Energy xxx (2010) 1e13
Contents lists available at ScienceDirect
Energy
journal homepage: www.elsevier.com/locate/energy
Improving bioethanol production from sugarcane: evaluation of distillation,
thermal integration and cogeneration systems
Marina O.S. Dias a, b, *, Marcelo Modesto c, Adriano V. Ensinas c, Silvia A. Nebra d, Rubens Maciel Filho a,
Carlos E.V. Rossell a, b
a
School of Chemical Engineering, University of Campinas, P.O. Box 6066, 13083-970 Campinas, SP, Brazil
Laboratório Nacional de Ciência e Tecnologia do Bioetanol (CTBE), P.O. Box 6170, 13083-970 Campinas, SP, Brazil
Centre of Engineering, Modeling and Applied Social Sciences, CECS, Federal University of ABC, Rua Santa Adélia, 166, 09210-170 Santo André, SP, Brazil
d
Interdisciplinary Centre of Energy Planning - NIPE, University of Campinas, P.O. Box 6192, 13400-970 Campinas, SP, Brazil
b
c
a r t i c l e i n f o
a b s t r a c t
Article history:
Received 1 March 2010
Received in revised form
9 September 2010
Accepted 10 September 2010
Available online xxx
Demand for bioethanol has grown considerably over the last years. Even though Brazil has been producing
ethanol from sugarcane on a large scale for decades, this industry is characterized by low energy efficiency,
using a large fraction of the bagasse produced as fuel in the cogeneration system to supply the process
energy requirements. The possibility of selling surplus electricity to the grid or using surplus bagasse as raw
material of other processes has motivated investments on more efficient cogeneration systems and process
thermal integration. In this work simulations of an autonomous distillery were carried out, along with
utilities demand optimization using Pinch Analysis concepts. Different cogeneration systems were
analyzed: a traditional Rankine Cycle, with steam of high temperature and pressure (80 bar, 510 C) and
back pressure and condensing steam turbines configuration, and a BIGCC (Biomass Integrated Gasification
Combined Cycle), comprised by a gas turbine set operating with biomass gas produced in a gasifier that uses
sugarcane bagasse as raw material. Thermoeconomic analyses determining exergy-based costs of electricity and ethanol for both cases were carried out. The main objective is to show the impact that these
process improvements can produce in industrial systems, compared to the current situation.
Ó 2010 Elsevier Ltd. All rights reserved.
Keywords:
Bioethanol
Process integration
Sugarcane
Exergetic analysis cost
Power generation
1. Introduction
Increased interest on alternative fuels has been observed in the
past few years, as a result of increasing energy demand and forecasted depletion of fossil resources [1,2]. Global warming and the
consequent need to diminish greenhouse gases emissions have
encouraged the use of fuels produced from biomass [3], which is
the only renewable carbon source that can be efficiently converted
into solid, liquid or gaseous fuels [4].
Bioethanol is presently the most abundant biofuel for automobile
transportation [5]. It is produced from fermentation of sugars
obtained from biomass, either in the form of sucrose, starch or
lignocellulose. Sugarcane is so far the most efficient raw material for
bioethanol production: the consumption of fossil energy during
sugarcane processing is much smaller than that of corn [6]. One of
the main by-products generated during sugarcane processing is
sugarcane bagasse, which is usually burnt in boilers for production of
steam and electrical energy, providing the energy necessary to fulfill
* Corresponding author. Laboratório Nacional de Ciência e Tecnologia do Bioetanol (CTBE), P.O. Box 6170, 13083-970 Campinas, SP, Brazil. Fax: þ55 19 3518 3164.
E-mail address: [email protected] (M.O.S. Dias).
the process requirements. The use of efficient technologies for
cogeneration coupled with optimization of bioethanol production
process allows the production of surplus bagasse [7], which may be
used as a fuel source for electricity generation or as raw material for
producing bioethanol and other biobased products [8,9]. Different
cogeneration systems, such as the Rankine Cycle with condensing
steam turbines and those based on gasification technologies, may
improve the amount of electricity produced from sugarcane bagasse,
thus allowing its sell to the grid.
Optimization of bioethanol production process can reduce energy
consumption, consequently increasing sugarcane bagasse availability. Thermal integration using Pinch Technology can reduce hot
and cold utilities requirements, thus optimizing energy consumption
of the bioethanol production process [5].
Separation is the step where major costs are generated in process
industry [10]. In the case of bioethanol production, distillation
process may be optimized to reduce steam consumption on column
reboilers, therefore reducing production costs. The distillation
process commonly employed in Brazilian distilleries is based on the
same configuration used for decades, on which atmospheric pressures are adopted; in a new configuration presented in this work,
0360-5442/$ e see front matter Ó 2010 Elsevier Ltd. All rights reserved.
doi:10.1016/j.energy.2010.09.024
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M.O.S. Dias et al. / Energy xxx (2010) 1e13
wine is fed to distillation columns operating under vacuum pressures, where ethanol-rich streams containing from 40 to 50 wt%
ethanol are produced; these streams are fed to atmospheric-pressure
rectification columns, where hydrous ethanol (approximately 93 wt
%) is produced. Reduction of steam consumption on distillation
column reboilers is then achieved by the use of a double-effect
distillation system, since different temperature levels are obtained in
column condensers and reboilers, allowing the thermal integration
of these equipments [11].
In order to evaluate the consequences of introducing a doubleeffect distillation system in ethanol production from sugarcane,
simulations of bioethanol production processes were carried out
using software UniSim Design from Honeywell. Steam demand was
evaluated for each case (bioethanol production using conventional
or double-effect distillation columns), after thermal integration
using Pinch Analysis was carried out. The results obtained were
used to assess two different cogeneration systems: Rankine Cycle
operating with steam at high temperature and pressure (480 C and
80 bar) and a BIGCC, on which sugarcane bagasse is gasified. Even
though the gasification technology is not yet commercially feasible,
the possibility of improving electricity production in this case
should be evaluated. Simulations of the cogeneration systems were
carried out using Gate Cycle and EES (Engineering Equation Solver).
Thermoeconomic analyses and determination of exergy-based
costs for ethanol and electricity were assessed for both cases.
2. Bioethanol production process
In this work an autonomous distillery, on which all sugarcane
processed is used to produce ethanol, for the production of one
million liters of anhydrous ethanol per day was considered. This is
a typical industrial scale unit, where around 12,000 ton of sugarcane (TC) per day are crushed, producing around 85 L of anhydrous
ethanol per ton of sugarcane.
Simulations were carried out using the commercial software
UniSim Design from Honeywell.
2.1. Chemical and physical properties
In order to represent sugarcane composition (displayed in
Table 1) more accurately, some hypothetic components that are not
part of UniSim database were created as described in a previous
work [12]: bagasse components (cellulose, hemicellulose and
lignin); sand, with properties considered equal to those of SiO2;
impurities, represented by potassium salts and aconitic acid; input
materials such as phosphoric acid and lime; calciumephosphate,
the main salt formed during liming operation, of great importance in
removal of impurities during juice settlement; minerals, represented by K2O [13]; and yeast. Properties for these hypothetic
components were obtained in Wooley and Putsche [14] and Perry
and Green [15]. All reducing sugars are considered dextrose; all
other components (water, sucrose, ethanol, carbon dioxide, glycerol,
succinic acid, acetic acid, isoamyl alcohol, sulfuric acid, MEG
(monoethyleneglycol)) are part of UniSim database.
Table 1
Composition of sugarcane received in the factory.
NRTL (Nom-Random Two Liquid) was the model chosen for
calculating activity coefficient of the liquid phase, and the equation of
state SRK (SoaveeRedlicheKwong) for the vapor model. It was
verified that the NRTL model was the one that calculated elevation of
the boiling point of sucrose solutions with larger accuracy, when
compared to UNIQUAC (Universal Quasi Chemical) or PengeRobinson equation of state, as represented in Fig. 1. For the
extractive distillation process, employed in ethanol dehydration, the
model UNIQUAC and equation of state SRK were used.
2.2. Main aspects of the production process
Production of anhydrous bioethanol from sugarcane is
comprised by the following main steps: reception and cleaning of
sugarcane, extraction of sugars, juice treatment, concentration and
sterilization, fermentation, distillation and dehydration. All these
steps were represented on the simulation [12]. Some process
improvements were considered in this work, such as: for sugarcane
cleaning, use of a dry-cleaning system, instead of one based on water
use, what decreases sugar losses as well as water consumption;
efficient juice treatment, considering addition of phosphoric acid
and lime, what increases juice purity and consequently improves
the fermentation stage; concentration of juice on multiple effect
evaporators, decreasing process steam consumption; decreased
fermentation temperature (28 C), which allows the production of
wine with higher ethanol content, consequently diminishing sugar
and ethanol losses, as well as vinasse generation; a double-effect
distillation process, which allows thermal integration between
columns condensers and reboilers; optimization of the extractive
distillation process with MEG for anhydrous bioethanol production,
considering feed of hydrous ethanol on the vapor phase (in the
industry, hydrous bioethanol is obtained in a condensed phase in the
conventional distillation process, followed by vaporization prior to
the extractive distillation process, what leads to increased and
unnecessary steam consumption). A block-flow diagram of the
bioethanol production process is depicted in Fig. 2. Description of
the different steps of the bioethanol production process from
sugarcane is presented in the following subsections.
2.3. Sugarcane reception and cleaning, extraction of sugars and
juice treatment
A dry-cleaning system is used to clean sugarcane received in the
factory, removing about 70% of dirt before it is fed to the mills.
Extraction of sugars is done using mills, where sugarcane juice and
110
Experimental
NRTL
Peng Robinson
UNIQUAC
108
B o ilin g P o in t ( ° C )
2
106
104
102
Component
Content (wt%)
Sucrose
Fibers
Reducing sugars
Minerals
Impurities
Water
Sand
13.30
11.92
0.62
0.20
1.79
71.57
0.60
100
0
10
20
30
40
50
60
70
80
90
Mass fraction of sucrose (%)
Fig. 1. Boiling point of sucrose solutions: experimental data [16] and values calculated
using the process simulator with NRTL, PengeRobinson and UNIQUAC.
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M.O.S. Dias et al. / Energy xxx (2010) 1e13
3
acids, glycerol and yeast. Fermentation gases are collected and
washed in an absorber for ethanol recovery, while wine containing
yeast cells are centrifuged to recover yeasts. The yeast milk stream
receives an acid treatment prior to be fed back to the fermentor, in
order to avoid contamination. Wine is mixed with alcoholic
solutions from the absorber before being fed to the distillation
columns.
Fermentation was carried at 28 C, in order to obtain wine with
higher ethanol content (around 10 wt%), decreasing ethanol and
sugar losses and energy consumption during the purification step
[17]. An alternative cooling method (absorption with lithiumebromide) was considered to supply cool water to maintain this
low fermentation temperature.
2.6. Distillation and dehydration
The conventional distillation process consists of a distillation
column comprised by columns A, A1 and D, and a rectification
column comprised by columns B and B1. Wine obtained in fermentation is pre-heated and fed to column A1, which is located above
column A and below column D, as shown in Fig. 3. Pressure on the
distillation columns range from 133.8 to 152.5 kPa, and on the
rectification columns from 116 to 135.7 kPa.
In the distillation columns ethanol-rich streams containing
around 40 wt% ethanol (phlegms) are obtained, as well as vinasse
and 2nd grade ethanol. In the rectification column hydrous ethanol
R
Gases
Fig. 2. Block-flow diagram of the bioethanol production process from sugarcane in an
autonomous distillery.
Top
D
Top D
recycle
bagasse are obtained. Water is used to improve sugars recovery, in
a process termed imbibition [13]. A physical treatment is used to
remove sand and fiber from the juice, in which screens and hydrocyclones are used [13]. Juice then receives a chemical treatment
consisting of addition of phosphoric acid, lime, heating and settlement to remove other impurities. In the settler, mud and clarified
juice are obtained. A filter is used to recover some of the sugars
contained in the mud, and the filtrate is recycled to the process prior
to the second heating operation.
2nd grade
ethanol
Liquid
phlegm
D
R
R
Hydrous
Ethanol
Warm
wine
2.4. Juice concentration and sterilization
Clarified juice contains around 15 wt% solids, so it must be
concentrated before fermentation in order for the product to contain
an adequate ethanol content that allows reduction of energy
consumption during purification steps. Concentration is done on
a 5-stage multiple effect evaporators (MEE) up to 65 wt% sucrose.
Vapor purges may be done if needed, in order to supply heat to other
operations. Only part of the juice is concentrated, and final juice
contains around 22 wt% sucrose. In order to avoid contamination in
fermentation, juice is sterilized and cooled till fermentation
temperature (28 C).
Fusel
oil
A1
R
R
R
Phlegmasse
B,
B1
Vapour
Phlegm
2.5. Fermentation
Simulations were based on the Melle-Boinot (feed-batch with
cell recycle) process. Sterilized juice is added to the fermentors
along with yeast stream. In the fermentor sucrose is inverted to
glucose and fructose, which are consumed by the yeast producing
ethanol, CO2 and other products, such as higher alcohols, organic
Vinasse
A
Fig. 3. Conventional distillation process configuration.
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M.O.S. Dias et al. / Energy xxx (2010) 1e13
Gas
2nd
grade
ethanol
Anhydrous
Ethanol
MEG
R
D
Liquid
phlegm
Hydrous
ethanol.
Warm
wine
R
Extractive
Water
R
R
R
A1
R
Fusel
oil
R
Phlegmasse
Vapour
phlegm
Recovery
B,
B1
Solvent
Bottom
A
A
Boilup-A
R
Vinasse
Fig. 4. Double-effect distillation e integration of column A reboiler to column B and extractive column condensers.
and residues like phlegmasse and fusel oil are obtained. Hydrous
ethanol in vapor phase is produced on top of column B.
Since water and ethanol form an azeotrope with concentration
of 95.6 wt% ethanol at 1 atm, an extractive separation process with
MEG was used to produce anhydrous bioethanol. In this process
two columns are used: an extractive column, where hydrous
ethanol and a suitable solvent is used, anhydrous bioethanol is
produced on the top and a solvent solution on the bottom, and
a recovery column, in which extractive solvent is recovered.
Extractive column operates at atmospheric pressure (101.325 kPa),
while the recovery column operates at 20 kPa, in order to avoid
high temperature and solvent decomposition. The solvent is cooled
before fed to the extractive distillation. Since the bottom temperatures of both extractive and recovery columns are relatively high
(136 and 150 C, respectively), both reboilers need to operate with
high pressure (6 bar) steam.
Table 2
Initial and final temperature, heat flow of streams considered for thermal integration.
Streams
Hot streams
Conventional distillation
Tinitial
Tfinal
( C)
( C)
Heat flow
(kW)
Double-effect distillation
Tinitial
Tfinal
( C)
( C)
Sterilized juice
Fermented wine
Vinasse
Anhydrous ethanol cooling
Vapor condensates
Condenser column B
Condenser extractive column
Condenser column D
Condenser recovery column
Cold streams
130
28
112
78
107
82
78
81
82
Tinitial ( C)
Tfinal ( C)
27,606
5440
27,426
9070
10,499
24,297
9224
662
290
Heat flow (kW)
130
28
65
78
109
82
78
41
78
Tinitial ( C)
Tfinal ( C)
Raw juice
Limed juice
Juice for sterilization
Centrifuged wine
Reboiler column A
Reboiler column B
Reboiler extractive column
Reboiler recovery column
30
76
96
28
112
108
111
150
70
105
130
82
112
108
137
150
21,062
20,535
9287
23,161
38,553
6725
8432
1704
30
74
96
28
65
108
111
150
70
105
130
48
65
108
137
150
28
24
35
35
50
82
78
30
64
28
24
35
35
50
82
78
26
63
Heat flow
(kW)
27,602
5063
10,547
8983
12,418
23,656
13,734
4482
295
Heat flow
(kW)
21,096
20,667
9309
8636
37,389
5248
12,881
1794
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5
Fig. 5. Composite Curves for the conventional distillation case.
In the double-effect distillation system, the distillation columns
operate under vacuum pressures (19e25 kPa), while the rectification
column operates at atmospheric pressures (101.325e135.7 kPa), what
makes temperature on the bottom of column A reach about 65 C and
temperature on the top of column B reach 78 C. Consequently,
integration of column B condenser and column A reboiler is possible,
and the rectification column condenser may work as the distillation
column reboiler, what provides a reduction in steam consumption on
distillation. Nevertheless, since phlegms are obtained in the vapor
phase in column A, they must be compressed before being fed to
column B, what is done by using steam compressors. Since column A
reboiler duty is greater than that of column B condenser, this integration cannot supply all the heat necessary to the adequate operation of the distillation columns. Given that temperature on top of the
extractive column from the extractive distillation process reaches
78 C, it is possible to use anhydrous bioethanol vapors to provide the
remaining heat necessary to an adequate operation of column A. This
integration in the double-effect distillation is shown in Fig. 4.
3. Process integration analysis
The Pinch Point Method described by Linnhoff et al. [18] was used
to analyze the streams of the process which are available for thermal
integration. The method developed by works of Hohmann [19],
Umeda et al. [20] and Linnhoff and Flower [21,22] uses enthalpyetemperature diagrams to represent the process streams and to
find the thermal integration target for them, considering a minimum
approach difference of temperature (Dtmin) for the heat exchange.
The thermal integration target indicates the minimum
requirements of external hot and cold utilities for the process. The
Fig. 6. Composite Curves for the double-effect distillation case.
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Some streams with less than 1000 kW of heat flow were not
considered in the analysis because of their low thermal integration
potential. The evaporation system was considered as a separated
system which is integrated with the background process in a second
step. This procedure previously applied by Ensinas [23] in sugarcane
mills simplifies the analysis, maximizing the vapor bleeding use and
reducing the total exhaust steam consumption in the process.
As it was previously mentioned, in this analysis a single-effect
lithium bromide absorption system, with COP 0.65, was considered
to provide the cold utility requirements of the fermentation step,
using for that vapor from the third effect of the evaporation system
as heating source to produce the necessary cold water.
Thus, following the procedure presented by Linnhoff et al. [18]
the CC (Composite Curve) (Figs. 5 and 6) could be drawn and the
targets determined. Fig. 7 shows the GCC (Grand Composite Curve)
of double-effect distillation case considering the integration of the
background with the evaporation system after the thermal integration. The Pinch Point temperature in this case was found at
113 C, considering a global Dtmin of 10 C for all the streams of the
background process, and a Dtmin of 4 C for the evaporation vapor
streams. The software presented by Elsevier [24] was used for
calculation of the targets and drawing of CC and GCC curves.
4. Cogeneration systems
Fig. 7. Grand Composite Curves for the double-effect distillation case (background/
foreground process integration).
analysis is useful to represent all the streams available for heat
exchange in one diagram, being possible to evaluate the effect of
modifications in the process parameters before the heat exchange
network design.
Thermal integration analysis of the bioethanol production
process considered the hot and cold streams parameters determined in the simulation of the autonomous distillery for both
conventional and double-effect distillation. Thermal integration
between rectification and distillation columns has an important
impact on the steam balance, promoting changes in the integration
possibilities. Data of the streams available for thermal integration
are presented in Table 2.
Traditionally, cogeneration systems employed in Brazilian sugarcane mills are based on the Rankine Cycle [25]. Sugarcane bagasse is
used as a fuel to supply thermal, mechanical and electrical demand of
the sugar and ethanol production process. These plants used to
operate with steam at low levels of pressure and temperature (20 bar
and 350 C) and back pressure steam turbines, resulting in low
energetic efficiency, high bagasse consumption, low bagasse surplus
and a small electricity surplus, or even none. After the Brazilian
electrical crisis in 2001, independent energy producers were allowed
to sell electricity to the grid, thus providing the sugarcane industry
the possibility of improving electricity generation from biomass. As
a consequence, new projects of sugarcane mills considered the use of
Rankine Cycles with steam at higher levels of temperature and
pressure, condensing steam turbines and the use of all the bagasse
generated in the mills as a fuel to produce electricity, selling the
surplus electricity to the grid. In the following sections a description
of the cogeneration systems studied in this work is made.
Fig. 8. Configuration of the cogeneration system based on Rankine Cycle: back pressure steam turbines.
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Fig. 9. Configuration of the cogeneration system based on Rankine Cycle: condensing steam turbines.
4.1. Traditional cogeneration system in sugarcane mills (Rankine
Cycle)
Figs. 8 and 9 show two basic configurations of the Rankine Cycle
used in cogeneration systems traditionally employed in Brazil. The
first configuration (Fig. 8) displays a boiler, a back pressure steam
turbine, deaerator and pumps; the second configuration (Fig. 9)
shows the same set of equipment; however, a condensing steam
turbine is employed, therefore a condenser is added to system. The
cogeneration system provides steam (2.5 and 6 bar of pressure) and
electricity to supply the anhydrous ethanol production process. The
ethanol production process receives sugarcane (stream 11 on Fig. 8)
producing anhydrous ethanol (stream 12), vinasse (stream 13) and
sugarcane bagasse (stream 14), and requires 6 bar (stream 5) and
2.5 bar steam (stream 3) besides electrical energy (We). The
bagasse produced in the mills has a moisture content of 50% and is
used to supply energy to the cogeneration system (stream 10). 7% of
the bagasse is separated to provide energy for start-ups (stream 15)
and the remaining fraction is available for other uses (stream 16).
Sugarcane bagasse flow was calculated using Eq. (1).
mbagasse ¼
x
mcane
w
(1)
where: x is the percentage of fiber in sugarcane; w is the moisture
content of the bagasse. The values used for fiber content of sugarcane and moisture of sugarcane bagasse are 12% and 50%,
respectively.
Nowadays, bagasse surplus is used to increase the production of
electricity, mainly in condensing systems as shown in Fig. 9.
However, when production of ethanol from lignocellulosic materials
(second generation ethanol) becomes feasible, bagasse surplus may
be used to increase ethanol production from sugarcane, using the
same cultivated area through pretreatment and hydrolysis processes.
The use of a Rankine Cycle was assessed through a thermoeconomic analysis using software EESÒ; the main parameters used in
the simulation are shown in Table 3.
4.2. Cogeneration systems based on BIGCC (Biomass Integrated
Gasification Combined Cycle) in sugarcane mills
The use of the BIGCC in sugarcane mills has been investigated
for the past 15 years [26e33]. Other works assess the combined use
of biomass gas and natural gas in cogeneration plants, in order to
overcome the problems found in the BIGCC applied in sugarcane
mills [34e39]. The use of gas turbines with biomass is an option as
well [40e46].
A configuration of the BIGCC system employed in sugarcane
mills is illustrated in Fig. 10. The system consists of a gasifier,
a dryer, a gas turbine set, an HRSG (Heat Recovery Steam Generator)
and back pressure steam turbines that provide steam to supply the
demands of the process at both 2.5 and 6 bar. Similarly to the
previous case, sugarcane bagasse is produced during anhydrous
ethanol production (stream 10 on Fig. 10) and fed to an atmospheric
gasifier, on which biomass gas production takes place. The biomass
gas that leaves the gasifier (stream 15) is compressed in a biomass
gas compressor and fed to a combustion chamber, on which it
reacts with compressed air (stream 2) producing stack gases
(stream 3) and supplying power for compressor and generator and,
consequently, producing electric energy. Stack gases from the gas
turbine (stream 4) are fed in the HRSG producing steam at 480 C
and 80 bar, which supplies the thermal demand of the ethanol
production process through steam turbines that produce low
pressure steam (stream 17, which consists of 6 bar steam and
stream 19, 2.5 bar steam). The stack gases that leave the HRSG
(stream 5) are used to dry the bagasse before it is fed in the gasifier.
One of the most important steps of the simulation of BIGCC is
the modeling of the gasifier. In order to simulate the behavior of the
BIGCC two important parameters must be determined: lower
heating value of biomass gas and adequate moisture content of the
bagasse after the drier. In order to obtain these parameters,
a computational tool kindly provided by Pellegrini and Oliveira Jr.
[47] is employed. The model is based on a study carried out by Fock
and Thomsem [48]. The input variables are sugarcane bagasse
composition (47.7% carbon, 5.8% hydrogen and 46.5% oxygen) [49]
Table 3
Main parameters adopted in the simulation of the Rankine Cycle.
Parameters
Value
Sugarcane processed (ton/h)
Anhydrous ethanol production (L/TC)
Vinasse production (L/L of anhydrous ethanol)
Lower heating value of bagasse (kJ/kg)
Isentropic efficiency of pumps
Isentropic efficiency of steam turbines
First law efficiency of boiler
Electric power demand in the process (kWh/TC)
Fraction of bagasse for start-ups (%)
Efficiency of electrical generators
493
85
11
7500
0.85
0.80
0.85
28
7
0.98
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Fig. 10. Configuration of the cogeneration system based on the BIGCC cycle.
and moisture. Gasification was considered to occur in the presence
of atmospheric air. In fact, the best gasification conditions are
present when a mixture of oxygen and steam is employed;
however, costs for separating oxygen from atmospheric air for use
in power generation are prohibitive.
The composition of biomass gas as a function of sugarcane bagasse
moisture is represented in Fig. 11. The products of bagasse gasification
are CO, CO2, H2O, CH4, N2 and H2. Molar fraction of CH4 is considered
constant and that of the other components vary with bagasse moisture. Decrease of bagasse moisture from 50 to 40% leads to an increase
of the H2 fraction; lower moisture values lead to a decrease of the H2
fraction. A different behavior is observed for the molar fraction of both
CO and N2, which increase when moisture content decreases. Molar
fractions of H2O and CO2 decrease along with moisture; this behavior
affects the lower heating value of biomass gas as well as the gasification temperature, as shown in Fig. 12.
Profiles of the gasification temperature and lower heating value
of the biomass gas as a function of bagasse moisture are depicted in
Fig. 12. Increase of the CO molar fraction and decrease of water
molar fraction gives rise to an increase on the lower heating value
of biomass gas. In addition, if the temperature is too low, there will
not be enough energy to start up the gasification process. According
to Reed and Gaur [50], the gasification process requires a minimal
temperature of approximately 800 C. Thus, the minimal value of
sugarcane bagasse moisture is 25%, as illustrated in Fig. 12.
Sugarcane bagasse produced during sugarcane processing has
a moisture content of 50%. Therefore, in order to be employed in
Fig. 11. Biomass gas composition as a function of bagasse moisture.
Fig. 12. Lower heating value and biomass gas temperature versus bagasse moisture.
Please cite this article in press as: Dias MOS, et al., Improving bioethanol production from sugarcane: evaluation of distillation, thermal integration and cogeneration systems, Energy (2010), doi:10.1016/j.energy.2010.09.024
M.O.S. Dias et al. / Energy xxx (2010) 1e13
a gasification cycle, the moisture content of bagasse must be
reduced to 25%. The bagasse drying process takes place at the drier
(equipment IV depicted in Fig. 10). The drier makes use of the stack
gases from the HRSG (stream 5), and its main parameters are
temperature and composition of the stack gases and the outlet
minimal temperature of the gases leaving the device (which is
assumed to be equal to the stack gases adiabatic saturated
temperature). The energy balance in the drier is shown in Eq. (2)
m5 h5 þ m10 h10 m6 h6 þ m14 h14 ¼ 0
(2)
where: m10 ¼ mbd þ mw10, mbd represents the dry bagasse flow and
mw10 the amount of water present in the bagasse; m14 ¼ mbd þ mw14,
mw14 corresponds to the amount of water in the bagasse that leaves
the drier and m6 ¼ m5 þ mw6, on which mw6 is the amount of water
removed from the bagasse.
It is convenient to define the moisture of bagasse through
streams 10 (bagasse fed to the drier) and 14 (dried bagasse), as
shown in Eq. (3).
w10 ¼
mw10
m
and w14 ¼ w14
m10
m14
(3)
The values adopted in this simulation consider the moisture
content of the stream 14 equal to 15% [47]. However, since the
temperature (75 C, considered constant) of stream 6 (stack gases
leaving the drier) depends on the composition of the stack gases,
and it is necessary to reach 15% moisture on stream 15 (bagasse
leaving the gasifier), energy must be provided to reach the required
moisture level through a convenient value of temperature of the
stack gas from the HRSG (stream 5). Usually, the value used in the
HRSG is close to 150 C [51]. However, this value is too low to assure
a convenient bagasse drying, as shown in Fig. 13. Increasing stack
gases temperature leads to suitable levels of moisture for gasification; however, decrease of this temperature implies reduction on
the capacity of steam production for the ethanol production
process. In order to assure bagasse moisture of 15%, the minimum
temperature of the stack gases must be 215 C for this simulation, as
depicted in Fig. 13.
In a previous study, Modesto et al. [52] showed three proposals to
use the BIGCC cycle in ethanol plants. The first two proposals used
a commercial gas turbine (GE LM 2500) adapted to use biomass gas
as fuel; however, the thermodynamic conditions adopted could not
supply the steam demand (340 kg/TC) without using a supplementary fuel (natural gas) in the HRSG. The third proposal, which
considers a gas turbine designed for biomass gas, allowed the
elimination of the supplementary fuel; however, the operating
conditions of the gas turbine penalize the performance of electric
9
generation. According to Modesto et al. [52], a suitable value of
steam demand to make a BIGCC cycle feasible is 230 kg/TC. Thus,
thermal integration and the use of a double-effect distillation
system, as described in this work, will improve the feasibility of the
BIGCC cycle by not penalizing electric generation.
In order to execute the thermoeconomic analysis of the BIGCC
cycle the software Gate CycleÔ (mass and energy balances) and
EESÔ (exergy and exergy cost balances) were used. The main
parameters used in the simulation of BIGCC cycle are shown in
Table 4. In this simulation, a gas turbine using only biomass gas is
simulated using the Gate CycleÒ Software, developed by GePower.
Palmer et al. [42] provided some suggestions for the use of biomass
gas in the operation of this gas turbine. The main parameters
analyzed were pressure ratio and temperature at the outlet of the
combustion chamber. The biomass gas from gasifier is compressed
in a compressor and enters a combustion chamber along with an air
flow at the same pressure. Both react in a combustion chamber and
the gases produced expand in a power turbine producing electrical
energy. It is important to mention that the exclusive use of biomass
gas in the gas turbine needs a catalytic combustion chamber
designed to operate with fuels of low heating values. Proposals of
catalytic combustion chambers can be found at Witton et al. [53]
and Forzatti [54]. Another aspect that needs to be mentioned is
that this study does not consider the energetic consumption of an
important step of the gasification process, the cleaning of gases that
leave the gasifier. The cleaning of stack gases is a very important
step to allow the use of these gases in a gas turbine, in order to
assure suitable levels of impurities to avoid corrosion and scorching
in the blades of the turbine.
5. Exergetic cost analysis
In order to assess the two cogeneration systems proposals, an
exergetic cost analysis was performed. This analysis employs mass,
energy, exergy and exergy costs balances in all components of the
plant. Thus, it is possible to determine energetic consumption,
irreversibility generation and the exergetic cost of the products,
through a convenient cost allocation method.
Eqs. (4)e(6) show mass, energy and exergy balances for
a generic control volume, respectively, according to Kotas [55].
X
_ in m
_ þ
Q_ W
X
X
_ out ¼ 0
m
_ in hin m
X
(4)
_ out hout ¼ 0
m
(5)
X
X
T
_ þ
_ in ein _ out eout ¼ I_
Q_ 1 W
m
m
To
(6)
Stack gases temperature from HRSG (ºC)
230
In order to determine steam, stack gases, air, biomass gas and
water exergy, Eq. (7) was used
210
190
Table 4
Parameters used in simulations of BIGCC cycle.
170
150
130
110
90
15
20
25
30
35
40
45
50
Bagasse moisture (%)
Parameter
Value
Production of biomass gas (kg/kg of bagasse)
Lower heating value (kJ/kg)
Pressure ratio
Isentropic efficiency of compressors
Isentropic efficiency of power turbine
Isentropic efficiency of steam turbine
Isentropic efficiency of pumps
Temperature of gases leaving the HRSG ( C)
Temperature of gases leaving the drier ( C)
2.23
5100
18
0.85
0.92
0.9
0.85
215
75
Fig. 13. Bagasse moisture as a function of stack gases temperature from HRSG.
Please cite this article in press as: Dias MOS, et al., Improving bioethanol production from sugarcane: evaluation of distillation, thermal integration and cogeneration systems, Energy (2010), doi:10.1016/j.energy.2010.09.024
10
M.O.S. Dias et al. / Energy xxx (2010) 1e13
Table 5
Hot utilities demand before and after thermal integration for the studied configurations (A: conventional distillation without thermal integration; B: double-effect distillation
without thermal integration; C: conventional distillation with thermal integration; D: double-effect distillation with thermal integration).
Operation
Saturated steam consumption (kg/h)
Configuration A
1st juice heating
2nd juice heating
Multi-effect evaporation
Juice sterilization
Reboiler column A
Reboiler column B
Reboiler extractive column
Reboiler recovery column
Total
Configuration B
6 bar
2.5 bar
6 bar
2.5 bar
6 bar
2.5 bar
6 bar
34,632
33,765
56,270
e
63,390
11,058
e
e
199,115
e
e
e
15,843
e
e
14,426
2915
33,184
34,687
33,981
60,706
e
e
8629
e
e
138,003
e
e
e
15,916
e
e
22,039
3069
41,024
e
e
59,254
4056
63,390
11,058
e
e
137,758
e
e
e
e
e
68,519
5427
e
8629
e
e
82,575
e
e
e
(7)
where: hi ¼ enthalpy at point “i”; ho ¼ enthalpy of reference;
si ¼ entropy at point “i”, so ¼ entropy of reference; ech ¼ chemical
exergy of water. The values used for the temperature and pressure
of reference are 25 C and 1 bar, respectively. The chemical exergy
values are determined by Szargut et al. [56].
The exergy values of sugarcane bagasse were calculated by the
methodology described in Sosa-Arnao and Nebra [57]; exergy of
the ethanolewater mixture employed the methodology described
in Modesto et al. [58].
The methodology used to perform the exergetic cost analysis
is the Theory of Exergetic Cost proposed by Lozano and Valero
[59]. This methodology can be used to determine the exergetic
and monetary cost of each flow that compose the system. The
evaluation of the exergetic and monetary costs of a cogeneration
system in a sugar plant evaluating the influence of the price of
the main fuel (sugarcane bagasse) in steam production and
electricity costs was presented by Sanchez and Nebra [60]. The
use of the thermoeconomic methodology to assess the exergetic
cost of sugar in the production process was described by Fernandez Parra [49].
The exergetic cost calculation is made through cost balance
equations in each component, as shown by Eq. (8).
kin Ein X
Configuration D
2.5 bar
e ¼ hi ho To ðsi so Þ þ e0ch
X
Configuration C
kout Eout ¼ 0
_ 17 e17 k17 þ m
_ 19 e19 k19 þ We ke m
_ 7 e7 k7
_ 11 e11 k11 þ m
m
_ 12 e12 k12 m
_ 13 e13 k13 m
_ 14 e14 k14 ¼ 0
m
14,426
2915
22,052
4729
e
e
22,039
3069
29,837
The set of additional equations were included following
the considerations proposed by Lozano and Valero [59]. For the
exergy-based costs of the inputs (sugarcane) a unitary value is
assigned, and therefore
k11 ¼ 1
(10a)
In BIGCC, stack gases that leave the drier have the exergy-based
cost null
k6 ¼ 0
(11)
All the irreversibility generation in the turbines must be carried
out by the exergy-based cost of electric power, and consequently
the exergy-based costs of the steam entering and leaving these
turbines are considered equal. Therefore, Eq. (12) (Rankine Cycle)
and Eq. (13) (BIGCC) are obtained
k1 ¼ k5 ¼ k2 ¼ k17 ¼ k18
(12)
k3 ¼ k4 and k9 ¼ k17 and k18 ¼ k19
(13)
In the splitters, where no irreversibility generation takes place,
flows entering and leaving the valves have the same exergy-based
cost. Thus, Eqs. (14) and (15) are obtained, for the Rankine Cycle and
BIGCC, respectively.
k2 ¼ k3 ¼ k4 ; k4 ¼ k10 ¼ k15 ¼ k16
(14)
k8 ¼ k9 ¼ k18
(15)
(8)
where: “k” defines the exergy-based cost and “E” the total exergy
flow. The subscript “in” and “out” indicate the flows that enter and
leave the control volume, respectively.
The application of Eq. (8) in all control volumes forms a linear
equations set, where the number of variables is greater than the
number of equations. In order to obtain a set with a unique solution
it is necessary to add some additional equations. Cerqueira and
Nebra [61] reported in a simple way the postulates of the methodology to define these additional equations.
In the case of anhydrous ethanol production process, the exergetic cost balance equation is written by Eq. (9) (for the Rankine
Cycle) and Eq. (10) (BIGCC):
_ 3 e3 k3 þ m
_ 5 e5 k5 þ We ke m
_ 12 e12 k12
_ 11 e11 k11 þ m
m
_ 13 e13 k13 m
_ 14 e14 k14 ¼ 0
m
4711
e
ð9Þ
ð10Þ
In the BIGCC, the exergy-based cost of stack gases that enter and
leave the HRSG (streams 4 and 5, respectively), is the same. Thus, all
the irreversibility generated in the HRSG is carried by exergy-based
cost of steam produced by HRSG (stream 8)
k4 ¼ k5
(16)
In the anhydrous ethanol production process, the following
considerations were made:
Table 6
Thermal and electric energy demand of anhydrous ethanol production process.
Configuration 2.5 bar steam
consumption
(kg/s)
6.0 bar steam
consumption
(kg/s)
Electric energy
consumption
(kW)
Steam
Demand
(kg/TC)
A
B
C
D
9.22
4.43
6.12
8.29
0
4328
0
4328
468
310
322
226
55.31
38.33
38.26
22.94
Please cite this article in press as: Dias MOS, et al., Improving bioethanol production from sugarcane: evaluation of distillation, thermal integration and cogeneration systems, Energy (2010), doi:10.1016/j.energy.2010.09.024
M.O.S. Dias et al. / Energy xxx (2010) 1e13
11
Table 7
Results of different configurations of cogeneration systems (STBP: Rankine Cycle with back pressure steam turbine; STCOND: Rankine Cycle with condensing steam turbine;
BIGCC: Biomass Integrated Gasification Combined Cycle).
Cogeneration System
Process configuration
Bagasse surplus (%)
Electricity surplus (kWh/TC)
Exergy-based cost
Steam
Electricity
Ethanol
STBP
STBP
STBP
STBP
STCOND
STCOND
STCOND
STCOND
BIGCC
A
B
C
D
A
B
C
D
D
5.00
37.05
34.65
54.03
0
0
0
0
0
62.66
23.86
34.53
6.23
62.62
71.74
78.8
79.61
144.3
3.300
3.300
3.300
3.300
3.402
3.536
3.525
3.611
1.833
3.823
3.824
3.823
3.817
3.841
3.966
3.956
4.036
1.676
2.018
1.861
1.875
1.790
2.238
2.093
2.017
2.007
1.806
i) The exergy-based cost of the steam that enters the process is
the same of the condensate obtained after heat exchange:
For the Rankine Cycle:
k3 ¼ k5 ¼ k6
(17)
For the BIGCC cycle:
k17 ¼ k18 ¼ k7
(18)
ii) The exergy-based cost of bagasse and vinasse (streams 14 and
13, respectively) is the same as that of the sugarcane (stream
11) that enters the process. Thus, all the irreversibility
generation is carried by the exergy-based cost of the ethanol.
As a result:
k12 ¼ k14 ¼ k11
(19)
Thus, with the set equations above, the number of equations is
equal to the number of variables. The system was solved using the
EESÒ software.
6. Results and discussions
6.1. Process utilities requirements
After integration of the background process and the distribution
of the vapor bleeding, the hot utilities requirements could be
determined for each case. Four different configurations for the
anhydrous ethanol production process steam demand were
analyzed, considering conventional and double-effect distillation
with and without thermal integration. Two levels of steam pressure
were considered for hot utilities, derived from extractions at 6.0 bar
and 2.5 bar of pressure in the steam turbine. Steam consumption of
the analyzed cases is shown in Table 5, which summarizes the final
results of hot utilities demand for each case.
As can be seen in Table 5, the thermal integration promoted
important steam reductions in the anhydrous bioethanol production. For the mills with conventional distillation system scheme the
reduction of the steam consumption at 2.5 bar can reach 31% with
the integration and 34% for the 6.0 bar steam.
When double-effect distillation columns are used in the mills
the reduction of steam consumption with the thermal integration is
40% for the 2.5 bar steam and 27% for the 6.0 bar steam.
6.2. Evaluation of the cogeneration systems
Information regarding steam consumption, gathered from Table
5, and electric energy demand from the anhydrous ethanol process
are shown in Table 6. Exergetic costs and summarized results of the
evaluation are showed in Table 7.
The exergetic cost analysis in the Rankine Cycle with back
pressure steam turbine shows that the decrease of steam demand,
through the use of thermal integration and double-effect distillation (configuration D), leads to an increase of bagasse surplus;
however, lower electricity surplus is produced, mainly due to the
increase on electricity consumption of compressors in the doubleeffect distillation system as well as to decreased steam consumption. In fact, this configuration is the one that makes the best use of
bagasse for other applications, such as bagasse hydrolysis for
second generation bioethanol production: the exergy-based cost of
process steam and electricity remains constant, while exergy-based
cost of ethanol decreases due to the lower steam demand of the
process.
The case on which the Rankine Cycle is employed along with
condensing steam turbines shows higher surplus electricity than
the previous system. The possibility to use all the bagasse available
to generate electricity allows the production of an electricity
surplus of 79.61 kWh/TC, which corresponds to an increase of about
27%. However, the exergy-based costs are different than those of
the conventional cogeneration system used in Brazil: for electricity,
they are 5% higher, while for ethanol they are 0.5% lower.
The exergetic cost analysis of the BIGCC cycle was performed
only on configuration D, since this configuration has the lowest
steam demand (226 kg/TC), a value lower than the minimal
(230 kg/TC) foreseen by Modesto et al. [52]. The main difference of
the BIGCC cycle is the expressive electricity surplus produced in this
configuration. The value of 144.3 kWh/TC is 130% higher than
traditional Rankine Cycle with condensing steam turbines. In
addition, the exergy-based cost of electricity and ethanol is lower
(56 and 19%, respectively) than those of the traditional configuration. Thus, the use of the BIGCC cycle leads to a higher electricity
surplus with low production exergy-based costs of electricity and
ethanol. Composition of costs of ethanol is found in Modesto et al.
[62]. All results are shown in Table 7.
7. Conclusions
A study of the bioethanol process modeling and thermal integration was described in detail in this paper. The use of the
commercial software UniSim Design as a modeling tool proved to
be useful for analysis of the energy integration opportunities,
particularly considering different distillation systems.
The thermal integration of the rectification and distillation
columns promoted important energy savings in the bioethanol
process as a whole. The analysis of the integrated plant with
a double-effect distillation system showed that this alternative
presents a lower process steam demand, when compared with
traditional distillation schemes. On the other hand, it requires
Please cite this article in press as: Dias MOS, et al., Improving bioethanol production from sugarcane: evaluation of distillation, thermal integration and cogeneration systems, Energy (2010), doi:10.1016/j.energy.2010.09.024
12
M.O.S. Dias et al. / Energy xxx (2010) 1e13
a more complex control system and higher investments, but this
analysis was not included in this paper. Another important issue
regarding this scheme is the high electric energy consumption of
the ethanol vapor re-compression step.
A comparison of different cogeneration systems for production
of steam and electricity in anhydrous ethanol production plants in
Brazil was also carried out. Three configurations of cogeneration
systems were assessed: traditional Rankine Cycles with back
pressure and condensing steam turbines and the BIGCC.
It was shown that the use of thermal integration techniques
allows a decrease of process steam demand and consequently
produces more surplus bagasse (in the case of back pressure steam
turbine) or increases electricity surplus (for the case where
a condensing steam turbine is considered), in both cases using the
traditional technology for cogeneration systems (Rankine Cycle).
The main differential of this study is that the feasibility of the
BIGCC in anhydrous ethanol production plants was investigated. A
previous study [52] showed that the BIGCC is feasible if the process
steam demand reaches about 230 kg/TC (the traditional value in
Brazilian plants is in the range of 380e450 kg/TC). Thermal integration and the use of a double-effect distillation system lead to
a significant reduction of steam demand, reaching values below
230 kg/TC, thus indicating that the BIGCC cycle is feasible and may
provide a considerable increase on electricity surplus of the ethanol
plant.
This work assesses the feasibility of the use of BIGCC in anhydrous ethanol production process. It is true that gasification and gas
turbine technologies developed specifically to use biomass gas are
not yet available commercially. However, this study shows that
from a thermal and exergetic analysis point of view, the use of this
technology allows an expressive increase of electricity generation
from sugarcane biomass.
Acknowledgments
The authors would like to thank CNPq, FAPESP and FINEP for
financial support, Dr. Juan Harold Sosa-Arnao for valuable information about modeling of the bagasse dryer and the reviewers for
their comments.
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