Chemical Engineering & Processing: Process Intensification 119 (2017) 81–92

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Chemical Engineering & Processing: Process Intensification

journal homepage: www.elsevier.com/locate/cep

A hybrid design combining double-effect thermal integration and heat pump MARK
to the methanol distillation process for improving energy efficiency
⁎
Chengtian Cuia, Jinsheng Suna, , Xingang Lia,b,c
a
School of Chemical Engineering and Technology, Tianjin University, Tianjin, 300072, PR China
b
National Engineering Research Center of Distillation Technology, Tianjin, 300072, PR China
c
Collaborative Innovation Center of Chemical Science and Engineering, Tianjin, 300072, PR China

A R T I C L E I N F O A B S T R A C T

Keywords: Despite the low energy efficiency of distillation, it remains the popular separation technology for methanol
Methanol distillation purification. Enlightened by progress in heat pump (HP) concepts, which have been proposed to upcycle waste
Hybrid design heat and reduce energy consumption, this work introduces a hybrid methanol distillation process, which ela-
Double-effect plus heat pump borately integrates the HP with double-effect thermal integration by designing an intermediate heater to shunt
Waste heat recovery
the heat load of the reboiler. Simultaneously, a corresponding optimization function and schematic solution
Process optimization
procedure based on pinch technology are proposed to minimize the operational expenditure. The calculation
Economic evaluation
results demonstrate the validity of the optimization method. Compared with the popular 4-column double-effect
methanol distillation scheme, the hybrid scheme can considerably reduce utility depletion as well as operating
costs, with an acceptable payback period for the compressor. As a result, the hybrid design that gets the ad-
vantage of both double-effect and HP is worth extending to the methanol community as well as to other in-
dustrial plants.

1. Introduction synthesis of methanol [6,7]. However, these crude methanol products

must be refined through a distillation scheme before further utilization
As one of the most important and widespread thermal separation [8–14]. To date, a 4-column double-effect methanol distillation scheme
methods in the modern process industry, distillation has been widely has been the most readily adopted and widely used in China [8,9]. By
applied in the petrochemical, chemical, metallurgic, food, and textile dividing the methanol refining column into a pressured column (PC)
industries. Representing a large part of the global energy usage, it is and an atmospheric column (AC), this double-effect scheme has been
estimated that approximately 43% of thermal energy is used for in- shown to considerably decrease energy consumption compared with
dustrial applications [1]. In particular, distillation alone is responsible earlier designs [12].
for approximately 40% of the thermal energy consumption in the Although the 4-column scheme (designated as the prototype
chemical process industry [2,3], which is the impetus for various en- scheme) has resulted in a significantly greater decrease in energy
ergy saving programs that have been launched for improving distilla- consumption than that expected through double-effect thermal in-
tion performance. One major drawback of distillation lies in its low tegration, it still consumes a considerable amount of hot utility in the
thermodynamic efficiency, requiring consumption of high-quality en- PC reboiler (approximately 80% of the total hot utility consumption)
ergy in the reboiler, while rejecting a similar amount of waste heat to and cold utility in the AC condenser (over 75% of all cold utility con-
the condenser at a lower temperature [4]. In order to improve energy sumption) [8,11]. The literature surveys [8–12] demonstrated that
efficiency in a distillation column, several heat pump (HP) candidates previous works are mainly focused on using sole double-effect config-
have been proposed to aid the upcycling of waste heat that is removed uration to achieve higher energy efficiency, ignoring the application of
from the condenser and to reduce the consumption of valuable utilities HPAD. As a continuation of our previous efforts [8,11], we propose a
[5]. It is predicted that under certain conditions, the margin of energy hybrid 4-column methanol distillation scheme, combining double-effect
savings of heat pump assisted distillation (HPAD) can be approximately and HP. The hybrid design attempts to make full use of the HP to cool
20–50% [1]. part of the AC top vapor in parallel with the condenser, and it can
In the methanol industry, many methods have been proposed for the upcycle the waste heat available for certain heat sinks at higher

⁎
Corresponding author.
E-mail address: jssun2006@vip.163.com (J. Sun).

http://dx.doi.org/10.1016/j.cep.2017.06.003
Received 17 October 2016; Received in revised form 3 May 2017; Accepted 3 June 2017
Available online 21 June 2017
0255-2701/ © 2017 Elsevier B.V. All rights reserved.
C. Cui et al. Chemical Engineering & Processing: Process Intensification 119 (2017) 81–92

Nomenclature Roman letters

Acronyms D Distillate molar flow rate

hD Enthalpy of distillate
AC Atmospheric column hL* Enthalpy of liquid stream
BF Bottom flash h V* Enthalpy of vapor stream
CGCC Column grand composite curve Lmin Minimum molar liquid flow rate
COP Coefficient of performance Q Heat duty
CS Cold stream T Temperature
CW Cooling water Vmin Minimum molar vapor flow rate
GCC Grand composite curve W Work
HEN Heat exchanger network xH* Liquid molar fractions of heavy key component
HP Heat pump xH,D Distillate molar fraction of heavy key component
HPAD Heat pump assisted distillation xL* Liquid molar fractions of light key component
HS Hot stream xL,D Distillate molar fraction of light key component
LEC Light ends column yH* Vapor molar fractions of heavy key component
MOE Minimum operational expenditure yL * Vapor molar fractions of light key component
MTC Minimum thermodynamic condition X Price of steam
MVR Mechanical vapor recompression Y Price of CW
NLP Nonlinear programing; Z Price of electricity
NRTL Non-random two liquid
OE Operational expenditures Greek letters
PBP Payback period
PC Pressured column ε Iteration calculation step size
PNMTC Practical near-minimum thermodynamic condition η AC overhead vapor distribution factor
T-H Temperature-enthalpy λ PC top vapor distribution factor
TIC Total investment cost ξ Side-reboiler fraction
VC Vapor compression
WC Water column

temperatures. Because the AC reboiler is driven by the PC top vapor in literature [17,18]. In considering HEN synthesis, the pinch technology
the prototype scheme, in the hybrid scheme, a side-reboiler is designed emphasizes that any heat leakage from a higher temperature zone
at the AC stripping section, which acts as a new heat receiver at suitable above the pinch to a lower temperature region below it will cause an
temperatures for the donator by shunting part of the reboiler duty. To increase in both hot and cold utilities, namely a double penalty.
better match the heat cascade, an optimization programming function Therefore, the main concern of pinch technology becomes preventing
based on pinch technology was developed as a shortcut for the de- any heat transfer from crossing the pinch, to achieve maximum energy
termination of the column parameters. Simultaneously, rigorous simu- recovery. Following this philosophy, HEN should be designed sepa-
lations were employed to judge the profitability through comparisons of rately above and below the pinch.
the 4-column schemes before and after the introduction of HP in terms Given a list of hot streams (HSs), cold streams (CSs) and a minimum
of energy consumption and operational expenditure. To the authors’ temperature difference ΔTmin, a conceptual temperature-enthalpy (T-H)
knowledge, the hybrid methanol distillation design has not yet been curve of the heat integration system can be formulated using pinch
reported. technology [4]. Besides, a further grand composite curve (GCC) can be
This work is organized as follows. The related theories are depicted plotted to show the net heat flow leaving each internal temperature on
in Section 2. In Section 3, a prototype scheme is derived from a typical T-H axes and to identify heating and cooling requirements at different
4-column setup running in Northern China. The energy consumption temperature levels [16]. In this work, pinch technology is used to
and operating cost of the hybrid schemes are expected to be lower than analyze heat cascade within the methanol distillation process and to
those of the prototype scheme. An economic evaluation of the payback synthesize HEN.
period (PBP) for HP auxiliary equipment is necessary to determine the
feasibility of the hybrid schemes. These contents are introduced in
2.2. The column grand composite curve
Sections 4 and 5. Finally, the conclusions are given in Section 6.

The main goal of distillation process optimization is to reduce its

2. Theory energy consumption. The principal approaches are either improving the
distillation equipment or strengthening heat integration [8]. For this
2.1. Pinch technology purpose, in-depth analysis about the distribution of energy along tray in
distillation column can provide useful information for column design
Energy saving, in the way of reduction in the use of fossil fuel, has and heat integration.
been under active consideration for many years as this leads to the The column grand composite curve (CGCC) can be constructed from
strengthening of competitiveness by saving cost in operation [13]. Heat the column internal stream mass flows and enthalpies [19]. The T-H
integration has been widely applied to oil refineries and chemical in- curve for a binary distillation column under the minimum thermo-
dustries for process energy saving through synthesizing heat exchanger dynamic conditions (MTCs) can be plotted by solving the coupled heat
network (HEN) [14,15]. To achieve maximum energy recovery, and mass balance equations for the reversible separation scheme. For
Linnhoff’s widely respected and accepted pinch technology for heat multicomponent systems, a simplification proposed by Dhole and
integration and HEN synthesis must be applied [16]. A comprehensive Linnhoff [20] uses the light key and heavy key components to ap-
bibliography of the HEN design procedure can be reviewed in the proximate the binary situation, thereby reducing the problem to that of

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a pseudo-binary distillation. The simplification still requires an infinite COP = Qh/ Wcomp (4)
number of stages of separation and an infinite number of side-ex-
The upper theoretical value of COP is related to the Carnot cycle,
changers. The authors described the simplification as a practical near-
which is obtained as:
minimum thermodynamic condition (PNMTC).
The procedure for generating a T-H curve from a converged simu- COPC = Th/(Th − Tc ) (5)
lation of a distillation column involves the determination of the net
where the temperature lift (Th − Tc) (K) is the temperature difference
enthalpy deficit at each stage by generating envelopes over a column
between heat source and sink. Apparently, HPAD is suitable for close-
section from a chosen stage to the top of the column (see Fig. 1).
boiling mixtures since they have a small temperature lift.
Compositions of key components on a stage are available from simu-
In order to check whether or not the use of a HP is suitable for a
lation results. The minimum molar vapor (Vmin) and liquid (Lmin) flow
particular distillation process, Plesu et al. [26] provided an easy-to-use
rates for the column when pinched at this stage can be calculated from:
approach in the preliminary design. The simplified equation is as
VminyL* − LminxL* = DxL, D (1) follow:
COPS = Qreb/ Wcomp = Tcond/(Treb − Tcond ) (6)
VminyH* − LminxH* = DxH , D (2)
where the Qreb, Wcomp, Tcond, and Treb are reboiler duty, compression
where yL*, yH* , xL*, xH* , xL,D, xH,D and D are vapor molar fractions of light work, condenser temperature and reboiler temperature, respectively.
and heavy key component, liquid molar fractions of light and heavy key When the COPS is higher than 10, the HP is clearly recommended.
component, distillate molar fraction of light and heavy key component Between 5 and 10 it should be evaluated in more detail, and if it is
and distillate molar flow rate, respectively. lower than 5, using a HP should not bring any benefits.
The net enthalpy deficits of the column section are estimated by:

Q = VminhV* − LminhL* − DhD (3) 2.4. Economic evaluation

where hV*, hL* and hD are enthalpies of vapor stream, liquid stream and In this study, the straight PBP, not including tax and depreciation, is
distillate, respectively. used in the calculations. The definition of the straight PBP is as follows
When the net heat requirement (Q) of column section is plotted [27]:
against the stage number, a close approximation of the CGCC is ob-
PBP = TIC /(OEPrototype − OEHybrid ) (7)
tained at the PNMTC for the column. The CGCC can provide useful
information about the potential for side heating or cooling operations, where TIC is the total investment cost. This study only focuses on the
which could be integrated with the background process, regardless of costly investment of the compressor since it is much more expensive
whether the reflux ratio and the feed location are appropriate in the than other auxiliaries. OEPrototype and OEHybrid represent the operational
column. Large changes in heat load indicate that there is a large amount expenditure of the prototype and hybrid schemes, respectively.
of heat transfer from the vapor to liquid, and there is the potential to The utility prices based on Chinese currency were determined from
provide the heat or remove the heat at these temperatures. consultations with local engineering companies and were transferred to
In this study, the CGCC of AC is presented to search the opportunity US currency based on the recent exchange rate. Table 1 shows the
for heat integration between the HP pressured stream and column side prices of steam, CW, and electricity. These prices are acceptable in
heat exchanger. Northern China, where methanol plants are commonly invested. The
cost estimation (k US$) of compressors is based on the correlation
2.3. Heat pump assisted distillation proposed by Couper et al. [28]:

TIC = 9.56W 0.62 (8)

HPAD technologies like vapor compression (VC), mechanical vapor
recompression (MVR) and bottom flash (BF) presented in Fig. 2 are
useful way to improve energy quality in spite of disadvantages like high
investment cost and process complexity [21–25].
VC is a classic HP design that has already been applied industrially
[5]. VC uses a specific working fluid as the heat transfer medium that
absorbs heat from a source and discharges it to a sink through a com-
pressor to provide the required work input, while an expansion valve is
used to close the cycle.
MVR is a state-of-the-art HP system that has been widely in-
dustrialized and especially benefits from close-boiling components
[22]. This system uses overhead vapor as a heat transfer medium and
feeds the compressor directly. In addition, it serves as a condenser,
saving one more heat exchanger than the VC alternative. Moreover, it
also avoids cooling the working heat medium to below the boiling point
of the overhead products, an important issue in the VC scheme for heat
transfer. Further, MVR has a slightly higher efficiency and lower in-
vestment expenditure than the VC scheme [24].
BF can be regarded as a variant of MVR. Instead of upcycling the
energy from condenser, BF depressurizes the bottom liquid through an
expansion valve and reuses it as cold utility for condenser. Then, the
liquid is evaporated at the condenser to cool down the overhead vapor
and compressed to feed the column as bottom vapor flow.
The usual measure of the effectiveness of HP in distillation process is
the coefficient of performance (COP), which is defined as the ratio of
the heat rejected at high temperature to the work input: Fig. 1. Envelope for CGCC generation.

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Fig. 2. The schematics of VC, MVR and BF.

Table 1 ensure a sufficient temperature difference between the PC top and AC

Cost data to calculate operational expenditures. bottom, the pressure difference between these two columns is main-
tained over 650 kPa. Refined methanol products are obtained from both
Utility Price US$/(MW*h)
PC and AC tops, with the majority of purified wastewater leaving from
Steam 50.433 the AC bottom. Additionally, a side stream is drawn from a point
CW 4.587 slightly lower than the AC feed stage to control organic impurities, such
Electricity 122.640 as ethanol, in the top product. The side stream containing fusel oil is
pumped into the water column (WC) to recover more methanol from
the purified wastewater remaining at the bottom.
where W(kW) is the actual work done by the compressor.
Apparently, the HP configuration is favored only if COP > 2.432
(the ratio of electricity price over steam price), but to achieve an ac- 3.2. Simulation specifications
ceptable PBP, the COP must be even higher.
The crude methanol feed data, including stream composition, tem-
perature, pressure, and mass flow rate, are presented in Table 2. NRTL
3. The prototype scheme (Non-random two liquid) equation is selected for phase equilibrium,
with binary interaction data fitted for systems of alcohols, water, and
3.1. Process flow description other polar compounds [8,11]. The theoretical (real) stages of the LEC,
PC, AC, and WC are 31 (48), 51 (90), 54 (90), and 46 (85), respectively,
The prototype scheme is an industrialized 4-column double-effect in accordance with the industrial setup [8,11]. The corresponding top
methanol distillation scheme optimized in our previous work by fo- pressures of the LEC, PC, AC, and WC are 150 kPa, 850 kPa, 110 kPa,
cusing on the energy saving, which is presented in Fig. 3 [11]. In this and 110 kPa with sub-cooled condensers and thermosiphon reboilers
particular process, a two-stage condenser system is designed for the without baffles for the PC, AC, and WC. The hot and cold utilities are a
light ends column (LEC) overhead vapor to be partially condensed 600 kPa saturated steam (∼159 °C) and CW working between
stepwise, first at 70 °C and then at 40 °C, with both condensates being 25 ∼ 40 °C, respectively.
collected in the reflux drum. Fresh water (10% wt., with respect to the The quality of the refined methanol meets the requirements of US
crude feed) is pumped into the reflux drum as an extraction agent to federal specification O-M–232 M Grade “AA”: methanol purity not
help remove light impurities soluble in methanol. Compared with the lower than 99.85% (wt.) and ethanol not more than 10 ppm (wt.). In
one-stage condenser [12], this configuration saves cooling water (CW) addition, the methanol content in wastewater is specified as no more
by preventing a considerable part of the liquid from deeper cooling. than 10 ppm (wt.) for water treatment requirements [11].
After removing the light ends from the top of the LEC, the crude me- Because the refined methanol products are obtained from PC, AC,
thanol feed is pumped into the PC from the LEC bottom. The PC bottom and WC overheads, certain product distributions are specified for re-
stream enters the AC for further separation. On the other hand, the PC ference, following the on-site date. These are approximately 46% for
overhead vapor drives the AC bottom, resulting in double-effect dis- the PC and 52.5% for the AC of pure methanol in feedstock, with the
tillation by combining the PC condenser with the AC reboiler, thus remaining amount of approximately 1% from the WC.
saving a condenser or reboiler along with heating or cooling utilities. To Rigorous simulations were carried out under these proposed

Fig. 3. 4-column double-effect methanol distillation scheme:

the prototype scheme.

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Table 2 Table 4
Crude methanol feed data. Operating conditions of the prototype scheme.

Component Mass fraction% Component Mass fraction% Items LEC PC AC WC

Hydrogen 0.0100 Pentanol 0.0338 Top pressure (kPa) 150 850 110 110
Nitrogen 0.0710 Formic acid 0.0030 Bottom pressure (kPa) 153 855 115 115
Methane 0.0100 Acetic acid 0.0030 Top temperature (°C) 81.9 130.5 66.6 66.6
Carbon monoxide 0.0360 Methyl formate 0.0092 Bottom temperature (°C) 78.6 136.8 102.9 102.2
Carbon dioxide 0.6640 Acetone 0.0193 Reflux ratio 0.44a 2.93 1.72 3.35
Methanol 94.946 Methyl ethyl ketone 0.0030 Number of theoretical stage 31 51 54 46
Dimethyl ether 0.0309 Diisopropylether 0.0056 Rectifying section 10 46 30 18
Trimethylamine 0.0001 Water 3.7100 Stripping section 21 5 24 28
Ethanol 0.0900 i-Pentane 0.0562 Condenser duty (MW) −28.6 −115.1 −107.5 −3.6
i-Propanol 0.0024 n-Pentane 0.0169 Reboiler duty (MW) 27.75 114.0 96.3 3.5
Propanol 0.0264 n-Hexane 0.0487
Total reboiler duty using steam (MW) 126.40
i-Butanol 0.0338 n-Heptane 0.0009
Purified methanol yield (%) 99.62
Butanol 0.0185 n-Octane 0.0050
3-Pentanol 0.1407 n-Nonane 0.0050 a
For LEC, the reflux ratio is ratio of reflux to feed.
Temperature 40 °C
Pressure 500 kPa
Flow rate 239,230.0 kg/h Table 5
Product compositions for the prototype scheme.

specifications in SimSci Pro/II™ environment. Components Refined methanol Purified wastewater Fusel oil (wt)
(wt) (wt)

3.3. Simulation results and discussion Hydrogen 0.0000e0 0.0000e0 0.0000e0

Nitrogen 0.0000e0 0.0000e0 0.0000e0
Tables 3–5 summarize the material balances, operating conditions, Methane 0.0000e0 0.0000e0 0.0000e0
Carbon monoxide 0.0000e0 0.0000e0 0.0000e0
and composition of the prototype scheme, respectively. The results are Carbon dioxide 0.0000e0 0.0000e0 0.0000e0
in agreement with the on-site data and show a mass yield distribution of Methanol 0.9998e0 9.9998e-6 0.3948e0
46.2%, 52.7%, and 1.1%, from the PC, AC, and WC, respectively, with Dimethyl ether 4.8949e-20 0.0000e0 0.0000e0
corresponding specific steam consumptions of 3926.7 kJ/kg, Trimethylamine 1.8524e-8 1.1388e-26 0.0000e0
Ethanol 1.0000e-5 9.5115e-5 0.2447e0
2906.9 kJ/kg and 5059.6 kJ/kg. This finding reveals that the AC yield
i-Propanol 1.8442e-9 1.5752e-8 6.6965e-3
is the most energetically effective among the three producers. The Propanol 9.1043e-13 2.3655e-4 0.0645e0
quality of refined methanol is better than the “AA” standard, with i-Butanol 9.7272e-15 2.4964e-4 0.0847e0
methanol approaching 99.98%, the expected methanol content in waste Butanol 8.2839e-22 1.3190e-3 6.9244e-4
water, and a reasonable fusel oil rate. 3-Pentanol 5.2235e-20 0.0102e0 4.4583e-4
Pentanol 2.7669e-23 2.4369e-3 2.2567e-4
Formic acid 5.9997e-15 2.1650e-4 1.2027e-5
3.4. Heat integration analysis Acetic acid 1.1191e-21 2.1652e-4 1.1147e-5
Methyl formate 1.2697e-9 2.1867e-26 0.0000e0
Acetone 1.0651e-4 6.4472e-20 1.4978e-16
HEN candidate streams in the prototype scheme are color-coded in
Methyl ethyl ketone 2.5413e-5 5.8785e-16 9.4732e-11
Fig. 3, and their detailed information, including the stream name, in- Diisopropylether 1.2046e-8 0.0000e0 0.0000e0
itial and final target temperatures, heat duty, and corresponding stream Water 8.0999e-7 0.9848e0 0.2031e0
position, is provided in Table 6. The minimum temperature difference i-Pentane 0.0000e0 0.0000e0 0.0000e0
ΔTmin of the HEN synthesis is herein specified as 15 °C. To show the n-Pentane 7.6003e-14 0.0000e0 0.0000e0
n-Hexane 0.0000e0 0.0000e0 0.0000e0
potential heat integration of streams in Table 6, the GCC is shown in n-Heptane 1.8467e-5 0.0000e0 5.6410e-24
Fig. 4. The GCC indicates a hot stream pinch temperature of 81.8 °C and n-Octane 4.7993e-5 3.3304e-5 5.2208e-8
a cold counterpart of 66.8 °C, along with energy targets of 139.89 MW n-Nonane 1.8337e-5 2.3581e-4 7.8260e-6
and 138.47 MW for heating and cooling, respectively. The circled areas
are the pockets, which represent the potential energy integration within
the heat cascade. The potential energy integration can be achieved by temperature, which is in accordance with Smith [4], showing that
using four heat exchangers (HEs) marked on Fig. 3. Table 7 summarizes HPAD is only feasible when the candidate stream is operated across the
the process-to-process HEs information for the prototype scheme. pinch. Under this circumstance, there is a possibility for additional heat
However, it must be noted that the utility consumption is still integration between the HP and AC top, an advantageous option to
considerable, even with the double-effect thermal integration. The re- reduce CW consumption. This water-saving proposal is especially ben-
sults in Table 6 clearly show that HS4 (AC condenser) consumes con- eficial to methanol manufacturers located in water-deficient areas.
siderable CW, corresponding to approximately 77% of the total CW Because only HS4 is considered in HPAD, the evaluation of the PBP,
consumption. Specifically, this HS is located below the pinch thus, the profitability, is simplified.

Table 3
Material balances of each column for the prototype scheme.

Column Input (kg/h) Output (kg/h)

Feed 1 Feed 2 Sum Overhead Side-draw Bottom Sum

LEC 239,230.0 23,923.0 263,153.0 2,921.5 – 260,231.5 263,153.0

PC 260,231.5 – 260,231.5 104,516.5 – 155,715.0 260,231.5
AC 155,715.0 – 155,715.0 119,262.6 3,939.7 32,512.7 155,715.0
WC 3939.7 – 3939.7 2490.3 859.5 589.9 3939.7

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Table 6 Table 8
Information regarding HEN candidate streams in the prototype scheme. Methanol product distribution in PC/AC and their corresponding top/bottom duty.

Stream Stream Initial Target Heat duty PC top AC top PC top duty PC bottom AC top duty AC bottom
name position temperature (°C) temperature (°C) (MW) product product (MW) duty (MW) (MW) duty (MW)
fraction fraction
CS1 Crude feed 40.00 85.00 8.160
preheated 0.3 0.685 −74.7449 73.8832 −138.2766 124.3853
CS2 PC feed 78.94 135.97 13.864 0.32 0.665 −79.7341 78.8643 −134.6571 121.1079
preheated 0.34 0.645 −84.7485 83.8368 −131.0369 117.8333
CS3 LEC reboiler 78.60 78.62 27.750 0.36 0.625 −89.7552 88.8415 −127.4020 114.5472
CS4 PC reboiler 136.10 136.18 114.064 0.38 0.605 −94.7955 93.8361 −123.7676 111.2428
CS5 AC reboiler 102.90 102.92 96.267 0.40 0.585 −99.8288 98.8702 −120.1163 107.9325
CS6 WC reboiler 102.20 102.22 3.544 0.42 0.565 −104.8918 103.9106 −116.4613 104.6138
HS1 First stage 81.76 70.00 26.858 0.44 0.545 −109.9904 108.9593 −112.8039 101.3019
condenser 0.46 0.525 −115.0968 114.0439 −109.1311 97.9743
HS2 Second stage 70.00 40.00 1.744 0.48 0.505 −120.2381 119.1593 −105.4513 94.6308
condenser 0.50 0.485 −125.4087 124.3064 −101.7601 91.2858
HS3 PC condenser 130.50 125.19 115.115
HS4 AC condenser 66.60 40.00 107.510
HS5 WC condenser 66.60 40.00 3.584 respect to methanol content in crude methanol feed, is obtained from
HS6 PC top stream 125.19 40.00 7.425
the PC overhead, while 0.525 of the share is derived from the AC top.
to tanks
The product outlet must be redistributed because the AC reboiler re-
duces the energy requirement of the PC top vapor when HP is involved,
maintaining an identical product distribution and leading to an energy
cross-pinch.
To facilitate the calculation, we specified a subtotal mass fraction
yield of 0.985 and set the corresponding PC top vapor distribution
factor λ for the part from the PC top. For example, if λ = 0.30, then PC
mass fraction yield is 0.30, while that of AC is 0.985-0.3 = 0.685.
When the yield and heat distribution between the PC and AC in Table 8
are graphically expressed in Fig. 5, the heat duties (MW) of the PC
condenser and AC reboiler are fitted linearly versus the PC top vapor
distribution factor λ:

QPC , Cond = 253.1λ − 1.332 (9)

QAC , Reb = − 165.5λ + 174.1 (10)

These relationships reveal proportionalities between column yields

and their heat duties. The lines of heat duty of the PC overhead and the
Fig. 4. GCC for HEN streams in the prototype scheme. AC bottom cross at λ = 0.42, which is a smaller value than that of the
prototype (λ = 0.46), indicating that less heating medium is required.
Table 7
Currently, in China, λ of the industrial prototype scheme ranges from
Process-to-process heat exchanger information in the prototype scheme. 0.45 to 0.50 [8,11]. For the hybrid scheme, because HP is allocated on
the AC with a side-reboiler, λ should be less than 0.42 because the
Heat exchanger name Hot stream-Cold stream LMTD (°C) Heat duty (MW) intermediate heater shunts considerable heat duty from the AC reboiler.
HE1 HS6-CS1 29.26 3.300
It is important to note that changes in λ in the hybrid scheme will cause
HE2 HS1-CS1 25.32 4.860 variations in the column diameter. Therefore, the hybrid scheme is only
HE3 HS3-CS3 47.02 18.848 suitable for grassroots design, not for retrofit tasks.
HE4 HS3-CS5 25.31 96.267

4. The hybrid scheme

Based on the discussion on feasibility, this study considers the HPs

of VC, MVR and BF patterns for the hybrid scheme. When HP is em-
bedded in the prototype, heat integration must be reanalyzed, as well as
the economic evaluation of the PBP of the additional compressor. This
is performed with the aid of rigorous simulations.

4.1. Heat integration reanalysis

4.1.1. Distribution of refined methanol products

During the design of the hybrid scheme, the prototype with the side-
reboiler was allocated in the AC stripping section, so that refined me-
thanol products from the PC and AC top must be redistributed for future
heat integration.
In the prototype scheme, 0.46 of the refined methanol product, with Fig. 5. Heat duty versus PC overhead vapor distribution factor diagram.

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4.1.2. Condensation strategy of AC overhead vapor

As mentioned, heat-across-pinch causes the import of an extra
amount of heat from the hot utility and the export of an equivalent
amount of heat to the cold utility, increasing the energy consumption.
Table 8 shows that the AC overhead vapor has a heat duty of over
100 MW in each case. If all the waste heat is upcycled, then it will cause
a double energy penalty because there are no sufficient heat sinks to
match the upgraded heat sources, and CW will be used for the heat
surplus above the pinch. A condensation strategy is required to prevent
this energy penalty and to distribute vapor going either to the HP
system (above the pinch) or to the CW (below the pinch). The AC
overhead vapor distribution factor η (0 ≤ η ≤ 1) is defined to calculate
the fraction of vapor distributed to the HP. If QAC,cond is the energy
requirement for cooling down the entire AC overhead vapor, then
ηQAC,cond and (1 − η)QAC,cond will be the energy shunts entering the HP
and CW, respectively.

4.1.3. Energy balance reanalysis of the total site

After introducing HP to the prototype scheme, the energy balance of
the total site should be reanalyzed through shortcut energy balance Fig. 7. The CGCC of AC.
equations to fulfill the aforementioned condensation strategy. As an
example, the shortcut energy balance in the VC scheme is included with
fraction ξ is defined to represent the side-reboiler duty over the entire
an explanation. The MVR and BF schemes are calculated based on si-
AC reboiler duty without the intermediate heater (QAC,reb). For ex-
milar equations.
ample, if the heat duty of the AC side-reboiler in the VC scheme is
After energy balancing of the total site and the corresponding heat
ξQAC, reb , then the corresponding AC reboiler duty will be (1 − ξ )QAC, reb .
integration in the VC scheme, the candidate streams include not only 12
In this case, the HS7 duty becomes the following:
streams (6 cold and 6 hot in Table 6) in the prototype scheme but also
HS7, with the heat duty QCond representing the HP condenser heat load QHS 7 = QCS1 + QCS3 + ξQAC, reb (11)
and CS7 representing heat duty of the side-reboiler (see Fig. 6). Note
The hot utility (600 kPa saturated steam, 159 °C) is used to heat CS2
that the HEN structure has to be reformulated since changing product
(PC feed preheated, from 78.94 °C to 135.97 °C) and CS4 (PC reboiler,
distribution will inevitably affect process streams.
approximately 136.1 °C):
To clarify the heat cascade in AC, its CGCC is shown in Fig. 7, where
the blue dashed lines represent the condensation strategy, and the red Qsteam = QCS 2 + QCS 4 (12)
dashed lines indicate the energy distribution in heat sinks (reboiler and
CW is mainly used to cool down HS1 (first stage condenser, from
side-reboiler).
81.76 °C to 70 °C), HS2 (second stage condenser, from 70 °C to 40 °C),
In the VC scheme, it is expected that HS7 can match CS1 (crude feed
HS4 (AC condenser, from 66.6 °C to 40 °C), and HS5 (WC condenser,
preheated, from 40 °C to 85 °C), CS3 (LEC reboiler, approximately
from 66.6 °C to 40 °C). Because HS6 is expected to match suitable heat
78.6 °C), and CS7 (AC side-reboiler, approximately 82 °C). Herein, the
sink in a rigorous simulation, it is excluded from heat balances during
COPS = 21.5, which is higher than 10, so the HP is clearly re-
preliminary design. Thus, the CW duty is as follows:
commended according to Plesu et al. [26]. It is interesting to note that if
only HPAD is selected without considering double-effect heat integra- QCW = QHS1 + QHS 2 + QHS 4 + QHS5 (13)
tion, the COPS reduces to 9.46, because methanol-water being sepa-
rated has widely different boiling points. Therefore, the hybrid design QHS 4 = (1 − η)QAC, cond (14)
with double-effect plus HP configuration can bring a synergistic effect In double-effect distillation, the PC top vapor HS3 drives the CS5
− double-effect creates lower temperature heat sink for HP and HP (AC reboiler, approximately 102.9 °C) and CS6 (WC reboiler, approxi-
recovers waste heat for double-effect. Because the AC side-reboiler mately 102.2 °C):
shunts heat duty from the AC reboiler (Fig. 7), it is important to de-
termine the fraction of the share. Consequently, the side-reboiler QHS3 = QCS5 + QCS6 (15)

Fig. 6. Total sites energy analysis of the VC scheme.

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Table 10
COP of working fluid under different degrees of superheating in the HPAD cycle.

Superheating Evaporation Compression Condensing COP

temperature (°C) duty (MW) duty (MW) duty (MW)

0 120.741 37.267 158.009 4.240

2.5 120.675 36.883 157.558 4.272
5.0 120.716 36.523 157.239 4.305
7.5 120.809 36.331 157.140 4.325
10.0 120.737 36.157 156.893 4.339
12.5 120.728 36.002 156.730 4.353

With using the COP:

COP = Qcond / Wcomp (19)
Combining (11), (15), (16) and (17), the following are obtained:
Qcond = QHS7 (20)

Wcomp = {1/(COP − 1)}ηQAC , cond (21)

Qcond = {COP /(COP − 1)}ηQAC , cond (22)

4.2. Economic evaluation

4.2.1. Operational expenditure

The utility prices of steam, CW, and electricity are assumed to be X,
Y and Z, respectively. These prices are shown in Table 1. A minimum
operational expenditure (MOE), determined by several factors, must be
calculated. This economic evaluation leads to nonlinear programing
(NLP). Before inducing the optimization programming function, the
assumption, which is not very restrictive, is postulated to simplify the
optimization in which the reboiler and condenser of the PC and AC
have identical values of heat duty, for example, as follows:
QAC, reb = QAC , cond = QAC (23)

QPC , reb = QPC , cond = QPC (24)

With the assumption and above shortcut energy balance equations,
the optimization programming function is listed below:

min MOE = QsteamX + QCWY + WcompZ (25)

s.t. QCS1 + QCS3 + ξQAC, reb = {COP /(COP − 1)}ηQAC

Qsteam = QCS 2 + QPC

QCW = QHS1 + QHS2 + QHS5 + (1 − η)QAC

Fig. 8. Schematic diagram of the solution procedure. QPC = (1 − ξ )QAC + QCS6

WComp = {1/(COP − 1)}ηQAC , cond

Table 9
Choices for working fluid for the VC scheme.

Working fluid Saturated temperature under 100 kPa QPC = 264.92 − 1.53QAC
Trimethylchlorosilane 57.39 °C
Methyldichlorosilane 40.89 °C 0≤η≤1
Dimethyldichlorosilane 69.69 °C
Methyltrichlorosilane 66.13 °C
0 ≤ ξ ≤ ξmax

In this optimization function (25), the heat duties of certain streams

QCS5 = (1 − ξ )QAC, reb (16) are not sensitive to variations of λ. These streams include CS1, CS2,
CS3, CS6, HS1, HS2, and HS5; substituting their values from the pro-
Electricity consumption is derived from the energy balance of the
totype scheme into the optimization function makes the NLP model
HP cycle:
solvable. Note that this function is only for a shortcut calculation of λ
Wcomp = Qcond − Qevap (17) and η, ignoring the constraints of HS6 heat duty and the minimum
temperature difference ΔTmin. A more accurate economic evaluation of
Qevap = ηQAC , cond (18) operational expenditure relies on a subsequent rigorous simulation,

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Table 11
Material balances of each column for the VC scheme.

Column Input (kg/h) Output (kg/h)

Feed 1 Feed 2 Sum Overhead Side-draw Bottom Sum

LEC 239,230.0 23,923.0 263,153.0 2931.9 – 260,221.1 263,153.0

PC 260,221.1 – 260,221.1 62,259.6 – 197,961.5 260,211.1
AC 197,961.5 – 197,961.5 162,682.1 2,501.9 32,777.5 197,961.5
WC 2501.9 – 2501.9 1458.1 681.6 362.2 2501.9

Table 12
Operating conditions of the VC scheme.

Items LEC PC AC WC

Top pressure (kPa) 150 850 110 110

Bottom pressure (kPa) 153 855 115 115
Top temperature (°C) 81.9 130.5 66.6 66.6
Bottom temperature (°C) 78.6 135.5 103.0 98.3
Reflux ratio 0.438a 2.913 1.787 6.593
Number of theoretical stage 31 51 54 46
Rectifying section 10 46 30 18
Stripping section 21 5 24 28
Condenser duty (MW) −28.57 −68.62 −58.09 −3.67
Side-reboiler duty (MW) – – 65.00 –
Reboiler duty (MW) 27.75 67.43 64.96 3.64
Compressor duty (MW) – – 21.71 –

Total reboiler duty using steam (MW) 67.43

Purified methanol yield (%) 99.67

a
For LEC, the reflux ratio is ratio of reflux to feed.
Fig. 9. GCC for HEN streams in the VC scheme.
Table 13
Information regarding HEN candidate streams in the VC scheme. 5. Case study
Stream Stream Initial Target Heat duty
name position temperature (°C) temperature (°C) (MW) In this part, three case studies, using HPs of VC, MVR and BF pat-
terns, respectively, were performed based on the schematic procedure
CS1 Crude feed 40.00 85.00 8.160 in Fig. 8. Based on rigorous simulations under SimSci Pro/II™ en-
preheated
vironment, comparisons between the prototype scheme and the hybrid
CS2 PC feed 78.94 135.97 13.864
preheated schemes were carried out, with an estimation of the operational ex-
CS3 LEC reboiler 78.60 78.62 27.750 penditures of all of the schemes and PBPs of the compressors. It is ex-
CS4 PC reboiler 135.00 135.50 67.430 pected that these case studies can provide suggestions to the methanol
CS5 AC reboiler 102.90 102.92 64.960
industrial community.
CS6 WC reboiler 98.30 98.32 3.640
CS7 AC side- 83.30 84.90 65.000
reboiler 5.1. The VC scheme
CS8 HP cycle cold 41.29 53.79 71.757
stream
The choice of working fluid in the VC scheme is based on our pre-
HS1 First stage 81.76 70.00 26.858
condenser vious works [29,30], where MVR was revamped on the organosilicon
HS2 Second stage 70.00 40.00 1.744 monomer distillation scheme. Dew point temperatures of these orga-
condenser nosilicon monomers are found to be suitable for recovering waste heat
HS3 PC condenser 130.50 125.19 68.620
from AC overhead vapor. Table 9 shows the saturated temperature of
HS4 AC condenser 66.60 40.00 58.086
HS5 WC condenser 66.60 40.00 3.667 these monomers under 100 kPa. Because the temperature of the AC top
HS6 PC top stream 125.19 40.00 4.423 vapor is 66.6 °C, only methyldichlorosilane matched the minimum
to tanks temperature difference of 15 °C, excluding other choices as the working
HS7 HP condenser 116.24 98.85 93.616 fluid.
HS8 AC overhead 66.60 66.60 71.757
The compressor is specified to have an 80% efficiency [13] and the
vapor to HP
outlet pressure is specified as 500 kPa, corresponding to a saturated
dew point temperature of 98.85 °C. Itard [31] reported that for a
which is performed based on a fully matching heat cascade within the working fluid with a positive saturated vapor line slope in a real cycle
process. A schematic diagram of the solution procedure can be re- (isenthalpic expansion and non-isentropic compression), a certain de-
viewed in Fig. 8. In the schematic procedure, a rigorous simulation is gree of superheating in evaporator favors COP. The simulation results
performed, considering the heat duty of all of the streams and the shown in Table 10 verify this conclusion. Consequently, 12.5 °C of su-
minimum temperature difference ΔTmin. The utility prices X, Y, Z, the perheating is specified in the evaporator, satisfying the minimum
COP, initial iteration fraction ξ0, and iteration calculation step size ε temperature difference. The side-reboiler is set at the 43rd theoretical
(e.g., ε = 0.05) are input as the initial values. The calculation is ex- stage with a duty of 65 MW (side-reboiler fraction ξ is 0.5). The cal-
pected to yield a maximum side-reboiler fraction ξ max under the con- culated PC top vapor distribution factor λ is 0.277, and the AC over-
straints in this specific model to minimize operating expenses. head vapor distribution factor η is 0.553 in a rigorous simulation. The
material balance and operational conditions are summarized in Tables

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Fig. 10. HEN in the VC scheme.

Table 14
Process-to-process heat exchanger information in the VC scheme.

Heat exchanger name Hot stream-Cold stream LMTD (°C) Heat duty (MW)

HE5 HS7-CS3 28.04 27.750

HE6 HS7-CS7 15.74 65.000
HE7 HS3-CS5 24.99 64.960
HE8 HS7-CS1 21.19 0.866
HE9 HS6-CS1 24.79 2.253
HE10 HS3-CS6 27.02 3.640
HE11 HS1-CS1 24.57 5.041
HE12 HS8-CS8 18.35 71.757

11 and 12, respectively.

The results indicate that in the VC scheme, the distributions of re-
fined methanol products obtained from the top of the PC, AC, and WC
are 27.50%, 71.86%, and 0.64%, respectively. The specific steam
consumptions of refined methanol in the PC, AC, and WC are Fig. 12. GCC for HEN streams in the MVR scheme.
3899.0 kJ/kg, 2878.9 kJ/kg, and 8987.0 kJ/kg, respectively. Compared
with the prototype, the majority of refined methanol is obtained at a demonstrated in Fig. 9. Apparently, the pinch point does not changed,
lower level of energy consumption. The purified methanol yield is but two more pockets appeared in the heat cascade compared to the
99.67%, which is nearly identical to the prototype (99.62%), while prototype, providing more opportunities for thermal integration. The
energy targets for heating and cooling are decreased to 81.28 MW and potential energy integration can be achieved by using eight HEs shown
87.49 MW, respectively. Compared with the prototype, these values in the HEN (see Fig. 10). Table 14 gives the process-to-process HEs
show a decrease of 41.9% and 36.8% for heating and cooling, respec- information for the VC scheme.
tively. However, the electricity consumption increases by 21.71 MW. The estimated operational expenditure of the prototype
Table 13 lists the candidate streams involved into HEN synthesis. To OEPrototype = 7690 US$/h, while that of the VC scheme OEVC = 7163 US
present the potential thermal integration of these streams, the GCC is $/h, which is equivalent to a 6.85% decrease in the operational

Fig. 11. The proposed hybrid MVR scheme.

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C. Cui et al. Chemical Engineering & Processing: Process Intensification 119 (2017) 81–92

Fig. 13. The proposed hybrid BF scheme.

5.3. The BF scheme

In the BF scheme (see Fig. 13), the basic HEN structure of the
prototype scheme is utilized. A side stream is drawn from the 43rd stage
of AC to apply BF. The relief-valve outlet pressure is designed as 20 kPa,
which corresponds to an outlet temperature of 49.75 °C, satisfying the
minimum temperature difference 15 °C for HEN synthesis. Herein, the
compressor reaches a power of 20.02 MW, which is equivalent to a COP
of 5.560. The GCC (see Fig. 14) shows the pinch temperature is the
same to the prototype, along with energy targets of 139.89 MW and
27.08 MW for heating and cooling, respectively. It can be seen from the
GCC pocket that the BF only provides low-temperature heat for the AC
condenser. Although the cold utility target can decrease by 80.44%, the
operating cost of the BF scheme increases to 9635 US$, which is much
higher than the prototype scheme (7690 US$). This is because the cost
of CW is much lower than that of electrical power, and the HP stream
Fig. 14. GCC for HEN streams in the BF scheme. cannot better match with the background process streams. As a con-
sequence, the BF scheme is not feasible for the task.
expenditure. The TIC of the compressor is estimated as approximately
4678 k US$. Thus, 8877 h of PBP is expected. In current Chinese me- 6. Conclusions
thanol plants, 8000 h/y of an operation period is acceptable. In other
words, the PBP is approximately 1.11 y. In this study, a hybrid methanol distillation scheme is proposed,
featuring double-effect plus HP technology. An intermediate heater,
driven by the upcycled energy, is introduced to the AC to shunt the heat
5.2. The MVR scheme load in the reboiler. To find the best match for the heat cascade, an
optimization programming function is proposed to perform the calcu-
In the MVR scheme (see Fig. 11), the basic HEN structure and op- lations. Significant improvements are achieved. Both the VC and MVR
erating conditions are the same to the VC scheme, except for the direct schemes can considerably reduce utility consumptions as well as op-
vapor recompression using in this process. Herein, the AC overhead erating costs, and their PBPs are acceptable. The BF scheme is observed
vapor is compressed to 400 kPa, achieving a saturated dew point tem- not suitable for the hybrid design because its upgraded stream cannot
perature of 103.99 °C as well as a sufficient temperature difference for better match with the background process. Therefore, the hybrid
heat integration in the side-reboiler. Note that in the MVR scheme, it is schemes with VC and MVR are worth introducing to the methanol
not necessary to unify the compressor outlet pressure with the VC community. Moreover, the authors encourage the use of the hybrid
scheme because of the higher compression ratio, higher power con- technology with double-effect plus HP and similar schematic optimi-
sumption, as well as capital investment of the compressor. From an zation procedure based on pinch technology in other industrial plants.
energy savings and economics perspective, it is better to specify a sui-
table pressure that is as low as possible while still satisfying the Acknowledgment
minimum temperature difference for heat integration.
In this case study, specifying a compressor efficiency of 80% and The authors are grateful for the industrial consultations from
equivalent operating conditions in the VC scheme, the compressor Shandong Haicheng Petrochemical Engineering Design Co. Ltd.,
reaches a power of 12.30 MW, which is equivalent to a COP of 7.611. Tianjin, China.
The GCC is shown in Fig. 12. It is interesting to noted that the number
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