was performed with those calculated internal flux densities The comparison of experimental and simulated temperature values are presented in the Figure Inner and external temperature measurements agree quite well but one can note that other simulated temperature are overestimated, suggesting that the internal heat flux density was also overestimated due to underestimated heat losses at both ends This method would be improved by a better estimation of end losses Method C The internal heat flux is optimised to fit test temperatures with the numerical simulation results Results obtained with the simulation and experimental temperatures are presented Figure The temperature distribution simulated within insulated pipe section is shown in Figure 10 Calculated U values and apparent thermal conductivities of syntactic PP are given in the Table Thermal properties estimated by method C are comparative to those given by method A, which validates the use of the thermo-mechanical numerical simulation to determine the OHTC and the apparent thermal conductivity of insulation materials Transient state Since the analysis of experimental data with the transient model approach requires that initial temperatures are fully stabilised, the evolution with time of the inner temperature was simulated only for the 120W thermal gradient The experimental values and simulated curve are compared on Figure 11 Optimisation input data and results are reported in the Table The very similar apparent thermal conductivity values obtained compared to the steady state approaches (methods A and C) validate the transient state analysis The apparent heat capacity of syntactic PP is lower than the value determined experimentally, but this later value should be considered with caution since heat capacity values are extremely difficult to measure Case 2- 300bar Steady state Experimental heat fluxes measured in the steady state under 300bar and 120W, then 300bar and 240W, are reported in the Table No external heat fluxes were measured in simulated ultra deep water conditions, therefore U value and apparent thermal conductivity of syntactic PP cannot be determined using method A Therefore, method C was used to evaluate the heat transfer coefficient of the structure The internal heat flux used in the thermomechanical modelling was optimised to fit the temperatures more closely Results obtained with the numerical simulation agree quite well with the experimental temperatures (Figure 12) U values and apparent thermal conductivities of syntactic PP calculated using analytical expressions (1) and (2) are reported in the Table One can observe that values obtained under 300bar 120W are similar to those obtained under bar But this is no longer true for 240W experiments The significant increase of the U value (+10%) and related increase of the apparent thermal conductivity could be explained by damage occurring in the foam microstructure, in particular in the vicinity of the pipe where material is subjected to coupled effect of high temperature and complex stress distribution Transient state The evolution with time of the inner temperature was also simulated during the establishment of the temperature distribution with a heating power of 120W The experimental values and simulated curve are compared on Figure 13 Optimisation of input data and results, reported in the Table 4, led to very similar apparent thermal conductivity value compared to the steady state approach and validate once again the transient state analysis From apparent thermal conductivity and heat capacity values obtained under 120W 1bar and 120W 300bar, there is no significant difference as stated previously from steady state analysis Thus, one can conclude that the coupling effect of pressure and thermal gradient induces no short term consequences on syntactic PP provided the inner effluents have a temperature around 60°C But in the case of temperatures of 80°C and above coupled with ultra deep service pressure, short term phenomena may occur in the syntactic PP leading to a lowering of insulation performance Conclusion Very demanding in service conditions in ultra deep water require specific test conditions and experimental equipment to perform full scale test Tests have been performed on insulated structures under service conditions (P=300bar, Tint=95°C) An original heating system was developed instead of the classical circulating oil to limit the convection effects inside of the pipe and simplify the boundary conditions of heat flux modelling Novel instrumentation has been developped in order to monitor different test parameters (internal and external temperatures, heat fluxes, …) These tests allow pertinent results to be obtained When there is no available external heat flux sensor, one effective way to determine the OHTC and the insulation material thermal properties is to perform numerical simulations and fit the temperature distributions in both steady and transient states Satisfactory agreements between the twodimensional numerical simulation results including thermal and mechanical coupling and tests results given by conventional instrumentation were obtained at 1bar Numerical simulations, on the other hand, may be used to guide the design of an insulated flowline test systems In the near future, the water diffusion into the insulation material will be taken into account in order to predict the long term insulation behaviour Nomenclature U = heat transfer coefficient of the structure relative to a reference surface [W.m-2.K-1] S = inner surface area, expressed as S=πLD1 [m²] Sext = external surface area [m²] Text = external surface temperature in steady state conditions [°C] Tint = internal surface temperature in steady state conditions [°C] Di = inner diameter of the layer i of the structure [m] Di+1 = external diameter of the layer i of the structure [m] D1 = inner diameter of the steel pipe [m] L = steel pipe length [m] λi = thermal conductivity of the layer i [W.m-1.K-1] hext = convective heat transfer coefficient at the interface between insulation coating and water [W.m-2.K-1] a = thermal diffusivity [m2.s-1] T0 = initial temperature [°C] T= temperature [K] Acknowledgments The authors wish to thank Socotherm for providing the insulated coated pipes, in particular G P Guidetti for his interest to this work and N Lacotte and A Deuff for the performing of hyperbaric tests References MATHIEU, Y., IFP technical note, October 2006 ROBERTSON, S., MACFARLAN, G., et SMITH, M., "Deep water expenditures to reach $20 billion/year by 2010", Offshore Magazine, 2005 McMULLEN N.D., "Flow-Assurance Field Solutions", Offshore Technology Conference - OTC 18381, Houston, Texas U S A., 1-4 May 2006 BOYE HANSEN A., JACKSON A., 'High performance polypropylene thermal insulation for high temperature and deep water applications", 16th International Conference on Pipeline Protection, Paphos, Cyprus, 2-4 November 2005 BERTI, E., "Syntactic polypropylene coating solution provides thermal insulation for Bonga risers", Offshore Magazine, 2004 HALDANE D., GRAAF F.v.d et LANKHORST A.M., "A direct measurement system to obtain the thermal conductivity of pipeline insulation coating systems under simulated service conditions", Offshore Technology Conference - OTC 11040, Houston, Texas U S A., 3-6 May 1999 OTC 18679 MELVE B., RYDIN C et BOYE HANSEN A., "Long term testing of high temperature thermal insulation for subsea flowlines at simulated seabed conditions", 15th International Conference on Pipeline Protection, Aachen, Germany, 29-31 October 2003 DAVALATH J., "Cool-down thermal performance of subsea systems based on Gulf of Mexico Field Experience", Offshore Technology Conference - OTC 17972, Houston, Texas U S A., 1-4 May 2006 CHALUMEAU A., FELIX-HENRY A., "Water absorption effect on syntactic foam thermal insulation of a flexible pipe", 25th International Conference on Offshore Mechanics and Arctic Engineering (OMAE), Hamburg, Germany, 4-9 June 2006 10 CHOQUEUSE D., CHOMARD A et BUCHERIE C., "Insulation materials for ultra deep sea flow assurance: Evaluation of the material properties", Offshore Technology Conference - OTC 14115, Houston, Texas (U.S.A.), 6-9 May 2002 11 CHOQUEUSE D., CHOMARD A et CHAUCHOT P., "How to provide relevant data for the prediction of long term behavior of insulation materials under hot/wet conditions ?", Offshore Technology Conference - OTC 16503, Houston, Texas U.S.A.,36 May 2004 12 GIMENEZ N., SAUVANT-MOYNOT V et SAUTEREAU H., "Wet ageing of syntactic foams under high pressure / high temperature in deionized and artificial sea water", 24th International Conference on Offshore Mechanics and Artic Engineering, Halkidiki, Greece, 12-17 June 2005 13 HALDANE D., SCRIMSHAW K.H., "Development of an alternative approach to the testing of thermal insulation materials for subsea applications", 14th International Conference on Pipeline Protection, Barcelona, Spain, 29-31 October 2001 14 SAUVANT-MOYNOT V., GIMENEZ N et SAUTEREAU H., "Hydrolytic ageing of syntactic foams for thermal insulation in deep water: degradation mechanisms and water uptake model", Journal of Materials Science, 2006, 41 (13), p 4047-4054 15 LEFÈBVRE X., SAUVANT-MOYNOT V., CHOQUEUSE D et CHAUCHOT P., "Durabilité des matériaux syntactiques d'isolation thermique et de flottabilité: des mécanismes de dégradation la modélisation des propriétés long terme", Matériaux 2006, Dijon, France, 13-17 November 2006 16 BOUCHONNEAU N et al., "Multilayer systems for thermal insulation: thermomechanical behaviour of prototypes for deep sea applications", Oilfield Engineering with Polymers, 29-31 March 2006 17 EYGLUNENT B., "Manuel de thermique - Théorie et pratique"; HERMES Science Publications, Paris, 1997 18 MAILLET D., ANDRÉ S., BATSALE J.-C., DEGIOVANNI A and MOYNE C., “Thermal quadrupoles : solving the heat equation through integral transforms”; John Wiley & Sons, Inc., 2000 Table 1-Thermal properties Thermal conductivity a) Heat capacity a) (W.m-1.K-1) (J.kg-1.K-1) Steel pipe 45 475 Fusion bonded Epoxy 0.3 2000 Adhesive PP 0.22 2090 PP 0.22 2000 Syntactic PP 0.165 + 10-4 * T 1506.6 + 6.26 * T Steel (cap) - APX4 19 460 PTFE (insulating end cap) 0.24 1050 a) Values are given at 20°C when the temperature dependence is not specified Material Table 2- Mechanical and thermomechanical properties Expansion coefficient Material (between 10 and 100 °C) (°C-1) Steel pipe 7850 218 0.33 1×10-5 Fusion bonded Epoxy 1200 0.4 5.3×10-5 Adhesive PP 900 1.3 0.4 1.6×10-4 PP 900 1.3 0.4 1.6×10-4 Syntactic PP 640 E = -0.94 * 10-3 T + 1.1 0.32 5×10-5 Steel (cap) - APX4 7700 211 0.33 1×10-5 PTFE (insulating end cap) 2200 0.4 46 1.3×10-4 a) Values are given at 20°C when the temperature dependence is not specified Density (kg.m-3) Elastic modulus a) (GPa) Poisson coefficient Table 3- Heat fluxes and temperatures in the steady state and related thermal properties (simulation results are in italics) Pressure (bar) Heating mat power (W) 1 1 1 300 300 120 120 120 240 240 240 120 240 Method Internal radial heat flux (W) External axial heat flux (W) External radial heat flux (W) Ti (°C) Te (°C) U (W.m-2.K-1) A B C A B C C C 99.7 114.6 109 177 226.3 223.7 97.5 195 5.4 13.7 15.5 89 96.6 91.9 178 191 188.8 84.3 172 56.4 56.4 56.4 95.8 95.8 95.8 55.6 88.5 16.4 16.4 16.4 17.6 17.6 17.6 18.3 20 3.88±0.12 4.21±0.13 4.00±0.12 3.97±0.12 4.26±0.13 4.21±0.13 3.94±0.12 4.37±0.13 Apparent thermal conductivity of syntactic PP (W.m-1.K-1) 0.152±0.005 0.166±0.006 0.157±0.005 0.155±0.006 0.168±0.006 0.166±0.006 0.154±0.005 0.173±0.006 Table 4- Thermal properties of syntactic PP determined by the transient state analysis Pressure (bar) Heating mat power (W)) External radial heat flux in steady state (W) Water temperature (°C) Mean external convection coefficient (W.m-2.K-1) 300 120 120 89 84.3 15.3 16.7 125 170 Apparent thermal conductivity of syntactic PP (W.m-1.K-1) 0.150 0.154 Apparent heat capacity of syntactic PP (J.kg-1.K-1) 1510 1519 OTC 18679 PP (2,5 mm) Syntactic PP (55 mm) PP (3 mm) Adhesive PP (0,25 mm) Fusion bounded Epoxy (0,25 mm) Steel pipe wall (18,26 mm) Figure 1-Section of the 5-multilayer insulation coating on steel pipe (thickness values are given in brackets) Electric connection to the data acquisition unit Pressure vessel Pmax = 1000 bar Ø=1m h=2m Coated pipe section Figure 2-Schematic representation of the insulated pipe section during test in the pressure vessel Figure 3-Instrumentation in progress of the insulated pipe section (1) PP multilayer insulation coating, (2) Steel cap with connector, (3) PTFE cap, (4) Heating mat, (5) Outer flux meter OTC 18679 Te T50 Tb φb TPTFE T100 Ti φe Heating system power input φi Tb Figure 4-Configuration of the fully instrumented test section Heating power (W) 240 120 Time Figure 5-Testing program under bar pressure Heating power (W) Pressure (bar) 240 300 120 Time Figure 6-Testing program simulating ultra-deep water 10 OTC 18679 z Displacement (r) blocked Thermal convection : Teau, h Thermal insulation Hydrostatic pressure Thermal heat flow O Displacement (z) blocked r Figure 7-Boundary conditions and mesh of insulated pipe section (a) Ti 110 Te 100 Tb Temperature (°C) 90 T50 80 T100 70 TPTFE 60 Ti-simulation 50 Te-simulation 40 Tb-simulation 30 T50-simulation 20 T100-simulation TPTFE-simulation 10 50000 100000 150000 200000 Time (s) (b) 110 Ti 100 Te Temperature (°C) 90 Tb 80 T50 70 T100 60 Ti-simulation 50 Te-simulation 40 Tb-simulation 30 T50-simulation 20 T100-simulation 10 40000 80000 120000 160000 Time (s) Figure 8-Comparison between experimental and simulated temperatures during test at bar (a) 120W, (b) 240W (method B) OTC 18679 11 (a) 110 Ti 100 Te Temperature (°C) 90 Tb 80 T50 70 T100 60 Ti-simulation 50 Te-simulation 40 Tb-simulation 30 T50-simulation 20 T100-simulation 10 50000 100000 150000 200000 Time (s) Temperature (°C) (b) 110 100 90 80 70 60 50 40 30 20 10 Ti Te Tb T50 T100 Ti-simulation Te-simulation Tb-simulation T50-simulation T100-simulation 40000 80000 120000 160000 Time (s) Figure 9-Comparison between experimental and simulated temperatures during test at bar (a) 120W, (b) 240W (method C) Figure 10-Temperature distribution within insulated pipe section in the steady state during test at bar 120W 12 OTC 18679 100 Temperature (°C) 90 Ti experience proto2 1bar 120W 80 70 60 50 Ti matlab optimisation proto2 1bar 120W 40 30 20 10 50000 100000 150000 200000 Time (s) Figure 11-Comparison between experimental and simulated temperatures with transient model during test at bar (a) 110 Ti 100 Te Temperature (°C) 90 Tb 80 T50 70 T100 60 Ti-simulation 50 Te-simulation 40 Tb-simulation 30 T50-simulation 20 T100-simulation 10 50000 100000 150000 200000 250000 Time (s) (b) 110 Ti 100 Te Temperature (°C) 90 Tb 80 T50 70 T100 60 Ti-simulation 50 Te-simulation 40 Tb-simulation 30 T50-simulation 20 T100-simulation 10 60000 120000 180000 240000 Time (s) Figure 12-Comparison between experimental and simulated temperatures during test at 300 bar (a) 120W, (b) 240W (method C) OTC 18679 13 Temperature (°C) 110 Ti experience proto2 300bar 120W 90 70 50 Ti matlab optimisation proto2 300bar 120W 30 10 50000 100000 150000 200000 250000 Time (s) Figure 13-Comparison between experimental and simulated temperatures with transient model during test at 300 bar View publication stats