Energy Conservation in Ethanol-Water Distillation Column with Vapour Recompression Heat Pump

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DOI: 10.5772/37310 · Source: InTech
In book: Distillation - Advances from Modeling to Applications

Figures

2
Energy Conservation in
Ethanol-Water Distillation Column
with Vapour Recompression Heat Pump
Christopher Enweremadu
University of South Africa, Florida Campus
South Africa
1. Introduction
Ethanol or ethyl alcohol CH
3
CH
2
OH, a colorless liquid with characteristic odor and taste;
commonly called grain alcohol has been described as one of the most exotic synthetic
oxygen-containing organic chemicals because of its unique combination of properties as a
solvent, a germicide, a beverage, an antifreeze, a fuel, a depressant, and especially because
of its versatility as a chemical intermediate for other organic chemicals. Ethanol could be
derived from any material containing simple or complex sugars. The sugar-containing
material is fermented after which the liquid mixture of ethanol and water is separated into
their components using distillation.
Distillation is the most widely used separation operation in chemical and petrochemical
industries accounting for around 25-40% of the energy usage. One disadvantage of
distillation process is the large energy requirement. Distillation consumes a great deal of
energy for providing heat to change liquid to vapour and condense the vapour back to
liquid at the condenser. Distillation is carried out in distillation columns which are used for
about 95% of liquid separations and the energy use from this process accounts for an
estimated 3% of the world energy consumption (Hewitt et al, 1999). It has been estimated
that the energy use in distillation is in excess. With rising energy awareness and growing
environmental concerns there is a need to reduce the energy use in industry. The potential
for energy savings therefore exists and design and operation of energy efficient distillation
systems will have a substantial effect on the overall plant energy consumption and
operating costs.
The economic competitiveness of ethanol has been heightened by concerns over prices and
availability of crude oil as well as greenhouse gas emissions which have stimulated interest
in alternatives to crude oil to provide for automotive power and also by the use of
bioethanol in the production of hydrogen for fuel cells. Therefore, there is the need to
explore ways of producing ethanol at competitive costs by the use of energy efficient
processes. To cope with the high energy demand and improve the benefits from the process,
the concept of polygeneration and hydrothermal treatment especially when dealing with
small scale ethanol plants is fast gaining interest. However, the analysis of the bioethanol
process shows that distillation is still the most widely used.
Distillation – Advances from Modeling to Applications
36
Over the years, there have been many searches for lower energy alternatives or improved
efficiencies in distillation columns. One such search led to the use of heat pumps, the idea
which was introduced in the 1950s. Also, Jorapur and Rajvanshi (1991) have used solar
energy for alcohol distillation and concluded that it was not economically viable. Heat
pumping, however, has been known as an economical energy integration technology for
reduction in consumption of primary energy and to minimize negative impact of large
cooling and heating demands to the environment. One of the heat pump cycles which have
been widely studied is the recompression of the vapours where the reboiler is heated by
adding a compressor to the column to recover some of the heat lost in the distillate.
Most studies have concluded that heat pumping is an effective means of saving energy and
reducing column size without estimating the actual energy consumption and the parameters
that are likely to have significant effect on energy consumption. Estimating the actual
energy consumption is an important aspect towards the determination of the viability of the
system in ethanol–water separation.
The purpose of this chapter was to study how previously neglected and/or assumed values
of different parameters (the pressure increase across the compressor was ignored, column
heat loss was assumed to be 10% of the reboiler heat transfer rate, and the overall heat
transfer coefcient was determined without considering it as an explicit function of
dimensionless numbers, and its dependence on uid viscosity and thermal conductivity
neglected) affect the process efficiency, energy consumption and the column size of a
vapour recompression heat pump.
2. Energy requirements in ethanol distillation
Ethanol distillation, like any other distillation process requires a high amount of thermal
energy. Studies carried out by several authors reveal that the distillation process in ethanol
distilleries consumes more than half of the total energy used at the distillery (Pfeffer et al.
2007). It has been estimated that distillation takes up about 70-85% of total energy consumed
in ethanol production. Pfeffer et al (2007) estimated that distillation consumes half of the
total production energy 5.6 MJ/Liter out of 11.1 – 12.5 MJ/Liter.
The energy requirements for ethanol production have improved markedly during the past
decade due to a variety of technology and plant design improvements. The energy needed
to produce a liter of ethanol has decreased nearly 50% over the past decade and that trend is
likely to continue as process technology improves ( Braisher et al, 2006).
3. Energy conservation schemes in distillation column
Distillation columns are usually among the major energy-consuming units in the food,
chemical, petrochemical and refining industries. According to Danziger (1979), the most
effective method of economizing energy in a distillation column is energy recovery of which
direct vapour recompression has been regarded as the best solution.
3.1 Heat pumping distillation systems
Basically, the heat pump can be regarded simply as reverse heat engine. The heat pump
requires either work input or external driving thermal energy to remove the heat from a low
temperature source and transform it to a higher level.
Energy Conservation in Ethanol-Water
Distillation Column with Vapour Recompression Heat Pump
37
The conventional heat pumps are electrically driven vapour recompression types, which
work on the principle that a liquid boils at a higher temperature if its pressure is increased.
A low-pressure liquid passes into the evaporator, where it takes in heat causing the liquid to
boil at low temperature. The low-pressure vapour is passed to the compressor where it is
compressed by the application of work to a higher pressure. The resulting high pressure
vapour flows to the condenser where it condenses, giving up its latent heat at a high
temperature, before expanding back to a low pressure liquid.
The heat pump cycle may be connected to a distillation column in three ways (Fonyo and
Benko, 1998) . The simplest alteration is to replace steam and cooling water with refrigerant
(closed system). The other two types of heat pump system apply column fluids as
refrigerant . When the distillate is a good refrigerant the vapour recompression can be used.
If the bottom product is a good refrigerant the bottom flashing can be applied.
In this work, the direct vapour recompression system is studied due to its good economic
figures ( Emtir et al, 2003). Also the vapour recompression is the most suitable as the boiling
points of both key components (ethanol and water) are close to each other (Danziger, 1979)
and the appropriate heat transfer medium (ethanol vapour) is available.
3.2 Use of vapour recompression in distillation columns
Vapour recompression system has been extensively studied since 1973, the year of drastic
rise in energy (Null, 1976). The vapour recompression system is accomplished by using
compressor to raise the energy level of vapour that is condensed in reboiler–condenser by
exchange of heat with the bottoms. The condensate distillate is passed into reflux drum
while the bottom product is vaporised into the column.
Vapour recompression consists of taking the overhead vapour of a column, condensing the
vapour to liquid, and using the heat liberated by the condensation to reboil the bottoms liquid
from the same column. The temperature driving force needed to force heat to flow from the
cooler overhead vapours to the hotter bottoms product liquid is set up by either compressing
the overhead vapour so that it condenses at a higher temperature, or lowering the pressure on
the reboiler liquid so it boils at a lower temperature, then compressing the bottoms vapour
back to the column pressure. While conventional column has a separate condenser and
reboiler, each with its own heat transfer fluid such as cooling water and steam, the vapour
recompression column has a combined condenser–reboiler, with external heat transfer fluids.
The advantage of vapour recompression lies in its ability to move large quantities of heat
between the condenser and reboiler of the column with a small work input. This results
from cases where there is only a small difference between the overhead and bottoms
temperature. Also, the temperature, and therefore the pressure, at any point may be set
where desired to achieve maximum separation. This effect is of particular importance where
changing the pressure affects the relative volatility. By operating at more favourable
conditions, the reflux requirement can be reduced and therefore the heat duties. These
advantages can reduce a large amount of energy.
4. Ethanol-water vapour recompression distillation column
Figure 1 shows a schematic illustration of the distillation column with direct vapour
recompression heat pump. An ethanol-water solution in a feed storage tank (FST) at
Distillation – Advances from Modeling to Applications
38
Fig. 1. Schematic Diagram of Column with Direct Vapour Recompression Heat Pump
ambient conditions, is preheated with bottom product and condensate in heat exchangers,
preheaters PH1 and PH2, and fed to the column. An auxiliary reboiler (AR) is used to start
the unit. This reboiler supplies the auxiliary heat duty, which is the heat of vaporization
because the main reboiler can work only if there is some compressed vapour already
available. The overhead vapours from the top are compressed in the compressor (CP) up to
the necessary pressure in such a way that its condensing temperature is greater than the
boiling temperature of the column bottom product. The vapour is then condensed by
exchanging heat within the tubes of the reboiler-condenser (RC). In a condenser, the inlet
temperature is equal to the outlet temperature. Ethanol vapour will only lose its latent heat
of condensation. At the same time, the cold fluid (ethanol-water mixture) in the reboiler will
absorb this latent heat and its temperature will increase to boil up the mixture to
temperature T
CEV.
The liberated latent heat of condensation provides the boil-up rate to the
column while the excess heat extracted from the condensate is exchanged with the feed in
preheater PH2. The condensate, which is cooled in the cooler (CL) up to its bubble point at
the column operating pressure, expands through the throttling valve (TV) at the same
pressure and reaches the flash tank (FT). After expansion, the output phases are a vapour
phase in equilibrium with a liquid phase. One part of the product in the liquid phase is
removed as distillate and stored in the tank (DST), while the remainder is recycled into the
column as reflux L1. The excess of vapour is recycled to the compressor.
4.1 Methodology
Like this work, nearly all publications in this field are based on modelling and simulation
(Brousse et al., 1985; Ferre et al., 1985; Collura and Luyben, 1988; Muhrer et al, 1990; Oliveira
Energy Conservation in Ethanol-Water
Distillation Column with Vapour Recompression Heat Pump
39
et al. 2001). The mathematical modeling of the distillation system is derived by applying
energy, composition and overall material balances together with vapour-liquid equilibrium
under some assumptions (see Muhrer et al, 1990 and Enweremadu, 2007). These and other
assumptions are aimed at simplifying the otherwise cumbersome heat-and mass-transfer,
and the fluid flow equations Mori et al (2002).
4.2 Calculation of the distillation column
In this system, there is a direct coupling between the distillation column and the rest of the
system, as the heat pump working fluid is the column’s own fluid which, is a binary mixture
of ethanol and water at composition X
D
. Therefore, the set of equations are not solved
separately as in distillation column assisted by an external heat pump.
The detailed calculation of the overall material and component material balance such as the
bottom ow rate, B and distillate ow rate, D; reux ratio, R
r
; the molar vapour ow rate
which leaves the column top and feeds the condenser, V
1
; feed vapour ow rate, V
F
; feed
vapour fraction, q; vapour molar ow rate remaining at the bottom of the column, L
2
are
given (see Enweremadu, 2007).
The overall (global) energy balance equation applied to a control volume comprising the
distillation column and the feed pre-heaters provides the total energy demand in the reboiler:
Q
reb
= Dh
D
+ Bh
B
+ L
1
h
LV, e
+ Q
losses
– Fh
F
– Q
1
– Q
2
(1)
where Q
reb
is the total heat load added to the reboiler, Q
losses
represents the heat losses in the
column, which are to be determined; Q
1
and Q
2
are the heat loads of the pre-heaters; h
LV,e
is
latent heat of vaporisation downstream of throttling valve; h
D
, is the enthalpy of the
distillate; h
B
is the enthalpy of the bottom product; h
F
, is the enthalpy of the feed. The details
of the mass balance variables are determined in Enweremadu (2007).
The first step in the design of a distillation column is the determination of the number of
theoretical plates required for the given separation. The theoretical trays are numbered from
the top down, and subscripts generally indicate the tray from which a stream originates
with n and m standing for rectifying and stripping sections respectively. The design
procedure for a tray distillation column consists of determining the liquid and vapour
composition or fraction from top to bottom, along the column. In calculating the
composition profile of the column two equations relating liquid mole fraction to
temperature and vapour mole fraction to the liquid fraction are used. The compositions at
the top (X
D
) and bottom (X
B
) of the column are previously pre-established data. In this work,
the minimum number of theoretical stages (N
min
) is calculated using Fenske’s equation:
min
1
log .
1
log
DB
DB
XX
XX
N



(2)
where α is the relative volatility in the column. The actual number of plates is given by:
min
T
N
N
(3)
Distillation – Advances from Modeling to Applications
40
where
T
is the tray efficiency.
4.2.1 Heat losses from distillation column
The heat loss from the distillation column is the main factor that affects heat added and
removed at the reboiler and condenser respectively. Most distillation columns operate above
ambient temperature, and heat losses along the column are inevitable since insulating
materials have a finite thermal conductivity. Heat loss along the distillation column increase
condensation and reduces evaporation. Thus, the amount of vapour diminishes in the
upward part of the column, where the flow of liquid is also less than at the bottom.
To prevent loss of heat, the distillation column should be well insulated. Insulation of
columns using vapour recompression varies with the situation. Where the column is hot and
extra reboiler duty is used, the column should be insulated (Sloley, 2001). The imperfect
insulation of the column causes some heat output.
In determining the heat loss from the distillation column, it is assumed that the temperature
is uniform in the space between two plates. The heat transfer between the column wall and
the surrounding is then determined from the well-known relationship for overall heat
transfer coefficient:
Losses o
PP
QUAT
(4)
where U
p
, the overall heat transfer coefficient is given by Gani, Ruiz and Cameron (1986), as
1
,, , , , ,p
p
oi o mins
UfhhKAAAt (5)
where the temperature difference,
p
T
, is given as
p
amb
p
TTT

h
o
, the heat transfer coefficient between the surroundings and the column external surface,
is given as
h
o
= f(Nu, K
ins
, d
o
, t
ins
) (6)
h
i
is the heat transfer coefficient inside the column; K
p
is the thermal conductivity of the tray
material; A
o
is the external area of heat exchange; A
i
is the internal area of heat exchange;
A
m
is the logarithmic mean area; t
ins
is the thickness of insulation.
The heat output is calculated with the general expression for convection around cylindrical
objects.

ln / ln /
11
wall wall ins wall
amb
P
loss
oi o
wall wall
ii oo
ins m
TT
Q
rr rr
hA hA
KA KA


(7)
The column inner surface heat transfer resistance is neglected as the heat transfer coefficient
for condensing vapor is large and therefore will have little effect on the overall heat transfer.
Based on the assumptions in Enweremadu (2007), the heat transfer due to free convection
between the surroundings and the external column wall and due to conduction through the
insulation materials is predicted.
Energy Conservation in Ethanol-Water
Distillation Column with Vapour Recompression Heat Pump
41
Also, from geometry of the insulated cylinder (Fig.2), the external diameter of insulation is
given as
d
ins
= d
o
+ 2t
ins
(8)
Details of how the logarithmic mean diameter of the insulating layer (d
ins,m
), external area of
heat exchange (A
o
) and the logarithmic mean area (A
m
) can be found in Enweremadu (2007).
From dimensional analysis,
2
ins
o
oins
KNu
h
dt
(9)
where, t
ins
is the thickness of insulation; K
ins
– thermal conductivity of the insulation
materials; Nu – Nusselt number; d
o
– external diameter of column; T
amb
– temperature of the
surrounding; T
p
– plate temperature.
Fig. 2. Hypotethical Section of the Distillation Column with Insulation
For vertical cylinders, the commonly used correlations for free convection are adapted from
Rajput (2002) as:
For laminar flow,

1/4
Nu 0.59 Gr.Pr for (10
4
<Gr.Pr<10
9
) (10)
For turbulent flow,

1/3
0.10 Gr.PrNu for (10
9
<Gr.Pr<10
12
) (11)
where Gr is the Grashof number and Pr is Prandtl.
Based on the assumptions of neglecting h
i
, A
i
and the effect of thermal resistance, equation
(5) reduces to:
r
i
r
ow
r
ins
hot
fluid
Cold fluid
(air)
Q
Lo
h
o
T
i
T
ins
T
o
T
surf
K
ins
K
wal
T
p
h
i
Distillation – Advances from Modeling to Applications
42
U
P
= f(h
o
, K
p
, A
o
, A
m
, t
ins
) (12)
while equation (7) is given as

()
ln /
1
ins wall
pamb s
losses
o
p
moo
TT PN
Q
rr
KA hA
(13)
where

1
ln
1
wall
P
o
ins
p
moo
U
rr
KA hA
The heat loss from the column trays is given by


loss from trays
,.
()2
ln
1
2.
()
2
2
ln 1
amb s
P
o
ins
sins ins
pinsm
s
ins
oins
o
TT PN
Q
rr
Pt K Nu
KdP
t
dt
d

(14)
The total heat loss from the column is expressed as
loss from trays
lossQ Q Heat loss from the two c
y
linder heads
(15)
Based on the assumptions made, heat loss through the cylinder heads is given by
2
pamb
o
loss at cylinder heads
ins
p
2(T - T ) r
Q
t1
K
oh
(16)
Therefore,


,.
()2
ln
1
2.
()
2
2
ln 1
amb S
P
loss
o
ins
sins ins
ins m
Ps
ins
oins
o
TT PN
Q
rr
Pt K Nu
KdP
t
dt
d

2
Pamb
ins
p
2(T - T )
t1
K
o
o
r
h
(17)
4.3 Calculation of heat pump and compressor parameters
The heat pump is thermodynamically linked to the column through the heat load from the
pump to the column Q
HPC
and from the column to the pump Q
CHP
, and reboiler–condenser
temperature. These parameters provide the basis for the heat pump calculation.
The calculation of the heat pump parameters begins with the estimation of the working fluid
condensation temperature obtained from the reboiler temperature and temperature drop
across the heat exchangers.
T
CHP
= T
CEV
+ T
CHP
(18)
Energy Conservation in Ethanol-Water
Distillation Column with Vapour Recompression Heat Pump
43
where T
CEV
is the column vapourization (reboiler) temperature and ΔT
CHP
, a pre-established
mean temperature difference across the heat exchangers (temperature drop in reboiler-
condenser). Next is the estimation of the relevant thermodynamic properties of the working
fluid. These are obtained from thermodynamic correlations.
The thermodynamic properties are determined as functions of temperature. The
relationships used for calculating the working fluid density, viscosity, thermal
conductivity and heat capacity for input at various locations are presented in
Enweremadu (2007). The condensation pressure, P
CHP
is expressed as a function of
condensation temperature as
P
CHP
= f(T
CHP
) (19)
while the condensation pressure is determined from ideal gas equation.
The latent heat of condensation from column to heat pump is numerically exactly equal to
the latent heat of vaporisation, but has the opposite sign: latent heat of vaporisation is
always positive (heat is absorbed by the substance), whereas latent heat of condensation is
always negative (heat is released by the substance). Latent heat of condensation is expressed
as a function of condensation temperature and is determined from the relationship
(Ackland, 1990):
CHP
CHP
C
C
LV,CHP
bp
C
vap
T
AB1-
T
T
1-
T
hH
T
1-
T








(20)
where
vapH is the heat of vapourisation at the boiling point of ethanol; T
C
is the critical
temperature;
A and B are constants.
The vapour specific volume at location “
a” (entrance to the compressor) is expressed as a
function of column condensation temperature, T
CC
and pressure at the top of the column,
P
TOP
:
ccT R
a
TOP
x
v
P
(21)
Since compression is polytropic, at location
“b” (compressor discharge), the vapour specific
volume is determined by:
b
1
v
TOP
a
CHP
n
P
PP
v




(22)
where n is the polytropic index and ΔP is the pressure increase across the compressor.
The vapour specific enthalpy at “a” is a function of top pressure,P
TOP
and top temperature,
T
TOP
.
h
a
=f(T
TOP,
P
TOP
) (23)
Distillation – Advances from Modeling to Applications
44
The vapour specific enthalpy at “b” may be determined as a function of compressor
discharge temperature T
b
and condensation pressure P
CHP
but in this study, it is determined
by the development of numerical computation with calculations utilizing the Redlich –
Kwong equation of state. The Redlich – Kwong equation of state is given as

1
2
)
RT a
P
Vb
TVV b

(24)
Where
a = 0.42747
5
2
2
c
c
RT
P




b = 0.08664
c
c
RT
P



and P = pressure (atm); V = molar volume (liters/g-mol); T = temperature (K); R = gas
constant (atm. Liter/g-mol.K); P
c
= critical pressure (atm).
Taking the reference state for the enthalpy of liquid ethanol
o
L
h
, temperature, T
o
and the
enthalpy of vaporisation Δ
o
va
p
H , then the enthalpy of ethanol vapour as an ideal gas at
temperature T can be calculated from
oo
b
hh
o
T
oo
vap
L
T
HCdT
p

(25)
Using the isothermal enthalpy departure and the Redlich-Kwong equation of state, the
enthalpy of ethanol vapour at T and P can be calculated from
1.5
1.5
1ln1
o
T
oo o
b
Lvap
T
ab
hh H CdTRTZ
p
V
bRT




(26)
where Z is the compressibility factor, Cp
o
is the molar specific heat capacities of gases at
zero pressure given as a polynomial in temperature.
Equations 24–26 are then solved with POLYMATH
(R)
Simultaneous Algebraic Equation
Solver (See Enweremadu, 2007).
The vapour specific heat at location “e” is calculated thus
(Oliveira et al, 2002):
e,h LLc SChCpT
(27)
The specific liquid enthalpies have been assumed to be simple functions of temperature.
The liquid specific enthalpy at location “c” is determined from EZChemDB Thermodynamic
Properties Table for Ethanol (AM Cola LLC, 2005) using the expression
L,c CHPh f(T )
(28)
Energy Conservation in Ethanol-Water
Distillation Column with Vapour Recompression Heat Pump
45
while the liquid specific enthalpy at location e (at the exit of the throttling valve) is
determined from
h
L,e
=f(T
CC
) (29)
The temperature at compressor discharge is determined from the knowledge of the
compressor efficiency. The ideal discharge temperature (the temperature that gives an
overall change in entropy equal to zero) is calculated before correcting with the compressor
efficiency.
CHP
TOP
TOP
bTOP
0.263
PP
T-1
P
TT
pol




 (30)
The dryness fraction after the isenthalpic expansion is given by
,
,
L
LV
ee
e
e
hh
h
(31)
where the molar latent heat of vaporization at location “e” is adapted from Ackland (1990):




, *1 1 1LV
bp
ed d
vap C bp C C
hH TT TTABTT
 (32)
Since this is a throttling process, T
d
= T
e
and h
d
= h
e
The molar vapour flow rate which is recycled in the flash tank and conveyed to the
compressor is calculated by
1
1
e
R
e
V
V
(33)
Therefore, the molar flow rate across the compressor is expressed as
1 R
M
VV
(34)
While the dryness fraction at condenser exit is determined by
,
23
L
LV CHP
PCHP d
c
QMCT T
h

(35)
where Q
23
is the distribution of excess heat between the pre-heater Q
2
and the cooler Q
3
and
Cp
L
is the molar specific heat of the working fluid in the liquid phase.
The energy balance, applied to the heat pump working fluid, yields the available energy for
exchange at the condenser, as follows:
,
1v
cd b CHP c LV CHP
QMCpTT h

(36)
Distillation – Advances from Modeling to Applications
46
A comparison is made between this energy available at the condenser, Q
cd
, with the
energy required by the column reboiler, Q
reb
. This brings about the following heat load
control.
i.
If the rate of energy available at the heat pump condenser, Q
cd
, is greater than the rate
of energy required by the reboiler Q
reb
, then the condenser gives up Q
reb
to the reboiler
and the remaining energy is conveyed to the preheaters (Q
2
) and cooler (Q
3
)
if
cd rebQQ then HPC rebQQ
(37)
ii.
But if Q
cd
is smaller than or equal to Q
reb
, then all energy available is transferred to the
reboiler and the auxiliary reboiler will provide the “extra” Q
reb
i.e.
if
cd rebQQ
then HPC cdQQ
(38)
where Q
HPC
is the energy yield by the heat pump to the distillation column. The factor by
which the heat pump contributes to the heat load of the reboiler is given as
HPC
reb
Q
f
Q
(39)
For a distillation column with vapour recompression, driving the compressor uses the most
energy. Thus, the power consumption must be known so as to assess the feasibility of such a
system. For a perfect gas, that is, a gas having a constant specific heat, Cp = Cp
o
, then the
specific enthalpy rise between the compressor inlet and outlet is
o
ba ba
hh h CT T
p
 (40)
And if the change of state is isentropic,
1
1
1
b
b
a
a
P
Ru T
hvdp
MP




(41)
In reality, ideal gases do not exist and therefore improvements are made on equation (41).
Therefore, compression is polytropic and the isentropic index γ, is replaced by the
polytropic index, n (see Enweremadu, 2007). The compressor polytropic efficiency
p
ol
= 0.7 - 0.8 is used.
Also, because a saturated vapour, especially at higher pressures, shows deviations from the
ideal gas behaviour, the compressibility factor, Z is used. Hence equation (41) becomes
1
1
1
n
n
ua b
eff
pol a
ZR T P
n
h
nMP






(42)
Therefore, the power input for driving the compressor is the energy that increase the
enthalpy of the gas
Energy Conservation in Ethanol-Water
Distillation Column with Vapour Recompression Heat Pump
47
1
1
1
n
n
b
cp a a
pol a
P
Mn
WP
nP



(43)
Equation (43) shows that the pressure ratio
b
a
P
P
is crucial to the power requirement. This
ratio or the pressure increase to be provided by the compressor of a column with vapour
recompression is influenced by the following (Meili, 1990; Han et al, 2003):
Pressure drop in vapour ducts (pipes) and over valves and fittings, P
p
.
Pressure drop across the column, P
cl
.
The difference in boiling points between the top and bottom products, P
b
.
Temperature difference in the reboiler, P
CHP
.
4.3.1 Determination of the pressure increase over the compressor
Pressure drops in the vapour ducts may be caused by frictional loss, P
f
; static pressure
difference, due to the density and elevation of the fluid, P
s
; and changes in the kinetic
energy, P
k
. Since, there are elbows, valves and other fittings along the pipes then the
pressure drop is calculated with resistance coefficients specifically for the elements.
Therefore, the pressure drop along a circular pipe with valves and fittings is given by
PsfK
PPPP
  
2
1
2
p
P
ul
d





(44)
and u is the fluid velocity; d
p
is the pipe diameter and ρ is the fluid density; is the Fanning
friction factor which is a function of Reynolds number; l
p
is the pipe length; is the dynamic
viscosity of the fluid and ξ is the resistance coefficient.
The pressure drop over the entire distillation column, P
cl
is caused by losses due to vapour
flowing through the connecting pipes and through pressure drop over the stages in
rectifying and stripping section. This depends mainly on the column internals, number of
stages, gas load and operating conditions. ΔP
cl
=0, if zero vapour boil up is assumed. But
constant pressure drop is assumed in this work. The pressure drop over a stage consists of
dry and wet pressure drop. The dry pressure is caused by vapour passing through the
perforation of the sieve tray. The aerated liquid (static head) on the tray causes the wet
pressure drop. Constant pressure drop per tray have been estimated from several authors to
be equal to 5.3mmHg per tray (Muhrer, Collura and Luyben
1990). The total column
pressure drop has been found by summing plate pressure drops ΔP
cl
0.13332 5.3 0.707
clPxxNN
 (45)
The top and bottom products have different compositions and boiling points. For a fixed
bottom temperature of the column, there is a vapour – pressure difference, P
b
due to the
difference in boiling points.
bP TOP BOTTOMPP
(46)
Distillation – Advances from Modeling to Applications
48
where,
TOP
TOP
TOP
TOP TOP
-
B
A
P
10
TC
(47)
BOTTOM
BOTTOM
BOTTOM
BOTTOM BOTTOM
-
B
A
P
10
TC
(48)
The temperature difference in the reboiler- condenser
is expressed by means of the vapour –
pressure equation as a pressure difference, P
CHP
. Temperature differences of 8 – 17
o
C are
quite common
for ethanol-water distillation (Gopichand et al, 1988; Canales and Marquez,
1992). Using the Clausius – Clapeyron equation for a two- point fit,
-
e
R.
CHP
CHP
CHP CEV
Hvap
T
P
TT





(49)
Therefore the total pressure increase over the compressor becomes
bclCHPpP P P P P
   
(50)
For this distillation system, the compression (pressure) ratio is
CHP
TOP
b
a
P
PP
PP
(51)
where P
TOP
is inlet pressure (vapour pressure at top temperature).
Other compressor parameters are calculated by the following equations:
i.
Compressor power input is determined from equation (43) and (51)
1
1
1
n
n
CHP
TOP
TOP
cp a
pol
Mn P P
WP
nP





(52)
ii.
Compressor heat load rate (energy balance)
pol
c
p
c
p
ba
QWMhh

(53)
iii.
Compressor volumetric efficiency
1
11
m
CHP
vv
TOP
PP
Cr
P






(54)
where Cv is empirical volumetric coefficient and r is the compressor clearance ratio.
iv.
Compressor nominal capacity or compressor displacement rate
c
v
a
M
V
(55)
Energy Conservation in Ethanol-Water
Distillation Column with Vapour Recompression Heat Pump
49
where V
c
is compressor displacement volume (m
3
) and ω is angular velocity (rad s
–1
).
4.3.2 Determination of the reboiler-condenser parameters
The overall heat transfer coefficient between condenser and reboiler is given by

HPC
HPC
CHP
Q
UA
T
(56)
However, a careful analysis reveals that the overall heat transfer coefficient U is an explicit
function of Prandtl, Reynolds and Nusselt numbers, and depends on other properties such
as viscosity and thermal conductivity. The overall heat transfer coefficient referenced to
inner surface is given by
11 1 ln(/)
(/)
U2
ioi
io
iowall
rrr
rr
hhK

(57)
As thermal resistance of the wall is negligible, (K
wall
is large and ln(r
o
/r
i
)) 0, it is then
compared with the inner tube diameter (r
i
/r
o
1)
Then
11 1
ex ex
ioUh h
 (58)
0.8 0.4 0.14
0.023
(Re ) (Pr ) ( )
ex
ex
mwall
omm
om
K
h
d
(59)
where μ
m
is the mean bulk fluid viscosity and μ
w all
is the viscosity of the liquid at the wall.
The expression for condensation at low velocities inside tubes is adapted from
(Holman,
2005).
0.25
3'
()
0.555
()
ll v
ex
CHP
ex
L
fg
i
Li wall
Kgh
h
dT T

(60)
where
fg fg p CHP
L
'
wall
h h 0.375C (T -T )
where K
L
is thermal conductivity of the liquid, di
ex
is the inside diameter of the reboiler-
condenser tubes and μ
L
is the density of the condensate (liquid).
Therefore, the overall heat transfer coefficient may be determined from
0.8 0.14
3'
0.4
11 1
()
0.023
0.555
()
HPC
LL vL fg
mm pm wall
ex
LCHP
mmm
ex
o
iwall
o
ex
UA
Kgh
Kud C
dT T
dK










(61)
Assuming adiabatic expansion at the throttling valve, then
Distillation – Advances from Modeling to Applications
50
h
e
= h
d
= h
L,c
– Cp
L
.ΔT
SC
(62)
From the condenser prescribed degree of sub-cooling, the temperature of the working fluid
after cooling and before throttling is given by
T
d
= T
CHP
T
SC
(63)
where
T
SC
is the degree of sub-cooling (K)
The corresponding latent heat (enthalpy) is given as
,
L
dLc SC
hh C T
P

(64)
4.4 Analysis of distribution of excess heat rate
The distillation system uses the column’s working fluid as refrigerant and does not execute
a closed cycle. Therefore the excess heat which may occur is not assessed by an overall
energy balance but by the method of
Oliveira et al (2001). When the energy available at the
condenser Q
cd
, is greater than the energy required by the reboiler Q
reb
, the column receives
the amount Q
cd
and the energy left over corresponds to the excess. But if Q
cd
is smaller than
or equal to Q
reb
, then all the energy available is transferred to the reboiler, i.e. there will be
no excess. Thus,
if
cd reb
QQ then
23 cd reb
QQQ
(65)
if
cd reb
QQ
then
23
0Q
where Q
23
is the excess heat due to energy interactions between the heat pump and the
reboiler.
The distribution of the excess heat rate, Q
23
, between the pre-heater (Q
2
) and cooler (Q
3
) is
accomplished by controlling the feed condition pre-heated by Q
2
. In other words, the value
of Q
2
should be such that the feed reaches a prescribed condition. The pre-heating of the
feed is carried out by Q
1
(heat exchanged between the bottom product and the feed) and Q
2
(heat exchange between the heat pump working fluid and the feed), in the heat exchangers.
The heat provided by the bottom product is determined as follows:
1
.
p
CEV BE
B
BC T T
Q
 (66)
where, T
CEV
is the column evaporation temperature. Cp
B
is the specific heat of the bottom
product. T
BE
is the temperature at the bottom product flow exit.
The best feed condition is that of saturated liquid (Halvorsen, 2001). The energy required to
pre-heat the feed to reach saturated condition is expressed as
,.( )FSL PF F FsatQFCT T
(67)
where, Tsat,
F
and T
F
are the saturation temperature of the feed source and the temperature
of the feed source respectively. Cp
F
is the specific heat of the feed source.
To make the feed a saturated vapour, the energy required is given as
Energy Conservation in Ethanol-Water
Distillation Column with Vapour Recompression Heat Pump
51
FSV FSL ,QQ LV FFh
(68)
It is important to verify whether Q
1
alone is capable of pre-heating the feed to reach the
desired condition, otherwise the amount of heat that should be withdrawn from the second
pre-heater Q
2
, will be determined as
,
,
.( )0.5p
FLVF
withdrawn sat F F
QFCTTFh
 (69)
The value of heat at the second pre-heater Q
2
should be, at the most equal to Q
withdrawn
to
prevent the feed reaching 50% dry. Therefore, a convenient heat load control could be made
as follows:
If
23 withdrawnQ Q , then 2 withdrawnQ Q
3232Q Q -Q
(70)
If
23 withdrawnQ Q ,
then 223Q Q ,
3 Q0
4.5 Thermodynamic analysis
Since vapour recompression uses a refrigeration cycle rather than a Carnot cycle, the
performance of the heat pump is defined according to the following relation;
23
HPC
h
cp
QQ
COP
W
(71)
The thermodynamic efficiency of a separation process is the ratio of the minimum amount
of thermodynamic work required for separation to the minimum energy required for the
separation (Olujic
et al, 2003). For a vapour recompression distillation column, the energy
required for separation process is composed of the reboiler heat load, Q
reb
, and the
compressor power input,
.
cp
W
T
.
QWreb
c
p
Q (72)
For the separation of a binary mixture by distillation the minimum thermodynamic energy
required to achieve complete separation is given by (Liu and Quian, 2000):
min
ln( ) (1 )ln(1 )TOP F F F FWRTXX X X (73)
Then the thermodynamic efficiency is expressed as:
T
minW
Q
VRC
(74)
4.6 Solution method and error analysis
The equations that model the system components were grouped together in one single
system. The analyses of the status of the variables were carried out to identify those that
were the input data and those which were the unknowns. The equations were then grouped
Distillation – Advances from Modeling to Applications
52
together, resulting in a set of non-linear algebraic equations, which were solved iteratively
based on the step by step use of the successive substitution method. Solution was obtained
when convergence was attained. The convergence was checked by using the criterion:
i1
i1
i
a
X
- X
x 100%
X
(75)
The model is coded in MATLAB environment and used to evaluate the unknowns. A
control programme for column VRC was written to compare the actual column (column in
which the parameters studied were considered, VRC
ΔP
).
5. Discussion of results
5.1 Effects of pressure increase over the compressor
Figure 3 shows how the compressor power input,
.
c
p
W varies with the pressure increase
over the compressor, ΔP. It is obvious from the plots that an increase in pressure over the
compressor increases the compression (pressure) ratio leading to increase in compressor
power input. The curve in Figure 3 shows the effect of pressure increase over the
compressor on coefficient of performance. As the pressure increase over the compressor, ΔP
increases, the compression ratio increases and the coefficient of performance, COP decreases
due to increase in compressor power input.
0.5
0.52
0.54
0.56
0.58
0.6
5
5.2
5.4
5.6
5.8
6
0 20406080100
Compressor power input (kW)
Coefficient of performance
Pressure increase across compressor (kPa)
COP
Power
Fig. 3. The variation of coefficient of performance and compressor power input with
pressure increase across compressor
In Figure 4, a negative non-linear relationship exists between the compressor volumetric
efficiency and pressure increase over the compressor. For a given compressor nominal
capacity, when the pressure increase across the compressor, ΔP increases, the pressure ratio
also increases resulting in the reduction in volumetric efficiency.
The influence of pressure increase over the compressor, ΔP on the required compressor
displacement rate (Vcω) is shown in Figure 4. The linear relationship that exists between the
Energy Conservation in Ethanol-Water
Distillation Column with Vapour Recompression Heat Pump
53
0.00374
0.00378
0.00382
0.00386
0.0039
0.72
0.725
0.73
0.735
0.74
0.745
0.75
0.755
0 20406080100
Compressor volume flow rate
(m³/s)
Compressor volumetric efficiency
Pressure increase across compressor (kPa)
Vol efficiency
Vol flow rate
Fig. 4. The variation of compressor volumetric efficiency and volume flow rate with
pressure increase across compressor
variables indicates that the compressor displacement rate is directly proportional to ΔP. A
reduction in compressor volumetric efficiency caused by increase in ΔP, increases the
compressor displacement rate. This implies an increase in the displacement volume required
although this cannot be observed in the figures. The compressor displacement required for a
given speed is related to compressor size. For the large specific volume obtained, 9.74m
3
kg
-1
ethanol as the heat pump working fluid will require a compressor of greater capacity i.e
large displacement rate.
It can be observed from Figure 5 that an increase in pressure increase across the compressor,
ΔP increases the total energy consumption.
0.114
0.134
0.154
0.174
9.24
9.26
9.28
9.3
9.32
9.34
0 20406080100
Heat load rate (kW)
Total energy consumption (kW)
Pressure increase across com
p
ressor (kPa)
Energy
Heat load
Fig. 5. The variation of total energy consumption and heat load rate with pressure increase
across compressor
For a given reboiler heat transfer rate, Q
reb
, it is obvious that as ΔP increases, the total
energy consumption increases. The total energy consumption and hence the energy
Distillation – Advances from Modeling to Applications
54
savings from the work of Oliveira et al (2001) in heat pump distillation gave lower values.
This could be attributed to non-consideration of the effect of pressure increase over the
compressor and the subsequent increase in compression ratio and in compression power
input respectively.
5.2 Effects of column heat loss
The direct effect of column heat loss could be seen from equation (17). From the equation, it
follows that, for a given reboiler heat load or heat expenditure, fewer trays are required for a
given separation if heat losses are reduced. Where heat loss occurs, more vapour has to be
produced in the reboiler, since the reboiler must provide not only the heat removed in the
condenser but also the heat loss. The effect of this is a decrease in process and energy
efficiency. Indirectly, heat loss affects the column size in terms of number of plates. In the
control system, the reflux ratio was as high as 7.5 compared with 5.033 obtained for the actual
system. Therefore, if heat losses are properly accounted for, there may not be any need for
downward review of the number of plates in order to reduce the reflux ratio (Enweremadu
and Rutto, 2010). Therefore, pressure drop across the column, ΔP
cl
and the difference in boiling
points between the top and bottom products, ΔP
b
which have the most profound effects on the
pressure increase across the compressor, ΔP will be properly predicted. The overall
implication of this is that the column size would be determined properly.
5.2.1 Overall heat transfer coefficient
Analysis of the overall heat transfer coefficient, U of the heat pump reboiler-condenser
revealed an increase in the value of U. This was expected as the value of U in boiling and
condensation processes are high. Also, the value of the heat transfer coefficient of the
condensing ethanol is dominated the relationship used in determining U. A low value of
overall heat transfer coefficient U will result in an increase in the heat exchanger surface
area which may be a disadvantage to ethanol-water system. But the results from this work
showed an increase in the value of U with the implication of a reduction in the reboiler-
condenser heat transfer area.
The variation of the reboiler-condenser thermal conductance (UA) with the heat pump
distribution factor, f, is shown in Figure 6. The plots show that the greater the heat load
taken by the heat pump i.e. larger f’s, the larger the thermal conductance, UA, and the larger
the heat exchanger area. However, for better performance of any heat transfer system, the
thermal resistance (R
th
) which is the inverse of thermal conductance (R
th
= 1/UA), should be
as low as possible. Therefore the value of the heat transfer area for the VRC
ΔP
system will be
smaller compared to the VRC system. Hence, the reboiler-condenser studied has a better
performance.
The relationship between the thermal conductance, UA and the reboiler-condenser
temperature difference, ΔT
CHP
shows that the higher the reboiler-condenser temperature
difference, the lower the thermal conductance. The implication of this is that a higher ΔT
CHP
causes a reduction of the necessary heat transfer area. However, beyond a certain limit of
the thermal driving force, the heat transfer area and the performance of the heat pump
reboiler-condenser reduces. This is expected as higher ΔT
CHP
leads to higher compression
ratio, higher compressor power input and higher energy consumption.
Energy Conservation in Ethanol-Water
Distillation Column with Vapour Recompression Heat Pump
55
0
0.2
0.4
0.6
0.8
1
1.2
0 0.2 0.4 0.6 0.8 1 1.2
Thermal conductance (kW/K)
Condenser distribution factor
Fig. 6. Variation of Reboiler-condenser thermal conductance with condenser distribution
factor
5.3 Comparison of the vapour recompression distillation systems
Table 1 summarises the comparison of some parameters of the two vapour recompression
columns studied, the control (VRC) and the actual (VRC
ΔP
). Both systems operate at 101.2
kPa and the effect of pressure drop effect is considered to account properly for variations
in heat duty. A pressure drop of 0.707kPa per tray is assumed here as a reasonable
estimate for the purposes of this study. The pressure ratio as used throughout this work is
the ratio of the condensation pressure, P
CHP
to the top pressure, P
TOP
, for the VRC system
and sum of the condensation pressure, P
CHP
and pressure increase over the compressor,
ΔP, to the top pressure, P
TOP
for the VRC
ΔP
system. Since the energy consumption changes
linearly with the feed flow rate, and as the present work makes a comparison of the
relative performances, the base case flow rate was taken to be 1.098x10
-4
kmol s
-1 .
Also the
column heat loss and the effect of such parameters as pressure increase across the
compressor, the overall heat transfer coefcient of reboiler–condenser as an explicit
function of Prandtl, Reynolds and Nusselt numbers which in turn depend on uid
properties are considered to account properly for variations in heat duty (Enweremadu,
2007; Enweremadu et al, 2009).
Table 1 shows that the VRC enables some energy savings when compared with VRC
ΔP
. The
VRC has a slightly lower compression ratio and consumes less energy than VRC
ΔP
system.
Although there was a marginal increase in ΔP, which increased the compressor power input
slightly, the performances of the two systems differ greatly by 27.5%. However, in addition
to the pressure increase, the energy consumption appears to depend more on the rate of heat
transfer in the reboiler. Hence the total energy consumption is indirectly related to column
heat loss and pressure increase across the compressor through reboiler heat transfer and
compressor power input respectively.
Also from Table 1, the heat transfer duties of the reboiler-condenser in terms of the overall
heat transfer coefficient for the two systems show that the VRC
ΔP
has a higher value when
compared with the VRC system. This implies that the VRC
ΔP
will require a smaller heat
Distillation – Advances from Modeling to Applications
56
Parameter Actual
column
Control
column
Compressor power input (kW) 0.518 0.495
Total ener
gy
consumption (KW) 9.26 7.26
Rate of column heat loss (kW):


,.
()2
ln
1
2.
()
2
2
ln 1
S
P
amb
loss
o
ins
sins ins
ins m
s
P
ins
oins
o
TT PN
Q
rr
Pt K Nu
dP
K
t
dt
d

P
2
amb
ins
p
2(T - T )
t1
K
o
o
r
h
2.6 0.7
Overall heat transfer coefficient (kW/m²K):
0.8 0.14
3'
0.4
11 1
()
0.023
0.555
(
)
HPC
LL L
f
mm pm wall
ex
LCHP
mmm
ex
v
o
iwall
o
ex
UA
K
g
h
Kud C
dT T
dK










1.4 0.3
(U=Q/A
ΔT)
Reboiler heat transfer rate (kW):
Q
reb
= Dh
D
+ Bh
B
+ L
1
h
LV,e
+ Q
losses
– Fh
F
– Q
1
– Q
2
8.7 6.8 (10%
of
reboiler
heat rate)
Coefficient of performance:
23HPC
cp
QQ
COP
W
5.85 6.15
Condenser heat distribution factor:
HPC
reb
Q
f
Q
0.3 0.5
Compression ratio 1.22 1.12
Compressor displacement rate (m³/s) 3.88x10
-3
3.74x10
-3
Compressor heat load rate (kW) 0.13 0.12
Vapour specific volume after compression (m³/k
g
) 9.49 9.93
Temperature at compressor dischar
g
e (K) 395 390
Compressor volumetric efficienc
y
0.749 0.756
Thermod
y
namic efficienc
y
15.8 20.2
Reflux ratio 5.033 7.5
Table 1. Model results of the actual column and control column (Enweremadu, 2007;
Enweremadu et al, 2008 & 2009)
Energy Conservation in Ethanol-Water
Distillation Column with Vapour Recompression Heat Pump
57
transfer area which is economical in terms of material conservation. Also, with higher U, the
VRC
ΔP
will have a better performance.
The column heat losses for the two vapour recompression distillation columns are shown in
Table 1. The heat losses in distillation columns with heat pumps have been assumed to
correspond to around 3% of the energy supplied to the reboiler (Danziger, 1979). Oliveira,
Marques and Parise (2002) assumed it to be as high as 10%. However, in this study, the heat
exchanged by the distillation column with the surroundings is considered and its effect
included in the balance equation (1). The results obtained for the VRC
ΔP
showed a marked
difference between the two systems.
A comparison of the main reboiler heat transfer rate for the two systems is presented in
Table 1. It is evident that neglecting and /or assuming a value for column heat loss instead
of determining it had a significant effect on the values of Q
reb
in both systems. The heat
pump distribution factor, f, for the VRC
ΔP
system is 0.346 which is slightly less than that
for the VRC system (0.451). Since low value of heat load taken by the heat pump, f implies
lower thermal conductance, then the value of the heat transfer area for the VRC
ΔP
system
is smaller when compared to the VRC
system. However, lower value of thermal
conductance, UA, for the VRC
ΔP
system indicates that the VRC system will have a better
performance.
The simulation results also show that the coefficient of performance and the thermodynamic
efficiency of the VRC
ΔP
system is lower when compared to the VRC system. Fonyo and
Benko (1998) have shown that the electrically-driven compression heat pump should work
with a COP not lower than 3-5. The value of 5.85 obtained from this study has shown that
although there is a decrease in the effectiveness of the VRC
ΔP
system when compared with
the VRC system with COP of 6.15, it is within the acceptable range. This may be due to the
fact that the compressor power input and the total energy consumption in the VRC
ΔP
system
is higher than in the VRC system.
6. Conclusions
From the outcome of the study, the following conclusions may be drawn:
1.
Pressure increase across the compressor, ΔP increases the compression ratio, the
compressor power input, temperature in the heat pump reboiler-condenser while the
compressor volumetric efficiency decreases. The effect of these is the reduction in
the heat pump coefficient of performance and the use of a compressor of greater
capacity.
2.
Neglecting the effects of pressure increase across the compressor, ΔP reduces the
compression ratio and hence maximizes the energy efficiency. However, this leads to a
substantial decrease in temperature in the heat pump reboiler-condenser. The overall
effect of this is a decrease in the overall heat transfer coefficient, U resulting in an
increase in heat transfer area.
3.
From the comparison between the VRC
ΔP
and VRC
systems, there was a profound
difference in the overall heat transfer coefficient while the column heat loss was
substantial.
The increase in the total energy consumption, reboiler heat transfer rate and the
thermodynamic efficiency were appreciable, while there was only a marginal increase
Distillation – Advances from Modeling to Applications
58
in the compressor power input resulting in a difference of 5.12% in the coefficient of
performance between the VRC
ΔP
and VRC systems.
4.
Calculation of the column heat loss in contrast to the assumed value of certain
percentage of the reboiler heat transfer rate gives a higher value resulting in higher
energy consumption and lower thermodynamic efficiency.
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  • [Show abstract] [Hide abstract] ABSTRACT: Vapour recompression system has been used to enhance reduction in energy consumption and improvement in energy effectiveness of distillation columns. However, the effects of certain parameters have not been taken into consideration. One of such parameters is the column heat loss which has either been assumed to be a certain percent of reboiler heat transfer or negligible. The purpose of this study was to evaluate the heat loss from an ethanol-water vapour recompression distillation column with pressure increase across the compressor (VRCAS) and compare the results obtained and its effect on some parameters in similar system (VRCCS) where the column heat loss has been assumed or neglected. Results show that the heat loss evaluated was higher when compared with that obtained for the column VRCCS. The results also showed that increase in heat loss could have significant effect on the total energy consumption, reboiler heat transfer, the number of trays and energy effectiveness of the column.
    Full-text · Conference Paper · Sep 2010