Novel Heat-Pump-Assisted Extractive Distillation for Bioethanol

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Novel heat pump assisted extractive distillation for bioethanol purification Hao Luo, Costin Sorin Bildea, and Anton Alexandru Kiss Ind. Eng. Chem. Res., Just Accepted Manuscript • DOI: 10.1021/ie504459c • Publication Date (Web): 06 Feb 2015 Downloaded from http://pubs.acs.org on February 10, 2015

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Industrial & Engineering Chemistry Research

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Novel heat pump assisted extractive distillation for bioethanol purification

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Hao Luo,a,b Costin Sorin Bildea,c Anton A. Kiss d,e* a

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State Key Laboratory of Multi-Phase Complex Systems, Institute of Process Engineering, Chinese Academy of Sciences, Beijing 100190, China b Sino-Danish Center for Education and Research, Chinese Academy of Sciences, Beijing 100190, China c University “Politehnica” of Bucharest, Polizu 1-7, 011061 Bucharest, Romania d AkzoNobel Research, Development & Innovation, Process Technology SRG, Zutphenseweg 10, 7418 AJ Deventer, The Netherlands. *E-mail: [email protected], [email protected] e Sustainable Process Technology Group, Faculty of Science and Technology, University of Twente, PO Box 217, 7500 AE, Enschede, The Netherlands

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Abstract

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The purification of bioethanol fuel involves an energy intensive separation process to concentrate the diluted streams obtained in the fermentation stage and to overcome the azeotropic behavior of ethanol-water mixture. The conventional separation sequence employs three distillation columns that carry out several tasks, penalized by high energy requirements: pre-concentration of ethanol, extractive distillation and solvent recovery.

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To solve this problem, we propose here a novel heat pump assisted extractive distillation process taking place in a dividing-wall column (DWC). In this configuration, the ethanol top vapor stream of the extractive DWC is recompressed from atmospheric pressure to over 3.1 bar (thus to a higher temperature) and used to drive the side reboiler of the DWC, which is responsible for the water vaporization. For a fair comparison with the previously reported studies, we consider here a mixture of 10%wt ethanol (100 ktpy plant capacity) that is concentrated and dehydrated using ethylene glycol as mass separating agent. Rigorous process simulations of the proposed vapor recompression (VRC) heat pump assisted extractive DWC were carried out in AspenTech Aspen Plus.

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The results show that the specific energy requirements drop from 2.07 kW.h/kg (classic sequence) to only 1.24 kW.h/kg ethanol (VRC assisted extractive DWC), thus energy savings of over 40% being possible. Considering the requirements for a compressor and use of electricity in case of the heat pump assisted alternative, about 24% decrease of the total annual costs is possible – in spite of the 29% increase of the capital expenditures – for the novel process, as compared to the optimized conventional separation process.

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Keywords: Bioethanol, extractive distillation, dividing-wall column, heat pump, energy savings

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1. Introduction

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Bioethanol is arguable the most promising renewable fuel, with an important advantage over other alternatives, namely that it can be directly integrated in existing fuel systems – typically as a 5-85% mixture with gasoline which does not require modifications of the current engines.1 The industrial production of bioethanol relies on various routes: corn-to-ethanol, sugarcane-to-ethanol, basic and integrated lignocellulosic biomass-to-ethanol.1 In all these cases, the raw materials go through pre-treatment steps before entering the fermentation stage, where bioethanol is actually produced. A common feature of these technologies is the production of diluted bioethanol (typically 5-12 %wt ethanol) that is further concentrated to reach the requirements of the international bioethanol standards.2-4 Depending on the standard requirements, the maximum water content in ethanol is 0.2 %vol (EN 15376, Europe), 0.4 %vol (ANP no. 36/2005, Brazil) or 1.0 %vol (ASTM D 4806, USA).4

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To reach the purity targets, an energy demanding separation is needed in practice, in order to break the binary azeotrope ethanol-water (95.63 %wt ethanol). The separation is usually carried out by distillation, the first step being a pre-concentration distillation column (PDC) which increases the ethanol content from about 5-12% up to 91-94 %wt.2,3,5 The second step consists of the bioethanol dehydration, up to concentration levels exceeding the azeotropic composition. Quite a number of separation alternatives are available as described in the literature: extractive distillation (ED), azeotropic distillation (AD), pressure-swing distillation, pervaporation, adsorption, and hybrid methods that combine some of these options.2,3 Among them, extractive distillation (ED) is the main option of choice in case of the large scale production of bioethanol fuel.3,4 Typically, ED is carried out in a sequence of two columns, one being the extractive distillation column (EDC) which separates ethanol, while the other one is the solvent recovery column (SRC) that recovers the mass separating agent (MSA) that is recycled back in the process. Further improvements to the distillation process were proposed, with the aim to increase the energy efficiency of bioethanol purification: e.g. ED process optimization;5 thermally coupled distillation columns;6 azeotropic and extractive dividing-wall columns.4,7

ACS Paragon Plus Environment

Industrial & Engineering Chemistry Research Novel heat pump assisted extractive distillation for bioethanol purification

Luo, Bildea, Kiss

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The present work proposes a novel heat pump assisted extractive distillation process that efficiently combines several technologies: extractive distillation, dividing-wall column (DWC) technology and vapor recompression (VRC) heat pump, thus allowing a significant reduction of the energy requirements for bioethanol purification. A 100 ktpy plant is considered, processing a feed with 10%wt ethanol, by pre-concentration and extractive distillation using ethylene glycol as heavy solvent. Rigorous Aspen Plus simulations were carried out and for a fair process comparison the same approach was used as in previous studies.8,9

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2. Problem statement

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For its use as biofuel or additive, bioethanol should have a purity of min. 99.0-99.8 %wt, but this can be reached only by paying a large energy penalty. Although most of the water present in the diluted bioethanol (5-12 %wt) from the fermentation step can removed by normal distillation, the ethanol purity is limited to max. 95.6 %wt which is the composition of the binary azeotrope ethanol-water. Using process optimization and distillation technologies such as azeotropic and extractive dividing-wall column, the energy requirements were reduced by 20% from about 2.5 down to almost 2 kWh/kg bioethanol.9 However, to put this into perspective, the energy needed for purification is still considerably high, especially when taking into account the energy content of the ethanol fuel (7.45 kWh/kg).

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To solve this issue we integrate VRC technology with an extractive dividing-wall column (E-DWC) capable of purifying bioethanol in a single distillation unit. As shown later, this integration unit allows significant energy savings, less CO2 emissions and a reduced footprint, in comparison to the conventional separation process.

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3. Results and discussion

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For all process alternatives described hereafter, Aspen Plus simulations were performed using the rigorous RADFRAC (Reactive Absorption, Distillation & FRACtionation) unit. NRTL (non-random two-liquid) property model was used as suitable model due to the presence of a non-ideal mixture containing polar components.

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3.1 Heat pump assisted distillation

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Notably, a distillation column can be also considered as a heat engine that produces separation (or entropy reduction) instead of work. In this case, the heat is provided in the reboiler (heat source) and removed at a lower temperature level in the condenser (heat sink). Heat pumps can be used in distillation to increase the energy efficiency of columns, by upgrading the low temperature energy from the top (condenser) to higher temperature levels, such that the recovered heat can be reused to heat up a lower column stage (e.g. reboiler). This can bring significant energy (utilities) savings, but the energy required to increase the pressure (and temperature level) is of higher quality/price than hot utilities. Heat pump assisted distillation can be applied to classic columns10 but also to dividing-wall columns, to a lesser extent.11 In case of DWC systems, the temperature span across the column is larger, thus the temperature lift required by the heat pump is also larger and hence less efficient.8

33 Q / W ratio 45.00 40.00-45.00

40.00

35.00-40.00

35.00

30.00-35.00 25.00-30.00

30.00

20.00-25.00 15.00-20.00

25.00

10.00-15.00

20.00

5.00-10.00

15.00

0.00-5.00

10.00 5.00

403

0.00

dT = Tr − Tc

353 180 190 200

10 20 30 40 50 60 70 80 90 100 110 120 130 140 150 160 170

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Tc

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Figure 1. Dependence of the Q/W ratio on the condenser temperature (Tc) and the temperature difference between reboiler and condenser (Tr – Tc).

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While Kiss et al.10 provided a useful scheme for the quick selection of appropriate heat pumps, Plesu et al.12 introduced a simple criterion depending on the Carnot efficiency to decide whether a heat pump is worth considering. This can be simplified to the following form:

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Q / W = 1/η = Tc / (Tr - Tc) > 10

2 Environment ACS Paragon Plus

(1)

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where Q is the reboiler duty, W the work provided, while η is the Carnot efficiency, Tr the reboiler temperature and Tc the condenser temperature. When the Q/W ratio exceeds 10, then a heat pump should be considered, while when the ratio is lower than 5 then using a heat pump will not bring any benefits (Plesu et al., 2014).

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Figure 1 illustrates the dependence of the Q/W ratio on the condenser temperature and the temperature span across the column (Tr-Tc). In case of ethanol separation, Q/W=16 (Tr=373 K, Tc=351 K) which is favorable.

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3.2 Classic extractive distillation process Figure 2 presents the process flowsheet, including the mass and energy balance along with the key parameters of this classic bioethanol purification sequence based on extractive distillation – the optimization of this process being described in details in our earlier work.5 The classic process employs three distillation columns: preconcentration distillation column (PDC), extractive distillation column (EDC) and solvent recovery column (SRC). The first column (PDC) separates water as bottom product and a near-azeotropic mixture as top distillate (91%wt ethanol being the optimal value as reported by Kiss and Ignat).5 The second column (EDC) makes use of ethylene glycol that is added at a solvent-to-ethanol ratio of 1.25 mol/mol, at a location above the feed stage of the ethanol-water mixture. The presence of the solvent alters the relative volatility of ethanol-water in such a way that their separation becomes possible. High purity ethanol is removed as top distillate stream of the EDC, while the bottom product consists of solvent and water. The third column (SRC) separates the remaining water as distillate and completely recovers the solvent as bottom product. Note that a key difference as compared to previous work5 is the use of the bottom product of SRC to pre-heat the feed in feed-effluent heat exchanger (FEHE) unit, and then recycle the cooled solvent stream to the extractive distillation column. Remarkable, this minor heat integration reduces the specific energy requirements from 2.11 to 2.07 kWh/kg bioethanol.

22 PDC – pre-concentration distillation column EDC – extractive distillation column SRC – solvent recovery column 1

Ethanol

Solvent

Make-up

78°C, 1.0 bar 12517.5 kg/h 99.8%w EtOH 00.2%w H2O

4 25°C, 1.2 bar

52.9°C, 1.04 bar 20793 kg/h 100%w EG

(Ethylene glycol) EDC

1

PDC-TOP

PDC

Q = 2336 kW

79.2°C, 1.05 bar 13729.4 kg/h 91.0%w EtOH 09.0%w H2O

1

11 Reflux ratio = 0.346 Boilup ratio = 0.994 Qcnd = –3972 kW Qreb = 5606 kW

FEHE

SRC

17

25°C, 1.2 bar 125000 kg/h 10%w EtOH 90%w H2O

EDC-BTM

21

Feed

Water 100°C, 1.0 bar 1225.3 kg/h 00.2%w EtOH 99.0%w H2O 00.8%w EG

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42.6°C Reflux ratio = 1.651 Boilup ratio = 0.264 Qcnd = –9858 kW Qreb = 18424 kW

153.2°C 1.04 bar 22004.8 kg/h 05.5%w H2O 94.5%w EG

Reflux ratio = 0.410 Boilup ratio = 0.359 Qcnd = –1085 kW Qreb = 1797 kW

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Water 103°C, 1.13 bar 111270.6 kg/h 100%w H2O

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16 197.8°C, 1.04 bar 20779.6 kg/h 100%w EG

Solvent (recycle)

Figure 2. Flowsheet including the mass balance and key parameters of the classic bioethanol purification sequence based on extractive distillation (the numbers on columns indicate the top, bottom and feed stage)

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3.3 Vapor recompression assisted extractive DWC

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As all distillation columns of the classic separation sequence (Figure 2) operate at ambient pressure, the use of a DWC was already explored as an attractive process alternative: e.g. combine EDC and SRC into an E-DWC standard configuration, or combine all three columns into a single E-DWC with non-standard configuration.4,7 Hereby we go further by combining the E-DWC technology with vapor recompression (VRC) in order to further increase the energy efficiency. The starting point is the optimized single step separation in an E-DWC described elsewhere.4 The VRC part is added on top of E-DWC based on sensitivity analysis. Figure 3 presents the flowsheet of the novel process for bioethanol purification based on a VRC assisted extractive distillation in a dividing-wall column. For the reader’s convenience, the mass and energy balance, as well as the key process



Luo, Bildea, Kiss

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parameters are also provided. Note that the liquid split ratio (rL) is defined as the ratio between the liquid flowrate going down to the pre-fractionator section (PDC section in Figure 3) over the total liquid flowrate available at the bottom of the top common section (EDC) just before the liquid split occurs. Similarly, the vapor split ratio (rV) is defined as the ratio between the vapor flowrate going up to the pre-fractionator section (PDC section in Figure 3) over the total vapor flowrate available at the top of the bottom common section (SRC).

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As shown in Figure 3 the feed side (prefractionator) has the role of the PDC unit of the classic sequence. Water is collected as liquid side stream, but an extra side-reboiler is needed to return to the column the required amount of water vapor. The diluted ethanol feed is fed as liquid stream on top of the prefractionator side, and therefore serves as liquid reflux to the PDC section. The vapor flow leaving the feed side (PDC) of the E-DWC is enriched in ethanol. Heavy solvent is fed at the top of the E-DWC unit, this common top section playing the role of the EDC unit of the classic sequence. The solvent-to-ethanol ratio considered is 1.0 mol/mol, due to the different ethanol concentration at top of PDC. Ethanol is separated in the top section (EDC) as high purity vapor distillate that is compressed to a higher pressure and temperature level, and then used to drive the side reboiler of the column and eventually being condensed and collected as main product. The liquid flowing down the top section (EDC) is collected and distributed only to the (SRC) side located opposite to the feed side. This complete redistribution of the liquid flow (liquid split ratio rL=0) is required to avoid the presence of solvent on the feed side (PDC) which would lead to a loss of solvent in the water side stream. In the bottom common section (SRC), the solvent is removed as bottom product, then cooled in a feed-effluent heat exchanger (FEHE) and recycled to the E-DWC unit. Due to the large difference of volatilities between water and EG, the separation is quite easy and therefore the recovery of solvent is practically complete, hence all the recovered solvent is recycled. The vapor coming from the bottom part of the E-DWC to the lower part of the dividing-wall consists mainly of water, but this amount is not sufficient for the PDC section hence the need for an extra side-reboiler – which can be effectively driven by a heat pump (VRC in this case). Note that in spite of the high degree of integration of DWC technology and extractive DWC, the controllability of such systems is satisfactory.13,14 149.7°C, 3.7 bar

78°C, 1 bar

Qcnd = –267 kW

W = 1808 kW

COMP

25°C, 1.25 bar

EDC rL

Q = 1947 kW

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Feed

18

39.7°C

S-REB Q = 12024 kW 115.3°C 3.7 bar 52542 kg/h 99.9%w EtOH 00.1%w H2O

115.3°C, 3.7 bar 12500 kg/h 99.9%w EtOH 00.1%w H2O

4

FEHE

25°C, 1.2 bar 125000 kg/h 10%w EtOH 90%w H2O

Ethanol

SRC

Solvent

1

52.5°C 16737 kg/h 100%w EG

PDC

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rV

Reflux ratio = 0.646 Boilup ratio = 2.587 rL=0.0, rV=0.548 Qreb = 10043 kW

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Water

25

104.7C, 1.19 bar 112750 kg/h 99.8%w H2O 00.2%w EG

E-DWC Solvent (recycle)

204.3°C, 1.25 bar 16487.3 kg/h 100%w EG

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Figure 3. Flowsheet including the mass balance and key parameters of the novel process for bioethanol purification based on VRC assisted extractive distillation in a DWC.

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Figure 4 plots the temperature and composition profiles in the E-DWC. The changes in the composition along the column are clearly illustrated by these profiles, being in line with the functional task of each column section: PDC on the diluted feed side, EDC in the common top part, and SRC on the bottom common part of the column. Notably, the temperature difference between the two sides (PDC vs SRC) of the wall is very low (less than 20 °C) hence no practical issues are expected. Also, high purity and recovery is possible for all three products: ethanol as top distillate (99.9 %wt), water as side stream product (99.8 %wt) and EG solvent (>99.9 %wt) as


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220

1

200

0.9

180

0.8

160 140 120

Prefractionator (PF)

100

Main column

- - - PF Side

DWC Side Ethylene Glycol

0.7

Ethanol

0.6 0.5 0.4

Water

0.3

80

0.2

60

0.1 0

40 0

10 11

Luo, Bildea, Kiss

recovered bottom product. Note that the column profiles are slightly different as compared to previous work.4 The difference comes mainly from using fewer stages, but a higher vapor split ratio (rV=0.548 instead of 0.4) which means that more vapor is distributed to the PDC side. This was needed in order to get a feasible match between the duty of the side reboiler and the heat available for recovery from the top vapor stream (VRC loop in Figure 3). Although we have started from an optimized system,4 it is worth noting that optimizing a chemical process is typically a mixed-integer nonlinear problem that is non-convex and likely to have multiple locally optimal solutions. Such problems are intrinsically very difficult to solve, and the solution time increases rapidly with the number of variables and constraints. A theoretical guarantee of convergence to the globally optimal solution is not possible for non-convex problems.

Mass fraction / [-]

1 2 3 4 5 6 7 8 9

Temperature / [°C]

5

10

15

20 25 Stage / [-]

30

35

40

0

5

10

15

20 25 Stage / [-]

30

35

40

Figure 4. Temperature (left) and liquid composition (right) profiles along the extractive dividing-wall column

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3.4 Sensitivity analysis

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Integrating a vapor recompression heat pump with an extractive dividing-wall column requires the setting of the appropriate discharge pressure from the compressor. The actual figure can be obtained by performing sensitivity analysis. Figure 5 shows the dependence of the compressor duty and the inverse of the log-mean temperature difference (LMTD) on the discharge pressure of the compressor involved by the VRC system. In terms of reducing the VRC costs, the compressor duty should be as low as possible (smaller and cheaper compressor), while the LMTD should be as high as possible (larger driving force and smaller heat exchanger).

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However, practical limits are imposed and these reduce the available range for the discharge pressure to 3.1-3.7 bar. For the heat exchanger (side reboiler of E-DWC) which is part of the VRC loop, the LMTD must exceed 5 K to obtain a reasonable sized and inexpensive side reboiler. Note that the area of the heat exchanger (A) depends proportionally with the inverse of LMTD (e.g. A = Q/U · 1/LMTD) – considering a constant heat duty (Q) and heat transfer coefficient (U). Similarly, the compressor is also limited not only by the compression ratio (typically up to 2.5 - 4.0), but mainly by the discharge temperature (DCT) which should not exceed 150°C for safety reasons – at higher temperatures the system may fail from worn rings, acid formations and oil breakdown. 2000

1.2

1900

1

← Compressor duty

1800

0.8

1700 1600 1500

Tcomp,out=150°C

1/LMTD →

0.6 LMTD=5

0.4

1/LMTD / [1/K]

Compressor duty / [kW]

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

1400 0.2

1300 1200

0 2.7 2.8 2.9 3.0 3.1 3.2 3.3 3.4 3.5 3.6 3.7 3.8 3.9 4.0 Pressure / [bar]

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Figure 5. Dependence of compressor duty and the log-mean temperature difference (LMTD) in the sidereboiler, on the discharge pressure of the compressor (VRC system)



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However, practical limits are imposed and these reduce the available range for the discharge pressure to 3.1-3.7 bar. For the heat exchanger (side reboiler of E-DWC) which is part of the VRC loop, the LMTD must exceed 5 K to obtain a reasonable sized and inexpensive side reboiler. Note that the area of the heat exchanger (A) depends proportionally with the inverse of LMTD (e.g. A = Q/U · 1/LMTD) – considering a constant heat duty (Q) and heat transfer coefficient (U). Similarly, the compressor is also limited not only by the compression ratio (typically up to 2.5 - 4.0), but mainly by the discharge temperature (DCT) which should not exceed 150°C for safety reasons – at higher temperatures the system may fail from worn rings, acid formations and oil breakdown.

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Further process optimization based on sensitivity analysis would be possible, but no spectacular improvements are to be expected compared to the energy savings already achieved – although other optimization objectives might be considered additionally: e.g. increase purity of water, reduce the amount of solvent recycle, minimize start-up time, maximize robustness during transient operating regimes, etc.

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3.5 Economic evaluation

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The total investment costs (TIC), total operating costs (TOC) and total annual costs (TAC) have been calculated according to the procedure described in our previous work.5 Note that the economic estimation considers a 20% extra-cost factor for DWC internals that are somewhat more complex as compared to conventional equipment. Concerning the sizing of DWC, each section can be sized for the particular liquid and vapor loads, using the available software. In this way, an equivalent cylindrical cross section area needed to accommodate the required vapor and liquid loads can be rather easily calculated. The latitudinal position of the wall is then set in such way that each column section has an equivalent cross-section area to that calculated by individual sizing.8

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The equipment costs are estimated using correlations, based on the Marshall & Swift equipment cost index (M&S=1468.6 in 2012). For columns and heat exchangers (e.g. condensers and reboilers) made of carbon steel, the estimated cost in US dollars is given by:

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Cshell = fp (M&S / 280) dc1.066 hc0.802 0.65

Chex = (M&S / 280) cx A

(2) (3)

Ccomp = (M&S/280) (664.1 P

0.82

Fc)

(4)

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where fp is the cost factor (2981.68 in this case), cx = 1609.13 (condensers) or 1775.26 (kettle reboilers), A is the heat transfer area (m2), P is the compressor power (kW), hc is the column height (tangent-to-tangent) and dc is the column diameter – calculated using the internals-sizing procedure from Aspen Plus, and in case of DWC as equivalent diameter based on cross sectional area of the two column sections. For compressors, the correction factor Fc varies with design type as follows: 1.00 for centrifugal (motor), 1.07 for reciprocating (steam), 1.15 for centrifugal (turbine), and 1.29 and 1.82 for reciprocating (motor and gas engine). A price of 600 US $/m2 was used for the sieve trays cost calculations. For the TAC calculations, a plant lifetime of 10 years was used. Furthermore, the following costs were considered for different types of utilities: 0.03 $/t cooling water and multiple levels of steam rated at 15, 17 and 20 $/t steam of low, medium and high pressure, respectively.15

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Note that the CO2 emissions were calculated according to method described in earlier studies:5

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[CO2]emissions = (Qfuel / NHV) × (C%/100) α

(5)

where α = 3.67 is the ratio of molar masses of CO2 and C, NHV is the net heating value, and C% is the carbon content – dependent on the fuel. For natural gas, NHV is 48900 kJ/kg and the carbon content is 0.41 kg/kg. Therefore, the total amount of fuel used can be calculated as follows: Qfuel = Qproc / λproc × (hproc – 419) × (TFTB – T0) / (TFTB – Tstack)

(6)

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where λproc (kJ/kg) and hproc (kJ/kg) are the latent heat and enthalpy of the steam, TFTB (K) and Tstack (K) are the flame and stack temperature, respectively. This leads to emissions of 139.84 kg/hr CO2 per MW thermal.

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Table 1 provides more details about the key performance indicators, including the total investment, operating and annual costs. Due to the use of a compressor and a larger side-reboiler required by the VRC system, the total investment cost of the proposed VRC E-DWC process is about 29% higher than for the classical process, but this is largely compensated by the significant energy savings, which exceed 60% at a direct comparison. However, one should note that such a high percentage value in energy savings is based on the direct comparison of thermal duty (heat) vs combined thermal and compressor duty (kW heat and power) of the VRC alternative. When taking into account the inefficiencies in power generation – e.g. by considering that the ratio of heat to


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electrical kW is about 3, then the equivalent energy requirements are 0.805 + 3×0.145 = 1.24 kWh/kg ethanol – the real savings in primary energy become 40% which is still a remarkable figure.

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Considering both the investment and operating costs, the total annual cost (TAC) is reduced by about 24%, while the CO2 emissions – which are closely linked to the primary energy requirements – are reduced by 40% (or even 60% when the electricity for the compressor comes from renewable sources). Note that in case of the VRC assisted E-DWC process, the investment cost may include also some additional equipment required for the startup procedure. An example is a trim reboiler – of a smaller size, working at larger temperature approach due to the use of steam – which could be used next to the side reboiler.

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Table 1. Process comparison in terms of key performance indicators Key performance indicator Equipment cost breakdown (k$) – column shells (incl. internals) – condensers (heat exchangers) – reboilers (heat exchangers) – process-process heat exchangers – compressor (VRC) Total investment costs, TIC (k$) Total operating costs, TOC (k$/yr) Total annual costs, TAC (k$/yr) CO2 emissions (kg CO2 / ton product)* Thermal energy use (kWh / kg product) Electrical energy use (kWh / kg product) Equivalent energy requirements (kWh / kg)

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Classic process

E-DWC process4

VRC E-DWC process

Difference vs. classic (%)

1103 1335 885 137 – 3462 5784

1111 1073 1442 – – 3626 5355

912 71 356 1503 1632 4477 4221

+29.3 –27.0

6130 288.94 2.07 n/a 2.07

5718 288.31 2.07 n/a 2.07

4668 173.04 (112.35) 0.80 0.14 1.24

–23.8 –40.1 (–61.1) –61.1 n/a –40.1

* Note: The values given in parenthesis are for the case when electricity is generated from renewable sources.

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4. Conclusions

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The novel heat pump assisted extractive distillation process proposed here is based on an efficient combination of vapor recompression heat pump and DWC technology. In this configuration, the ethanol top vapor stream of the extractive DWC is recompressed from atmospheric pressure to 3.1-3.7 bar (thus to a higher temperature) and used to drive the side reboiler responsible for water vaporization. The results show that the specific energy requirements drop from 2.07 kW.h/kg (classic sequence) to only 1.24 kW.h/kg ethanol (VRC assisted extractive DWC), thus energy savings of 40% being possible. Considering the requirements for a compressor and use of electricity in case of the heat pump assisted alternative, about 24% decrease of the total annual costs is possible – in spite of the 29% increase of the capital expenditures – for the novel process, as compared to the classic ED.

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Acknowledgment

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The authors thank Kim Dam-Johansen and Hao Wu (Technical University of Denmark) for helpful discussions, and gratefully acknowledge the support provided to Hao Luo and his colleagues – by the Sino-Danish Center for Education and Research, Technical University of Denmark, Haldor Topsøe A/S, Hempel A/S, and Novozymes A/S – to attend the Grøn Dyst 2014 (Green Challenge) at Technical University of Denmark.

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Novel Heat-Pump-Assisted Extractive Distillation for Bioethanol

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