Heat transfer engineering an international journal, tập 32, số 1, 2011

Heat Transfer Engineering, 32(1):1–13, 2011 Copyright C Taylor and Francis Group, LLC ISSN: 0145-7632 print / 1521-0537 online DOI: 10.1080/01457632.2010.505127 Heat Transfer Fouling: 50 Years After the Kern and Seaton Model ă HANS MULLER-STEINHAGEN Technische Universităat Dresden, Dresden, Germany Fouling of heat exchangers is a chronic problem in processing industries In addition to the appropriate selection of operating conditions and exchanger geometry, there are numerous chemical and mechanical methods to mitigate fouling and to remove deposits from the heat transfer surfaces However, all methods to reduce fouling require some understanding of the mechanisms of the deposition process and of the structure and adhesion of deposits on the heat transfer surfaces Almost exactly 50 years ago, D Q Kern and his co-author, R E Seaton, published a paper attempting to describe the growth of fouling deposits in terms of an unsteady-state heat and mass balance for the heat transfer surface More or less at the same time, the TEMA fouling resistances were published based on operational and anecdotal evidence of fouling for a range of heat exchanger applications These two approaches have since formed the basis for most heat transfer fouling models and heat exchanger designs Increased costs of energy, raw materials, and production downtime have contributed to the growing interest in heat transfer fouling More recently, environmental legislation has put additional pressure on fouling-related CO2 emissions and disposal of cleaning chemicals Despite these efforts, fouling of heat exchangers is still far from been understood in its whole complexity The present paper documents the 2009 D Q Kern Award Lecture in which some selected aspects of fouling research to date have been presented and areas have been identified where significant research and development activities are still required INTRODUCTION confronted with fouling problems, as found in a detailed study for New Zealand [2] To date, the formation of deposits on heat transfer surfaces is the least understood problem in the design of heat exchangers Well-proven codes and correlations are now available for standard heat exchanger design, and computational fluid dynamics simulation can be performed for complex single-phase flow conditions However, all these sophisticated calculations are offset by the current practice of using constant, crudely estimated, experience- or imagination-based fouling resistances or safety margins Even worse is the situation for the prediction of pressure drop While more heat exchangers are taken out of operation due to excessive, fouling-related increase in flow restriction [5], there is virtually no information about the potential effects of deposits on pressure drop Considering the fact that heat exchangers are the workhorse of most chemical, petrochemical, food processing, and power generating processes, this situation is most unsatisfactory The costs of heat exchanger fouling due to oversizing of equipment, maintenance, fluid treatment, additional hardware, additional fuel consumption, and loss of production have been estimated as about 0.25% of the gross domestic product (GDP) of industrialized countries in several studies from the 1980s and early 1990s [2–4] Even today, where “a billion dollars” seems In most industrial processes, heat-exchanging fluids contain certain amounts of dissolved or suspended material or provide conditions favorable for the growth of biological organisms Design and operation of heat exchangers are still to a major extent determined by the process-related formation of deposits on the heat transfer surfaces, i.e., fouling A typical example of a fouled heat exchanger is shown in Figure for the water side of a gas cooler Since the thermal conductivity of such deposits is low, their resistance to heat transfer may well exceed that of the process fluids, resulting in significantly reduced heat exchanger performance [1] As a result, substantial safety margins in the design, pretreatment of hot/cold fluids, and regular cleaning of equipment are usually required Several surveys [2–4] have reported that more than 90% of industrial heat exchangers suffer from fouling problems and must be designed with some allowance for the resulting reduction in thermal and hydraulic performance This is also indicated in Figure 2, which shows the percentage of operating heat exchangers Address correspondence to Dr Hans Măuller-Steinhagen, Technische Universităat Dresden, 01062, Dresden, Germany E-mail: rektor@tu-dresden.de ¨ H MULLER-STEINHAGEN so-called fouling resistances in the calculation of the overall heat transfer coefficient U = U Figure Fouled heat exchanger Courtesy of Hong Kong Towngas to become a common order of magnitude in terms of public expenditure or debts, the costs due to fouling are an excessive burden on industry and economy For a long time, fouling was treated as an incomprehensible and unavoidable curse of any heat exchanger operation Empirical knowledge was developed with respect to the beneficial aspects of additives and operating conditions, but no systematic approaches have been developed to understand the mechanisms of fouling and to affect these mechanisms in a beneficial way Typical examples for this were the addition of potatoes or sawdust to the boiler feedwater to mitigate fouling in early steam generators The present paper documents the 2009 D Q Kern Award Lecture in which some selected aspects of fouling research to date have been presented and areas have been identified where significant research and development activities are still required It has deliberately been focused on fouling during heat transfer to single-phase, liquid fluids, as this represents the majority of investigations to date, and in order to limit the paper to an acceptable number of pages STANDARD DESIGN PROCEDURE FOR FOULING The possibility of deposition on heat transfer surfaces is generally considered in the design of heat exchangers by using + Rf,1 α1 A2 + Rwall + + Rf,2 A1 α2 (1) In Eq (1), α, A, and Rf are the heat transfer coefficients, the heat transfer areas, and the fouling resistances, respectively, for the two heat-exchanging fluids; Rwall is the thermal resistance of the separating wall It is obvious that the frequently used expression “fouling factor” is incorrect, as the effect of fouling is to create an additional thermal resistance The fouling resistance reduces the overall heat transfer coefficient U, and hence leads to the reduction of the heat duty of an existing heat exchanger or to additional surface requirements in the design of new heat exchangers The results of this procedure are heat exchangers with excess heat transfer surface that may (or may not) allow plant operation for an acceptable period of time In the early 1950s the first compilation of fouling resistances was published in the Standards of the Tubular Exchangers Manufacturers Association (TEMA), based on operational and anecdotal evidence of fouling for a range of heat exchanger applications Even though additional proprietary data are available within specialist companies, the TEMA values still form the basis for the design of most heat exchangers, worldwide However, there are several problems with respect to the uncritical use of the TEMA fouling resistances, such as: Their origin and operating conditions are not known The majority of values are for flow of water or hydrocarbons They apply for shell and tube heat exchangers only They not provide any information on the effect on the deposition rate of operating parameters such as flow velocity, fluid temperature, heat flux, and fluid composition They not indicate after which operating time the given fouling resistances are reached They not provide for time-dependent management of fouling resistance In 1990 Chenoweth and co-workers [6] critically reviewed the original TEMA fouling resistances However, only minor modifications have been included in the later editions of the TEMA Standards, mainly due to the lack of suitable industrial data THE KERN AND SEATON MODEL Figure Fouling problems in various heat exchanger types [2] heat transfer engineering While several attempts were made before the 1950s to correlate the fouling-related reduction of heat transfer, none of these equations was based on first principles The decisive change with respect to the analysis of fouling came with the model developed by D Q Kern and his co-author R E Seaton Almost exactly 50 years ago, they published a paper attempting to describe the growth of fouling deposits in terms of an unsteady-state heat and mass balance for the heat transfer surface [7] Together with the vol 32 no 2011 ă H MULLER-STEINHAGEN Figure Predicted fouling resistance as a function of time and the parameter b in Eq (5), according to the Kern and Seaton model [8] layer: Figure Typical fouling resistance versus time curves ˙ r = K2 τw S m TEMA fouling resistances, this approach formed for the next 20 years the basis for most heat transfer fouling models and practical heat exchanger designs Having observed that fouling in industrial heat exchangers often followed a decreasing or even asymptotic trend, as depicted in Figure 3, Kern and Seaton suggested modeling the fouling processes as a balance between opposing transport processes to and from the heat transfer surface, namely, deposition and removal [7], as shown in Figure Therefore, the accumulation of the deposited mass of fouling material with time was written as dRf dm ˙d−m ˙r = =m ρd λd dt dt (2) assuming that the thermal conductivity λd and the density ρd of the deposit remain constant with time and deposit thickness The deposition rate was modeled with a simplified mass transfer correlation as being proportional to the bulk flow velocity and the foulant concentration: ˙ d = K1 V C m (3) Modeling of the removal of already deposited material due to shear forces from the bulk flow was significantly more difficult, and it was assumed that this may be proportional to the wall shear stress and to the thickness of the deposit, which may be a measure for the presence of structural weaknesses in the deposit Figure Deposition and removal of deposit heat transfer engineering (4) Both modeling approaches have since been frequently criticized, extended, and improved [8], as they are obviously based on very simplistic assumptions and ignore several mechanisms that may be responsible for the accumulation of dirt on the heat transfer surfaces Nevertheless, depending on the values of constants K1 and K2 in Eqs (3) and (4), curves (a), (b), and (c) in Figure can be reproduced Combining Eqs (3) and (4) and integrating with respect to time leads to Eq (5): Rf (t) = K3 C V (1 − e−K4 τw t ) = Rf∗ (1 − e−bt ) λd K4 τw (5) which includes the so-called asymptotic fouling resistance Rf ∗ , a value that will be obtained after some period of operation if the removal rate becomes equal to the deposition rate, i.e., the deposit is not very hard and adherent While the value of this asymptotic fouling resistance is approximately inversely proportional to the flow velocity for turbulent flow, the rate at which it is approached increases strongly with flow velocity as shown in Figure THE DEVELOPING YEARS OF FOULING RESEARCH While the Kern and Seaton model was a significant step forward and provided a physically meaningful description of the effects of velocity on deposition and removal, and an equation to model the increase of fouling resistance with time, it nevertheless included two parameters that had to be fitted to the actual fouling problem (i.e., required real operational data) No information was available on how these two parameters may depend on the materials of fouling and their concentration, the structure of the deposit, and operating conditions such as surface temperature and flow conditions Furthermore, the Kern and Seaton model has obvious deficiencies in that it does not include the chemical reactions that are the basis of most fouling processes such as scale formation, crude oil fouling, or food fouling vol 32 no 2011 ă H MULLER-STEINHAGEN It was, therefore, not surprising that little use was made of the Kern and Seaton model in terms of actual heat exchanger design The majority of heat exchangers continued to be designed using the TEMA or proprietary fouling resistances for tubular heat exchangers For compact heat exchangers, the use of the TEMA values would lead to excessive overdesign, making them ineffective and uneconomic Compact heat exchangers are hence generally designed with 15–25% excess surface, to accommodate the fouling-related drop in heat transfer capacity [1] Frequently, design engineers try to compensate for their lack of accurate physical properties of the heat-exchanging fluids or the limited reliability of correlations for the clean heat transfer coefficients (for example, for multiphase and/or multicomponent applications) by arbitrarily increasing the fouling resistance or by multiplying the calculated overall heat transfer coefficients with a “safety factor,” which also increases the fouling resistance It has been reported [4] that the practice of specifying fouling resistances increases the heat transfer surface calculated for clean conditions by 20–300% These findings have been confirmed by a study from Heat Transfer Research, Inc (HTRI), plotting the fouling-related excess area of 2000 recently designed heat exchangers (see Figure 6) In addition to increased equipment cost, oversizing of heat exchangers may even accelerate the rate of deposit formation if it results in low flow velocities or high surface temperatures This unsatisfactory procedure would probably have continued if Taborek et al [9] had not reminded the heat transfer community in 1972 that fouling is the major unresolved problem This important paper triggered a range of investigations, most notably the systematic investigations on cooling water fouling by HTRI together with J Knudsen from Oregon State University, by N Epstein and A P Watkinson at the University of British Columbia, by T R Bott at Birmingham University, and by M Bohnet at the University of Braunschweig In this work, typical fouling processes such as scale formation, particulate deposition, and the growth of biological matter have been investigated using synthetic model fluids under controlled conditions Significant differences have been found with respect Figure Impact of fouling resistance on the design of 2000 shell-and-tube heat exchangers Courtesy of HTRI heat transfer engineering Figure Effect of (a) flow velocity and(b) surface temperature on coolingwater fouling to the effect of the main operational parameters flow velocity and surface temperature on the fouling behavior of the different types of fouling, as exemplified in Figure In 1985 Epstein summarized findings to-date in his famous × matrix [10], which has been adopted to plot Figure Here, for the first time, the different mechanisms of fouling and the different steps in the net deposition process have been brought together and analyzed This has led to a much more systematic and focused approach to the investigation and mitigation of heat transfer fouling, for both practical and fundamental problems For example, fouling is now generally modeled as a consecutive process made up from transport, reaction/attachment, and removal The transport rate is determined according to Eq (6) Figure Epstein’s × matrix: perceived level of understanding (increasing from to 5) versus fouling mechanism and type of fouling [10] vol 32 no 2011 ă H MULLER-STEINHAGEN with the mass transfer coefficient β obtained from the appropriate Sh-Re-Sc relationships: ˙ t = β(cb − cs ) m (6) The subsequent attachment or reaction rate is obtained from Eq (7), with the reaction rate constant kR and the reaction order n: ˙ a = kR (cs − c∗ )n m −E kR = Ke RTS (7) (8) For a second-order reaction such as the formation of CaSO4 and assuming that the reaction rate must be equal to the transport rate, Bohnet and co-workers derived Eq (9) [11]: ⎡ ⎤ β β β ˙ d = β⎣ +(cb −c∗ )− (cb − c∗ )⎦ m + kR kR kR (9) which subsequently has been applied in many investigations [e.g., 12, 13] Research and development efforts during these years have shed considerable light into the most common fouling mechanisms, such as crystallization, particulate, biological, corrosion, and chemical reaction fouling Numerous models for fouling during convective heat transfer have been derived based on these approaches, to correlate available data It is not the aim of this paper to summarize in detail the vast area of heat transfer fouling research, or to provide a historical treatise of it The latter has already been done in a laudable way for the period up to 1990 by Somerscales [14] Significant progress has been made in this period of time by following several approaches in parallel, such as: • Detailed investigation of fouling mechanisms, increasingly also for gas-side fouling and for fouling during boiling • Empirical development of mechanical and chemical on-line fouling mitigation techniques, such as sponge ball systems, wire brush systems, and chemical additives [15] • Development of advanced mechanical and chemical cleaning systems and procedures for heat exchangers [15] • Development of heat exchanger types with reduced fouling rates, for example, the fluidized bed heat exchanger [16] • Development of guidelines for heat exchanger design, e.g., by HTRI It is, however, noteworthy that most of the investigations published during this period have been obtained for ideal (or “model”) fluids, and not many research results, and hardly any of the numerous deposition models, have found their way into practical heat exchanger design and operation heat transfer engineering FOULING BECOMES AN INTERNATIONALLY ACCEPTED RESEARCH TOPIC Following up on the increasing academic interest in heat exchanger fouling, the first conferences targeted specifically at this topic were organized in Guildford (1979) [17], Troy (1981) [18], and Alvor (1987) [19] These pioneering meetings contributed much to the formation of a “fouling research community” with significant coherence and interaction Consequently, United Engineering Foundation Conferences (now Engineering Conferences International) decided to initiate a series of international meetings on fundamental and technological aspects of heat exchanger fouling Seven highly successful meetings have been held in San Luis Obispo, CA (United States, 1994), Castelvecchio Pascoli (Italy, 1997), Banff (Canada, 1999), Davos (Switzerland, 2001) [20], Santa Fe, NM (United States, 2003), Kloster Irsee (Germany, 2005), and Tomar (Portugal, 2007) These conferences attracted an increasing number of participants from industry, research organizations, and universities Papers presented at each of these conferences have been published in the respective conference proceedings and probably provide the most comprehensive overview of the state of the art of this complex subject The full proceedings of the 2003–2007 meetings can be downloaded from http://services.bepress.com/eci/heatexchanger For organizational reasons, the ECI fouling conference series was continued from 2009 onward as the EUROTHERM Seminar Series, starting with the 2009 conference in Schladming (Austria) Proceedings of and information about this conference can be found at http://www.heatexchanger-fouling.com The next conference in this series will be held in Crete (Greece, June 2011); see the website just given For engineers working in the area of food processing, an excellent series of bi-annual conferences at Cambridge University (England), organized by Wilson, Fryer, and Hastings, provides current developments in fouling and cleaning in that industry FOULING RESEARCH REACHES MATURITY Increasing costs of energy, raw materials, and production downtime have contributed to the growing interest in heat transfer fouling More recently, environmental legislation has put additional pressure on fouling-related CO2 emissions and disposal of cleaning chemicals [21] For immediate benefits, fouling task forces including representatives from major international process engineering companies, heat exchanger design and construction companies, and suppliers of chemical and mechanical fouling mitigation measures have been established by ESDU and HTRI to compile best practice guides for heat exchanger design and operation To date, very detailed reports have been prepared for crude oil [22], seawater [23], and freshwater [24] Based on almost 50 years of experience, HTRI has developed a design methodology that yields smaller, more cost-effective vol 32 no 2011 ă H MULLER-STEINHAGEN shell-and-tube heat exchangers with extended run times between cleanings [1, 25] While this methodology has, so far, only been validated for crude oil processing, its rigorous approach can be taken as an example for other fluids and heat exchangers types Using this methodology, only a small design margin may be added to the design to address design uncertainties Rarely is this margin in excess of 30% These experience-based approaches are extremely useful for appropriate design and operational mitigation of standard fouling problems However, they cannot be extrapolated to individual fouling problems or lead to a general solution for the reduction or even elimination of fouling For this, more fundamental research and development are required Some of these efforts are described in the following It is obvious that the general approach to improved understanding of deposition mechanisms has been moving from macro-scale to micro-scale to molecular level, and that advanced computational tools are increasingly finding their way into fouling analyses Whole Plant Modeling Heat exchangers are rarely stand-alone units unaffected by upstream and downstream processes Hence, the conditions leading to and resulting from fouling are the result of complex interactions within a range of equipment, including, e.g., heat exchangers, settling tanks, reactors, mixers, and evaporators In many bulk material processes, addition of chemicals to mitigate fouling is not possible due to product requirements, and significant changes to the existing hardware cannot be afforded However, it may still be possible to reduce the formation of deposits or improve the economy of the process, if appropriate operating conditions and/or operating schedules are selected This requires understanding of both the local fouling conditions and the overall plant operation Ideally the optimization processes should include the following steps: Analysis of plant operating data Analysis of deposits and of foulant solubility behavior Laboratory experiments to determine the effect of operating conditions (concentration flow velocity, bulk and heat transfer surface temperature) Laboratory experiments to determine possible fouling mitigation methods (e.g., seeding, turbulence promoters, fluidized bed) Modeling of fouling process Limited number of plant measurements on a slipstream to confirm the validity of the laboratory data and of the fouling model for actual plant operating conditions Heat exchanger model to predict local and overall temperatures and fouling rates Comparison of heat exchanger model with plant operating data Overall plant model with/without fouling related deterioration of heat transfer heat transfer engineering 10 Use of model to determine optimum operating conditions/procedures and to investigate the effects of plant modifications to maximize throughput or minimize operating costs This model can also be used for model-based process control and environmental impact studies An effective approach to provide the information required for optimizing plant operation and plant layout requires a combination of fundamental and industrial studies It is unlikely that all the information just specified can be collected, due to financial and/or time constraints The important criterion is, however, that some plant verification for the developed fouling/operating model is available This general approach has been applied successfully in several comprehensive studies: • • • • Kraft black liquor in the pulp and paper industry [26, 27] Bayer liquor in bauxite refineries [28, 29] Phosphoric acid plants for fertilizer production [30, 31] Sulfuric acid recovery plant in a titanium oxide extraction process [32] • Crude oil preheat train [33, 34] In the processes just listed, heat exchangers generally suffer from severe fouling problems, leading to significant limitation in plant operation and high additional costs The investigations have been performed in close collaboration with industry and resulted in significant gain in knowledge and industrial benefits Results of these investigations have been implemented into design and operation of the investigated plants, or are further investigated in pilot-plant studies Actual and potential future savings are in the order of many millions of dollars, providing a significant payback on the investments for the detailed studies A typical example is shown in Figure 9, indicating the predicted extension of operating time of a sulfuric acid concentration unit if operated at higher temperatures and with variable flow velocity [32] The dashed line shows the original operation with a constant acid flow velocity of 2.5 m/s and increasing heating steam temperature to overcome the effects of fouling The solid line shows the suggested operation with constant maximum steam temperature of 200◦ C and variable flow velocity from m/s to 2.5 m/s With the second option, the run time could be increased from 150 hours to 275 hours, with only minor plant modifications Neural Networks Despite increased attention during the past decades, correlations recommended for heat exchanger fouling can only be applied to a limited number of idealized deposition processes, while they lead to massive uncertainties and inaccuracies for industrial fluids These drawbacks may be the result of: • Nonlinearity of the fouling process • The character of the fouling process, which is unsteady-state with potentially high fluctuation vol 32 no 2011 ă H MULLER-STEINHAGEN 204 202 2.5 200 2.0 st T [°C] v [m/s] * 198 1.5 11 11 additional operation time with the new process control configuration -ln (Rf), experiment Velocity Steam Temperature 12 12 -ln (Rf), experiment 3.0 10 * 196 Q=22000 W 5 10 -ln (R*f), prediction 1.0 6 10 11 12 10 11 12 -ln (R*f), prediction Figure 11 Measured versus fitted asymptotic fouling resistances, using a mechanistic model (left, mean average error 38%) and a neural network (right, mean average error 15%) 194 0.5 25 50 75 100 125 150 175 200 225 250 275 Time [h] date: Figure Operational time before shut-down for cleaning of a sulfuric acid evaporator, operating either with constant flow velocity (dotted line) or constant steam temperature (solid line) • The large number of variables and different mechanisms • The lack of rigorous understanding of the underlying mechanisms • The inherent inadequacy of conventional regression methods to correlate experimental data with an ill-distributed parameter variation The use of artificial neural networks is a pragmatic alternative to address many industrial fouling problems with significantly better accuracy than conventional parametric regression models This can be done by using neural networks as an interpolation tool within a range of experimental results (black-box approach) [35], or as a hybrid approach where the neural network is used in combination with prior knowledge (PK) of the process [36, 37], as shown in Figure 10 This “prior knowledge” may, for example, be the experience that the fouling rate generally increases with surface temperature and/or decreases with flow velocity The results of the second method are found to be more reliable than those provided by the first method The following promising results have been found in the very limited number of investigations that have been published to • Experimental data could be correlated significantly better with a suitable neural network than with the models recommended by the original authors This is clearly demonstrated in Figure 11 for cooling water fouling data • Satisfactory capability of the network for those areas (i.e., induction period and high surface temperature) where not enough information about the underlying phenomena and/or insufficient experimental data are available • The reliability of the resulting networks was confirmed when they were applied to those data that had not been used before • Once converged, the resulting network is a simple and small program that even inexperienced users can apply or that can be embedded into any heat exchanger design software Despite these promising results, several questions still remain unanswered, which will have to be addressed if such techniques are to be pursued for industrial applications: • Validation for process fluids where the number of input variables is large, and with poorer understanding of the basic phenomena which govern the fouling process A typical example for this could be crude oil fouling • Application to cases where the dominant mechanisms change with operating conditions and/or time One such example is crystallization fouling, which is diffusion-controlled at very low velocities and reaction-controlled at higher velocities • The predictability of the network may severely deteriorate if data bases are ill-distributed and much weight of the data is concentrated only in specific domains • Poor extrapolation of the resulting network beyond the range of learning data • Inclusion of discrete variables such as heat exchanger geometries into neural network modeling CFD Modeling Figure 10 Hybrid neural network (PK = prior knowledge) heat transfer engineering Fouling in industrial heat exchangers is strongly dependent on local concentrations, temperatures, and shear rates This is exemplified in Figure 12, which shows the inlet zone of a gas vol 32 no 2011 ă H MULLER-STEINHAGEN Figure 12 Negative effect of excessive inlet baffle spacing on deposit formation cooler with cooling water flowing on the shell side Very severe deposit formation was found in the area between the last baffle and the tube sheet; further away, fouling was significantly less Looking at the pictures on the left side of Figure 12, one recognizes the large spacing between tube sheet and first baffle, as compared to the subsequent baffle–baffle spacing The large gap leads to a significant reduction in flow velocity and to significant flow maldistribution, both reducing local shear rates and increasing local wall temperatures, even to the extent that undesirable local nucleate boiling may have occurred, which significantly increases the deposition rate [38] It is obvious that the relatively simple analytical models developed for heat exchanger design and fouling not provide the required information about local conditions However, numerical simulation of flow and temperature distribution using commercial computational fluid dynamics (CFD) software has now reached a quality where it is possible to identify critical areas in industrial heat exchangers in terms of hot spots or low velocity zones Detailed modeling of shell-side flow of large shell-and-tube heat exchangers has been performed, including leakage streams between baffles, tubes, and shell [39] This is an area of work with tremendous potential, not only for shell-and-tube heat exchangers, but also for compact heat exchanger types [40] First attempts have been undertaken to model the local growth of deposits in addition to the local temperatures and shear rates [41] The inclusion of additional mass transfer mechanisms and heat transfer engineering reaction kinetics increases the computational effort enormously, but this will be resolved with the advent of increasingly powerful microprocessors More importantly, such modeling approaches depend on the quality of models for the local deposit formation, which are still under investigation Heat Transfer Surface–Deposit Interaction Numerous methods have been developed to remove depositforming constituents from heat exchanging fluids, to increase their solubility in these fluids, or to clean heat transfer surfaces once they have fouled While the first are highly specific to the composition/chemistry of the fluids, the last of these only deals with a problem after it has occurred From a technical point of view, it would be much more desirable if heat transfer surfaces could be developed on which deposits not stick at all In general, maximum adhesion occurs in interacting systems that undergo a maximum decrease in surface energy, and poorest fouling adhesion should occur on materials that have low surface energies Surface coatings with organic polymers such as polytetrafluoroethylene (PTFE) and Saekaphen have a very low surface energy, but they are mainly used to avoid corrosion as the coatings themselves provide a significant additional resistance to heat transfer While the durability of the coatings increases with thickness, this has the inverse effect on heat transfer vol 32 no 2011 ă H MULLER-STEINHAGEN Therefore, the coating thickness should be kept as thin as possible These conditions can be met with modern surface-coating techniques such as ion beam implantation, magnetron sputtering, and autocatalytic Ni–P–PTFE coatings However, results obtained with a wide range of surface coatings have been contradictory [42], indicating increased or decreased deposit formation on surfaces with low surface energy, as compared with standard stainless steel Hence, there is a lack of understanding of the principal interacting forces between depositing material and metallic substrate Since the effect of gravitational forces on deposition is usually negligible, these forces consist of a Lifshitz–van der Waals (LW) interaction component, electrostatic double-layer component (EL), Lewis acid–base component (AB), and Brownian motion component (Br) Equations to predict these interactions energies may be found in [43] The total interaction energy ETOT between a deposit and a metal surface can be written as the sum of the respective interaction terms: E TOT = E LW + E EL + E AB + E Br (10) It has been suggested, e.g., by Visser [44] that the balance between all possible interactions between a deposit and a metal surface determines whether a system will foul or not; i.e., adhesion/fouling will take place when ETOT is negative However, experimental evidence has shown that under certain conditions some systems may foul, even though the total interaction energy E132 TOT between deposit (1) and metal surface (2) in fluid (3) is positive—for example, if the initial E132 TOT is positive (i.e., repulsive), but the substantial cohesive energy E131 TOT between the foulant particles leads to coagulation into larger particles It is also not necessarily correct that the system will foul if the total interaction energy E132 TOT is negative For example, if the cohesive energy E131 TOT exactly equals the adsorption energy E132 TOT, the energy characteristics of the heat transfer surface will be the same as that of the foulant particles This means that the colloidal particles in the wall-near boundary layer may not attach to the surface or coagulate to each other, but remain suspended in the solution in some sort of dynamic equilibrium Therefore, the cohesive energy E131 TOT will have to be taken into account in the investigation of fouling behavior, and particularly during the fouling induction period Based on these findings, the following criterion to determine whether a system will foul or not has been suggested in [45]: T OT E 131 − T OT E 131 − T OT > 0, E 132 T OT E 132 =0 fouling possible either, immediately or later (11) no or minimal fouling (12) If the Lifshitz–van der Waals forces are dominant, the surface free energy γs,min at which fouling is minimal can be calculated from: √ γ S,min = (1/2) γ1L W + γ3L W (13) heat transfer engineering Figure 13 Surface free energy versus asymptotic fouling resistance for various surface materials Experimental data from Făorster et al [46] with 1L W and γ3L W being the surface free energies of deposit and the liquid, which can be determined by standard measurements For the deposition of CaSO4 from aqueous solutions, the free surface energies of crystalline CaSO4 and of water are γ1L W = 35.5 mN/m and γ3L W = 21.8 mN/m According to Eq (12), a heat transfer surface with a surface free energy γ2L W = γs,min = 28 mN/m should have minimum fouling This is confirmed by comparison with the experimental data by Făorster et al [46] shown in Figure 13 The location of the fouling minimum coincides with the experimental findings for the DLC-F sputtered surface, for which also a surface free energy of 28 mN/m was measured Similar agreement between predicted minimum fouling surfaces and measurements has also been found for deposition of milk and microbes While the findings just described are promising, they are nevertheless only a first step forward There is experimental evidence that other effects, such as surface roughness, aging, and temperature, will also have significant effects on deposition rate and asymptotic fouling resistance Molecular Modeling While the modeling of surface–deposit interaction may provide some information about the sticking propensity of various deposits on various surfaces, it still depends on the measurement of lumped parameters taking into account several molecular effects happening on the interface between surface and deposit The weakness of this approach is evidenced by the poor correlation of surface energy and fouling rate or fouling delay time, as reported in [42] Here it was found that stainless-steel surfaces implanted with hydrogen ions suffered considerably more from fouling than the original steel surfaces, which in turn fouled significantly faster than stainless-steel surfaces implanted with fluorine ions In both cases, the surface energies of the implanted surfaces as well as their polar components are almost identical This phenomenon has been analyzed by Rizzo et al [47], who found that the induction period of the nucleation process of CaSO4 crystallization fouling could not be correlated with results from surface energy measurements Instead, a linear relation between the slope of the ln(induction period) versus ln–2(supersaturation ratio) plots and the electronegativity of the implanted ions was observed, as shown in Figure 14 This empirical result sheds some light on the contradictory results reported in [42]; it nevertheless does not allow vol 32 no 2011 70 X SUN ET AL foam fracturing fluid was proportional to the decrement of foam quality, which peaked at a quality value of 25% and minimized at a quality value of 75% This phenomenon was completely reversed in the Watkins et al [4] study Another substantial distinction between these two research reports is that the viscosity in the Harris and Heath study [1] was two to three times higher than that in the Watkins et al study [4] Further investigations, especially those where there is urgent need for field application, should be carried out, for the simple reason that the viscosity of cross-linked foam fracturing fluids may be influenced by numerous factors Convective heat transfer is another important property of foam fluid, as lots of chemical and physical reactions would occur under the condition of heating However, little literature about this research area has been published Actually, heat transfer between the stratum and the foam fracturing fluid exists during the fracturing process Considerable temperature variations of foam fracturing fluid may occur from the well head to the bottom hole Harris and Heath [1] and Freeman et al [5] found that the stability of foam texture and the network structure of the external cross-linked guar could be influenced by the temperature They also found that the viscosity of cross-linked foam fracturing fluid was inversely significantly proportional to the temperature For that reason, it is especially important to study the temperature field and the convective heat transfer coefficients of cross-linked foam fracturing fluids to obtain their accurate features and improve the fracture treatment design To date, more than 40 heat transfer correlations and manifold experimental data concerning forced convective heat transfer during gas–liquid two-phase flow in vertical and horizontal pipes have been proposed and published [6]; however, they are mainly confined to two-phase flow of gas and Newtonian liquid The study of cross-linked foam fracturing fluid of a supercritical gas and non-Newtonian two-phase fluid is still in the rudimental stage Thus, there is an urgent need to clarify the heat transfer mechanisms of two-phase fluid In this article, the rheology and heat transfer characteristics of the borate cross-linked guar and the borate cross-linked foam fracturing fluid have been investigated and analyzed through the experimental study under downhole conditions Figure Large-scale test loop nitrogen cylinder gas flowmeter gas booster air compressor air container guar generator gear wheel pump flowmeter plunger pump 10 check valve 11 foam generator 12 electrical heater 13 EJA differential pressure transmitter 14 segregator 15 valve 16 manometer 17 thermocouple nitrogen cylinder gas flowmeter gas booster air compressor air container 6.guar generator gear wheel pump flowmeter plunger pump 10 check valve 11 foam generator 12 electrical heater 13 EJA differential pressure transmitter 14 segregator 15 valve 16 manometer 17 thermocouple loop The whole system pipeline is about 20 m long, the heating section is about 12 m before fluid enters the test section, and the time of residence in the loop is about 35 to 75 s The pH value of water was initially adjusted to by adding acetic acid Base guar fluids were batch mixed with 0.55% hydroxypropyl guar powder, 0.5% surfactant, 0.1% fungicide, and 0.12% sodium carbonate, thus raising the pH value to 10 due to cross-linking formation The 0.5% cross-linker and 0.1% foaming agent were then mixed into the guar, which was circulated for in order to reach equilibrium Then the cross-linked guar fluids were pumped into the foam generator by a plunger pump Meanwhile, nitrogen was compressed and pumped into the foam generator by a gas booster operated by a valve and measured by a flowmeter Therefore, nitrogen foam fracturing Table Instruments and transducers used in the experiment Classification EXPERIMENTAL SETUP The experimental setup of the study is illustrated in Figure The flow loop could generate foam fracturing fluid at temperature above 120◦ C and pressure above 60 MPa The equipment consists of the test sections of rheology and convective heat transfer The pipes of both sections have mm inner diameter, 16 mm outer diameter, and 1000 mm length Pressure drop and temperature are monitored by an EJA differential pressure transmitter and K-type thermocouples, respectively, and the accuracy and resolution of each sensor are illustrated in Table As shown in Figure 1, the flow loop is not a real recirculating heat transfer engineering Differential pressure transmitter Manometer Thermocouple Type Remark EJA differential pressure Measure range: 0–100 kPa transmitter Current output: 0–20 mA Accuracy: 0.5 Diaphragm manometer Measure range: 0–60 MPa Accuracy: 0.5 K-type thermocouple Maximum temperature: 900◦ C Maximum cause error: ±0.25% Maximum allowable error: ±2.25◦ C vol 32 no 2011 X SUN ET AL fluids were generated when the mixture of cross-linked guar fluids and compressed nitrogen passed through the foam generator The foam quality of 55%, 65%, or 76.5% was generated by displacing the appropriate volume of liquid from the volume of the loop The temperature of the foam was increased by the electrical heater, and the pressure was adjusted by regulating the back pressure valve at the end of the flow loop After the initial equilibration, the shear rate was adjusted to a required value by regulating the rotate speed of the plunger pump All data were saved on the computer for later analysis THEORETICAL ANALYSIS or µ = K γ˙ n−1 (1) The flow pattern in the pipes is supposed to be laminar flow, and the effect of wall slip [7, 8] is neglected, and the fluid is in a stable state, and fully developed flow has been ensured in these pipes, and through using a tube viscometer, the shear rate at the wall can be calculated by the following equation: 8U 3n + D 4n And the shear stress at the wall is obtained by: γ˙ w = (2) D p (3) 4l Insert Eq (2) and Eq (3) into Eq (1) and adopt logarithmic form, yields: τw = lg pD 4l = lg K the thermal conductivity of the pipe wall separating them, and the specific heat of the working fluid In this system, if the convective heat transfer coefficient of coolant h is given, the convective heat transfer coefficient of the working fluid is then: h1 = 3n + 4n n + n lg 8U D (4) By measuring the pressure drop p and flow flux Q, the flow curve representing the measured stress Lp D4 versus the Newtonian shear rate 8U can be obtained Through calculating D the line slope tan θ and intercept B, we can obtain the parameters n and k: n = tan θ k = 10 B (5) n ( 3n+1 4n ) (6) The characteristics of convective heat transfer were investigated in the heat exchanger As the convective heat transfer coefficient of the working fluid hinges on the inlet and outlet temperatures of the working fluid and coolant, the amount of heat loss of the working fluid and that gained by coolant can be measured based on the temperature difference The amount of transferred heat is determined by factors such as the temperature difference between the working fluid and coolant, the area and heat transfer engineering A1 k − A2 h (7) where A1 is the internal surface area of the pipe, A2 is the external surface area, and k is the overall heat transfer coefficient, which is calculated by the following formula: k= The behavior of cross-linked guar and foam fracturing fluids can be described as non-Newtonian fluid and power-law fluid at a high shear rate The constitutive equation for a power-law fluid is: τ = K γ˙ n , 71 t1in +t1out − t2in +t2out (8) where, t1in and t1out are the inlet and outlet temperatures of the working fluid, respectively; t2in and t2out are the inlet and outlet temperatures of the coolant, respectively; and is the heat flux According to our research, the convective heat transfer coefficient of coolant is determined by the water which was used first as the working fluid In order to describe the rehology and convective heat transfer properties of foam fracturing fluid, it is necessary to measure the temperature, pressure drop, and flow flux in this experiment Owing to measurement errors and some other factors, the uncertainties of measurement are unavoidable We used nonparametric statistics to evaluate the standard uncertainty of measurements in this experimental research The main causes of uncertainty that affect experimental measurements include indicative error, instrument lag, repeatability error, environment impact, and instrument resolution [9] The standard synthetic uncertainties of the temperature, pressure drop, and flow flux are 2.79◦ C, 0.3%, and 0.23%, respectively RESULTS AND DISCUSSION Since the viscosity of the fluid is significantly influenced by the rheology of the external phase of the fluid [1], the viscosity of single-phase liquid and foam fracturing fluid and their correlations were investigated Cross-linked guar is a macromolecule polymer containing a cross-linked structure Unlike conventional lineal guar, which only suffers thermal thinning, cross-linked guar suffers both chemical and physical degradation The rheology of crosslinked guar is deeply dependent on the pH value of fluid and polymer concentration In order to study the relationship between viscosities of the external phase and foam and the characteristics of heat transfer, the experiment was conducted by stabilizing the concentration of polymer and cross-linker and altering temperature under electrical heating at a given shear rate Figure shows the apparent viscosity of borate cross-linked guar versus temperature at shear rates of 276.3 s−1, 414.5 s−1, and 552.6 s−1, respectively Obviously, borate cross-linked guar vol 32 no 2011 72 X SUN ET AL Figure Apparent viscosity of borate cross-linked guar at different shear rate and temperature is a shear thinning fluid In the shear field, the cross-linked structure and the molecular weight of borate cross-linked guar are in rapid exchange equilibrium The guar molecular structure becomes tighter and smaller as shear rate rises, resulting in a decrease in the energy dissipation and viscosity When the temperature was below 50◦ C, the viscosity decreased sharply with temperature increment as shown in Figure 2, which suggested that the cross-linking reaction was substantial and would result in a great chemical degradation caused by temperature increment However, when the temperature was higher than 50◦ C, apparent viscosity was almost the same as that of non-cross-linked guar, implying that the guar was no longer cross-linked and the viscosity had reverted to the conventional value During the borate-cross-linking reaction process, the boron compound dissociated in alkaline solutions and formed two primary matters: boric acid, B(OH)3 , and monoborate ions, B(OH)4 −, which are only present in low-concentration solution of the compounds As the concentration of boron compound increases, the cross-linker solution may dissociate further to form polyborate ions: triatomic borate, B3 O3 (OH)4 −, ions and pentatomic borate, B5 O6 (OH)4 −, ions The boron compounds dissociated based on the following equilibrium reactions [10]: B(O H )3 + O H − → B(O H )− pK = 9.23 Figure Effect of temperature on fluid behavior index of borate cross-linked guar the dissociation of boric acid into borate ions is an exothermic reaction The dissociation of equilibrium of Eq (9) slows down as the temperature increases, which contributes to the decreasing concentration of the monoborate ions at a given boron concentration In this way, the temperature increment would cause the decrease of the number of cross-linked nodes that link the guar polymer chains, of the molecular weight, and of the viscosity of the borate cross-linked guar According to our research findings, the cross-linker almost lost its function when the temperature was higher than 50◦ C However, this turning point of temperature would vary with different pH values of fluid, with higher final fluid pH value [1] requiring a higher temperature At a high shear rate, the borate cross-linked guar behaves like a power-law fluid The fluid behavior index of borate crosslinked guar, n g , and fluid consistency index of borate crosslinked guar, K g , can be deduced from Eq (4), which are shown in Figures and 4, respectively As is shown in Figure 3, the fluid behavior index of borate cross-linked guar, n g , increased with the temperature increasing and reached the maximum value of 0.52 The correlation between the fluid behavior index of borate cross-linked guar and (9) − 2B(O H )3 + B(O H )− → B3 O3 (O H )4 + 3H2 O pK = −2.33 (10) − 4B(O H )3 + B(O H )− → B5 O6 (O H )4 + 3H2 O pK = −2.28 (11) The cross-linking ion in the reaction between the boron compound and the guar polymer is the monoborate ion However, heat transfer engineering Figure Effect of temperature on the fluid consistency index of borate crosslinked guar vol 32 no 2011 X SUN ET AL 73 temperature can be expressed as: ng = n g1 − n g2 + n g2 + (T /T0 )m (12) where n g1 is the lower limit value of n g for decreasing temperature, n g2 is the upper limit value of n g for increasing temperature, m is a constant showing the extent to which the temperature influences n g , and T0 is a reference temperature In this research, n g1 = 0.272, n g2 = 0.519, m = 10.5, and T0 = 44.2◦ C According to Figure 4, the fluid consistency index of borate cross-linked guar, K g , decreased with the temperature increasing and reached a minimum value of 0.516 The correlation between the fluid consistency index of borate cross-linked guar and temperature can be expressed as: K g = K g2 + K g1 e−(T −T1 )/m (13) where K g2 is the lower limit value of K g for increasing temperature, and K g1 is the initial limit value of K g at an initial low temperature,T1 ; m is a constant showing the extent to which the temperature influences K g In this research, K g1 = 83.73, K g2 = 0.52, m = 8.5, and T1 = 10◦ C Thus, by substituting Eq (12) and Eq (13) into Eq (1), the viscosity of borate cross-linked guar at different temperatures can be expressed as: 0.27−0.52 µg = K g γ˙ n g = [0.516 + 83.73e−(T −10)/8.5 ]γ˙ 1+(T /44.17)10.5 +0.52 (14) As a gas phase dispersed in an aqueous phase that contains surfactants, foams exhibit both elastic and viscous behaviors The volume fraction of gas phase is designated as foam quality When the surface-active substance is added in the liquid phase, it is adsorbed spontaneously at the foam surface, reducing the surface energy and surface tension, generating colloidal structures and preventing coalescence of bubbles even if they are approaching one another [11] Moreover, with sufficient concentration the large yield stress forces small bubbles to disperse The elastic behavior of foams can be attributed to the surface tension existing in the thin liquid films [12] The viscoelasticity of foams can be described by the capillary number, Ca = dµγ˙ σ [13] The higher the capillary number is, the more deformable the interfaces become When Ca = 0, foams are purely elastic and solid-like and can be considered to be dry foams Under the condition of sufficient large foam quality, except a thin bubble-depleted layer in the vicinity of the wall, a relatively uniform distribution of bubbles in the radial direction of the pipe can be realized because of the interaction of the bubbles The thickness of this layer is on the order of the bubble diameter At a high shear rate, the bubble diameters are small compared to the pipe diameter Within the limit of small d/D, the presence of the particle-depleted layer can be neglected In addition, the high viscosity of the liquid phase can prevent turbulence and reduce the effect of secondary flow resulting from free convection heat transfer engineering Figure Effect of temperature on fluid behavior index of borate cross-linked foam fracturing fluid at different foam qualities Thus, the foam fracturing fluid can be treated as a single-phase fluid Foam rheology has been widely studied in the past years Because of the broad range of foam qualities, foam fluid has been described as a yielding pseudoplastic fluid Mitchell [14] studied aqueous foam in small capillary tubes and identified it as a Bingham plastic-type fluid with a substantial dependence on foam quality Reidenbach et al [15] used a recirculating loop to conduct experimental tests with aqueous and guar foam fluids, using N2 and CO2 as internal phase They used the Herschel–Bulkley model to describe guar foam fluids, and the classic Bingham plastic model for aqueous foam fluids In their research, the flow behavior index of foam fluids was assumed to be the same as that of the liquid phase All of these studies, however, were carried out at a low pressure of 6.84 MPa Actually, fluid contains a high percentage of gas, which greatly influences the rheology of the fluid The compression dependence and temperature dependence of gases imply that the rheology of foam fluids must be measured at high temperature and high pressure, the ordinary condition in field application In this research, the pressure was 30 MPa It was found that the borate cross-linked foam fracturing fluid could be described by a power-law model at high shear rate The fluid behavior index of foam, nf , and fluid consistency index of foam, Kf , are shown in Figures and 6, respectively, for different temperatures and foam qualities According to Figure 5, the fluid behavior index of borate cross-linked foam fracturing fluid, nf , was proportional to the temperature and inversely proportional to the foam quality In contrast, the fluid consistency index of borate cross-linked foam fracturing fluid, Kf , was inversely proportional to the temperature and proportional to the foam quality, as shown in Figure When foam quality is lower than 52%, bubbles will flow adequately and will not interfere with each other to generate a significant increase in viscosity until they are adequate to occupy approximately 52% of the total volume [16] As the internal phase volume reaches 52%, more bubble interference occurs, causing substantial increases of foam viscosity, yield vol 32 no 2011 74 X SUN ET AL where: Y0 = 1.178 − 0.00855T T < 50◦ C Y0 = 0.685 − 0.00617T T > 50◦ C R0 = 5.543 − 0.023T Figure Effect of temperature on fluid consistency index of borate crosslinked foam fracturing fluid at different foam qualities points, and further changes in the packing arrangement to a hexagonal structure at 74% of the total volume As more bubbles are added and the internal phase volume reaches 74%, the packing arrangement cannot be altered since the monodisperse rigid spheres cannot occupy more than 74% of the volume At this point, the spherical bubbles undergo shape distortion and polyhedron arrangement, causing further increase in the foam viscosity It is also necessary to investigate the correlation between the viscosities of the external phase and the borate cross-linked foam fracturing fluid For borate cross-linked guar, the viscosity can be expressed as: µg = K g γ˙ n g −1 (15) And for borate cross-linked foam fracturing fluid, the viscosity can be expressed as: µ f = K f γ˙ n f −1 The rheology of borate cross-linked foam fracturing is influenced by the temperature, which mainly affects the rheology of the external liquid phase In addition, the liquid drainage in the lamella can gradually be accelerated by the temperature, causing a decrease in the film thickness and therefore bubble destabilization [16] Nevertheless, based on the preceding equations for A, R1 , and Y0 , there should be two different mechanisms concerning the influence of temperature upon the rheology of borate cross-linked foam fracturing fluid When the temperature is lower than 50◦ C, surfactant diffusivity in the cross-linked guar is trivial, surfactant concentration at the bubble surface is unequal, and the resulted intensified Marangoni effect will cause a great rise of foam viscosity In consequence, the first term of Y0 for temperature lower than 50◦ C is 1.178 However, when the temperature is higher than 50◦ C, the cross-linker is almost disabled Surfactant in the non-cross-linked guar diffuses freely, and the Marangoni effect is weakened Equations (18) and (19) can be used for borate cross-linked foam fracturing fluid in the whole range of foam quality But when the value of foam quality is above 93% to 97%, there is a tendency that the two-phase fluid would be inverted into mist [17] The liquid in such a fluid is an internal phase while the gas is an external one, which implies that the preceding equation is not proper any more The Reynolds number for power-law fluid in pipes could be calculated by: (16) Re = Dividing Eq (16) by Eq (15): µf K f γ˙ n f −1 K f n f −n g γ˙ = = µg K g γ˙ n g −1 Kg (17) where n f − n g can be written as: n f − n g = Ae( − )/R1 (18) In our research, = 0.74, which is the foam quality where the spherical bubbles undergo shape distortion and polyhedron arrangement Also: D n U¯ 2−n ρ K (20) Values of the flow Reynolds number for borate cross-linked foam fracturing fluid are shown in Figure with a pressure of 30 MPa and a shear rate of 414.5 s−1 From Figure 7, we can see that the values of the flow Reynolds number are no more than 5.5, which indicates that the flow pattern of borate cross-linked foam fracturing fluid must be laminar flow Convective Heat Transfer ◦ A = 0.049 + 0.00094e T /11.46 T < 50 C A = 0.116 + 19.5174e−T /7.67 T > 50◦ C R1 = 0.13 T < 50◦ C R1 = 0.13 + 0.001T T > 50◦ C K f /K g can be written as: Kf = Y0 + (1 − Y0 ) e R0 Kg (19) heat transfer engineering The convective heat transfer coefficient of borate cross-linked guar is shown in Figure Differing from Newtonian fluid, the convective heat transfer coefficient of borate cross-linked guar decreased with the increased temperature at first, and then increased with the temperature increment after reaching a minimum point Two factors are essential in the process of convective heat transfer of borate cross-linked guar: the shear rate and thermal conductivity of the borate cross-linked guar near the wall vol 32 no 2011 X SUN ET AL 75 For non-Newtonian fluid in laminar flow, the average convective heat transfer coefficient can be calculated by [18]: h = 1.61Re 1/3 Pr × Figure Reynolds number of borate cross-linked foam fracturing fluid at different temperatures and foam qualities with the shear rate of 414.5 s−1 For convective heat transfer, the most important heat conduction process occurs near the wall The convective heat transfer coefficient of fluid will be considerably influenced by the velocity gradient of fluid on the wall, especially for a nonNewtonian fluid The shear rate on the wall can be calculated by Eq (2) Figure shows that the fluid behavior index of cross-linked guar is within the range of 0.272 to 0.519, where the shear rate at the wall is highly sensitive Below 50◦ C, n g increased and γ˙ w decreased rapidly with increasing temperature, which will weaken the heat conduction process near the wall and the total heat transfer However, for temperatures higher than 50◦ C, n g varied extremely slowly, which would not be higher than the value of 0.55, and almost no heat transfer reduction could happen with respect to the fluid behavior index The heat conduction process near the wall and the total heat transfer would be improved as the thermal conductivity of borate cross-linked guar was proportional to the temperature 3n + 4n 1/3 D l 1/3 1/3 Ka Kw 0.14 λ D (21) However, the convective heat transfer coefficient calculated by Eq (21) is much smaller than the experimental value Naturally, it is assumed that thermal conductivity of fluid is constant or temperature-dependent In fact, most of the previous measurements of thermal conductivity were conducted in stationary flow fields and the dependence on temperature was well considered and tabulated Lee and Thomas [19], however, recently measured the thermal conductivity of both aqueous Separan solutions and CMC solutions with a concentric rotating cylinder device, through which he concluded the linear increase of the thermal conductivity with increasing shear rate for both solutions Chitrangad and Picot [20] also measured the thermal conductivity of aqueous carboxymethyl cellulose (CMC) solutions in a cone and plate geometry They suggested that the thermal conductivities of solutions were proportional to temperature and shear rate Wallace et al [21] found the dependence of thermal conductivity of polymers on shear rate and molecular weight, and concluded that the thermal conductivity of the high polymer in molten state would increase with the increasing shear rate The shear-dependent thermal conductivity of non-Newtonian fluid is essential to the enhancement of heat transfer In order to consider the ununiformity of physical properties field, a multib b plier, such as µ f /µw or Pr f /Prw , is introduced Therefore, a multiplier is also introduced to indicate the ununiformity of shear-dependent thermal conductivity There are several empirical models of the shear ratedependent thermal conductivity The simplest one is the linear function of the shear rate [19] and [22]: λe = λ0 + cγ˙ (22) Since the most important heat conduction process occurs near the wall, the behavior of heat transfer must be chiefly sensitive to the thermal conductivity of the wall Thus, the multiplier may be (λw /λo )b , and the convective heat transfer coefficient can be calculated by: h g = 1.61Re × Figure Convective heat transfer coefficient of borate cross-linked guar at different temperatures and shear rate heat transfer engineering 1/3 Pr λ0 + cγ˙ w λ0 1/3 b D l λ0 D 1/3 3n g + 4n g 1/3 Ka Kw 0.14 (23) where the parameters b and c are constants, obtained by fitting experimental data In our research, b = 0.14, and c = 0.00015 vol 32 no 2011 76 X SUN ET AL Figure Convective heat transfer coefficient of borate cross-linked foam fracturing fluid at different temperature and foam quality with the shear rate of 276.3 s−1 According to Figure 8, the convective heat transfer coefficient of borate cross-linked guar calculated by Eq (23) has good agreement with the experimental data, with the maximum relative error being only 3.1%, which is very small in the engineering field Though temperature has not been an explicit independent variable, the parameters in Eq (23) are all temperature-dependent, except D and l Therefore, Figure shows the heat transfer coefficient as a function of temperature Figures to 11 show the relationship between convective heat transfer coefficients of borate cross-linked foam fracturing fluid and temperature with different foam quality values of 55%, 65%, and 76.5%, respectively Generally, the convective heat transfer coefficient decreases with increasing foam quality, and increases with increasing shear rate As temperature increases the convective heat transfer coefficient decreases For temperatures lower than 55◦ C, the convective heat transfer coefficient changes greatly However, the change is much milder for temperatures higher than 55◦ C The heat transfer mechanism of foam fracturing fluid is sophisticated and includes heat conduction, convective heat trans- Figure 10 Convective heat transfer coefficient of borate cross-linked foam fracturing fluid at different temperatures and foam qualities with the shear rate of 414.5 s−1 heat transfer engineering Figure 11 Convective heat transfer coefficient of borate cross-linked foam fracturing fluid at different temperatures and foam qualities with the shear rate of 552.6 s−1 fer of gas in the bubble, and radiation When the temperature of the foam fracturing fluid is low, radiation can be neglected Whether or not the convection of gas in bubble happens is determined by the Jeffery number L [23]: L= gα T d aσ (24) The convection of gas in bubbles happens when Jeffery number is higher than the critical value of 1709 The Jeffery number of foam fracturing fluid, however, is much lower than the critical value, suggesting that no convection of gas in bubbles can happen The thermal conductivity of static foam could be deduced from a Fourier law and foam thermal conduction model, and could be expressed as [23]: (1 − f ) + λλNg f λf = λ λg (1 − f − ) λ Ng + ( f + ) (25) ) , and is the foam quality where f = 2(1− Based on the preceding equation, the contribution of gas thermal conductivity to foam fracturing fluid is substantial, despite the fact that the thermal conductivity of borate cross-linked guar is much higher Consequently, the convective heat transfer coefficient of borate cross-linked foam decreases with the increasing foam quality The fluid behavior index of foam fracturing fluid was within the range of 0.17 to 0.52 as shown in Figure 5, where the shear rate at the wall is highly sensitive to the fluid behavior index Since the rheology of the fluid is dominated by the external phase, the temperature has the same heat transfer mechanism for foam and borate cross-linked guar Nevertheless, the surface tension will decrease with increasing temperature, which will not only make the bubble much easier to rupture in the shear field but also enhance the nonlinear interaction between the bubbles and the surrounding liquid Nevertheless, all of these results caused by increasing temperature only exist near the axis, while having almost no effect on the vol 32 no 2011 X SUN ET AL heat conduction near the wall compared with those of the shear effect When we compared the experimental heat transfer coefficients with the ones calculated with Eq (21), the same problem concerning borate cross-linked guar arises: The convective heat transfer coefficients calculated with Eq (21) are much smaller than the data from the experiment Therefore, it is reasonable to believe that shear effect must have taken place in heat transfer process It is well known that the heat transfer rates in two-phase systems are usually higher than those in single-phase fluids under comparable flow conditions A number of mechanisms can be responsible for enhancing transport processes These mechanisms are effects of latent heat, effects associated with particle-scale or bubble-scale microconvection, effects of increased velocity gradient at the wall, the contribution of turbulent eddies, etc The microconvection effect and turbulent eddies in continuous phase mentioned here could be caused by many factors, such as shear field, velocity difference between the two phases, and rotating particles For common two-phase fluids, the microconvection effect and turbulent eddies enhance the heat transfer by introducing an additional transport mechanism at the molecular scale However, for foam fluid at large foam quality, the liquid phase exists only in two ways: lamella and Plateau borders With relatively sufficient surfactant and highly viscous liquid phase, almost no film rupture and drainage will happen instantaneously If this is the case, foam fluids will be highly structured fluid The liquid phase and gas phase will be held together as a whole bubble by surface tension and capillary forces This means that the highly viscous liquid phase cannot flow freely in the microchannels between bubbles, and no turbulent eddies can be created in the liquid phase at the molecular scale Due to foam fluids being highly structured, the foam at the bubble scale is worth investigating and our study should be focused on the entire bubbles instead of on the two isolated gas or liquid phases Yuan and Edwards [12] studied foam fluid numerically by using a vertex model When shear strain rises, they found the rearrangement of foam, the transition of bubbles, and turbulent eddies at bubble scale in the high-shear regions near the wall Since the most important heat transfer process takes place near the wall, these shear-induced bubble-scale microconvections are expected to play an important role in the heat transfer enhancement of foam fluid Moreover, the structures of foam cell in the high-shear region are strongly deformed while the cells in the low-shear region are much less deformed from the equilibrium structure In addition, bubbles in flowing foam will deform under the shear stress The critical value of bubble deformation is 0.5 according to Taylor’s law [24–26] When deformation is more significant than the critical value, a bubble will rupture into two smaller ones As shear stress increases with flow rate, more and smaller bubbles will come into being, which will result in a rapid increase in the number of bubble and the interphase contacting area between bubbles Moreover, as the number of heat transfer engineering 77 bubbles increases, more and more bubbles will collide with one another, which will lead to enhancement of heat transfer Applying the same method to borate cross-linked guar, a multiplier of (λw /λo )b is introduced to account for the sheardependent thermal conductivity Hence, the convective heat transfer coefficient of borate cross-linked foam fracturing fluid is caculated by the following equation: h f = 1.61Re × 1/3 Pr λo + cγ˙ w λo 1/3 b D l λo D 1/3 3n f + 4n f 1/3 Ka Kw 0.14 (26) where b = 0.14 and c = 0.00019 in our research As illustrated in Figures to 11, the convective heat transfer coefficient of borate cross-linked foam fracturing fluid calculated with Eq (26) is highly consistent with the experimental data, and the maximum relative error is no more than 10% CONCLUSIONS The rheology and convective heat transfer characteristics of borate cross-linked guar and borate cross-linked foam fracturing fluid were investigated in this study at 30 MPa Both of these two fluids are non-Newtonian fluid and can be described by a power-law model at high shear rate At temperatures lower than 50◦ C, the viscosity of borate cross-linked guar decreased sharply with increasing temperature, indicating a great chemical degradation was caused by the temperature increment However, at temperatures higher than 50◦ C, the apparent viscosity was almost the same as that of non-cross-linked guar The guar was no longer cross-linked and had reverted back to base guar The viscosity of foam fluid increased with increasing foam quality, and decreased with increasing temperature It was mainly dependent on the viscosity of the external phase Temperature influenced the rheology of borate cross-linked foam fracturing, mainly via exerting an influence on the rheology of the external liquid phase and the liquid drainage However, there could be two different additional mechanisms: When the guar was cross-linked, the Marangoni effect would cause a huge rise of the viscosity; when the cross-linker was almost disabled at a high temperature and the guar was non-cross-linked, the Marangoni effect was insignificant Different from Newtonian fluid, the convective heat transfer coefficient of borate cross-linked guar decreased with increasing temperature, reached a minimum point, and then increased with increasing temperature The convective heat transfer coefficient of borate cross-linked foam fracturing fluid decreased with increasing foam quality, but increased with increasing shear rate, and decreased with increasing temperature, but displayed a tendency to increase with increasing temperature In the heat transfer of a non-Newtonian fluid, the velocity gradient at the wall vol 32 no 2011 78 X SUN ET AL was greatly influenced by fluid behavior index, which would enhance or reduce heat transfer Finally, in order to calculate the convective heat transfer coefficient of these two fluids, a multiplier of (λw /λo )b should be introduced to account for the shear-dependent thermal conductivity and heat transfer enhancement for matching the experimental data well It should be noted that shear-induced bubblescale microconvection must play an important role in the heat transfer enhancement of foam fluid NOMENCLATURE A B a Ca cp D d f g h h1 h2 K Ka Kg Kf Kw L l n ng nf Pr Pr Re Re R0 R1 T T U¯ Y0 constant intercept thermal diffusivity, m2/s ˙ capillary number, dµg γ/σ heat capacity at constant pressure, J/(kg-K) diameter of pipe, m diameter of bubble, m surface fraction, (1 − )/3 acceleration of gravity, m/s2 convective heat transfer coefficient, W/(m2-K) convective heat transfer coefficient of working fluid, W/(m2-K) convective heat transfer coefficient of coolant, W/(m2-K) fluid consistency index fluid consistency index under the average temperature fluid consistency index of borate cross-linked guar fluid consistency index of borate cross-linked foam fracturing fluid fluid consistency index under the wall temperature Jeffery number, gα T d /aσ length of pipe, m fluid behavior index fluid behavior index of borate cross-linked guar fluid behavior index of borate cross-linked foam fracturing fluid Prandtl number, ν/a generated Prandtl number, c p KD1−n /λU 1−n Reynolds number, UD/ν generated Reynolds number, D n U 2−n ρ/K constant constant temperature, ◦ C temperature difference, ◦ C average velocity, m/s constant Greek Symbols λ0 λ λw λf unenhanced thermal conductivity, W/(m2-K) thermal conductivity, W/(m2-K) thermal conductivity under the wall shear rate, W/(m2-K) thermal conductivity of foam fluid, W/(m2-K) heat transfer engineering λg λN α θ σ ρ µ µg µf ν τw τ γ˙ gas thermal conductivity of foam fluid, W/(m2-K) liquid thermal conductivity of foam fluid, W/(m2-K) gas expansion coefficient K−1 oblique angle, degrees surface tension, N/m Density, kg/m3 viscosity, Pa-s viscosity of borate cross-linked guar, Pa-s viscosity of borate cross-linked foam fracturing fluid, Pa-s dynamic viscosity, m2/s shear stress at the wall, Pa shear stress, Pa shear rate, s−1 reference foam quality Subscripts e f g N w effective foam guar nitrogen wall REFERENCES [1] Harris, P C., and Heath, S J., Rheology of Crosslinked Foams, SPE Production & Facilities, vol 11, no 2, pp 113–116, 1996 [2] Khade, S D., and Shah, S N., New Rheological Correlations for Guar Foam Fluids The 2003 SPE Production and Operations Symposium, Oklahoma City, OK, 2003 [3] Herzhaft, B., Rheology of Aqueous Foams: A Literature Review of Some Experimental Works, Oil & Gas Science and Technology, vol 54, no 5, pp 587–596, 1999 [4] Watkins, E K., Wendorff, C L., and Alnley B R., A New Crosslinked Foamed Fracturing Fluid, 58th Annual Technical Conference and Exhibition, San Francisco, CA, 1983 [5] Freeman, E R., Bilden, D M., and Hossaini, M., Delayed Crosslinked Gels: Their Role in Aqueous Foam Fracturing, 56th California Regional Meeting of the Society of Petroleum Engineers, Oakland, CA, 1986 [6] Kim, D., Ghajar, A J., Dougherty, R L., and Ryali, V K., Comparison of 20 Two-Phase Heat Transfer Correlations With Seven Sets of Experimental Data, Including Flow Pattern and Tube Inclination Effects, Heat Transfer Engineering, vol 20, no 1, pp 15–40, 1999 [7] Thondavadi, N N., and Lemlich, R., Flow Properties of Foam With and Without Solid Particles, Industrial and Engineering Chemistry Process Design and Development, vol 24, no 3, pp 748–753, 1985 [8] Harris, P C., and Reidenbach, V C., High-Temperature Rheological Study of Foam Fracturing Fluids, Journal of Petroleum Technology, vol 39, no 5, pp 613–619, 1987 vol 32 no 2011 X SUN ET AL [9] Sha, D G., Error Analysis and Evaluation of Measurement Uncertainty, China Metrology Publishing House, Beijing, China, pp 142–185, 2003 (in Chinese) [10] Naval, G., Viscoelastic Measurements of Fracturing Fluids for Proppant Transport Application, Ph.D thesis, University of Oklahoma, Norman, 2001 [11] Langevin, D., Influence of Interfacial Rheology on Foam and Emulsion Properties, Advances in Colloid and Interface Science, vol 88, no 1, pp 209–222, 2000 [12] Yuan, X F., and Edwards, S F., Flow Behaviour of TwoDimensional Random Foams, Journal of Non-Newtonian Fluid Mechanics, vol 60, no 2, pp 335–348, 1995 [13] Pozrikidis, C., Numerical Investigation of the Effect of Surfactants on the Stability and Rheology of Emulsions and Foam, Journal of Engineering Mathematics, vol 41, no 2–3, pp 237–258, 2001 [14] Mitchell, B J., Rheology of Foams, Ph.D thesis, University of Oklahoma, Norman, 1970 [15] Reidenbach, V G., Lee, Y N., and Lord, D L., Rheological Study of Foam Fracturing Fluids Using Nitrogen and Carbon Dioxide, SPE Production Engineering, vol 1, pp 31–46, 1986 [16] Cawiezel, K E., and Niles, T D., Rheology Properties of Foam Fracturing Fluids Under Downhole Conditions, SPE Hydrocarbon Economics and Evaluation Symposium, Dallas, TX, 1987 [17] Deshpande, N S., and Barigou, M., The Flow of Gas–Liquid Foams in Vertical Pipes, Chemical Engineering Science, vol 55, no 19, pp 4297–4309, 2000 [18] Wang, K., The Flow, Mixing and Heat Transfer Process of Non-Newtonian Fluid, Zhejiang University Press, Zhejiang, China, pp 164–170, 1988 (in Chinese) [19] Lee, D L., and Thomas, F I., Shear Rate Dependent Thermal Conductivity Measurements of Non-Newtonian Fluids, Experimental Thermal and Fluid Science, vol 15, no 1, pp 16–24, 1997 [20] Chitrangad, B., and Picot, J C., Similarity in Orientation Effects on Thermal Conductivity and Flow Birefringence for Polymers—Polydimethylsiloxane, Polymer Engineering and Science, vol 21, no 12, pp 782–786, 1981 [21] Wallace, D J., Moreland, C., and Picot, J C., Shear Dependence of Thermal Conductivity in Polyethylene Melts, Polymer Engineering and Science, vol 25, no 2, pp 70–74, 1985 [22] Sehyun, S., The Effect of the Shear Rate-Dependent Thermal Conductivity of Non-Newtonian Fluids on the Heat Transfer in a Pipe Flow, Heat and Mass Transfer, vol 23, no 5, pp 665–678, 1996 [23] Qian, Z P., Foam Rubber, China Petrochemical Publishing Company, Beijing, China, 1998 [24] Triplett, K A., Ghiaasiaan, S M., Abdel-Khalik, S L., and Sadowski, D L., Gas–Liquid Two-Phase Flow in Microchannels Part I: Two-Phase Flow Patterns, International Journal of Multiphase Flow, vol 25, no 2, pp 395–410, 1999 heat transfer engineering 79 [25] Park, C B., and Sub, N P., Filamentary Extrusion of Microcellular Polymers Using a Rapid Deeompressive Element, Polymer Engineering and Science, vol 36, no 1, pp 34–52, 1996 [26] Mu, W J., and Wu, S Y., Influence of Shear Factor on Bubble Nucleation Using Super Critical Fluid CO2 as Foaming Agent, China Plastics, vol 18, no 3, pp 71–76, 2004 Xiao Sun is a Ph.D student in the Energy and Power Engineering Department at Xi’an Jiaotong University, Xi’an, China He worked on how to make use of the energy of oilfield-produced water for his bachelor’s thesis and received his bachelor’s degree in College of Storage & Transportation and Architectural Engineering from China University of Petroleum in 2006 He obtained his master’s degree in 2008 from Energy and Power Engineering from Xi’an Jiaotong University, where he mainly studied the rheology and convective heat transfer properties of foam fracturing fluid His doctoral research is focused on the rheology and sand-carrying ability of sand-carrying foam fracturing fluid Shuzhong Wang is a professor and doctoral supervisor of the Energy and Power Engineering Department at Xi’an Jiaotong University His research interests include two-phase and multiphase thermophysics, the flow and heat transfer of non-Newtonian fluid, waste treatment of supercritical fluid, and resource recycling He has authored more than 80 research papers He serves on the editorial board of Environmental Pollution & Control He is a fellow of the Chinese Society of Toxicology and Chinese Mechanical Engineering Society He was the director of or participant in more than 40 national and provincial programs, including the programs of National Natural Science Foundation of China, 973 and 863 Program, and Science and Technology Support of China He is the first inventor of 20 national patents, containing 17 national invention patents In 1994, he successfully developed STPHD, the first simulation software in China that is applied to calculate the property of mixed transportation in multi-flow of oil, water, and gas He explored the continuous SCWO/SCWG processing system, which aims to process waste organic matter with highly efficient harmless treatment and resource recycling, in 2005 His efforts in research and engineering education have been recognized with the New Century Excellent Researcher Award Program from the Ministry of Education of China in 2007 and the Scientific or Technical Award Yu Bai is a postgraduate in the Energy and Power Engineering Department of Xi’an Jiaotong University, Xi’an, China He received his bachelor’s degree in energy and power engineering from Xi’anJiaotong University in 2005 His present work for the master’s degree is focused on the sand-carrying characteristics of the foam fracturing fluid Shuangshuang Liang is a graduate student in the Energy and Power Engineering Department at Xi’an Jiaotong University, Xi’an, China She worked on PC and biomass co-firing and the consequent hightemperature corrosion behavior in water walls for her bachelor’s thesis and received her bachelor’s degree in energy and power engineering from Xi’an Jiaotong University in 2008 Her present research for the master’s degree is focused on carbon dioxide foam flooding and numerical simulation of multiphase flow vol 32 no 2011 Heat Transfer Engineering, 32(1):80–81, 2011 Copyright C Taylor and Francis Group, LLC ISSN: 0145-7632 print / 1521-0537 online DOI: 10.1080/01457632.2010.517701 new products and services continue to expand our product line with new solutions for our customers” says Kevin Castonguay, Product Marketing Manager for Gems Sensors & Controls For more product information, contact Gems Sensors & Controls, One Cowles Road, Plainville, Connecticut 06062 Phone: (800) 378-1600 www.GemsSensors.com GEMS SENSORS & CONTROLS INTRODUCES VERSATILE NEW LINE OF PROXIMITY SWITCHES R Gems Sensors & ControlsTM (Gems) announces the launch of a new line of magnetically actuated proximity switches for the most basic to the most complex customer applications The 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and reuse, lowering consumption rates and overall costs Meg Grant, Sales & Service Manager of Zebra Skimmers Corporation, comments, “The Oasis was engineered to meet our customers’ objectives and many aspects of this design address specific features customers utilizing our SnapdragonTM Deluxe or ADAPTTM models wished for us to incorporate Their generous 80 new products and services feedback has allowed us to design and offer a more solid system, eliminating the two former models altogether.” New & Improved Features of the Oasis include: • • • • • ON/OFF lever activates power to the air pump Indicator gages placed for easy viewing Gages and filters labeled for easy reference Odor Control Capsule eliminates foul coolant odors Incorporation of SidewinderTM Tube Skimmer for most effective oil removal Zebra Skimmers invents new technology to save its customers money on fluid while reducing waste, and sells and supports its equipment worldwide Most products are guaranteed for at least a year, and some for life For a free catalog, contact Zebra Skimmers Corporation, Phone: 888-249-4855; Email: Meg@ZebraSkimmers.com; or request online: www.ZebraSkimmers.com SUMP LEVEL REGULATOR (SLR) TO MAINTAIN TANK FLUID LEVELS Zebra Skimmers Corporation, a full-line manufacturer of oil skimmers and coolant maintenance equipment, recently introduced its Sump Level Regulator (SLR) to maintain tank fluid levels “Given the costs of labor in this economic climate, it is important to decrease the amount of time spent on maintenance and redirect it to productive tasks We find it important to develop costsaving equipment to reduce labor as well as material expenses The Sump Level Regulator will so by activating the customer’s water supply or fluid pump to maintain the correct fluid operating level in each machine tool sump or central tank, thereby limiting time otherwise spent by the operator,” comments Meg Grant, Sales & Service Manager The Zebra R Sump Level Regulator Features • • • • WELDING END PREP TOOL, SELF-CENTERING WITH ROBUST GEAR DRIVE An I.D clamping welding end prep tool for boiler tube work on stainless steel and other highly alloyed tubes with overlays is being re-introduced with an optional 1.25 HP motor from ESCO Tool of Holliston, Massachusetts R The ESCO Mongoose MILLHOG is a self-centering right angle drive I.D clamping air-powered end prep tool for boiler tubes ranging from 3/4” I.D to 3” O.D and is only 2-1/4” W to fit a “Dutchman.” Featuring a robust gear drive with dual tapered roller bearings, it has a standard 3/4 HP motor or an optional 1.25 HP motor for end prepping stainless steel and other highly alloyed tubes with overlays, and doesn’t need cutting oils Providing chatter-free operation, the ESCO Mongoose R MILLHOG utilizes the EscoLockTM rigid blade lock system and T-15 tool steel TiN coated cutter blades with a radical chip breaker that directs the chip away from the tube and minimizes heat generation Wrenches are permanently attached 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Meg@ZebraSkimmers.com; or request online: www.ZebraSkimmers.com heat transfer engineering 81 vol 32 no 2011 Heat Transfer Engineering, 32(1):82–84, 2011 Copyright C Taylor and Francis Group, LLC ISSN: 0145-7632 print / 1521-0537 online DOI: 10.1080/01457632.2010.517701 hot dates International Symposium on Innovative Materials for Processes in Energy Systems 2010 (IMPRES2010) – For Fuel Cells, Heat Pumps and Sorption Systems – November 29–December 1, 2010 Furama Riverfront Hotel, Singapore Organizer: IMPRES2010 Organizing Committee, c/o Prof Bidyut Baran Saha, Kyushu University, Japan E-mail: saha@mech.kyushu-u.ac.jp Co-organizers: Faculty of Engineering, National University of Singapore Faculty of Engineering, Kyushu University, Japan School of Mechanical & Aerospace Engineering, Nanyang Technological University, Singapore Sponsor: Mayekawa Manufacturing Co Ltd., Japan Symposium web-site: www.impres2010.org/ CONFERENCE OUTLINE This international symposium welcomes participants of professionals dedicated to theories, experiments, simulations, on the development of functional materials for fuel cells, heat pumps, sorption systems and their applied aspects Attendees will include consulting engineers, design engineers, contractors, architects, manufacturers, researchers and academics The IMPRES held in every three years and this is the second international conference after the successful completion of IMPRES2007 at Kyoto, Japan We hope to have stimulating and lively discussions in the heart of the city state Singapore Venue Furama Riverfront Singapore 405 Havelock Road Singapore 169699 Tel: +65-67396420, Fax: +65-67327025 http://www.furama.com/riverfront/ Email: salesadmin.riverfront@furama com SCOPE The conference is concerned with the application of novel materials in the field of energy systems with special focus on the gas-solid reaction processes in various energy conversion systems Materials for fuel cells, heat pumps, sorption systems and other energy conversion and storage devices will be discussed Common concerns include material reactivity, heat and mass transfer characteristics, durability, stability under high-temperature and severe conditions and cost This conference aims to bring workers focusing on different aspects of gas-solid reactions in energy conversion and promote an interchange of ideas across subjects TOPICS The following list of topics illustrates the scope of the conference I Materials application Materials for electric energy conversion and storage a) Solid oxide fuel cells b) Polymer electrolyte membrane and direct methanol fuel cells c) Thermoelectric devices d) Hydrogen production, storage and carrier systems e) Batteries and materials for thermal energy conversion and storage f) Chemical heat pumps g) Adsorption heat pumps h) Phase change material heat storage i) Desiccant systems j) Materials Performance k) Improved performance of materials l) Reactivity m) Heat and mass transfer n) Energy storage and conversion density o) Durability in repetitive operation including high temperatures and severe conditions p) Materials design q) Composite materials r) Membrane reactors s) System reliability (corrosion, reactor sealing, reaction selectivity, etc.) t) Others CONFERENCE FEE* Registration Fees (On or before September 15, 2010): Delegate: S$ 750 Student: S$ 450 Accompanied person**: S$ 300 82 Late Registration (including onsite registration): Delegate: S$ 850 Student: S$ 550 Conference proceedings in book and CD-ROM, reception party, conference banquet, lunches and coffee breaks are included in the registration fees *Each paper requires registration of at least one participant before September 15, 2010, who will present the paper during the conference **Accompanying person will not receive conference proceedings GENERAL ORGANIZING COMMITTEE Chairman Prof Bidyut Baran Saha Kyushu University, Japan E-mail: saha@mech.kyushu-u.ac.jp General Secretary Dr Anutosh Chakraborty Nanyang Technological University, Singapore E-mail: AChakraborty@ntu.edu.sg Accounting Coordinator Mr Kyaw Thu National University of Singapore, Singapore E-mail: mpekt@nus.edu.sg LOCAL ORGANIZING COMMITTEE Chairman Prof Kim Choon Ng National University of Singapore, Singapore E-mail: mpengkc@nus.edu.sg Executive Committee Members Prof Arun Majumder National University of Singapore, Singapore Prof Cristopher Yap National University of Singapore, Singapore Dr Kandadai Srinivasan National University of Singapore, Singapore Prof Fei Duan Nanyang Technological University, Singapore Dr Hideharu Yanagi National University of Singapore, Singapore Dr Abdul Halim National University of Singapore, Singapore INTERNATIONAL SCIENTIFIC COMMITTEE MEMBERS Prof Yukitaka Kato (Chairman, IMPRES2007) Tokyo Institute of Technology, Japan Prof Lua Aik Chong Nanyang Technological University, Singapore hot dates Prof Afshin J Ghajar Oklahoma State University, USA Prof Srinivasa Murthy Indian Institute of Technology, Madras, India Prof Seong Ho Yoon Kyushu University, Japan Prof Atsushi Akisawa Tokyo Univ of Agr & Tech., Japan Prof Yasuyuki Takata Kyushu University, Japan Prof Shigeru Koyama Kyushu University, Japan Prof Felix Ziegler Technische Universitat Berlin, Germany Dr Elisa Boelman European Commission, Brussels Prof Min-Soo Kim Seoul National University Prof Hideo Mori Kyushu University, Japan Prof Robert E Critoph Warwick University, UK Prof Yuriya I Aristov Boreskov Institute of Catalysis, Novosibirsk, Russia Prof Yong Tae Kang Kung Hee University, Korea Dr Belal Dawoud Viessmann GmbH & Co KG, Germany Prof Ruzhu Wang Shanghai Jiao Tong University, China Prof Pradip Dutta Indian Institute of Science, India Dr Uli Jacob dr Jakob Energy Research, Germany Dr Keiko Fujioka Functional Fluids Co Ltd., Japan Prof Kuenman Cho Sungkyunkwan University, Korea Prof Akio Kodama Kanazawa University, Japan Prof Agami T Reddy Arizona State University, USA Prof Mike Tierney Bristol University, UK ATTRACTIVE PLACES IN SINGAPORE Sentosa Nature beckons everywhere on Sentosa, where a lazy holiday of sun, sand and sea awaits one and all The fun paradise of Singapore is MERLION standing 37 m tall The MERLION (the shape of a lion’s head and the body of a fish) offers vantage viewing point of Sentosa, Singapore’s city skyline and the surrounding islands 83 Jurong Bird Park is a 20.2 hectare openconcept park It is the largest in the Asia Pacific and one of the finest bird parks in the world Jurong Bird Park, collection of more than 8,000 birds from 600 species, offers visitors an experience that is entertaining and educational Clark Quay Clark Quay features five blocks of restored warehouses and is the home to hip entertainment, dining outlets and shops of all kinds In the evening, theme pubs and bars become alive with classic rock, hard rock, the blues and music from the 60s Conference website: www.kalasalingam.ac.in/mech/ incotee2011.html For further details contact INCOTEE 2011 Organizing Secretary: Dr P Rajesh Kanna Department of Mechanical Engineering Kalasalingam University (Kalasalingam Academy of Research and Education) Krishnan koil, Srivilliputhur(via) Tamilnadu-626190 India Tel: +91 4563 289042; Fax: +91 4563 289322 E-mail: incotee2011@klu.ac.in Heat Exchanger Fouling and Cleaning 2011, June 05–10, 2011, Crete, Greece http://heatexchanger-fouling.com/ INCOTEE 2011 The International Conference on Thermal Energy and Environment (INCOTEE 2011) is scheduled on 24– 26, Mar 2011 at Kalasalingam University, India Possible topics under these themes, but are not limited to • Applications of Heat and Mass Transfer • Process Heat Transfer and Instrumentation • Heat and Mass Transfer in BioChemical Applications • Combustion • Energy Management and Recovery • Clean Development Mechanism and Environment Original unpublished contributions are invited to present at INCOTEE 2011 Selected papers from this conference will be published as a special issue of Heat Transfer Engineering (HTE) Journal after careful peer review as per the journal norms Important dates are, Last date for submitting full length paper Notification of acceptance for full length paper Last date for registration Selection of papers for possible publication in the special issue of HTE December 31, 2010 February 15, 2011 March 9, 2011 June 15, 2011 Jurong Bird Park heat transfer engineering vol 32 no 2011 CALL FOR PAPERS Conference chair: ă Professor H Muller-Steinhagen University (TU) of Dresden, Germany E-mail: rektor@tu-dresden.de Conference co-chairs: Dr M.R Malayeri, University of Stuttgart, Stuttgart, Germany E-mail: m.malayeri@itw.uni-stuttgart.de Professor A.P Watkinson The University of British Columbia, Canada e-mail: apw@chml.ubc.ca Fouling, i.e the formation of deposits on the heat transfer surfaces is one of the most severe problems in the design and operation of heat exchangers in most industries In most industrial situations, the inefficiency of heat transfer resulting from fouling has a direct link to excess fuel consumption in the process The aim of this conference is to facilitate innovative thinking and to explore new theoretical and practical approaches to address the tremendous challenges associated with fouling of heat exchangers It will also provide an opportunity for experts from industry, academia and research centers from around the world to present their latest research and technological developments in fouling mitigation and cleaning strategies In addition to academic research we, therefore, particularly welcome industrial case studies, hot dates 84 whether there have been successful solutions to fouling or not Scope of the Conference The key themes of Heat Exchanger Fouling and Cleaning 2009 are: • Crude oil and hydrocarbon fluid fouling • Fouling in cooling water, thermal desalination units, power plants, dairy and food industries • Fouling in automotive industry • Mechanisms of heat transfer fouling (crystallisation, particulate, reaction, corrosion, solidification and biofouling) • Surface and chemical treatment • Modeling and CFD studies • Design of heat exchangers • Micro and compact heat exchanger fouling • Fouling mitigation, monitoring and cleaning of heat exchangers Important Dates Abstracts due Notification of Acceptance Full Manuscript due November 30, 2010 December 15, 2010 March 31, 2011 Abstracts with a maximum of 250 words are to be submitted electronically to: Dr Reza Malayeri, Institute heat transfer engineering vol 32 no 2011 of Thermodynamics and Thermal Engineering, University of Stuttgart, Stuttgart-Germany, m.malayeri@itw uni-stuttgart.de T: +49-711-6856-7656; F: +49-711-6856-3503 Upon acceptance, guidelines for preparation of the full manuscript will be provided For industrial contributions, the option of submitting an extended abstract rather than a full manuscript is offered Following the conference, papers will be peerreviewed for publication in the postconference proceedings and selected papers will also appear on the special issue of the international journal of Heat Transfer Engineering