Energy & Fuels 1994,8, 179-187
179
Applicability of Furnace Analysis in Determination of the Performance Characteristics of a Hot Wall Furnace Fired with Sorbent-Loaded Coal-Water Fuel Robert H. Essenhigh,*p+James J. Obloza, and Thomas K. Hammondt Department of Mechanical Engineering, The Ohio State University, 206 West 18th Avenue, Columbus, Ohio 43210 Received April 2, 1993. Revised Manuscript Received October 18, 199P
A hot-wall furnace of 40 ft3fired with sorbent-loaded coal-water fuels at 1MMBtu/h was found to generate firing and efficiency (performance) curves in functional agreement with the equations developed in furnace analysis. (In the FPS system, M = lo00 and MM = 1000 000. It is based on the Latin prefix, M, as abbreviation for mille.) The present study includes a theoretical correction for the dependence of the firing parameters on excess air due to water in the fuel, and this correction is shown to be supported by test against the experimental results. The complete results show the variation with output of: thermal input (firing curve), thermal (operational) efficiency,heat utilization factor, specific gas enthalpy, and wall loss. The furnace analysis firing constants were obtained from the experimental results for the variation with output of the heat utilization factor, and crosschecked against the data for the variation of specific gas enthalpy with output. With corrections to the excess air values for water in the fuel, the intrinsic efficiency and maximum output values obtained are then shown to fall and rise linearly with, respectively, effective excess air and equivalence ratio, in support of prediction. The values of the firing parameters were used to calculate the firing and efficiency curves at 0%,155% ,30 % ,and 45 % excess air, with acceptable agreement between the predictions for these conditions and the experimental results for data within bands defined by the four excess air levels. With a data base of about 7000measurements, these experiments have permitted better evaluation and test of the excess-air function in the furnace analysis equations, including the correction for the effective excess air addition required to account for water in the fuel mixture.
Introduction In this paper we evaluate the applicability of furnace analysis1t2for determining the performance characteristics of a hot-wall furnace fired with sorbent-loaded coal-water fuels (SLCWF). The measurements used in the evaluation were made to provide supporting data in a study3 of the potential of SLCWF for controlling sulfur oxide emissions from coal flames by addition of limestone sorbents to a coal-water formulation of the fuel (CWF). Interpretation of these sorption result^,^ however, depended crucially on an accurate description of the overall furnace-performance characteristics. This was obtained from application of furnace analysis, and presentation of that information is the subject of this paper. This presentation includes comparison with CWF measurements made earlier.4 The furnace performance characteristics obtained, using the procedures of furnace analysis, are in the form of a set
* Address correspondence to this author.
E. G. Bailey Professor of Energy Conversion. Now with Carbogel, Inc. * Abstract published in Advance ACS Abstracts, December 1,1993. (1) Essenhigh, R. H.; Thekdi, A. C.; Malhouitre, G.; Tsai, Y. Furnace Analysis A Comparative Study. In Combustion Technology: Some Modern Developments; Academic Press: New York, 1974; Chapter 13. (2) Eesenhigh,R. H. Comparative Thermal Behavior of Furnace and Engines: Prediction of Thermal Efficiency in Real Time. In Thermodynamic Analysis and Improvement of Energy Systems; International Academic PublisherslPergamonPress: New York, 1989; pp 116-125. (3) Obloza, J. J.; Hammond, T. K.; Essenhigh, R. H. Control of SOX Emissions by Sorbent-LoadedCoal-Water Fuel Mixtures. h o c . IOWA Znt. Conf. High-Sulfur Coals, Iowa State University 1989. (4) Li, K.; Han, K.-I.; Essenhigh, R. H. Performance Characteristics of a Hot- Wall Furnace Fired with Coal Water Slurry (CWS)Using Gas/ Air Atomization;AGA UniversityMonographs,Vol. 1 (No. l),Paper No. 2; American Gas Association: Washington, DC, 1985. t
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of functional equations governing the overall relationship between thermal input, Hf, and useful output, H,. The equation derived between Hfand Hgis the firing equation, and it describes a firing curve. From the firing equations, the overall furnace performance can be readily described in principle. By extension, the equations also lead to needed relationships showing the dependence on useful output of the operational thermal efficiency and of the heat utilization factor (a)as defined by Thring and Reber: with the intrinsic efficiency obtained at the zero-load limit of a. The development of the procedures to obtain these relationships have a history of about a century but, as they are still not widely known, some of that past history is outlined in the section Furnace Analysis: Origins and Background later in this paper. Since the furnace analysis method is phenomenological, the conceptual pattern described by the results is correspondingly general and is broadly obeyed by furnaces and engines, or by energy conversion devices in general, as detailed in the cited However, on account of that phenomenological basis, the procedures (nominally) require demonstration of the structural validity of the relevant equations in each particular case. Such validation for the experimental results obtained in the sorbent experiments is part of the objective of this present paper. The experimental measurements also then provide the data for a unique test of the applicability of the furnace analysis equations to a hot-wall furnace with a thermal load at different levels of excess air, including an extension (5).Thring,M. W.; Reber, J. W. The Effect of Output on the Thermal Efficiency of Heating Appliances. J . I n s t . Fuel 1946, 18, 12.
1994 American Chemical Society
180 Energy &Fuels, Vol. 8, No. 1, 1994
of t h e excess air relations to account for the effects of water in the fuel, treated as equivalent excess air.
Experimental Section Furnace and Burner. The CWFs were fired in a hot wall furnace which has been described in detail el~ewhere.~~' The following summary is presented for context. The furnace dimensions are 2 ft X 2 f t X 10 ft, with 30 water-cooled tubes on the floor acting as a thermal load to remove 20-30% of the heat released, dependingon the firingconditions. This thermalloading provided a reasonable match of local heat flux and temperature with conditions typically found at the wall in a large water wall boiler. The slurry was supplied in 17 55-gal drums, and the furnace was fired at 0.9-1.5 MMBtu/h. (In the FPS system, M = 1000 and MM = 1000000. It is based on the Latin prefix, M, as abbreviation for mille.) At rates below 0.75 MMBtu/h, behavior tended to be erratic as the flow rates were then outside the operational range of the burner. Most experiments were carried out near 15% excess air, but with a range of 6 4 5 % . A high-swirl burner was used with a two-fluid, nozzle-mix, oil burner gun abstracted from a Cleaver-Brooks CB 125 burner, and refitted with a Delavan inside-mix SL701-1 "Swirl-Air" Tipintle atomizer. Atomizingair pressure was normally in the range 100-120 psi. In the previous experiments' the pintle was Sic. A steel pintle was eroded in about 1 h. A total of 17 sets of experiments were carried out (that is, one set per drum). In one set, the slurry was fired with oil support with the oil co-atomized with the slurry. In all other (16) sets, it was fired without either supplementary fuel support or air preheat. In the earlier work? gas support was used in the majority of the tests, using the gas/ air support mixture as the atomizing agent. Fuel Characteristics and Delivery Equipment. The slurrieswere prepared from two Ohio coals. medium-sulfurLower Kittaning No. 5 seam (2.23% S and 44.6% VM daf) from the Sands Hill Coal Co., and high-sulfurPittsburgh No. 8 seam (5.45 % S and 47.5% VM daf) from the Cravat Coal Co. The respective coals were made up into eight and nine drums (55gal) of slurry with sorbent levels at Ca/S ratios of 0,1.75,2.0, and 2.25. With the Sands Hill coal the sorbent used was a dolomitic limestone, and with the Cravat coal it was a calcitic limestone. The calcitic limestone was found to be superior to the dolomite, with up to 65% of the sulfur removed in the flame3. Water levels as fired were mostly in the range 40-45 % . The target figure was 3540% but the increased water was found experimentally to be necessary for fluidity in pumping and atomizing. Combustion efficiency and flame temperatures were good, nevertheless, and it is possible that the increased water may have aided dispersion in the flame. In preparing the slurries, the coals were ground to a median size of about 50 pm, and the sorbent to a median in the range 5-10 pm. Atomization appeared to be unaffected by the sorbent. The slurry pumping system consists of two Moyno progessing cavity pumps in series loops. The first pump is a high-volume, low-pressure pump that, to maintain fluidity, overpumps the slurry out of a 55-gal dry tank and back into it through a loop circuit. The second pump is a low-volume,high-pressure pump with a maximum delivery at 280 psi but normally operatad at 70-100 psi. This line taps the primary loop to deliver the slurry through a Micromotion flow meter to the burner. The line includes an overpressure switch set at 150 psi that turns off the second pump if there is an overpressure due to blockage. Instrumentation. Principal instrumentation consists of recording flow meters for fuel and air, thermistors for the water with asuction tubes, recording gas analyzers for CO, COZ,and 02, pyrometer for the exhaust gas temperature. The flow meter for the slurry is a Micromotion No. C25. The meters for the main and atomization air are Kurtz (mass flow) Models 505-13 and 505-11, respectively. Flow meters are not used for the water tubes: constant flow orificesare used instead with periodicchecks of the calibrations by direct measurement at flow rates of 1.5 gpm for the first 10 tubes, and 1 gpm for the last 20. The
Essenhigh et al. temperature rise in the tubes is measured by YSI-701thermistors at exit and entry. The exhaust gas is conditioned by washing and drying (in a Drierite tower) and analyzed dry by MSA-NDIR instruments, using Model 4000 for 02,and Model 303 for both CO and COZ. SO2 was also measured as reported earlier3but the results are not relevant to this paper. Other instrumentation includes an array of 25 Type K TC's along the roof center line, mostly set with tips flush with the flame-side face of the roof: five are set at different depths to measure the roof temperature profile. These measurements were used to monitor overall furnace response. For the same purpose, and also to control in-leaking air, furnace draft is measured using a Setra Model No. 239 pressure transducer with digital read-out to 1/1000-in. water column (1thou wc). Measurements from these instruments were used to calculate mass and enthalpy flows as identified below. Data Compilation and Reduction. Energy balances were used to check the accuracy of measurement. The enthalpy input was determined from the slurry firing rate (from the Micromotion flow meter) and the calorific value of the slurry. The heating values varied but were in the range 4000-6000 Btu/lb of mixture. These low values were due to the added water and sorbent, and the valuea were calculated from the dmmf analysesof the relevant coals corrected for the water content. The effect of such high water loadings in firing other carbonaceous fuels is discussed elsewhere? The exhaust gas loss was determined from its composition analysis, flow rate, and temperature. The composition analysis included measurement of the dry CO, COz, and 02,with calculated corrections for the water content. The flow rate was based on the slurry flow rate using the calculated stoichiometric ratio (from the fuel analysis) and the excess air (determined from the dry gas analyses). The exhaust temperature was measured with the suction pyrometer. The thermal load output was determined from the water flow rates in the cooling tubes and the temperature rise across each of the 30 tubes. The net balance between thermal input and measured output (cooling tubes plus exhaust) gave the wall loss, determined by difference, and was typically 10-15 % of input. Closure was used to determine overall measurement accuracy as discussed below. Most of the required measurements were recorded by a Digitec Datalogger 3000 and transmitted periodically to a VAX 8550. The slurry firing rate was entered manually. In the data treatment and reduction, the consistency of the gas analyses was first checked using a computer formulation of the Ostwald triangular chart;' then all enthalpies were calculated to determine the overall closure on the heat balance. Other measurements included SOX, NOX, and temperature profiles inside the flame using a suction pyrometer and a two-color pyrometer; further details are given elsewhere.s* Accuracy. Accurate measurements on coal firing systems are notoriously difficult to make and, with the mix of components in this system, accuracy was proportionately more difficult but, nevertheless, was acceptable. This was determined, in principle, by a closure calculation on the energy balance, using wall loss values determined by difference obtained from previous wall loss measurements' as well as in the present work. The actual procedure used was to calculate and compare the wall loss values. This is discussed in more detail in the Results section. The criterion for acceptable closure on the energy balance was then set by requiring agreement with the previous data on the wall loss to within f 3 0 MBtu/h. This represents a closure of 95% or better. Using this closure criterion, a total of 97 acceptable data sets were obtained during the experiments, representing a data base of about 7000 measurements. Individual measurements on (6) Essenhigh, R. H. Burning Rates in Incinerators (Part I1 The Influence of Moisture on the Combustion Intensity). R o c . 3rd Natl. Incin. Conf. 1968,94-100. (7) Salisbury,J. K., Ed.KENTS Mechanical Engineers Handbook, 12th ed.; John Wiley Inc.: New York, 1962; Vol. 1, pp 2-06, 2-07. (8) Essenhigh, R. H.; Hammond, T. K.; Kennedy, L. A.; Obloza, J. J. "Organic and Inorganic Sulfur Reduction by Physical Cleaning and InFlameSulfurFbduction";FinalReporttoStateofOhioCoalDevelopment Office on Project No: R-86-034-01-OH, May 1989.
Performance Characteristics of a Hot Wall Furnace input and output enthalpiesare estimated to be within f25MBtu/h which mostly represents between 5 and 10% error on those values or better.
Furnace Analysis: Origins and Background Furnace analysis as it exists at this time is a phenomenologicaltreatment of thermal systems in which the focus is on obtaining relations for the total performance of the systems by relating the variation of thermal input (Hd with thermal output (Ha).The basic input/output relation is commonly known as the firing equation and it defines the firing curve. Characteristically, the firing curve is a concave-upwards curve with the firing rate increasing nonlinearly with increasing output up to a maximum output determined, theoretically, by the adiabatic flame temperature, and at this point the firing rate, theoretically, goes to infinity. With the firing equation relationship established, it is then a simple matter as we show below to obtain dependent relations for the variation with output of the thermal (operational) efficiency (q); the Thring and Rebel.5heat utilization factor (a); and the exhaust enthalpy (H& The complete suite of relations thus represents the effective integration of the overall interactive effects of heat transfer, reaction kinetics, aerodynamics, and the system process (heat treatment, melting, boiling, power generation, etc.), with the overall results combined in only three performance equations. Although this approach is unable to provide information on detailed local behavior inside the device, it does provide overall performance behavior in a general format that (nominally) applies to all thermal conversion systems, generally designated as furnaces and engines2, and it also defines bounds within which actual or predicted operation must be constrained. The origins of furnace analysis are generally attributed to Hudsongin 1890. Hudson developed an empirical input/ output equation to describe steam boiler performance that, in 1926,was modified and extended by Orrockloto become the Hudson-Orrock equation. This was substantially the origins and driver that initiated the theoretical developments of the 1930s,notably by Wohlenberg and by Hottel (see, e.g., McAdams and Hottel).ll At the same time, there were alternative empirical studies of other systems, notably of furnaces, and an earlier landmark study byArmstrong12 in 1927was the first to show the influence and importance of output (Ha)as a major factor in determining furnace efficiency. This was followed by other contributions (see ref 1for citations) relating to forge furnaces, melting and reheating furnaces, oil retorts, blast furnaces, and, particularly, glass tanks. Independent studies of glass tanks led to essentially the same linear input-output equation, known in one context as the SGT (Society of Glass Technology) equation,’3 and in another context as the HAPF (Hartford Average Practice Formula) re1ati0n.l~ This evident linearity of the firing curve is, of course, at variance with the general nonlinear rise predicted by the (9) Hudson, J. C. Heat Transmission in Boilers. The Engineer 1890, 70,449, 483, 523. (10) Orrock, G. A. Radiationin Boiler Furnaces. Tram. A.S.M.E. 1926, 48,218. (11)McAdams, W. H.; Hottel, H. C. Heat Transmission, 2nd ed.; McGraw-Hill: New York, 1942; pp 77-84. (12) Armstrong, H. C. Characteristicsof Furnaces Curves as an Aid to Fuel Control. The Engineer 1927, 144,445. (13) Seddon, E. The Assessment of the Thermal Performance of Tank Furnaces for Melting Glasa. J. SOC. Glass Technol. 1944, 28, 33. (14) Lester, W. R. Average Practice Formula for Glass Melting Furnaces. Glass Ind. 1963, 44,260.
Energy &Fuels, Vol. 8, No. 1, 1994 181 firing equation. This is now known to be the result of (regenerative)heat recovery, and the linearity of the firing curve for glass tanks is, in fact, a first approximation to a flat curve.15 The limitations of the empirical approach were recognized at an early stage, and increasingly complex mechanistic schemes were developed in attempts to predict observed behavior and to account for the empirical relationships from a more fundamental basis. These approaches, developed from the 1930s through the 1950s and also known as internal furnace analysis, typically focused on heat transfer, mostly radiative, inside the furnace or boi1er;ll but, in the absence at that time of adequate means of solving the formulated equations, numerically or otherwise, substantial approximations in the theoretical constructs were necessary to obtain solutions in any form, particularly when a radiation model was used. One such example was the invention by Hottel of the “speckled furnace” in which the thermal (useful) load is distributed uniformly on the furnace ~ a l l . ’ ~ J ~ Alternatively, simplifications were obtained either by assuming uniform temperatures (Le., stirred reactor conditions), or by limiting heat transfer to conduction/ convection, as in the treatment by Thring and Reber in 19445(see alsoThring).18 It may be noted that, even today, the numerical models generally require the assumption of substantial geometric symmetry since otherwisethe models become intractable. This limitation does not apply to the furnace analysis approach. The realism of assumptions such as the speckled furnace is clearly questionable in most applications. Nevertheless, the equations obtained, particularly by Thring and Reber? accurately described the general pattern of the overall performance characteristics in terms of the input/output and efficiency behavior. More significantly, from examination of experimental records of operating furnaces, Thring and Reber interpreted the idle heat, HfO, (or firing rate at zero output) as “lost” input due to wall or friction losses that, to a first approximation, was essentially constant. Consequently, the “net” or effective heat available for the process was only (Hf - HfO) and, using this, they introduced the concept of the heat utilization factor (HUF) defined as a = H$(Hf - HfO),as compared with the operational (thermal) efficiency with the standard definition, 1 = Ha/Hf. The intrinsic efficiency is then obtained from this definition of a as the maximum value of the HUF at zero output. This was a major contribution. Emphasis on the idle heat was an important element in the derivation of the furnace analysis equations as it is responsible for a peaking behavior in the operational efficiency curve that, typically, is not predicted in most of the other analyses. This is shown in Figure 3 of ref 1 which is a comparison of efficiency behavior with output predicted from four sources: the Thring and Reber a n a l ~ s i swhich , ~ parallels actual performance, the HudsonOrrock equation, the SGT/HAPF relations, and the Hottel treatment. (15)Essenhigh, R. H. Energy Use, Efficiency, and Conservation in Industry. Proceedings of the Symposium on Enuironment and Energy Comeruation; EPA Technical Series: Report EPA600/2-76-212(ERDA 47); E P A Washington, DC, 1976; pp 161-188. (16) Hottel, H. C. Radiative Transfer in Combustion Chambers. J. Inst. Fuel. 1961, 34, 220. (17) Hottel, H. C.; Sarofim, A. F. Radiatiue Transfer; McGraw-Hill: New York, 1967; Chapter 14. (18)Thring, M. W. The Science of Flames and Furnaces, 2nd ed.; Wiley: New York, 1962; Chapter 7.
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These theoretical approaches were still limited, however, by a consequence of the quasi-mechanistic basis of the analyses used since this generated equations that generally required an arbitrary closure postulate or heuristic assumption to obtain solution. In the Hottel model, for example, a relation between a mean "flame temperature" ( T f )and the exhaust gas or furnace outlet temperature ( T,)was needed, and Hottel's heuristic solution17based on examination of experimental data was to assume that (Tf- Tg) = constant, with the constant generally set at 200 "C. An additional and broader disadvantage of such an approach is that the inclusion of mechanisms in any model typically imposes additional limitations, typically due to need to specify precise physical boundaries, with the consequence that analyses can lose generality. At best, an analytical model would then apply to a group of furnaces of similar type and shape; and at worst, it would be necessary to develop a separate model for each individual furnace. This is something of the problem being faced at this time, for example, by the development of boiler models by ACERCle and similar groups. If heuristic assumptions are to be adopted, however, the adoption at the start of the analysis instead of at the end then, potentially, allows for development of a totally phenomenological,and thus totally general analysis, hence bypassing the need for introducing mechanisms. This was the approach takenzOinanalysis, nominally, of glass tanks, but applicable generally' as shown, for example, by successful application to automobilesz1 and engines;z it has also been extended and successfully used to predict the effect of enhanced oxygen levels in firing an aluminum reheat furnace.zz Moreover, as shown elsewhereF3the joint and parallel use of both the phenomenological and mechanistic approaches can be used to obtain understanding of behavior that may not be possible using only one. This general approach, of course, largely lost favor in the last one or two decades as computing facilities grew in power, with the prospect of being able to model any given furnace, boiler, or engine in complete detail from the full array of relevant equations. This is extension of the approach taken by Wohlenberg and Hottel, and by Thring and Reber. More recently, however, with fuller recognition of the complexity of such a task, even using massively parallelmachines, there is a parallel recognition of the joint value of the very simple (relatively speaking) approach of furnace analysis, as described in this paper, in reducing behavior to descriptions of performance characteristics that are more manageable from the practical or applied engineering point of view. (19) Smoot, L. D., Ed. Fundamentals of Coal Combustion; Elsevier: New York, 1993. (20) Essenhigh, R. H.; Tsai, Y.-W. Furnace Analysis Applied to Operation and Design of Glass-Melting Tanks. Glass Ind. 1969,50,278, 333. Furnace Analysis Applied to Glass Tanks at High Output: The Heat Transfer 'Bottleneck' Effect. Glass Ind. 1970, 51, 68, 108. (21) Essenhigh, R. H. Evaluation of Fuel Consumption Rates and Thermal Efficiency of Automobiles by Application of Furnace Analysis. J. Transportation Res. 1974,8, 457-464. (22) Kuhn, J. E.; Wiesner, J. J.; Yu, H.: Paper No. 9 in Proceedings of the 4th Aluminum Industry Energy Conservation Workshop; Aluminum Association of America: Washington, DC, 1984. (23)Enomoto, H.; Tsai, Y.; Essenhigh, R. H. Heat Transfer in a Continuous Model Furnace. ASME Heat Trans. Conf., Paper No. 75HT-5; ASME: New York, 1975. (24) Cole, J. A,; Clark, W. D.; Heap, M. P.; Kramlich, J. C.; Samuelson, G. S.; Seeker,W. R. Draft Task Final Report, Energy and Environmental Research Corporation, EPA Contract No. 68-02-3633, October 1983.
Theoretical Summary 1. Basis. The basis for the analysis is the primary heat balance between thermal input from fuel and/or other sources,Hf,and the three outputs, useful output, H,, stack gas loss, H,, and wall loss, H,, giving
Hf = Fh, = H, + H,+ H,= H ,
+ Hf(h,/hf)+ Ha ( 1 )
where F is the firing rate (fuel supply rate), hf is the heat of combustion (specific enthalpy) of the fuel, and h, is the specific enthalpy of the exhaust gas per unit of fuel fired (F):i.e., H , = Fh,. At zero output, or idle, defined by zero superscripts, eq 1 becomes
Hf" = HWo+ H,"
(2) Using the heuristic assumption of linear increase of specific exhaust gas enthalpy, h,, with useful output, H,, and a similar relation based on experiment for wall loss, H,, then we have'sz
H , = H2[1 + m(H,/H,m)l (4) where m is an empirical linearization parameter whose value at this time has to be determined experimentally but is a target for prediction by mechanistic analysis. 2. The firing equation is obtained by subtraction of eq 2 from eq 1, and rearranging using eqs 3 and 4. It describes the variation of the dependent thermal input, Hf, as a function of the independent useful thermal output, H , and has the form (5) Hf = H,O + H,/a$(l - H,/H,m) This is the equation of a concave upward curve (the firing curve) starting at the idle heat, Hfo, at zero output, with firing rate going to infinity at a maximum output, H," At this condition of H , = H,m;eq 3 shows that h, = hf, implying thereby that the furnace and the exhaust are at the adiabatic flame temperature.lI2 The third parameter, CYEO, is an intrinsic efficiency defined by eq 7 and is a function of excess air. 3. The Operational Efficiency, 9, is as normally defined, as the ratio of useful output to total thermal input, and it yields the variation of 9 with output, H,, in the form
This is a curve with an asymmetric inverted U-shape starting at the origin, and returning to 9 = 0 at a maximum output: H , = Ham;the maximum output, of course, is a theoretical limit and is not realizable experimentally. This peaking characteristic is, commonly, predicted and experimentally mapped in utility boiler operation for electricity generation as it has substantial influence on generation costs under overload conditions. 4. The Heat Utilization Factor and Intrinsic Efficiency. The heat utilization factor, a,is based on the Thring and Reber6 definition as the useful utilization of the net available heat: (Hf - Hfo). Thus, combining this definition and eq 5 we get, for any given value of excess air, E % : = H,/(Hf- Hp)= CY$(^ - H,/H,")
(7)
This is the effective efficiency of the furnace if the wall losses at idle are subtracted or if the furnace was totally
Performance Characteristics of a Hot Wall Furnace insulated and thus adiabatic. It is also the equation of a straight line that can be used as a test of the validity of eq 5, with the abscissa and ordinate intercepts providing values of the coefficients, Hamand EO. The intercept CYEO defines the intrinsic efficiency of the furnace under the conditions of use. 6. Parameter Relations. Reduction of the eqs 1-4 to the three performance equations, (5), (6),and (7), involves reduction of parameters,lI2 with definitions from the reductions as follows:
which shows the dependence of the intrinsic efficiency, LYE'), on the idle heat, HP, and on the wall-loss increment coefficient, m; and where
which defines the maximum intrinsic efficiencyfor a given set of operational conditions. 6. Excess Air Relations. The three performance equations, (3,(6), and (7), hold at any given value of excess air ( E% ), where the three coefficients, Hfo,Ham,and are functions of excess air, and also of water content which can be treated as equivalent to excess air as shown below. The idle wall loss, HWo,uniquely, is evidently independent of excess air so that we can write
where the subscripts m,s and m,E are values at stoichiometric (subscript s) and at excess air E % (subscript E). The excess air dependencies of the remaining parameters are then given by112
Energy &Fuels, Vol. 8, No. 1, 1994 183
G, is the mass stoichiometric ratio of air to (dry) fuel; cp is the POC specific heat (mean of constituents); and T f is the flame temperature. For a water percent in the CWF of w % , then m, = w%/(lOO - w % ) , For the moisture effect after evaporation, also treated as additional excess air, both the volume generation and the thermal loading due to a given mass are approximately twice that of an equivalent mass of air. The proportional equivalent volume is the ratio of the two molecular weights: MW(POC)/MW(water vapor), which has avalue of about 1.8. The thermal loading is the product of (mass X specific heat X temperature rise). On both counts, the equivalent excess air due to the moisture can be treated as additional air by weight times a multiplier factor, f , of value approximately or slightly in excess of 2. Firing unit mass of dry fuel with m,mass of water, the total equivalent mass on a thermal equivalent basis can be written: (fuel + air + moisture) = [1+ G,(1+ E%/100)+ fm,] Equating to an equivalent excess air value,Ez % ,we obtain
E,% -E% = AE,% = 100fm,/G,
(16)
Combining both factors, we get
AE% = MI%+ AE,% = (f+ L/cPA7')*100~,/G,(17) Of the two factors includedin this expression, the direct thermal loading (represented by f, is the more important. Evaluating eq 17, taking w % = 45% and f = 2, and using standard values for L and cp,then we get AE% = 20 f 2 5%. This is an increment to be added to the true excess air value so that the suite of actual values, 0,15,30, and 4594, must be read as 20,35,50, and 65%. The corresponding equivalence ratios are 0.83,0.74,0.67, and 0.61, and these are the values used in the graphs to evaluate the excess air behavior.
Results I: Performance Characteristics
@)/4,,I
H,$H,E = (f$,a/@,8)/[1 - (1( 14) 7. Excess Air Equivalence for Moisture. Moisture acts jointly as a diluent and as an additional thermal load. In both aspects, it can be equated to an equivalent excess air factor. As a diluent after evaporation, it increases the velocity of the flame gases through the furnace on account of increased volume, thus reducing the time for heat exchange between the flame and load and, consequently, reducing the thermal efficiency. At the same time, the moisture is added thermal loading to the products of combustion (POC) at exhaust which, likewise, reduces the thermal efficiency. An additional thermal load is provided by the latent heat (enthalpy) of evaporation (L). To calculate the effect of the enthalpy of evaporation, this is first treated as a reduction in the enthalpy of combustion (hf). For mass of moisture, m,,per unit mass of dry coal, the effective enthalpy of combustion is (hf m,L); so the fractional reduction in enthalpy per unit mass of dry coal is (m,L/hf), and there would be a corresponding proportional reduction in flame temperature. If, instead, we obtained the same reduction in flame temperature by increasing the excess air, the equivalence can be written (m&/hf) [G8(~,%/1OO)IcP(TfTo) (15) where A,?&% is the equivalent increase in excess air required to produce the same reduction in temperature;
The focus of the results evaluation was first, on the closure for the energy balance, using eq 1,to establish the accuracy of the measurements and to establish the overall performance characteristics, as described in this section; and second, to use the data to establish the extent to which the furnace analysis equations, (51, (61,and (71,adequately predicted behavior, as described in the section following. General Performance Characteristics. Slurry firing was started after the furnace was preheated to near-steady state on No. 1oil. After the slurry firing was started, the firing rate and air rate were set to the desired target thermal-input value between 0.75 and 1.5 MMBtu/h at a nominal 15% excess air. Because of vagaries in the firing conditions and notably in the slurry water content, actual excessair levels varied, as described below, and calculations were based on measured conditions. Each drum of slurry would provide 3-4 h firing at the target rates. Initially, firing problems were experienced due to low fluidity at the original water level of 35-40 % . Accordingly,the water content was increased to about 40-45%, with samples taken during firing for later analysis. At the time of firing, the slurries were up to 12 months old and had set almost solid, and all drums had to be redispersed with a power mixer, a procedure that took 1-2 days. With the increased water, the mixtures were found to fire with relatively little further trouble. This established that settled slurries of considerable age could be satisfactorily redispersed and
Essenhigh et al.
184 Energy &Fuels, Vol. 8,No. 1, 1994
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Figure 1. Variation of wall loss (H,)with thermal output (HB): center line is linear regression;bounds are fl standarddeviation.
successfully fired without either air preheat or other thermal support. This is an important practical consideration. One important characteristic of the sorbent behavior in the flame gases was that part of the sorbent dropped out of suspension (a well-known problem) and deposited on the furnace floor, partly on the water tubes. As the tube were progressively covered, the whole energy balance changed continuously, as described below, though sufficiently slowly that the behavior was essentially quasisteady state, and we found that the effective rate of deposition could be quantitatively monitored by the changes in wall "loss". Overall, there was no evident difference in the firing and flame characteristics of the different coals, or with different sorbents, or sorbent ratios, except for the differences in SOP emissions, as described elsewhere.3~~ Although firing was at a nominal 15% excess air, the actual values varied quite extensively because of the vagaries of firing this type of fuel. In particular, the exact level of water in the slurry was imprecisely known until after the fact when the samples taken during a run had been analyzed, and the measured mass flow of slurry was corrected to dry content. Energy Balance Closure. The basis used for determining closure was to calculate the wall loss (H,)by difference using eq 1, and to compare the values obtained with those determined previously4 and in other test firings using both oil and gas. In doing this, it was necessary to restrict the data used in the comparison to those obtained at the start of firing when the furnace was reasonably clean and there was believed to be relatively little sorbent fallout, except for the tests when the time dependence was being followed. The results of the comparison are illustrated in Figure 1. The solid points are those obtained in this work; the open points are those obtained by Li et aL4 The average line is the linear least-squares average of all data points, and the bounding lines are the range in ordinate (H,) values at 1 standard deviation. The SD in the ordinate values is about 30 MBtu/h, corresponding to 2-3% of the total thermal input, depending on firing rate. As these results show,the trend is alinear rise, in support of eq 4, with the intercept and slope giving the required experimental values of Hwo and slope, m. Characteristically, as found previously, the wall loss was primarily or totally a function of output; no evident dependence on excess air or slurry type has been identified, and the data appear to be randomly scattered in regard to those factors.
Figure 2. Variation of thermal efficiencywith time for run CC7. Falling efficiency is due to progressive sedimentationof sorbent on thermal-load tubes.
t
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Figure 3. Variation of thermal input (Hi); stackloss(ITg);thermal output (Ha); and wall loss with time (run CC7).
Time Dependence. The influence of sorbent sedimentation appeared, as noted above, as a time-dependent factor. The dependence on time is illustrated in Figure 2 using the data in run CC7. This figure shows the progressive fall in thermal efficiency over a period of 1h as the sorbent settled on the thermal-load tubes. The thermal input rate was effectivelyconstant over this period as shown in Figure 3, which also shows the proportional enthalpy variations of load, exhaust, and wall loss. Since the load output, H,, was falling when& was constant then, had to be rising. This is by eq 4, the balance (H,+ Hg) seen in Figure 3 as nearly a 10% rise in the (larger) exhaust gas loss, Hg, but about a 100% "rise" in the smaller wall loss, H,. In absolute values, the "increase" inH, was about of 50 100 MBtu/h which was twice the increase in Hg, MBtu/h. This increase in wall loss, however, was only nominal since a true doubling would require a doubling of the internal wall temperature which was essentially physically impossible and was not shown by the temperature measurements. This pattern was interpreted by examination of the temperature profiles in the roof. Although these showed an increase over the 1-hperiod, they were relatively small, with an increase at the midpoint of the roof of about 50 "C and a maximum increase of less than 100 OC (about 10% ) at the roof interface. For a furnace weight of 5000 lb, and specific heat of 0.25 Btu/lb-OF, with an average temperature rise of 80-100 OF, this corresponded to an increased wall storage of about 100MBtu/h. These figures show that the real increase in true wall loss was only about l/lOth of the 100% nominal increase, and the balance of the apparent increase was due to the increased storage in the wall. Thus, although the wall and roof temperatures did rise, this rise was essentially second order compared
Performance Characteristics of a Hot Wall Furnace
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Energy & Fuels, Vol. 8, No. 1, 1994 185
15
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Figure 4. (a, top) Variation of exhaust gas specific enthalpy ratio (H,/Hf) with thermal output (Ha).Bounding lines at 0,15, 30, and 45% excess air were obtained by inspection (see text). (b, bottom) Variation of exhaust gas specific enthalpy ratio (Hd Hf) with thermal output (H,) for data values between 0 and 15 %
excess air.
to the (impossible) value that would have been needed without the wall storage effect. In summary, the reduced heat removal in the load by mechanical screening of the depositing sorbent was offset and jointly by an increase in both exhaust gas loss (Hg) wall "loss" (I&), reflecting some increase in furnace temperature. The increased wall "loss", however, was mostly (90 5% ) increased wall storage, and only the minor balance (10%) was increase in true wall loss. This is an important result as it shows the buffering effect of the masonry on temperature swings.
Performance Curves and Excess Air Level. The experimental data for the performance characteristics, comprising the heat utilization factor, the (normalized) exhaust gas enthalpy, the firing curves, and the thermal (operational) efficiency, are given in Figures 4-7. These plots show bands of data divided into three groups according to (actual) excess air level in the ranges 0-15 % , 15-30%, and 30-45%. The reason for dividing the data into bands was the result of the difficulties already described of determining the true excess air until after the end of a run. Although any one run was consistent, the variation in excess air levels from one run to another made it impossible to obtain a range of firing rate values at a fixed excess air. The lines dividing the data seta are performance curves, calculated at 0,15,30, and 45% excess air accordingto the equations given above using coefficients obtained by the procedure described below. In all the curves it will be seen that there are a number of outliers, mostly corresponding to the "high-firing" side of the sets of curves. Examination of these outliers and location on
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Figure 5. (a, top) Variation of heat utilization factor (a)with thermal output (H& Boundinglinesat 0,15,30, and 45%excess air were obtained by inspection(seetext). (b, bottom)Variation of heat utilization factor (a)with thermal output (Ha)for data values between 0 and 15% excess air.
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Figure 6. Firing curve: variation of thermal input (firingrate) with thermal output.
Figure 7. Efficiency curve: variation of thermal efficiency with thermal output (Ha).
Figure 1showed that the corresponding wall loss was above average. The implication was that these data points corresponded to conditions where there had been some
Essenhigh et al.
186 Energy &Fuels, Vol. 8, No. 1, 1994
I
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0 15 30 45
Table 1. Values of Firing Coefficients 6 9(eff)" amgo EO Hf 1 0.870 0.769 0.690
0.833 0.741 0.667 0.606
0.790 0.754 0.733 0.740
0.555 0.520 0.495 0.470
124 130 136 145
Ha," 655 553 474 411
Effective equivalence ratio, includes the equivalent effect of the moisture in the CWF.
Figure 8. Variation of maximum intrinsicefficiency with excess air: for CWF (thiswork) and for No. 2 oil. Excess air values for CWF include correction for moisture in fuel (plot in accordance with eq 12). 990
; 500.
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Figure 9. Variation of maximum output with equivalence ratio (9) for CWF (this work) and for No. 2 oil. Equivalence ratio values for CWF include correction for moisture in fuel (plot in accordance with eq 13). deposition on the water tubes. The points were not deleted from the data set on that account, however. Determination of Firing Equation Coefficients. To calculate the Performance Curves shown in the Figures 6 and 7, from the eqs 5,6, and 7, it was necessary to obtain the values of three coefficientsused in the equations. Those coefficients are the idle firing requirement, Hfo, the maximum output, Hem,and the intrinsic efficiency, (YEO, which is the limit heat utilization factor at zero output. The dependency of these parameters on excess air is as defined in eqs 11-14. The procedure used to obtain the required parameter values was to use the experimental results shown in Figures 4 and 5 to obtain "best-estimate" values by inspection, and then to compare the values for the different excess air or equivalence ratio conditions using plots (Figures 8 and 9) to test eqs 12 and 13. In the data displays in Figures 4 and 5, Figures 4a and 5a are the complete set of all data values. Figures 4b and 5b are the data values in the range 0-15s excess air. Sets for the other two data bands show similar characteristics. In fitting the (straight) lines shown for the four excess air levels, these were drawn in by inspection, as the best-fit boundary to each relevant data group in each figure, and iterated within the following constraints: (1) that the maximum outputs, H,", in the Figures 4 and 5 were the same in each plot at the four excess air levels; and (2) that the specific enthalpy ratios in Figure 5, of value HBOJHfO [=(I- (Yk,E)], were consistent at each excess air level with the value of CYEO, the y-axis intercept on Figure 4. These were determined using eq 11, with Figure 1as the source
of the values HwO and m, the idle wall loss, and slope of the wall loss line, respectively. Table 1lists all the values obtained. Output Dependence on Excess Air. The values of the maximum output and maximum intrinsic efficiency were then used to test eqs 12 and 13, by the plots shown as Figures 8 and 9, as a measure of validation of the data reduction procedures and determination of the firing parameters. The two plots include data for No. 2 oil firing as well as for the CWF data of this work. Figure 8 shows a linear decline of the maximum intrinsic efficiency with the actual or equivalent excess air as appropriate, in accordance with eq 12. The ordinate intercept and slope of the line, as shown, to give the best-fit according to eq 1 2 are 0.825 and 1.75 X 10-3; these are close to the least squares values of 0.820 and 1.72 X 10-3, Figure 9 is, likewise, the plot according to eq 13of maximum output (see Table 1)against equivalence ratio, 4. Again, the result supports the linearity of the equation, also with acceptable agreement between the line specified by the equation parameters: the slope of the line shown, of value 988, is close to the least squares value of 963; the intercept values at 4 = 1, for the stoichiometric value of the maximum output are almost identical, at 810 and 813 Mbtu/h. These results provide significant support as validation of the firing parameters, and of the equations on which they are based. I t may also be noted that since the plots include data for oil firing as well as for CWF, the maximum output is shown to be dependent primarily on the thermal firing values and not on the fuel type. This is also an important independent result. Firing and Efficiency Curves. Using the values of the three parameter coefficients: Hfo,HSm,and CYEO, listed in Table 1, obtained by the procedures described above for the different levels of excess air, the firing and efficiency curves were then calculated, as shown in Figures 6 and 7. The curves are based on the two equations, eqs 5 and 6. These are the most important and significant of all the results, and they show the expected trends. The firing curves are a set of concave upwards curves, starting from the idle heat values at zero output, and rising to (theoretically) infinite values at finite values of output, given by the determined values of HSm. The efficiency curves are a set of inverted, asymmetric, U-shaped curves, passing through a maximum at an optimum output value and for which both the maximum efficiency and the optimum output increase with decreasing excess air, thus showing the importance of knowing the location of the optimum output for maximum efficiency. These results generally confirm the agreement between these present experiments and the prior work of Li et al.4 The agreements on the excess air ranges in the Figures 6 and 7 are considered adequate. For these curves, the bounding lines contain about 80% of the points in the 0-1596 range, and about 65% of the points in the 15-30% range. A probable reason for some of the outliers is given above. Within this element of uncertainty, the agreement
Performance Characteristics of a Hot Wall Furnace
between the experimental data points and the bounding excess air lines is considered to be good, particularly in view of the experimental difficulties involved and the nature of the fuel being fired.
Discussion and Conclusions The principal results of this paper may be considered to be largely self-evident: primarily, they show that the overall performance of a hot-wall furnace with an asymmetric thermal load, fired with an unusual fuel, can be quantitatively described by the three fairly simple, theoretically-based, mathematical equations. These equations are: the firing equation (eq 51, the operational efficiency equation (eq 6),and the heat utilization factor equation (eq 7). The associated result is the finding that the effect of moisture can be adequately included as an additive, equivalent excess air factor. The clear limitation at this time is the need to obtain the firing coefficients (Hfo,HSm,and EO) empirically from the experimental results, at least at zero excess air. If initial parameter values of the firing coefficients can be determined with sufficient accuracy at any fixed value of excess air, the results then show that it is possible to calculate the parameter values at other excess air levels with acceptable accuracy. This also means that, even if the numerical values of the firing coefficientsare uncertain or unknown, the relative effects of excess air changes can be estimated with some reliability. This alone is a result of substantial value. The secondary result, of significant practical interest nevertheless, is the effect of sorbent deposition on the tubes. The calculations show that the majority of the heat diverted by the sorbent screening enters the wall, with a doubling of the nominal or apparent wall "loss". The majority of this diverted heat, however, was found to be accommodated by increased wall storage, with a relatively small increase in wall and flame temperatures. The drift is sufficiently slowthat the furnace is in quasi-steady state, noting that this is a common condition both in furnace experiments and in many industrial furnace operations. In addition, the increased wall "loss" was found to correlate very well with apparent changes in SO2 sorption efficiency3sdue evidently to the correlation between time of flight of the sorbent in the flame gases and the rate of sorbent deposition on the load tubes. This is discussed more fully else~here.~~~ Also discussed in those previous reports are measurements of internal temperature profiles and evaluation of their significance. I t will sufficeto summarize those results here, noting that suction pyrometer and two-color py-
Energy & Fuels, Vol. 8, No. 1, 1994 187
rometer measurements showed roughly parallel axial and paraxial temperature profiles, but up to 300-400 "C apart. The higher, two-color pyrometer measurements were interpreted as due to the burning coal particles, and the suction pyrometer measurements were interpreted as the flame gas and sorbent particle temperatures, with coal particles at temperatures high enough to sustain good combustion with (essentially) total burn-out while the sorbent particles remained within the preferred temperature bound (850-1250 "C) for good sorption efficiency" (up to 65% in these experiments3). In conclusion, these experiments support the validity and applicability of the furnace analysis procedure, even for fuels with high moisture loadings, and demonstrate its value for interpreting performance with respect both to the furnace and to the emission-control experiments that were the stimulus for the work.
Acknowledgment. We appreciatively acknowledge support of this work by the Stateof Ohio Coal Development Office (D. A. Burger, Director) under contract CDO/R86-34with partial support from the Combustion Laboratory Industrial Fund. Glossary CP
E, % hi
G, L mv
specific heat of POC (products of combustion) excess air specific enthalpy stoichiometric air/fuel ratio (mass basis) enthalpy of combustion enthalpy latent heat (enthalpy) of evaporation of water mass ratio of water to dry coal in CWF = w%/(lOO - w%) empirical linearization parameter: see Eq. 4 et seq flame temperature intrinsic efficiency thermal efficiency equivalence ratio = 1/(1 + E%/100)
Subscripts i=f fuel =g exhaust gas 1= s useful output wall loss i=w i=m maximum value i,s value at stoichiometric i,E value at excess air E %
i
Superscripts k =0 value at idle k =m maximum value
