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Energy conservation in alcohol distillery with the application of pinch technology
Energy Conversion & Management 40 (1999) 1495±1514
Energy conservation in alcohol distillery with the application of pinch technology Antonio Ficarella, Domenico Laforgia* UniversitaÁ di Lecce, Dip. Scienza dei Materiali, Via Arnesano, 73100 Lecce, Italy Received 11 July 1998; accepted 22 January 1999
Abstract The energy audit of an operating distillery producing ethyl alcohol from low quality wine and wine dregs is presented. Three dierent processes were analyzed: the production of raw (unre®ned) alcohol using stripping columns working at pressure lower than atmospheric, the same production using columns at higher pressure and the production of neutral (re®ned) alcohol. The operational and design data of these three processes were used to compute mass and energy balances. The liquid streams in a distillery are multicomponent nonideal solutions. The data reported in the literature for ethyl and methyl alcohol±water mixtures were utilized, together with purposely developed correlations for density, speci®c heat and vapor±liquid equilibrium and simpli®ed exergy formulas. The methodologies used to study the alcohol production are based on the pinch technology approach: the detailed energy balances of the three industrial processes are presented. The main step was to present the heat sources and sinks of the production processes in the grand composite curve, laying out all the thermodynamic opportunities of any heat recovery. Thermodynamic analysis methods were used to minimize the heating energy needed by the production processes when using heat pump systems. The use of heat pump systems, mainly based on mechanical vapor re-compression, has proven to be eective in energy saving and pro®table in other applications, as the concentration of gelatine, distillation of organic vapors, pulp drying and beer production. Dierent heat pump systems were investigated and compared with respect to fuel utilization and capital expenditures: electrical engine driven compressors, gas engine driven compressors, steam turbine driven compressors, gas absorption chillers, steam absorption chillers and thermo-vapor re-compression. It showed that the cogeneration of mechanical energy and heat to drive vapor compression (the so-called thermodynamic heating method of Frutschi et al.) is superior to other types of systems. Then, for mechanical vapor compression systems, two dierent applications to dierent production processes were analyzed: (a) a system using commercially available refrigerants and (b) a heat pump cycle using water from the bottom of the distillation columns or steam condensate as
* Corresponding author Tel.: +390-832-32-02-39; fax: +390-832-32-05-25. E-mail address: [email protected] (D. Laforgia) 0196-8904/99/$ - see front matter # 1999 Elsevier Science Ltd. All rights reserved. PII: S 0 1 9 6 - 8 9 0 4 ( 9 9 ) 0 0 0 5 1 - 5
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Nomenclature COP T
coecient of performances temperature (K)
Subscripts c condenser C Carnot e evaporator f fuel W mechanical, electrical power
working ¯uid. The thermodynamic analysis, based on performance coecients and fuel utilization, as well as the economic pro®tability in terms of costs, bene®ts and payback period, were discussed in detail. # 1999 Elsevier Science Ltd. All rights reserved. Keywords: Pinch technology; Thermodynamic analysis; Heat pumps; Distillation
1. Introduction In recent years, alcohol has become an important alternative to Diesel and fuel oil. Consequently, agrofood industries have been encouraged to diversify their businesses with nonfood products. Also, the European Community (EC) has recently reduced expenditures on intervention buying and storage of surplus products, thus favouring diversi®cation of energy suppliers. In view of this, energy conservation has become increasingly important due to depletion of natural resources and concern for the environment. It is also a must for the enterprise's competitiveness in the international market, as well as for the factory under investigation. One way to prevent unnecessary waste of primary energy and environmental damage is the installation of heat pump systems, as proved by Stachel et al. [11], who investigated cogeneration systems and heat pump systems with respect to fuel utilization for realistic site conditions. Moreover, several European Community Demonstration Projects were developed, applying heat pumps to agro-food industries, chemical plants, beet pulp drying and technical gelatine production (see EC [3±7]). The integration of a new process into the existing facility provides signi®cant improvements in the design of process plants that would minimize the net cost of energy purchase. The most useful tool that enables this design advance is pinch technology. Based on the representation of the entire process on a temperature±enthalpy diagram, pinch technology represents the cumulative heat sources and heat sinks in the process which allows optimization of the heat exchanger networks and, consequently, minimization of the costs [9,10]. Also, exergy analysis is a valuable tool for achieving substantial improvements in process plant energy use [1,8,10].
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Exergy, or second law, analysis of an alcohol distillery was performed by Castier and Rajagopal [2], who evaluated the second law eciency in important sectors of the plant and in the whole process. The present paper presents a pinch technology analysis of a distillery, aiming at optimization of the heat exchangers network for maximum energy recovery and the application of a combined cogeneration power plant and heat pump. The analysis has been performed in a distillery which produces ethyl alcohol from low quality wine and wine dregs. In the production process of ethyl alcohol, a solution of 5±10 vol% alcohol in water must be concentrated and re®ned to about 92 vol% (raw or unre®ned alcohol) or 96 vol% (re®ned or so-called neutral ethyl alcohol) by means of multiple stage stripping columns. 2. Plant description Figs. 1±3 show the outlines of an existing industrial site, located in the South of Italy, where unre®ned (raw) alcohol and ethyl (re®ned or `neutral') alcohol are produced using low quality wine and wine dregs, saturated steam at 0.7 MPa for heating and water at ambient temperature for cooling. The case studies and processes described in this paper are based on
Fig. 1. Schematic of raw alcohol production cycle (1st technology).
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Fig. 2. Schematic of ethyl alcohol production cycle (1st technology).
simplifying assumptions. Some process data remain ill-de®ned in the absence of full information. In the ®rst process (Fig. 1), where raw alcohol is produced, a solution of 5±10 vol% alcohol in water (wine or dregs) feeds the distillation (stripping) column C210, operating at a pressure of 0.06 MPa abs, where fractional distillation takes place, 0.7 MPa abs saturated steam is used in a heat exchanger at the bottom of the column to provide the heat necessary to alcohol vaporization. While water is discharged from the bottom, some low boiling temperature products are sent to column C220, where fractional distillation is repeated to separate useful products, which are sent back to the main column, from by-products. In the middle of C210, low concentration (40 vol% with water) alcohol vapors are produced, which feed the concentration column C230. After a second fractional distillation, the bottom product (water) is discharged into C210, while the distillate is upgraded in a two-stage condenser. Some partially condensed low grade product is sent back to C210, while the high grade (90±92 vol%) raw alcohol is ®nally condensed. In this process, S1-1 and C4-1 are the same heat exchangerÐ the wine is heated by the water discharged from C210. In the second process (Fig. 2, the ®rst two columns are the same as the ®rst process), the raw alcohol produced in C210 and C230 is sent to the scrubber C250, operating at atmospheric pressure, where the liquid from the bottom of the rectifying column C240 is sprayed countercurrent to the alcohol vapors. Some low boiling products, discharged from the top of
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Fig. 3. Schematic of raw alcohol production cycle (2nd technology).
C250, are re®ned in column C260, while the liquid discharged from the bottom is sent to the rectifying column C240. In this column, operating at higher pressure (0.2 MPa abs), a high grade alcohol (96 vol%) is obtained, feeding the heat exchanger at the bottom of C260 with saturated steam. In the last column, C270, a fractional distillation takes place to separate the methyl alcohol at the top of the column from the re®ned ethyl alcohol at the bottom (96 vol% with water). In this process, S1-1 and C4-1 are the same heat exchanger, as well as S2 and C10-1 and C6-1 and S11. In the third process (Fig. 3), raw alcohol is produced with the same process described in Fig. 1, but in this case, a higher operating pressure is used (0.2 MPa abs), while 0.7 MPa abs saturated steam directly feeds the bottom of the distillation column C10 (no heat exchanger is used). The liquid discharged from the bottom (consisting of the water in the wine or in the dregs plus the steam condensate) is ¯ashed at a pressure of 0.04 MPa abs. The produced steam is compressed, using an ejector fed by saturated steam at 0.7 MPa abs and returned to the bottom of C10. The concentration column C30 works in the same way as C230. Some high boiling temperature products (oils) are discharged from the bottom. Heat exchangers S1-1 and C5 are the same unit. Tables 1±3 list the details of the heat and mass balances in the existing processes. Selected instantaneous operational data, complemented with operational and design values, are used to
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Table 1 Raw alcohol production Heat exchanger (a) First process S1-1 S1-2 S2 S2A C3 S3 C4-1 C4-2 S4 Heating (in) Cooling (out)
Mass
Flow (kg/s)
Wine Wine Water Water Low-boiling Low-boiling Raw alcohol Raw alcohol Raw alcohol
(b) First process optimized S1-1 Wine S1-2 Wine S2 Water S2A1 Water S2A2 Water C3 Low-boiling S3 Low-boiling C4-1 Raw alcohol C4-2 Raw alcohol S4 Raw alcohol Heating (in) Cooling (out)
Tin (8C)
Tout (8C)
Heat out/in (kW)
Recovery
2.315 2.315 0.353 2.083 0.062 0.021 0.633 0.633 0.211
15 60 87 87 79 79 65 65 65
60 79 87 15 79 15 65 65 15
494 220 729 ÿ631 ÿ132 ÿ6 ÿ494 ÿ160 ÿ35 1443 ÿ1458
C4-1
2.315 2.315 0.353 2.083 2.083 0.062 0.021 0.633 0.633 0.211
15 64 87 87 25 79 79 65 65 65
64 79 87 25 15 79 15 65 65 15
544 170 729 ÿ544 ÿ87 ÿ132 ÿ6 ÿ654 0 ÿ35 1443 ÿ1458
S1-1 949.15 ÿ963.71
S2A1 S1-1
899.28 ÿ913.84
calculate the mass and energy balances. The formulas for the energy of the multicomponent nonideal mixtures used by Castier and Rajagopal [2] have been employed for density data, speci®c heat and formation data.
3. Pinch technology analysis The ®rst step of this analysis is to present the complete process on a temperature±enthalpy diagram by a grand composite curve, which represents the cumulative heat sources and heat sinks in the process. A full description of pinch technology for the solution sequence in question is provided by Linnho et al. [9], and Linnho and Alanis [10]. Figs. 4±6 represent the composite curves of the three processes analysed: raw alcohol production ethyl alcohol production and raw alcohol production using steam obtained by ¯ashing the condensate. The curves are named as follows:
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GCINHE
GCINCO GCACHE GCACCO GRCOOT
1501
which describes the temperature±enthalpy diagram for the heat sinks in the process, corresponding to the mass ¯ows to be heatedÐnot considering heat recovery from some other ¯ow. The heat need of the process in the temperature range T1 ±T2 is H2 ±H1 ; which describes the temperature±enthalpy diagram for the heat source in the process, corresponding to the mass ¯ows that need to be cooledÐnot considering heat recovery to some other ¯ow; same as above for heat sinks but considering heat recovery in the actual processes; same as above for heat sources but considering heat recovery in the actual processes; optimal grand composite curve; represents the solution obtained at the so-called targeting stage, calculated considering an unfeasible DT 0 (no heat across the pinch) in the heat exchangers.
The pinch divides the process into two thermodynamically separate systems, as depicted in Table 2 Raw alcohol production Heat exchanger (a) First process S1-1 S1-2 S2 S2A C3 S3 C4-1 C4-2 S4 S5 C6-1 C6-2 S7A C8 S8 S9 C10-1 C10-2 S11 S11A C12 S12 Heating (in) Cooling (out)
Mass
Flow (kg/s)
Wine Wine Water Water Low-boiling Low-boiling Raw alcohol Raw alcohol Raw alcohol Ethyl alcohol Low-boiling Low-boiling Water Low-boiling Low-boiling Ethyl alcohol Ethyl alcohol Ethyl alcohol Ethyl alcohol Ethyl alcohol Methyl alcohol Methyl alcohol
1.042 1.042 0.168 0.935 0.028 0.009 0.293 0.293 0.098 0.069 0.067 0.067 0.002 0.012 0.002 0.652 0.652 0.652 0.034 0.091 0.012 0.002
Tin (8C)
Tout (8C)
Heat out/in (kW)
Recovery
15 60 87 87 79 79 65 65 65 79 79 79 101 79 79 121 98 98 65 65 52 52
60 79 87 15 79 15 65 65 65 79 79 79 15 79 15 121 98 98 65 15 52 15
222 99 347 ÿ283 ÿ59 ÿ3 ÿ222 ÿ92 0 66 ÿ32 ÿ40 ÿ1 ÿ13 ÿ1 625 ÿ347 ÿ278 32 ÿ14 ÿ14 0 1392 ÿ1399
C4-1 C10-1
S1-1
S11
S2 lC6-1
790.32 ÿ797.18
(continued on next page)
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Table 2 (continued ) Heat exchanger
Mass
(b) First process optimized S1-1 Wine S1-2 Wine S2 Water S2A1 Water S2A2 Water C3 Low-boiling S3 Low-boiling C4-1 Raw alcohol C4-2 Raw alcohol S4 Raw alcohol S5 Ethyl alcohol C6-1 Low-boiling C6-2 Low-boiling S7A Water C8 Low-boiling S8 Low-boiling S9 Ethyl alcohol C10-1 Ethyl alcohol C10-2 Ethyl alcohol S11 Ethyl alcohol S11A Ethyl alcohol C12 Methyl alcohol S12 Methyl alcohol Heating (in) Cooling (out)
Flow (kg/s) 1.042 1.042 0.168 0.935 0.935 0.028 0.009 0.293 0.293 0.098 0.069 0.067 0.067 0.002 0.012 0.002 0.652 0.652 0.652 0.034 0.091 0.012 0.002
Tin (8C)
Tout (8C)
Heat out/in (kW)
Recovery
15 64 87 87 25 79 79 65 65 65 79 79 79 101 79 79 121 98 98 65 65 52 52
64 79 87 25 15 79 15 65 65 65 79 79 79 15 79 15 121 98 98 65 15 52 15
244 77 347 ÿ244 ÿ39 ÿ59 ÿ3 ÿ314 0 0 66 ÿ32 ÿ40 ÿ1 ÿ13 ÿ1 625 ÿ347 ÿ278 32 ÿ14 ÿ14 0 1392 ÿ1399
S2A1 C10-1 S1-1
S1-1
S11
S2 C6-1
768.61 ÿ775.47
Table 3 Raw alcohol production Heat exchanger (a) Second process S21-1 S21-2 S22A S22B S22A-1 S22A-2 C23 S23 S24A C25 S25 Heating (in) Cooling (out)
Mass
Flow (kg/s)
Wine dregs Wine dregs Water 0.6 MPa steam Condensate Water Low-boiling Low-boiling High boiling Raw alcohol Raw alcohol
3.345 3.345 0.056 0.056 0.056 3.070 0.110 0.033 0.033 0.686 0.208
Tin (8C)
Tout (8C)
Heat out/in (kW)
15 61 15 165 76 76 100 100 121 98 98
61 118 165 165 15 15 100 15 15 98 15
709 969 35 115 ÿ14 ÿ788 ÿ241 ÿ14 ÿ17 ÿ709 ÿ61 1828 ÿ1843
Recovery C5
S1-1 1119.55 ÿ1134.23
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Table 3 (continued ) Heat exchanger
(b) Second process S21-1A S21-1B S21-2 S22A S22B S22A-1A S22A-1B S22A-2A S22A-2B C23 S23 S24A C25A C25B S25 Heating (in) Cooling (out)
Mass
optimized Wine dregs Wine dregs Wine dregs Water 0.6 MPa steam Condensate Condensate Water Water Low-boiling Low-boiling High boiling Raw alcohol Raw alcohol Raw alcohol
Flow (kg/s)
Tin (8C)
Tout (8C)
Heat out/in (kW)
15 58 93 15 165 76 25 76 25 100 100 121 98 98 98
58 93 118 165 165 25 15 25 15 100 15 15 98 98 15
671 574 433 35 115 ÿ12 ÿ2 ÿ659 ÿ129 ÿ241 ÿ14 ÿ17 ÿ574 ÿ135 ÿ61 1828 ÿ1843
3.345 3.345 3.345 0.056 0.056 0.056 0.056 3.070 3.070 0.110 0.033 0.033 0.686 0.686 0.208
Recovery
S2A-1A-2A C25A
S1-1 S1-1
S21-1B 583.53 ÿ598.21
Figs. 4±6. The heat sinks curves (heated ¯ows) are presented in the positive heat region, while the heat sources curves (cooled ¯ows) are located in the negative heat region. This calculated target GRCOOT represents the minimum heat requirement for the process, an ideal solution that, however, violates the second law of thermodynamics, but it is very useful for a ®rst evaluation of possible improvements in the process under investigation. 4. Process pre-optimization using the pinch method The raw alcohol production process (Fig. 1 and Table 1) requires 949 kW of heat and 963 kW of cooling, see also curves GCACHE and GCACCO of Fig. 4a. The total heat requirement of the process is 1443 kW, but 494 kW are recovered from some cooled ¯ows (heat is recovered at 658C during condensation in C4-1 and used to heat the wine from 158C in S1-1)Ðthe total cooling power available is 1458 kW. The optimal target, considering only the heat exchangers network optimization, is shown by the GRCOOT curve of Fig. 4a, and it is achieved. The (unfeasible) target for heating is 728 kW, with a dierence of 30% with respect to the actual heating power (949 kW). The dierence is mainly due to the calculation of the optimal target considering DT 0 in the heat exchangers. The excess of heat to cool the discharged water in S2A and to complete the condensation in C4-2 is available at temperatures lower than 878C, so that the boiler S2 must be fed by steam. Note that in the optimized process, the wine in S1-1 and -2 is heated by water in S2A, while all the heat of condensation in C4-1 and -2 is available, but useless. In the actual process, matching S1-1 with C4-1, a big
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Fig. 4 (a) and (b). Grand composite curves of raw alcohol production.
dierence in temperature exists. A smaller heat exchanger could be employed, but considering the use of a heat pump, as subsequently discussed, it is better to match S1-1 and -2 with S2A to make available the heat in C4-1 at a higher temperature. In continuation, this last con®guration will be considered (see Table 1b and Fig. 4b). The ethyl alcohol production process of Fig. 2 (see also Table 2a) requires 790 kW of heating power. The total heating need is 1392 kW, but 602 kW are recovered by matching, in the same unit, S1-1 and C4-1, S2 and C10-1 and C6-1 and S11. The heat exchangers network
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Fig. 5 (a) and (b). Grand composite curves of ethyl alcohol production.
is practically in the optimal con®guration. The 625 kW at 1218C necessary in the boiler S9 cannot be recovered from the cooling streams at lower temperatures. Only the ¯ow of 277 kW at 988C in C10-2 could be recovered to heat some ¯ows, as in S1-2 (99 kW) and S5 (66 kW), but some plant complexities and low power of exchangers made the optimization economically unfeasible. The ideal optimal grand composite curve (GRCOOT of Fig. 5a con®rms the previous conclusions, showing an optimal target for heating of 625 kW (the actual value is 26% higher than the ideal one). As previously noted, considering the use of a heat pump, it is better to match S1-1 and -2 with S2A to make available the heat in C4-1 at a higher temperature. In continuation, this last con®guration will be considered (see Table 2b and Fig. 5b).
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Fig. 6 (a) and (b). Grand composite curves of raw alcohol production (2nd technology).
The last production process for raw alcohol, based on ¯ashing of the discharge condensate (Fig. 3), shows a heat consumption of 1119 kW (Table 3): Considering the total amount of heating power, 709 kW are recovered during the condensation in C25 (C25 and S21-1 are the same unit), but as shown by the ideal grand composite curve of Fig. 6a, it is possible to recover (ideally) a further 679 kW, leaving the minimum heating required as 440 kW by matching S22A-1/2 with S21. Considering a reasonable DT for the heat exchanger, it is possible to recover 670 kW from S22A-1/2 and 573 kW from C5 (see Table 3b and Fig. 6b). In this case, the heat requirement is 583 kW, only 32% higher than the target one.
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5. Thermodynamic analysis in presence of heat pumps Fig. 7 presents the representative design point performance data of several commercially available heat pumps for industrial use (heating power > 100 kW) and considering some constraints deriving from their application in the production processes previously seen. Data were supplied by known manufacturers. The points are named as follows: ELECOM GASCOM
STECOM GASABS STEABS GCONCO STEEJE
mechanical type heat pump with the compressor driven by an electrical motor; combined mechanical type heat pump and cogeneration plant, with the compressor driven by an internal combustion engine fueled by natural gas and with recovery of heat from o engine exhaust gasesÐthe heat recovery from engine cooling was neglected; combined mechanical type heat pump and cogeneration plant, with the compressor driven by a steam turbine (inlet steam at 4.5 MPa/3508C, discharge at 0.9 MPa) and with recovery of heat from o turbine discharged steam; absorbing type heat pump fueled by natural gas; absorbing type heat pump fed by steam; mechanical type heat pump with the compressor driven by an internal combustion engine fueled by natural gas, no heat recovery considered; ejector type heat pump fed by saturated steam at 0.8 or 4.0 MPa with ¯ashing water at 908C and heat recovery at 1208C.
For the analyzed heat pumps, the fuel consumption was considered in the calculations of the COP, boiler eciency was considered in the steam production, as well as the average eciency of the national grid production for the equivalence fuel±electricity (2.67 kJfuel/kJelectricity). The traditional COP is evaluated as:
Fig. 7. COP referred to fuel consumption.
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COPW E=W
1
where E is the heat recovered and W, the work employed. The COP referred to the fuel will be written as COPf E=Mf
2
where Mf is the fuel used in the prime mover (mechanical or thermal one). Moreover, the COP referred to Carnot's cycle is written as COPC Tc = Tc ÿ Te Tc =DT
3
where Tc is the temperature at which heat is available (temperature of the external heated ¯uid in the heat pump condenser) and Te is the temperature of the cooled external ¯uid in the heat pump evaporator. On an average, it is useless to have temperature levels in the condenser and in the evaporator such that COPC is less than 6 (for example, Tc 1008C and DT 608C), because at this level, COPf is equal to 1 (useless heat pump). This can be explained by the fact that in commercially available heat pumps, the ratio COPW to COPf is, on average, equal to 0.5, and then, when COPC is equal to 6, COPW will be equal to 3. Considering that the normal eciency of traditional systems in the production of work is of the order of 25±35%, without cogeneration, as much heat is recovered as fuel is burned. Fig. 7 shows that the combined cogeneration power plant and heat pumps (thermodynamic heating method analyzed by Stachel et al. [11] ÐGASCOM in the ®gureÐshows better performances for low COPC , while for COPC higher than 8, the electrical compressors (ELECOM) show better performances. On an average, the national grid electrical eciency is higher than the engine eciency in the commercially available GASCOM. Gas compressors performances worsen if the exhaust gas heat is not recoveredÐGCONCO points in Fig. 7. The absorption heat pumpsÐGASABS and STEABSÐshow good performances. Unfortunately, the typical COPC is highÐthe maximum working DT is of the order of 408C and the maximum absolute condensing temperature is approximately 1008C. The steam compressors (STECOM) show poor performances because of the very low eciency in the work conversion in industrial counterpressure steam turbines. Moreover, they require superheated steam, usually not used in the agrofood industries under investigation. The same behavior is shown by the ejector type compressors because of the low mechanical eciency of ejectors and because the steam condensate is mixed with the sewage of the process (see Fig. 3), so that its enthalpy content is lost. The costs of the recovered heat are presented in Fig. 8. They include only the cost of the energy used to move/operate the heat pump and not the cost of the recovered heat. The calculations were made with the following common characteristics: . cost of electricity: 22.92 ECU/GJ (0.0825 ECU/kWh) . cost of fuel for thermal use: 3.79 ECU/GJ . cost of fuel for mechanical energy production (considering some exemptions from taxes): 3.51 ECU/GJ. The heat reference cost was assessed equal to 4.74 ECU/GJ considering a traditional boiler
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Fig. 8. Heat costs.
with 80% eciency. Because of the high costs of electricity, the electrical compressors (ELECOM) do not compete with the gas compressors (GASCOM) even without considering cogeneration (GCONCO points) which show a heat cost lower than the reference one. Also, the absorption systems (GASABS and STEABS) and the steam turbine compressors (STECOM) show satisfactory economical performances. 6. Heat pumps performances The analysis of the alcohol production processes, performed using pinch technology, shows the possibility to apply heat pump technology to reduce the energy consumption. The three production processes previously examined show pinch temperatures of 87, 98±121 and 1008C, respectively (Figs. 4±6). Energy is saved overall in the process only if the heat pump operates across the pinch, i.e., when it pumps from the process source to the process sink, as stated by Linnho et al. [9]. It should be pointed out that no utilities optimization was consideredÐthat means, steam pressure optimization. In fact, the larger amount of heat in each process is needed at one temperature level only, whereas negligible heat is required at lower temperatures. Moreover, no heat pumps based on alcohol vapor compression were considered because the complexity and risks of such technology (alcohol is easily in¯ammable, hot alcohol vapors are corrosive) are not suitable for low and medium size agrofood industries. All processes were pre-optimized for the installation of heat pumps, making available heat at higher temperatures and matching the process heat exchangers with the lower technically feasible DT, making best use of the heat recovery network driving forces. Heat pumps with gas engine driven compressor have been considered, with heat recovery from the gas engine (used to produce steam directly) and, if possible, from engine cooling (at
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lower temperature, about 808C). As previously seen, the electrical compressors have high operating costs, and the absorbing systems or steam compressors show low performances or useless operating temperature ranges. Moreover, the steam compressors require superheated steam which is normally not available. 7. Traditional heat pumps A traditional closed cycle heat pump was considered. Calculations were made with the following common characteristics: . heat losses: 10% . ratio of heat recovery from engine discharged gases to heat pump heat: 0.25 . ratio of COPW to COPC : 0.50. The heat losses include losses in the heat pump, in the piping system and in the heat exchangersÐconsidering that the heat sinks and sources may be distant in the production plant lay-out. The recovery from engine o gases was considered constant because the heat is used to produce saturated steam at 0.7 MPa. In any case, for environmental reasons, the gas temperature at the stacks cannot be lower than 1508C. Moreover, smaller heat exchangers (less than 100 kW) match will not be considered for economic unfeasibility, and for the same reason, only one temperature level for condensation and evaporation has been considered. Table 4 Traditional heat pump Process
Table 1
Table 1
Table 2
Table 2
Table 3
Evaporators Condensers Tc (8C) Evaporator 1 E1 (kW) Te1 (8C) Evaporator 2 E2 (kW) Te2 (8C) Gas recovery E (kW) Te3 (8C) Performances COPC COPm COPf Fuel consumption (kW) Heat recovered (kW)
C3, C4-1/2 S2 87
C4-1/2 S2 87
C10-2, C4-1/2 S9 121
C4-1/2 S9 121
C23, C25B S1-2 118
131.69 79
131.69 79
277.59 98
277.59 98
240.63 100
653.61 65
0.00 65
313.90 65
0.00 65
134.75 98
785.30 500
131.69 500
591.49 500
277.59 500
375.38 500
16.36 7.36 2.56 399.54 1022.29
45.00 20.25 7.04 22.15 155.85
7.04 3.17 1.10 884.11 972.63
17.13 7.71 2.68 133.97 358.84
19.55 8.80 3.06 155.86 476.46
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Analyzing the ®rst production process (Table 1b and Fig. 4b), it is clear that the heat pump condenser will match the heat exchanger S2 at 878C, while the evaporators will match C4-1/2 at 658C and C3 at 798C. Table 4 shows very interesting performances of the heat pump, with a recovery of 1022 kW of heat at 878C (or more, including the hot gases heat recoveryÐnote that some heat will be used for some other process) using 399 kW of fuel. In the same way, the analysis of the second production process (Table 2b and Fig. 5b) shows that the heat pump condenser matches S9 at 1218C, while the evaporators match C10-2 at 988C and C41/2 at 658C. Of course, the increase of the condensing temperature reduces the system performances, reducing COPW to 3.17 and COPf to 1.10, making the heat pump practically useless. Finally, the third production process (Table 3b and Fig. 6b) behavior suggests to match the heat pump condenser to S21-2 at 1188C (S22 is the boiler and anyway, it is neglected for its high temperature) and the evaporators to C23 and C25B at temperatures, respectively, of 100 and 988C. 470 kW of heat are recovered (some heat will be used elsewhere) using 155 kW of fuel. The traditional closed cycle heat pumps show a major drawback for the condenser high temperature. For a condenser at 908C (actual refrigerant temperature of 1008C), the HCFC-22 (as its ecological substitutes R047C or R134a) and the HCFC-717 (NH3) show a condensation pressure higher than 5 MPa and about 1.4 MPa for the HCFC-21, making the system technically and economically unfeasible (especially for the compressor discharge high temperatures and the lubrication problems). 8. Heat pumps using water ¯ash Water presents a more favorable link between temperature and pressure. A heat pump cycle (Fig. 9) can be realized using water ¯ash in a boiler at low temperature and pressure to
Fig. 9. Schematic of the heat pump with sewage ¯ash.
1512
Table 5 Heat pump with ¯ash Table 1
Table 2
Table 2
Table 2
Table 3
Table 3
Table 2 and 3
Evaporators
C3, C4-1/2
C10-2, C4-1/2
C10-2, C4-1/2
C10-2, C4-1/2
C23, C25B
C23, C25B
S2 87
S9 121
S9 121
S9 121
S1-2 118
S1-2 118
C10-2, C4-1/2 C23, C25B S9, S1-2 121
55 0.0160
55 0.0160
70 0.0288
88 0.0583
70 0.0288
90 0.0630
88 0.0583
131.69 79
277.59 98
277.59 98
277.59 98
240.63 100
240.63 100
240.63 100
653.61 65
313.90 65
0.00 65
0.00 65
134.75 98
0.00 98
412.34 98
192.26 80
363.37 80
110.72 80
0.00 80
148.46 80
0.00 80
0.00 80
192.26 192
363.37 363
110.72 111
48.17 48
148.46 148
38.59 39
113.31 113
97 0.0829 256
131 0.3139 484
131 0.3139 148
131 0.3139 64
128 0.2791 198
128 0.2791 51
131 0.3139 151
11.25 4.92 1.71 737.74 1260.48
5.97 3.26 1.13 1394.31 1578.84
7.73 3.85 1.34 424.85 568.70
11.94 5.37 1.87 184.84 345.21
8.15 3.87 1.34 569.67 765.40
13.96 5.69 1.98 148.07 292.97
11.94 5.37 1.87 434.80 812.03
107,636 57,243 50,393 174,757 3.47
134,821 108,187 26,634 271,845 10.21
48,563 32,965 15,598 116,505 7.47
29,478 14,342 15,136 67,961 4.49
65,360 44,201 21,159 145,631 6.88
25,018 11,489 13,529 58,252 4.31
69,342 33,737 35,605 121,359 3.41
Condensers Tc 8C Boiler parameters Flashing temperature (8C) Flashing pressure (MPa) Evaporator 1 E1 (kW) Te1 (8C) Evaporator 2 E2 (kW) Te2 (8C) Engine cooling recovery Ecooling (kW) Tcooling (8C) Gas recovery Egases (kW) Tgases (8C) Compressor parameters Condensation temperature (8C) Condensation pressure (MPa) Mechanical energy (kW) Thermal performances COPC COPm COPf Fuel consumption (kW) Heat recovered (kW) Economical performances Energy savings (ECU/year) Operational expenditures (ECU/year) Total savings (ECU/year) Investment cost (ECU) Pay-pack period (year)
A. Ficarella, D. Laforgia / Energy Conversion & Management 40 (1999) 1495±1514
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produce enough driving forces to recover heat from the process ¯ows to be cooled. The produced low pressure steam will be compressed to a temperature level suitable to feed the energy sinks in the process. The gas engine driven compressor will be of centrifugal type, as used in some other similar applications (see EC [4±7]). Calculations were made with the following common characteristics: . . . . . . . . .
heat losses: 10% compressor thermodynamic eciency: 70% compressor mechanical eciency: 90% minimum driving force (DT): 108C natural gas consumption in the engine: 2.88 kJ/kJ of work ratio of heat recovery from engine discharge gases to engine power: 0.75 ratio of heat recovery from engine cooling to engine power: 0.75 plant usage: 5000 h/year maintenance and other operational expenditures (fuel excluded): 2.7 ECU/GJ power
Refer to the previous case for engine performance. Summarizing, the ®rst process (see Table 5) presents good thermal and economical performances: 1260 kW of heat recovered, using 737 kW of fuel power, with an investment of 175,000 ECU and savings, of 50,000 ECU/year. The pay back period (PBP) will be 3.5 years. For the second process, dierent ¯ashing temperatures can be chosen. To recover all the available heat (from engine cooling at 808C, from the two process evaporators at 98 and 658C and from the hot gases at 5008C), a temperature of 558C has to be achieved in the main boiler. The high condenser temperature (1218C) implies poor performances (more than 10 years of PBP). Also when the ¯ash is increased to 708C (C4-1/2 not matched to the boiler), the plant presents poor performances (PBP equal to 7.5 years). Only the ¯ash at 888C, with recovering heat from C10-2 and from the hot gases, implies a recovery of 345 kW of heat requiring 184 kW of fuel energy. The PBP is 4.5 years with an investment cost of 68,000 ECU. Analyzing the third production process, a ¯ashing temperature of 708C, allowing for heat recovery from engine cooling, involves poor performances (PBP is 6.9 years), but with the increase of such temperature to 908C, good thermal and economical performances can be obtained. 291 kW can be recovered using 148 kW of fuel with a PBP of 4.3 years and 58,000 ECU of investment.
9. Concluding remarks Pinch technology analysis was applied to processes to produce alcohol from low quality wine and wine dregs. The results clearly identify the heat sinks and sources of the process and the possible installation of heat pumps. Moreover, it was clear that optimization of the heat exchangers network can imply the maximum heat recovery and the improvement of heat pump application can optimize the driving forces in the network. Heat pumps based on a water cycle were analyzed, showing good thermal and interesting economical performances. A heat pump plant, working with ¯ashing temperature of 88±908C
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and 118±1218C for heat recovery, can be used at the same time for ethyl alcohol production and raw alcohol production (the last one using the third production process)while achieving interesting overall performances: 812 kW of heat recovered using 434 kW of fuel power and with an investment of 121,000 ECU and savings of 36,000 ECU/year, the pay back period (PBP) will be 3.4 years (see Table 5). References [1] Bidini G, Stecco SS. A computer code using exergy for optimizing thermal plants. Journal of Engineering for Gas Turbine and Power 1991;113:145±50. [2] Castier M, Rajagopal K. Thermodynamic analysis of an alcohol distillery. Energy 1988;13:455±9. [3] EC. Pompe aÁ chaleur sur la vapeur produite par un proceÂde thermomeÂcanique. Demonstration Projects for Energy Saving and Alternative Energy Sources. N. EE/193/79-F, 1979. [4] EC. Mechanical compression of organic vapors. Demonstration Projects for Energy Saving and Alternative Energy Sources. N. EE/012/82-F. 1982. [5] EC. Superheated steam beet pulp drier with mechanical vapor recompression. Demonstration Projects for Energy Saving and Alternative Energy Sources. N. EE/244/84-F, 1984. [6] EC. Concentration of (technical) gelatine using a falling ®lm evaporator with mechanical vapor compression. Demonstration Projects for Energy Saving and Alternative Energy Sources. N. EE/349/85-NL, 1985. [7] EC. Concentration of (technical) gelatine using a falling ®lm evaporator with mechanical vapor compression. Demonstration Projects for Energy Saving and Alternative Energy Sources. N. EUR 12108 EN, 1989. [8] Gaggioli RA, Sama DA, Quian S, El-Sayed YM. Integration of a new process into an existing site: a case study in the application of exergy analysis. Journal of Engineering for Gas Turbine and Power 1991;113:170± 83. [9] Linnho B, Townsend DW, Boland D, Hewitt GF, Thomas BEA, Guy AR, Marsland RH. Process Integration for the Ecient Use of Energy. Rugby, Warks, England: The Institution of Chemical Engineers, 1991. [10] Linnho B, Alanis FJ. Integration of a new process into an existing site: a case study in the application of pinch technology. Journal of Engineering for Gas Turbine and Power 1991;113:159±69. [11] Stachel K, Frutschi HU, Haselbacher H. Thermodynamic heating with various types of cogeneration plants and heat pumps. Journal of Engineering for Gas Turbine and Power 1995;117:251±8.
Energy conservation in alcohol distillery with the application of pinch technology Antonio Ficarella, Domenico Laforgia* UniversitaÁ di Lecce, Dip. Scienza dei Materiali, Via Arnesano, 73100 Lecce, Italy Received 11 July 1998; accepted 22 January 1999
Abstract The energy audit of an operating distillery producing ethyl alcohol from low quality wine and wine dregs is presented. Three dierent processes were analyzed: the production of raw (unre®ned) alcohol using stripping columns working at pressure lower than atmospheric, the same production using columns at higher pressure and the production of neutral (re®ned) alcohol. The operational and design data of these three processes were used to compute mass and energy balances. The liquid streams in a distillery are multicomponent nonideal solutions. The data reported in the literature for ethyl and methyl alcohol±water mixtures were utilized, together with purposely developed correlations for density, speci®c heat and vapor±liquid equilibrium and simpli®ed exergy formulas. The methodologies used to study the alcohol production are based on the pinch technology approach: the detailed energy balances of the three industrial processes are presented. The main step was to present the heat sources and sinks of the production processes in the grand composite curve, laying out all the thermodynamic opportunities of any heat recovery. Thermodynamic analysis methods were used to minimize the heating energy needed by the production processes when using heat pump systems. The use of heat pump systems, mainly based on mechanical vapor re-compression, has proven to be eective in energy saving and pro®table in other applications, as the concentration of gelatine, distillation of organic vapors, pulp drying and beer production. Dierent heat pump systems were investigated and compared with respect to fuel utilization and capital expenditures: electrical engine driven compressors, gas engine driven compressors, steam turbine driven compressors, gas absorption chillers, steam absorption chillers and thermo-vapor re-compression. It showed that the cogeneration of mechanical energy and heat to drive vapor compression (the so-called thermodynamic heating method of Frutschi et al.) is superior to other types of systems. Then, for mechanical vapor compression systems, two dierent applications to dierent production processes were analyzed: (a) a system using commercially available refrigerants and (b) a heat pump cycle using water from the bottom of the distillation columns or steam condensate as
* Corresponding author Tel.: +390-832-32-02-39; fax: +390-832-32-05-25. E-mail address: [email protected] (D. Laforgia) 0196-8904/99/$ - see front matter # 1999 Elsevier Science Ltd. All rights reserved. PII: S 0 1 9 6 - 8 9 0 4 ( 9 9 ) 0 0 0 5 1 - 5
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Nomenclature COP T
coecient of performances temperature (K)
Subscripts c condenser C Carnot e evaporator f fuel W mechanical, electrical power
working ¯uid. The thermodynamic analysis, based on performance coecients and fuel utilization, as well as the economic pro®tability in terms of costs, bene®ts and payback period, were discussed in detail. # 1999 Elsevier Science Ltd. All rights reserved. Keywords: Pinch technology; Thermodynamic analysis; Heat pumps; Distillation
1. Introduction In recent years, alcohol has become an important alternative to Diesel and fuel oil. Consequently, agrofood industries have been encouraged to diversify their businesses with nonfood products. Also, the European Community (EC) has recently reduced expenditures on intervention buying and storage of surplus products, thus favouring diversi®cation of energy suppliers. In view of this, energy conservation has become increasingly important due to depletion of natural resources and concern for the environment. It is also a must for the enterprise's competitiveness in the international market, as well as for the factory under investigation. One way to prevent unnecessary waste of primary energy and environmental damage is the installation of heat pump systems, as proved by Stachel et al. [11], who investigated cogeneration systems and heat pump systems with respect to fuel utilization for realistic site conditions. Moreover, several European Community Demonstration Projects were developed, applying heat pumps to agro-food industries, chemical plants, beet pulp drying and technical gelatine production (see EC [3±7]). The integration of a new process into the existing facility provides signi®cant improvements in the design of process plants that would minimize the net cost of energy purchase. The most useful tool that enables this design advance is pinch technology. Based on the representation of the entire process on a temperature±enthalpy diagram, pinch technology represents the cumulative heat sources and heat sinks in the process which allows optimization of the heat exchanger networks and, consequently, minimization of the costs [9,10]. Also, exergy analysis is a valuable tool for achieving substantial improvements in process plant energy use [1,8,10].
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Exergy, or second law, analysis of an alcohol distillery was performed by Castier and Rajagopal [2], who evaluated the second law eciency in important sectors of the plant and in the whole process. The present paper presents a pinch technology analysis of a distillery, aiming at optimization of the heat exchangers network for maximum energy recovery and the application of a combined cogeneration power plant and heat pump. The analysis has been performed in a distillery which produces ethyl alcohol from low quality wine and wine dregs. In the production process of ethyl alcohol, a solution of 5±10 vol% alcohol in water must be concentrated and re®ned to about 92 vol% (raw or unre®ned alcohol) or 96 vol% (re®ned or so-called neutral ethyl alcohol) by means of multiple stage stripping columns. 2. Plant description Figs. 1±3 show the outlines of an existing industrial site, located in the South of Italy, where unre®ned (raw) alcohol and ethyl (re®ned or `neutral') alcohol are produced using low quality wine and wine dregs, saturated steam at 0.7 MPa for heating and water at ambient temperature for cooling. The case studies and processes described in this paper are based on
Fig. 1. Schematic of raw alcohol production cycle (1st technology).
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Fig. 2. Schematic of ethyl alcohol production cycle (1st technology).
simplifying assumptions. Some process data remain ill-de®ned in the absence of full information. In the ®rst process (Fig. 1), where raw alcohol is produced, a solution of 5±10 vol% alcohol in water (wine or dregs) feeds the distillation (stripping) column C210, operating at a pressure of 0.06 MPa abs, where fractional distillation takes place, 0.7 MPa abs saturated steam is used in a heat exchanger at the bottom of the column to provide the heat necessary to alcohol vaporization. While water is discharged from the bottom, some low boiling temperature products are sent to column C220, where fractional distillation is repeated to separate useful products, which are sent back to the main column, from by-products. In the middle of C210, low concentration (40 vol% with water) alcohol vapors are produced, which feed the concentration column C230. After a second fractional distillation, the bottom product (water) is discharged into C210, while the distillate is upgraded in a two-stage condenser. Some partially condensed low grade product is sent back to C210, while the high grade (90±92 vol%) raw alcohol is ®nally condensed. In this process, S1-1 and C4-1 are the same heat exchangerÐ the wine is heated by the water discharged from C210. In the second process (Fig. 2, the ®rst two columns are the same as the ®rst process), the raw alcohol produced in C210 and C230 is sent to the scrubber C250, operating at atmospheric pressure, where the liquid from the bottom of the rectifying column C240 is sprayed countercurrent to the alcohol vapors. Some low boiling products, discharged from the top of
A. Ficarella, D. Laforgia / Energy Conversion & Management 40 (1999) 1495±1514
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Fig. 3. Schematic of raw alcohol production cycle (2nd technology).
C250, are re®ned in column C260, while the liquid discharged from the bottom is sent to the rectifying column C240. In this column, operating at higher pressure (0.2 MPa abs), a high grade alcohol (96 vol%) is obtained, feeding the heat exchanger at the bottom of C260 with saturated steam. In the last column, C270, a fractional distillation takes place to separate the methyl alcohol at the top of the column from the re®ned ethyl alcohol at the bottom (96 vol% with water). In this process, S1-1 and C4-1 are the same heat exchanger, as well as S2 and C10-1 and C6-1 and S11. In the third process (Fig. 3), raw alcohol is produced with the same process described in Fig. 1, but in this case, a higher operating pressure is used (0.2 MPa abs), while 0.7 MPa abs saturated steam directly feeds the bottom of the distillation column C10 (no heat exchanger is used). The liquid discharged from the bottom (consisting of the water in the wine or in the dregs plus the steam condensate) is ¯ashed at a pressure of 0.04 MPa abs. The produced steam is compressed, using an ejector fed by saturated steam at 0.7 MPa abs and returned to the bottom of C10. The concentration column C30 works in the same way as C230. Some high boiling temperature products (oils) are discharged from the bottom. Heat exchangers S1-1 and C5 are the same unit. Tables 1±3 list the details of the heat and mass balances in the existing processes. Selected instantaneous operational data, complemented with operational and design values, are used to
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Table 1 Raw alcohol production Heat exchanger (a) First process S1-1 S1-2 S2 S2A C3 S3 C4-1 C4-2 S4 Heating (in) Cooling (out)
Mass
Flow (kg/s)
Wine Wine Water Water Low-boiling Low-boiling Raw alcohol Raw alcohol Raw alcohol
(b) First process optimized S1-1 Wine S1-2 Wine S2 Water S2A1 Water S2A2 Water C3 Low-boiling S3 Low-boiling C4-1 Raw alcohol C4-2 Raw alcohol S4 Raw alcohol Heating (in) Cooling (out)
Tin (8C)
Tout (8C)
Heat out/in (kW)
Recovery
2.315 2.315 0.353 2.083 0.062 0.021 0.633 0.633 0.211
15 60 87 87 79 79 65 65 65
60 79 87 15 79 15 65 65 15
494 220 729 ÿ631 ÿ132 ÿ6 ÿ494 ÿ160 ÿ35 1443 ÿ1458
C4-1
2.315 2.315 0.353 2.083 2.083 0.062 0.021 0.633 0.633 0.211
15 64 87 87 25 79 79 65 65 65
64 79 87 25 15 79 15 65 65 15
544 170 729 ÿ544 ÿ87 ÿ132 ÿ6 ÿ654 0 ÿ35 1443 ÿ1458
S1-1 949.15 ÿ963.71
S2A1 S1-1
899.28 ÿ913.84
calculate the mass and energy balances. The formulas for the energy of the multicomponent nonideal mixtures used by Castier and Rajagopal [2] have been employed for density data, speci®c heat and formation data.
3. Pinch technology analysis The ®rst step of this analysis is to present the complete process on a temperature±enthalpy diagram by a grand composite curve, which represents the cumulative heat sources and heat sinks in the process. A full description of pinch technology for the solution sequence in question is provided by Linnho et al. [9], and Linnho and Alanis [10]. Figs. 4±6 represent the composite curves of the three processes analysed: raw alcohol production ethyl alcohol production and raw alcohol production using steam obtained by ¯ashing the condensate. The curves are named as follows:
A. Ficarella, D. Laforgia / Energy Conversion & Management 40 (1999) 1495±1514
GCINHE
GCINCO GCACHE GCACCO GRCOOT
1501
which describes the temperature±enthalpy diagram for the heat sinks in the process, corresponding to the mass ¯ows to be heatedÐnot considering heat recovery from some other ¯ow. The heat need of the process in the temperature range T1 ±T2 is H2 ±H1 ; which describes the temperature±enthalpy diagram for the heat source in the process, corresponding to the mass ¯ows that need to be cooledÐnot considering heat recovery to some other ¯ow; same as above for heat sinks but considering heat recovery in the actual processes; same as above for heat sources but considering heat recovery in the actual processes; optimal grand composite curve; represents the solution obtained at the so-called targeting stage, calculated considering an unfeasible DT 0 (no heat across the pinch) in the heat exchangers.
The pinch divides the process into two thermodynamically separate systems, as depicted in Table 2 Raw alcohol production Heat exchanger (a) First process S1-1 S1-2 S2 S2A C3 S3 C4-1 C4-2 S4 S5 C6-1 C6-2 S7A C8 S8 S9 C10-1 C10-2 S11 S11A C12 S12 Heating (in) Cooling (out)
Mass
Flow (kg/s)
Wine Wine Water Water Low-boiling Low-boiling Raw alcohol Raw alcohol Raw alcohol Ethyl alcohol Low-boiling Low-boiling Water Low-boiling Low-boiling Ethyl alcohol Ethyl alcohol Ethyl alcohol Ethyl alcohol Ethyl alcohol Methyl alcohol Methyl alcohol
1.042 1.042 0.168 0.935 0.028 0.009 0.293 0.293 0.098 0.069 0.067 0.067 0.002 0.012 0.002 0.652 0.652 0.652 0.034 0.091 0.012 0.002
Tin (8C)
Tout (8C)
Heat out/in (kW)
Recovery
15 60 87 87 79 79 65 65 65 79 79 79 101 79 79 121 98 98 65 65 52 52
60 79 87 15 79 15 65 65 65 79 79 79 15 79 15 121 98 98 65 15 52 15
222 99 347 ÿ283 ÿ59 ÿ3 ÿ222 ÿ92 0 66 ÿ32 ÿ40 ÿ1 ÿ13 ÿ1 625 ÿ347 ÿ278 32 ÿ14 ÿ14 0 1392 ÿ1399
C4-1 C10-1
S1-1
S11
S2 lC6-1
790.32 ÿ797.18
(continued on next page)
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Table 2 (continued ) Heat exchanger
Mass
(b) First process optimized S1-1 Wine S1-2 Wine S2 Water S2A1 Water S2A2 Water C3 Low-boiling S3 Low-boiling C4-1 Raw alcohol C4-2 Raw alcohol S4 Raw alcohol S5 Ethyl alcohol C6-1 Low-boiling C6-2 Low-boiling S7A Water C8 Low-boiling S8 Low-boiling S9 Ethyl alcohol C10-1 Ethyl alcohol C10-2 Ethyl alcohol S11 Ethyl alcohol S11A Ethyl alcohol C12 Methyl alcohol S12 Methyl alcohol Heating (in) Cooling (out)
Flow (kg/s) 1.042 1.042 0.168 0.935 0.935 0.028 0.009 0.293 0.293 0.098 0.069 0.067 0.067 0.002 0.012 0.002 0.652 0.652 0.652 0.034 0.091 0.012 0.002
Tin (8C)
Tout (8C)
Heat out/in (kW)
Recovery
15 64 87 87 25 79 79 65 65 65 79 79 79 101 79 79 121 98 98 65 65 52 52
64 79 87 25 15 79 15 65 65 65 79 79 79 15 79 15 121 98 98 65 15 52 15
244 77 347 ÿ244 ÿ39 ÿ59 ÿ3 ÿ314 0 0 66 ÿ32 ÿ40 ÿ1 ÿ13 ÿ1 625 ÿ347 ÿ278 32 ÿ14 ÿ14 0 1392 ÿ1399
S2A1 C10-1 S1-1
S1-1
S11
S2 C6-1
768.61 ÿ775.47
Table 3 Raw alcohol production Heat exchanger (a) Second process S21-1 S21-2 S22A S22B S22A-1 S22A-2 C23 S23 S24A C25 S25 Heating (in) Cooling (out)
Mass
Flow (kg/s)
Wine dregs Wine dregs Water 0.6 MPa steam Condensate Water Low-boiling Low-boiling High boiling Raw alcohol Raw alcohol
3.345 3.345 0.056 0.056 0.056 3.070 0.110 0.033 0.033 0.686 0.208
Tin (8C)
Tout (8C)
Heat out/in (kW)
15 61 15 165 76 76 100 100 121 98 98
61 118 165 165 15 15 100 15 15 98 15
709 969 35 115 ÿ14 ÿ788 ÿ241 ÿ14 ÿ17 ÿ709 ÿ61 1828 ÿ1843
Recovery C5
S1-1 1119.55 ÿ1134.23
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Table 3 (continued ) Heat exchanger
(b) Second process S21-1A S21-1B S21-2 S22A S22B S22A-1A S22A-1B S22A-2A S22A-2B C23 S23 S24A C25A C25B S25 Heating (in) Cooling (out)
Mass
optimized Wine dregs Wine dregs Wine dregs Water 0.6 MPa steam Condensate Condensate Water Water Low-boiling Low-boiling High boiling Raw alcohol Raw alcohol Raw alcohol
Flow (kg/s)
Tin (8C)
Tout (8C)
Heat out/in (kW)
15 58 93 15 165 76 25 76 25 100 100 121 98 98 98
58 93 118 165 165 25 15 25 15 100 15 15 98 98 15
671 574 433 35 115 ÿ12 ÿ2 ÿ659 ÿ129 ÿ241 ÿ14 ÿ17 ÿ574 ÿ135 ÿ61 1828 ÿ1843
3.345 3.345 3.345 0.056 0.056 0.056 0.056 3.070 3.070 0.110 0.033 0.033 0.686 0.686 0.208
Recovery
S2A-1A-2A C25A
S1-1 S1-1
S21-1B 583.53 ÿ598.21
Figs. 4±6. The heat sinks curves (heated ¯ows) are presented in the positive heat region, while the heat sources curves (cooled ¯ows) are located in the negative heat region. This calculated target GRCOOT represents the minimum heat requirement for the process, an ideal solution that, however, violates the second law of thermodynamics, but it is very useful for a ®rst evaluation of possible improvements in the process under investigation. 4. Process pre-optimization using the pinch method The raw alcohol production process (Fig. 1 and Table 1) requires 949 kW of heat and 963 kW of cooling, see also curves GCACHE and GCACCO of Fig. 4a. The total heat requirement of the process is 1443 kW, but 494 kW are recovered from some cooled ¯ows (heat is recovered at 658C during condensation in C4-1 and used to heat the wine from 158C in S1-1)Ðthe total cooling power available is 1458 kW. The optimal target, considering only the heat exchangers network optimization, is shown by the GRCOOT curve of Fig. 4a, and it is achieved. The (unfeasible) target for heating is 728 kW, with a dierence of 30% with respect to the actual heating power (949 kW). The dierence is mainly due to the calculation of the optimal target considering DT 0 in the heat exchangers. The excess of heat to cool the discharged water in S2A and to complete the condensation in C4-2 is available at temperatures lower than 878C, so that the boiler S2 must be fed by steam. Note that in the optimized process, the wine in S1-1 and -2 is heated by water in S2A, while all the heat of condensation in C4-1 and -2 is available, but useless. In the actual process, matching S1-1 with C4-1, a big
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Fig. 4 (a) and (b). Grand composite curves of raw alcohol production.
dierence in temperature exists. A smaller heat exchanger could be employed, but considering the use of a heat pump, as subsequently discussed, it is better to match S1-1 and -2 with S2A to make available the heat in C4-1 at a higher temperature. In continuation, this last con®guration will be considered (see Table 1b and Fig. 4b). The ethyl alcohol production process of Fig. 2 (see also Table 2a) requires 790 kW of heating power. The total heating need is 1392 kW, but 602 kW are recovered by matching, in the same unit, S1-1 and C4-1, S2 and C10-1 and C6-1 and S11. The heat exchangers network
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Fig. 5 (a) and (b). Grand composite curves of ethyl alcohol production.
is practically in the optimal con®guration. The 625 kW at 1218C necessary in the boiler S9 cannot be recovered from the cooling streams at lower temperatures. Only the ¯ow of 277 kW at 988C in C10-2 could be recovered to heat some ¯ows, as in S1-2 (99 kW) and S5 (66 kW), but some plant complexities and low power of exchangers made the optimization economically unfeasible. The ideal optimal grand composite curve (GRCOOT of Fig. 5a con®rms the previous conclusions, showing an optimal target for heating of 625 kW (the actual value is 26% higher than the ideal one). As previously noted, considering the use of a heat pump, it is better to match S1-1 and -2 with S2A to make available the heat in C4-1 at a higher temperature. In continuation, this last con®guration will be considered (see Table 2b and Fig. 5b).
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Fig. 6 (a) and (b). Grand composite curves of raw alcohol production (2nd technology).
The last production process for raw alcohol, based on ¯ashing of the discharge condensate (Fig. 3), shows a heat consumption of 1119 kW (Table 3): Considering the total amount of heating power, 709 kW are recovered during the condensation in C25 (C25 and S21-1 are the same unit), but as shown by the ideal grand composite curve of Fig. 6a, it is possible to recover (ideally) a further 679 kW, leaving the minimum heating required as 440 kW by matching S22A-1/2 with S21. Considering a reasonable DT for the heat exchanger, it is possible to recover 670 kW from S22A-1/2 and 573 kW from C5 (see Table 3b and Fig. 6b). In this case, the heat requirement is 583 kW, only 32% higher than the target one.
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5. Thermodynamic analysis in presence of heat pumps Fig. 7 presents the representative design point performance data of several commercially available heat pumps for industrial use (heating power > 100 kW) and considering some constraints deriving from their application in the production processes previously seen. Data were supplied by known manufacturers. The points are named as follows: ELECOM GASCOM
STECOM GASABS STEABS GCONCO STEEJE
mechanical type heat pump with the compressor driven by an electrical motor; combined mechanical type heat pump and cogeneration plant, with the compressor driven by an internal combustion engine fueled by natural gas and with recovery of heat from o engine exhaust gasesÐthe heat recovery from engine cooling was neglected; combined mechanical type heat pump and cogeneration plant, with the compressor driven by a steam turbine (inlet steam at 4.5 MPa/3508C, discharge at 0.9 MPa) and with recovery of heat from o turbine discharged steam; absorbing type heat pump fueled by natural gas; absorbing type heat pump fed by steam; mechanical type heat pump with the compressor driven by an internal combustion engine fueled by natural gas, no heat recovery considered; ejector type heat pump fed by saturated steam at 0.8 or 4.0 MPa with ¯ashing water at 908C and heat recovery at 1208C.
For the analyzed heat pumps, the fuel consumption was considered in the calculations of the COP, boiler eciency was considered in the steam production, as well as the average eciency of the national grid production for the equivalence fuel±electricity (2.67 kJfuel/kJelectricity). The traditional COP is evaluated as:
Fig. 7. COP referred to fuel consumption.
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COPW E=W
1
where E is the heat recovered and W, the work employed. The COP referred to the fuel will be written as COPf E=Mf
2
where Mf is the fuel used in the prime mover (mechanical or thermal one). Moreover, the COP referred to Carnot's cycle is written as COPC Tc = Tc ÿ Te Tc =DT
3
where Tc is the temperature at which heat is available (temperature of the external heated ¯uid in the heat pump condenser) and Te is the temperature of the cooled external ¯uid in the heat pump evaporator. On an average, it is useless to have temperature levels in the condenser and in the evaporator such that COPC is less than 6 (for example, Tc 1008C and DT 608C), because at this level, COPf is equal to 1 (useless heat pump). This can be explained by the fact that in commercially available heat pumps, the ratio COPW to COPf is, on average, equal to 0.5, and then, when COPC is equal to 6, COPW will be equal to 3. Considering that the normal eciency of traditional systems in the production of work is of the order of 25±35%, without cogeneration, as much heat is recovered as fuel is burned. Fig. 7 shows that the combined cogeneration power plant and heat pumps (thermodynamic heating method analyzed by Stachel et al. [11] ÐGASCOM in the ®gureÐshows better performances for low COPC , while for COPC higher than 8, the electrical compressors (ELECOM) show better performances. On an average, the national grid electrical eciency is higher than the engine eciency in the commercially available GASCOM. Gas compressors performances worsen if the exhaust gas heat is not recoveredÐGCONCO points in Fig. 7. The absorption heat pumpsÐGASABS and STEABSÐshow good performances. Unfortunately, the typical COPC is highÐthe maximum working DT is of the order of 408C and the maximum absolute condensing temperature is approximately 1008C. The steam compressors (STECOM) show poor performances because of the very low eciency in the work conversion in industrial counterpressure steam turbines. Moreover, they require superheated steam, usually not used in the agrofood industries under investigation. The same behavior is shown by the ejector type compressors because of the low mechanical eciency of ejectors and because the steam condensate is mixed with the sewage of the process (see Fig. 3), so that its enthalpy content is lost. The costs of the recovered heat are presented in Fig. 8. They include only the cost of the energy used to move/operate the heat pump and not the cost of the recovered heat. The calculations were made with the following common characteristics: . cost of electricity: 22.92 ECU/GJ (0.0825 ECU/kWh) . cost of fuel for thermal use: 3.79 ECU/GJ . cost of fuel for mechanical energy production (considering some exemptions from taxes): 3.51 ECU/GJ. The heat reference cost was assessed equal to 4.74 ECU/GJ considering a traditional boiler
A. Ficarella, D. Laforgia / Energy Conversion & Management 40 (1999) 1495±1514
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Fig. 8. Heat costs.
with 80% eciency. Because of the high costs of electricity, the electrical compressors (ELECOM) do not compete with the gas compressors (GASCOM) even without considering cogeneration (GCONCO points) which show a heat cost lower than the reference one. Also, the absorption systems (GASABS and STEABS) and the steam turbine compressors (STECOM) show satisfactory economical performances. 6. Heat pumps performances The analysis of the alcohol production processes, performed using pinch technology, shows the possibility to apply heat pump technology to reduce the energy consumption. The three production processes previously examined show pinch temperatures of 87, 98±121 and 1008C, respectively (Figs. 4±6). Energy is saved overall in the process only if the heat pump operates across the pinch, i.e., when it pumps from the process source to the process sink, as stated by Linnho et al. [9]. It should be pointed out that no utilities optimization was consideredÐthat means, steam pressure optimization. In fact, the larger amount of heat in each process is needed at one temperature level only, whereas negligible heat is required at lower temperatures. Moreover, no heat pumps based on alcohol vapor compression were considered because the complexity and risks of such technology (alcohol is easily in¯ammable, hot alcohol vapors are corrosive) are not suitable for low and medium size agrofood industries. All processes were pre-optimized for the installation of heat pumps, making available heat at higher temperatures and matching the process heat exchangers with the lower technically feasible DT, making best use of the heat recovery network driving forces. Heat pumps with gas engine driven compressor have been considered, with heat recovery from the gas engine (used to produce steam directly) and, if possible, from engine cooling (at
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lower temperature, about 808C). As previously seen, the electrical compressors have high operating costs, and the absorbing systems or steam compressors show low performances or useless operating temperature ranges. Moreover, the steam compressors require superheated steam which is normally not available. 7. Traditional heat pumps A traditional closed cycle heat pump was considered. Calculations were made with the following common characteristics: . heat losses: 10% . ratio of heat recovery from engine discharged gases to heat pump heat: 0.25 . ratio of COPW to COPC : 0.50. The heat losses include losses in the heat pump, in the piping system and in the heat exchangersÐconsidering that the heat sinks and sources may be distant in the production plant lay-out. The recovery from engine o gases was considered constant because the heat is used to produce saturated steam at 0.7 MPa. In any case, for environmental reasons, the gas temperature at the stacks cannot be lower than 1508C. Moreover, smaller heat exchangers (less than 100 kW) match will not be considered for economic unfeasibility, and for the same reason, only one temperature level for condensation and evaporation has been considered. Table 4 Traditional heat pump Process
Table 1
Table 1
Table 2
Table 2
Table 3
Evaporators Condensers Tc (8C) Evaporator 1 E1 (kW) Te1 (8C) Evaporator 2 E2 (kW) Te2 (8C) Gas recovery E (kW) Te3 (8C) Performances COPC COPm COPf Fuel consumption (kW) Heat recovered (kW)
C3, C4-1/2 S2 87
C4-1/2 S2 87
C10-2, C4-1/2 S9 121
C4-1/2 S9 121
C23, C25B S1-2 118
131.69 79
131.69 79
277.59 98
277.59 98
240.63 100
653.61 65
0.00 65
313.90 65
0.00 65
134.75 98
785.30 500
131.69 500
591.49 500
277.59 500
375.38 500
16.36 7.36 2.56 399.54 1022.29
45.00 20.25 7.04 22.15 155.85
7.04 3.17 1.10 884.11 972.63
17.13 7.71 2.68 133.97 358.84
19.55 8.80 3.06 155.86 476.46
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Analyzing the ®rst production process (Table 1b and Fig. 4b), it is clear that the heat pump condenser will match the heat exchanger S2 at 878C, while the evaporators will match C4-1/2 at 658C and C3 at 798C. Table 4 shows very interesting performances of the heat pump, with a recovery of 1022 kW of heat at 878C (or more, including the hot gases heat recoveryÐnote that some heat will be used for some other process) using 399 kW of fuel. In the same way, the analysis of the second production process (Table 2b and Fig. 5b) shows that the heat pump condenser matches S9 at 1218C, while the evaporators match C10-2 at 988C and C41/2 at 658C. Of course, the increase of the condensing temperature reduces the system performances, reducing COPW to 3.17 and COPf to 1.10, making the heat pump practically useless. Finally, the third production process (Table 3b and Fig. 6b) behavior suggests to match the heat pump condenser to S21-2 at 1188C (S22 is the boiler and anyway, it is neglected for its high temperature) and the evaporators to C23 and C25B at temperatures, respectively, of 100 and 988C. 470 kW of heat are recovered (some heat will be used elsewhere) using 155 kW of fuel. The traditional closed cycle heat pumps show a major drawback for the condenser high temperature. For a condenser at 908C (actual refrigerant temperature of 1008C), the HCFC-22 (as its ecological substitutes R047C or R134a) and the HCFC-717 (NH3) show a condensation pressure higher than 5 MPa and about 1.4 MPa for the HCFC-21, making the system technically and economically unfeasible (especially for the compressor discharge high temperatures and the lubrication problems). 8. Heat pumps using water ¯ash Water presents a more favorable link between temperature and pressure. A heat pump cycle (Fig. 9) can be realized using water ¯ash in a boiler at low temperature and pressure to
Fig. 9. Schematic of the heat pump with sewage ¯ash.
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Table 5 Heat pump with ¯ash Table 1
Table 2
Table 2
Table 2
Table 3
Table 3
Table 2 and 3
Evaporators
C3, C4-1/2
C10-2, C4-1/2
C10-2, C4-1/2
C10-2, C4-1/2
C23, C25B
C23, C25B
S2 87
S9 121
S9 121
S9 121
S1-2 118
S1-2 118
C10-2, C4-1/2 C23, C25B S9, S1-2 121
55 0.0160
55 0.0160
70 0.0288
88 0.0583
70 0.0288
90 0.0630
88 0.0583
131.69 79
277.59 98
277.59 98
277.59 98
240.63 100
240.63 100
240.63 100
653.61 65
313.90 65
0.00 65
0.00 65
134.75 98
0.00 98
412.34 98
192.26 80
363.37 80
110.72 80
0.00 80
148.46 80
0.00 80
0.00 80
192.26 192
363.37 363
110.72 111
48.17 48
148.46 148
38.59 39
113.31 113
97 0.0829 256
131 0.3139 484
131 0.3139 148
131 0.3139 64
128 0.2791 198
128 0.2791 51
131 0.3139 151
11.25 4.92 1.71 737.74 1260.48
5.97 3.26 1.13 1394.31 1578.84
7.73 3.85 1.34 424.85 568.70
11.94 5.37 1.87 184.84 345.21
8.15 3.87 1.34 569.67 765.40
13.96 5.69 1.98 148.07 292.97
11.94 5.37 1.87 434.80 812.03
107,636 57,243 50,393 174,757 3.47
134,821 108,187 26,634 271,845 10.21
48,563 32,965 15,598 116,505 7.47
29,478 14,342 15,136 67,961 4.49
65,360 44,201 21,159 145,631 6.88
25,018 11,489 13,529 58,252 4.31
69,342 33,737 35,605 121,359 3.41
Condensers Tc 8C Boiler parameters Flashing temperature (8C) Flashing pressure (MPa) Evaporator 1 E1 (kW) Te1 (8C) Evaporator 2 E2 (kW) Te2 (8C) Engine cooling recovery Ecooling (kW) Tcooling (8C) Gas recovery Egases (kW) Tgases (8C) Compressor parameters Condensation temperature (8C) Condensation pressure (MPa) Mechanical energy (kW) Thermal performances COPC COPm COPf Fuel consumption (kW) Heat recovered (kW) Economical performances Energy savings (ECU/year) Operational expenditures (ECU/year) Total savings (ECU/year) Investment cost (ECU) Pay-pack period (year)
A. Ficarella, D. Laforgia / Energy Conversion & Management 40 (1999) 1495±1514
Process
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produce enough driving forces to recover heat from the process ¯ows to be cooled. The produced low pressure steam will be compressed to a temperature level suitable to feed the energy sinks in the process. The gas engine driven compressor will be of centrifugal type, as used in some other similar applications (see EC [4±7]). Calculations were made with the following common characteristics: . . . . . . . . .
heat losses: 10% compressor thermodynamic eciency: 70% compressor mechanical eciency: 90% minimum driving force (DT): 108C natural gas consumption in the engine: 2.88 kJ/kJ of work ratio of heat recovery from engine discharge gases to engine power: 0.75 ratio of heat recovery from engine cooling to engine power: 0.75 plant usage: 5000 h/year maintenance and other operational expenditures (fuel excluded): 2.7 ECU/GJ power
Refer to the previous case for engine performance. Summarizing, the ®rst process (see Table 5) presents good thermal and economical performances: 1260 kW of heat recovered, using 737 kW of fuel power, with an investment of 175,000 ECU and savings, of 50,000 ECU/year. The pay back period (PBP) will be 3.5 years. For the second process, dierent ¯ashing temperatures can be chosen. To recover all the available heat (from engine cooling at 808C, from the two process evaporators at 98 and 658C and from the hot gases at 5008C), a temperature of 558C has to be achieved in the main boiler. The high condenser temperature (1218C) implies poor performances (more than 10 years of PBP). Also when the ¯ash is increased to 708C (C4-1/2 not matched to the boiler), the plant presents poor performances (PBP equal to 7.5 years). Only the ¯ash at 888C, with recovering heat from C10-2 and from the hot gases, implies a recovery of 345 kW of heat requiring 184 kW of fuel energy. The PBP is 4.5 years with an investment cost of 68,000 ECU. Analyzing the third production process, a ¯ashing temperature of 708C, allowing for heat recovery from engine cooling, involves poor performances (PBP is 6.9 years), but with the increase of such temperature to 908C, good thermal and economical performances can be obtained. 291 kW can be recovered using 148 kW of fuel with a PBP of 4.3 years and 58,000 ECU of investment.
9. Concluding remarks Pinch technology analysis was applied to processes to produce alcohol from low quality wine and wine dregs. The results clearly identify the heat sinks and sources of the process and the possible installation of heat pumps. Moreover, it was clear that optimization of the heat exchangers network can imply the maximum heat recovery and the improvement of heat pump application can optimize the driving forces in the network. Heat pumps based on a water cycle were analyzed, showing good thermal and interesting economical performances. A heat pump plant, working with ¯ashing temperature of 88±908C
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and 118±1218C for heat recovery, can be used at the same time for ethyl alcohol production and raw alcohol production (the last one using the third production process)while achieving interesting overall performances: 812 kW of heat recovered using 434 kW of fuel power and with an investment of 121,000 ECU and savings of 36,000 ECU/year, the pay back period (PBP) will be 3.4 years (see Table 5). References [1] Bidini G, Stecco SS. A computer code using exergy for optimizing thermal plants. Journal of Engineering for Gas Turbine and Power 1991;113:145±50. [2] Castier M, Rajagopal K. Thermodynamic analysis of an alcohol distillery. Energy 1988;13:455±9. [3] EC. Pompe aÁ chaleur sur la vapeur produite par un proceÂde thermomeÂcanique. Demonstration Projects for Energy Saving and Alternative Energy Sources. N. EE/193/79-F, 1979. [4] EC. Mechanical compression of organic vapors. Demonstration Projects for Energy Saving and Alternative Energy Sources. N. EE/012/82-F. 1982. [5] EC. Superheated steam beet pulp drier with mechanical vapor recompression. Demonstration Projects for Energy Saving and Alternative Energy Sources. N. EE/244/84-F, 1984. [6] EC. Concentration of (technical) gelatine using a falling ®lm evaporator with mechanical vapor compression. Demonstration Projects for Energy Saving and Alternative Energy Sources. N. EE/349/85-NL, 1985. [7] EC. Concentration of (technical) gelatine using a falling ®lm evaporator with mechanical vapor compression. Demonstration Projects for Energy Saving and Alternative Energy Sources. N. EUR 12108 EN, 1989. [8] Gaggioli RA, Sama DA, Quian S, El-Sayed YM. Integration of a new process into an existing site: a case study in the application of exergy analysis. Journal of Engineering for Gas Turbine and Power 1991;113:170± 83. [9] Linnho B, Townsend DW, Boland D, Hewitt GF, Thomas BEA, Guy AR, Marsland RH. Process Integration for the Ecient Use of Energy. Rugby, Warks, England: The Institution of Chemical Engineers, 1991. [10] Linnho B, Alanis FJ. Integration of a new process into an existing site: a case study in the application of pinch technology. Journal of Engineering for Gas Turbine and Power 1991;113:159±69. [11] Stachel K, Frutschi HU, Haselbacher H. Thermodynamic heating with various types of cogeneration plants and heat pumps. Journal of Engineering for Gas Turbine and Power 1995;117:251±8.