Ind. Eng. Chem. Res. 1992,31,2547-2555
2147
Operation and Experimental Results on a Vapor Recompression Pilot Plant Distillation Column Edgardo R. Canales* and Fernando E. Mdrquez Department of Chemical Engineering, University of Concepcibn, P.O.Box 53-C, Correo 3, Concepcibn, Chile
The open-loop steady-state and dynamic operation of a vapor recompression pilot-plant distillation column is studied under direct digital control of the loads to the process. Data acquisition, process control, and operator interface are programmed in a distributed control system. An ethanol-water mixture is separated. The manipulated inputs are cooling water flow rate and heat transferred a t the reboiler-condenser. Energy savings were of the order of 50% as compared to conventional distillation. The process presents quasi-linear stationary and dynamic behavior when subjected to small departures in heat input a t the reboiler condenser. Time lags and delays are higher than in conventional distillation due to process interactions a t the reboiler-condenser.
Introduction Vapor recompression or heat pump systems in distillation column have been extensively studied in the past decades, due mainly to the importance of energy conservation techniques that chemical and petrochemical industries are implementing after the 1973 oil crisis. A number of articles have appeared in the literature related to this topic, especially dealing with steady-state simulation, optimization, and economic aspects. Vapor recompression or heat pumps have been well described during the past two decades by Null (1976), Danziger (1979),Roffel and Fontein (1979),Finelt (19791, Menzies and Johnson (1981), Meili (1990), and other authors. In the past decade there has been additional work in control and experimental equipments, as mentioned in Moetafa (1981),Quadri (l98l4b), Carta et aL (19821,Florea et al. (1984), Ferre et al. (1985), Brousse et al. (19851, Omideyi et al. (1985), Meili and Stuecheli (19871, Collura and Luyben (1988), Miihrer et al. (1990), and Papastathopoulou and Luyben (1991). Most of these papers present vapor recompression or heat pumps as an alternative techique to conventional distillation. For some difficult separations of close-boiling mixtures the heat pump system ia often more convenient due to lower energy consumption. Omiyedi et al. (1984a,b, 1985) and Gopichand et al. (1984) in a series of four papers presented the analysis of a heat pump miatad distillation system (external recompression), where they compared different variables in the economics of the separation of mixtures of methanolwater, ethanol-water, and methanol-ethanol. Brousee et al. (1985) presented a simulation and optimization program which allows the analysis of the economid viability of substituting the reboiler and condenser of a conventional distillation column by a vapor recompression heat pump. They used an experimental pilot plant with a packed column working at total reflux (no feed, no distillate), and they adapted a data processing system to the vapor recompression unit that allows the identification of mass-tranfer parameters of the column, the heat exchange in the reboiler-condenser, and the isentropic yield of the compressor. Collura and Luyben (1988) presented a thorough study of distillation designs and operating conditions to produce a fuel-grade ethanol product, using several options like vapor recompression, multiple-effect columns, and other combinations. The analysis for the different schemes were *Towhom correspondence should be addressed.
shown as well as an economic evaluation of the investment and operating costs. Miihrer et al. (1990) reported the results of a detailed, quantitative simulation study of the dynamics of different schemes of vapor recompression columns for two specific systems (propylene-propane and ethanol-water). As shown from the literature there are not many experimental works done on the subject, but the following may be mentioned Danziger (1979), from the Swiss company Sulzsr Brothers Ltd., used a pilot plant using a vapor recompression column with special packing and a centrifugal turboblower to separate a mixture of cis-/trans-decalin. They concluded that the heat pump system is a feasible way of economizing energy in a distillation plant. Meili, also from Sulzer Brothers, in 1990 showed different applications of vapor recompression to several industrial commodity chemicals, and he mentioned the savings in operating cost compared with the conventional distillation columns. Carta et al. (1981) reported the application of direct vapor recompression to an existing plant, fractionating an ethylbenzene-xylene mixture, and concluded that the use of the heat pump system, in the usual range of operating conditions, seems to be of limited advantage as a way of economizing energy. However, they also included the possibility of optimizing the operating pressure to reassess the use of vapor recompression. Papastathopoulou and Luyben (1991) presented a comprehensive study of the dynamics and control of a large industrial distillation column that separates a binary mixture of propylene and propane and with a two-stage vapor recompression system used for energy conservation. They used plant data to validate steady-state and dynamic models of the column and also evaluated several control structuresof the column. Besides the literature review already presented, there are other papers related to the subject such 88 the works of Lynd and Grethlein (1986),Loken (1985),and Bjbrn et al. (1991). The aim of this paper is to report some experimental results on steady-state operation and dynamic behavior of a laboratory vapor recompression distillation column.
Experimental Procedure A scheme of the pilot plant is shown in Figure 1. The distillation column E-1 is an %in.-diameter copper tower with 20 trays and 3 bubbling caps per tray and a total column height of 4.57 m. Table I summarizes the main features of auxiliary equipment. An ethanol-water solution stored in tank D-3 at ambient conditions is preheated with the bottom product in the
0888-5885/92/2631-2547$03.00/00 1992 American Chemical Society
2548 Ind. Eng. Chem. Res., Vol. 31, No. 11, 1992
W Figure 1. Pilot-plant vapor recompression distillation column. Table I. Pilot Plant Auxiliary Equipment item eauipment description C-1 preheater 1-2 shell and tube heat exchanger, horizontally mounted; 1.046-m2 heat-transfer surface C-2 internal reboiler 24 vertical boiling tubes (1/2 in.) steam through shell; 0.525-m2 heat-transfer surface C-3 condenser 1-1 shell and tube heat exchanger, vertically mounted; 0.963-m2 heat-transfer surface C-4 reboiler-condenser 1-4 shell and tube heat exchanger, horizontally mounted; 4.65-m2 heat-transfer-surface C-5 cooler cooling water coil; 0.71-m2 heat-transfer surface J-1 compressor reciprocating cornpressor, 2 single-stage cylinders, 5 HP, 50 m3/h suction capacity at 1 atm
heat exchanger C-1 and fed to the column at the desired tray. The vapors from the top are compressed in the reciprocating compressor J-1 and then condensed within the tubes of the reboilexwondenser C-4. The liberated latent heat provides the boil-up rate to the column. The heat-transfer area of the reboilexwondenser is oversized, thus assuring a liquid seal at the bottom tubes. The saturated liquid expands to the column pressure in the control valve VC-7, and the flashing liquid-vapor mixture flows to the reflux drum F-1. The liquid from the bottom of the tower circulates through the shell of the reboilercondenser C-4 by free convection (thermosiphon),and the partially vaporized mixture reenters the column below the last tray (plate 20). As the capacity of the compreseor J-1is limited, in some operations it is necessary to condense the excess vapors from the top in the condenser C-3.Also, the heat-transfer rate Q,, in the reboiler-condenser C-4 is not sufficient to
provide the necessary boil-up rate to the tower, and the internal reboiler C-2 supplies the auxiliary heat duty 8,. A PDP 11/23 minicomputer (2 MBRAM, 16 A/D inputs, 6 D/A outputs, D.O.S. RSX ll/M, FORTRAN 77) was first programmed to perform data acquisition and pmaxxdng, process control, and operator interface with the pilot plant. All these computer functions were later distributed in a PS-IBM and an OPTOMUX A/D/A interface as shown in Figure 2. The measured process variables transmitted to the computer were feed flow rate, feed temperature, steam pressure, reflux flow rate, cooling water rate, column pressure, bottom level, top and bottom temperatures, temperature of tray number 6 (from the top), distillate and bottom product flow rates, and compressed vapor temperature and pressure. The following variables were manually registered and not transmitted to the data acquisition system: feed concentration, inlet and outlet cooling water temperatures, steam flow rate, compressed condensate flow rate, and compressor work. The outputs from the computer governed the openings of all control valves except valves VC-1 (feed) and VC-7 (liquid expansion), which were manually manipulated. Calibration testa carried out in all orifice-plate flowmeters showed a measurement error of approximately 1%. The noise from temperature transmitters was reduced to 0.05 "C by installing analog and first-order digital filters. The data was later smoothed by means of a smoothing subroutine available at the VAX mainframe computer. The column is started up under conventional distillation (without vapor recompression) keeping constant the loads to the towevfeed flow rate and feed concentration, reflux flow rate, cooling water rate, and downstream steam pressure-which is accomplished by activation of the corresponding control loops. Reflux drum level and bot-
Ind. Eng. Chem. Res., Vol. 31, No. 11,1992 2649
OPTO-22 DATA ACQUISITON AND LOCAL CONTROL 32 K B RAM 16 A/D INPUTS 8
D/A OUTPUTS
BASIC 2.0/3.0 SERIAL RS-232-C
Table 11. Experimental Conditions for Atmospheric Distillation feed flow rate, F 62-80 kg/h feed concn, Xp 35-42 w t % feed temp, TF 42-47 O C reflux flow rate, Lo 35-45 kg/h steam pressure, P, 108.2-151.6 Wa (15.7-22 psia) steam flow rate, W, 34-40 kg/h cooling water rate, F, 275-340 kg/h inlet cooling water temp, T, 15-22 "C feed tray 12 and 14 (from the top) 15-26 kg/h distillate product flow rate, D bottom product flow rate, B 36-63 kg/h top concn, XD 87-91 wt % 7-22 wt % bottom concn, XB reboiier heat-transfer rate, Q, 20.1-25.5 kW condenser heat-transfer rate, Q, 13.4-29.3 kW Table 111. Heat and Mass Balance Steady-State DiscresPancies ~~~~~
PS/2 50-2 IBM DATA STORAGE AND PROCESSING CONTROL CONFIGURATION OPERATOR DISPLAY 640 KB RAM
DM =
F-D-Bxlw
0.7-6.59b
F
FXF - DXD - BXB x 100 FXF FHp + 8,-DHD- BHB- Q c DQ = FHF + Qr DMX =
1.4-6.58 X 100
3.2-8.4%
40 MB HARD DISK
DOS 3.30 TURBO BASIC
I
T I
OPERATOR
Figure 2. Distributed control system.
tom level are also controlled as indicated in Figure 1. The open-loop operation of the pilot plant is conducted for about 2 h up to steady-state conditions. The heat and mass balances are then verified to check for proper operation. Next, step changes are carried out for each of the manipulated variables (steam pressure, reflux flow rate, and cooling water rate), resetting the set points to the controllers. The dynamic responses of the flow and pressure control loops are very fast (3U-60 8 ) and may be considered as step inputs in comparison to the slower dynamics of the tower. The column responses are then freely allowed to evolve to the corresponding conditions of the new steady regime, in order to observe the open-loop dynamic behavior of the process. Distillation under vapor recompression is started once the column has attained steady conventional operation. Previously the reboiilccondenser is preheated, circulating the warm exit cooling water through the shell, and the distillate product through the tubes,by means of auxiliBIy lines not shown in Figure 1. The bypass line around the compressor allow for vapor recirculation during compressor start-up, in order to avoid vacuum conditions in the column. The complete operation takes 3-4 h to reach steady-state conditions. Dynamic runs are then carried out following the same procedure described for conventional distillation. This time the bod-up rate to the tower is changed by varying the reboilemndenser area through manipulation of control valve VC-7. Experimental Results and Discussion A. Conventional Distillation. The stationary and
dynamic behavior of the pilot plant for conventional distillation was first studied, as a means of comparison with vapor recompression distillation. A.l. Atmospheric Distillation. The simplest case to consider is distillation at constant pressure, in order to eliminate the interactive effect of this variable on column responses. Operation at atmospheric pressure is obtained by opening the vent valve at the top of the reflux drum. Pressure fluctuations (relative to ambient pressure) measured at this place were of the order of 20-mm water column. The pressure drop per tray was 26-mm water column, and the total pressure drop on the tower was close to 500-mm water column. Several steady-state runs were carried out in the work of Leiva (1987)and SepiYveda (1989).Working conditione are listed in Table 11, and some results testing the achievement of steady regime are shown in Table 111. The discrepancies from steady state are all positive and higher than the maximum expected experimental errors (2% for mass balance, 4 % for ethanol balance, and 6% for heat balance). It is concluded that the discrepancies are mainly due to measurement errors and to mass and heat losses to the surroundings, as was experimentally observed. In particular the heat balance is very sensitive with the heat-transfer rates at the condenser and the reboiler. Nevertheless, the above results were considered satisfactory for the purpose of pilot plant experiments. Distillation overall plate efficiency calculated from experimental data was 37% average, which compares very favorably with the 40% figure estimated from available correlations (Henley and Seader, 1981). The dynamic identification of the process was obtained from step response data (SepiYveda, 1989). Figures 3 and 4 show bottom temperature responses for step inputs in steam pressure and reflux flow rate, respectively. The methods of Sundaresan and Krishnaswamy (1978)and Sundaresan et al. (1978)were used to fit first- and second-order dynamic transfer functions. The resulting model is the following:
2550 Ind. Eng. Chem. Res., Vol. 31, No. 11,1992
1Pol
I
'"I
u
STEP
INPUT
-0. 16e-2.7S
5.1s + 1
(2.7s
+
1)(0.8s + 1)
lk 1
0.0 17e-3.29
lls + 1
,-.
All time constants and time delays are expressed in minutes, and the gains are given in consistent units. At atmospheric pressure top temperature response becomes indistinguishable from noise, and the temperature of tray 6 was better selected for dynamic characterization. The temperature of tray 6 is also indistinguishablefrom noise for step inputs in feed temperature and feed concentration, and the correspondingtransfer functions are left unknown. The qualitative nature of the dynamic responses are in agreement with the results of dynamic simulations of the process (Rodriguez, 1985;Shchez, 1988) and with the experimental results of Ogunnaire et al. (1983).Response times are of the same order of magnitude as those of simulated responses. The effect of the reflux flow rate and steam pressure on the final evolution of top and bottom temperatures are in accordance with fundamental principles of distillation: as the internal reflux ratio L/V increases, separation in the rectification section becomes higher and separation in the stripping section diminishes. Therefore, an increase in reflux flow rate lowers the temperature profile (negative gains), and an increase in steam pressure raises the temperature distribution (positive gains). The relative gain matrix was determined from the steady-state gains: RGM = 1.27
-0.27 1 4 . 2 7 1.27
J
Therefore dual control is designed with two control loops pairing top temperature with reflux flow rate and bottom temperature with steam pressure. Figure 5 shows succ888ful bottom temperature control performance,cascading the slave pressure controller with the master temperature controller. Top temperature control was not tried at this early experimental stage. A.2. Variable-PressureDistillation. The vent valve at the top of the reflux drum was closed, and the control agents (reflux flow rate, cooling water flow rate, and steam pressure) adjusted to induce operation at variable pressure. Pressure variations were small (13.8-15.8psia)-since the column is not designed to operate at high overpressure or underpressure-but high enough for the tower to exhibit different operation trends as compared to atmospheric distillation. Operation conditions were like those of Table 11,except that feed flow rate changed from 50 to 65 kg/h, feed concentration changed from 16 to 25 wt 9% ,and the feed was introduced at tray number 11 (from the top). The discrepancies from steady state were higher than those of Table 111, being 1.3-6.4% for DM, 6.1-16.8% for DMX,and 15.7-27.6% for DQ. These larger errors may be attributed to several reasons: mass losses increased for overpressure operation, heat losses increased for a higher temperature profile, and experimental errors increased if both F and XFare lowered. De la Fuente (1991)studied the transient behavior of the column for step inputs in reflux flow rate and cooling
r
0
io
10
30
50
LO
60
TIME. min
Figure 3. Bottom temperature response for step input in steam pressure (16.2-18.7 psia).
E
I l
-
2 5
80-
I
m
I 0
Figure 4. Bottom temperature response for step input in reflux flow rate (44-37 kg/h). W
a
20 d
100
O.
0
0
a,*
1
1
3
t
0
I 1
16
21
20
T I M E , min
Figure 5. Bottom temperature control for set-point change (89.7-92.5 "C).
water rate. Figures 6 and 7 show column responses for an increase in reflux flow rate. The top temperature originauy diminished as expected, but later increased, and finally settled at a higher value, in a mode which is opposed to that encountered for constant-pressure operation. The pressure column also increased, but the bottom temperature decreased following the expected response. An analysis of Figures 6 and 7 may explain the occurring phenomena as follows: because of a higher reflux rate the temperature profile decreased, in particular the bottom temperature came down rapidly; the heat-transfer rate at the reboiier Q, = UA(T, - TB)consequently rose, since the condensing steam temperature and heat-transfer surface are kept constant through steam pressure control and level control. As a result, the boil-up rate increased causing the pressure to build up. This last effect was dominant. A dynamic identification of top temperature and pressure responses indicated equal lags (7.4min), but temperature delayed pressure by 6.3 min (dead times are 8.3 and 2 min,
Ind. Eng. Chem. Res., Vol. 31, No. 11, 1992 2661 82.4
r
I
\
f X
I
;:I
I
Y
I
2 8 2 2 4
8
L
I
62.0
I
c
40
3
36
@
34
IA.
,
32
15.8
I I
1
1
I
I
1
I I
2
14.8
1 0
loo0
2000
3000
1
0
1
I
956!
0
TIME, seconds
I
1
I '
I
1000
I
1
2000
1
I
3000
1
LOO0
TIME, seconds
Figure 6. Pressure reapom for step input in reflux flow rate (35-42 43th).
Figure 7. Top and bottom temperature responses for step input in reflux flow rate (35-42 kg/h).
respectively). It is concluded that thermodynamic p-T equilibrium was not instantaneous and that pressure reacted first followed by equilibrium temperature. The experiment was repeated by Sepiilveda (1991) but keeping column pressure under control with cooling water flow. For a sudden decrease in reflux rate, top temperature rose while pressure remained constant, a result which is consistent with the observed trend at atmospheric pressure. Figure 8 shorn the experimental results of De la Fuente (1991) for a positive step change in cooling water flow rate. The presence of minimum responses exhibiting overshoots is due to two opposing effects: a decrease in pressure caused by a higher vapor rate condensation, which in turn diminishes the temperature profile; and an increase in boil-up rate at the reboiler because of bottom temperature decrease, as previously explained. The cooling water effect was faster and stronger than the second, and both variables finally decreased. The fitted dynamic models are
small (0.1 OC/psi and 0.05 psi/psi). Again, boil-up rate increased vigorously at the beginning, but due to bottom temperature rise, the heat-transfer rate at the reboiler was later reduced. Time constants were about 1-2 min for the fmt effect and 6-8 min for the second effect. The bottom temperature showed first-order dynamics and high gain (2 OC/psi). Accordingly, steam pressure may be selected as an effective agent for bottom temperature control. The selection of steam pressure control as a mean to set heat input in the reboiler has proven troubleaome. Staam pressure was chosen because it is the easiest and fastest variable to be measured and controlled. In order to maintain thia advantage and to properly set the heat duty, a feed-forward scheme is being considered for future work the temperature difference T, - T Bwould be set and, by measurement of bottom temperature TB, the steam temperature would be determined, converted to steam pressure, and used aa set point to the pressure controller. Singleloop control runs for top temperature or pressure were carried out (Septilveda, 1991) manipulating the reflux flow rate or the cooling water rate. The corresponding master controller cascaded the respective flow controller shown in Figure 1. Of all four pairinge the pressure-cooling water loop performed the best (Figure 9) and again instantaneous p-T equilibrium was observed. The other three control combinationswere not satisfactory, as long as the second variable was floating and interacting with the controlled variable. Temperature control with reflux flow rate was very poor since temperature lagged pressure, as already discussed. Mdtilmp distillation control was done, pairing pressure to cooling water and top temperature to reflux rate. This pairing has a negative relative gain, g = -13.6, therefore
-1
T&) = 4.01[ 5 - 4 FJs) 3s+l 5 s + 1
OC
kg/h
(3)
Both transfer functions have equal lags,which indicatea that pressure and top temperature are in phase; that is, thermodynamic p T equilibrium is almost instantaneous, as can be observed from Figure 8. The effect of steam pressure changes on column transients was studied by Sepiilveda (1991). Maximum responses with large overshoots were obtained for top temperature and pressure, but the steady-state gains were
2552 Ind. Eng. Chem. Res., Vol. 31, No. 11, 1992 16.0 1 1 I
I
.-
0
I
I I
rD
*a
80
79-
W
a c
a
n
i r
E %
w 82.4 -
>. 8 2 . 3 -
a
82.2-
b
1000
I
2000
I
I
1
3000
I
I
!
18-
0
ih
rD
-I
3
8
!I
c
- lI
E
8
-
I
I
02.1-
I
I I
4 3
1I ME , seconds
Figure 8. Pressure and top temperature responses for step input in cooling water flow rate (150-325 kg/h).
81.% !
0
I
I
I
1
I
SO0
1000
.
I
1500
I
2 000
TlME, seconds 15.6
1
I
14.6
0
500
1000 1CAO 2000 TI M E , seconds
290
3000
Figure 9. Single-loop pressure control with cooling water flow rate (set-point changes).
was highly interactive, and care was taken to change the action of the proportional band to the controllers when the second loop was opened or closed. However, this pairing proved to perform very satisfactorily as can be seen in Figure 10. Temperature control improved when pressure was tracked. Also the pairing proved to be robust for steam pressure disturbances. Dual temperature control was not attempted. A 3 X 3 control pairing, adding the third loop TB-P,, has been dynamically simulated (OyarzCm, 1991) and the configuration has been shown to be feasible.
Figure 10. Multiloop pressure and temperature control (set-point change in pressure). Table IV. Experimental Conditions for Vapor Recompression Distillation feed flow rate, F 69-80 kg/h 30-41 wt % feed concn, XF reflux flow rate, Lo 36-40 kg/h distillate product flow rate, D 19-23 kg/h bottom product flow rate, B 49-52 kg/h top concn, XD 81-85 wt % 10-17 wt % bottom concn, XB column pressure, P 101.3-128.9 kPa (0-4 psig) feed tray 11 (from top) pressure of compressed vapors, P,, 253-294.3 kPa (22-28 psig) 101.5-103.5 "C temp of compressed vapors, T,, compressed vapors flow rate, V, 51.6-53.3 kg/h 9.7-17 "C temp drop in reboiler-condenser, AT,, = T, - TB reboiler-condenser heat duty, Q,, -15.2 kW compressor work, W, 1.8-3.75 kW 8.7-13.2 kW reboiler heat-transfer rate, 8, 0-3.7 kW condenser heat-transfer rate, Q, 1.4-6.2% mass balance discrepancy, DM ethanol balance discrepancy, DMX 2.1-13.1% energy balance discrepancy, DQ 10-2095
B. Vapor Recompression Distillation. The experimental procedure described in a preceding section was followed to set up vapor recompression operation. Cooling water flow rate-if necessary-was adjusted to keep column pressure under a prescribed value, normally 2 or 4 psig. Steady-state and transient responses were observed. B.l. Steady-StateResults. In the works of G o W e z (1988)and A& (1989),several trial rum were conducted in order to collect some preliminary results. A summary
Ind. Eng. Chem. Res., Vol. 31, No. 11, 1992 2853 Table V. Mass Balances in Steady-State Vapor Recompression Distillation expt A at % lift VC-7 expt B at % lift VC-7 63 80 80 60 57 57 61 61 B, kdh 25.4 33.0 44.4 F*kgJh 35.3 16.4 29.9 27.0 20.2 D,kgJh 32 32 32 LmkgJh 32 3.63 2.98 1.7 1.1 DM, % TF,OC 44 44 44.2 44 "8, "C 80.3 81.4 82.3 81.2 TB,O C 92.4 94.3 93.1 90.8 Xp,wt % 38 38 39 39 XD, wt % 81.6 78.0 77.4 81.0 XB, wt % 14.0 6.1 7.5 17.0 P,PSk 1 2 2.3 1.1 DMX, % 1.1 -14.8 1.8 12.4 feed tray 11 11 11 11
of experimental conditions is given in Table IV. Alcohol losses through the crankcase of the compressor were noticed (as high as 0.7 L/h), which in addition to losses in the column contributed to mass and energy balance discrepancies. The energy balance discrepancy is defined as in Table 111,except that the compressor work W,is added to the energy inputs. The energy consumption for vapor recompression distillation is given by Q, + W,.If the operation is assumed to be conducted under conventional distillation-by opening the recompression loop in Figure 1 and assuming that the heat duty Q,, is provided from an external source-and processing the same feed, then the energy requirement would be Q, + Q, in order to reach equivalent separation conditions. Hence, the energy saving due to vapor recompression as compared to conventional distillation becomes (Q,+ Q,.J - (8,+ WJ, or in a percent basis, Qrc - Wc % energy saving = 100 (4) Q,, + Qr Vapor recompression operation is achieved, applying little work at the compressor and a heat input at the internal reboiler of the order of magnitude of Q, Therefore, the energy saving accounts for about 50% of the total heat requirement. In fact, the reduction of energy consumption varied from 47 to 56%. This is one of the most relevant results because of its economical importance in future applications. It was very difficult to design vapor recompression runs that reproduce conventional distillation data (flow, temperature and composition of the feed stream, and flow and cornposition of the top and bottom products) as a way of real comparison of the two configurations. However, Acufla (1989)compared some of his reported vapor recompression data with similar separations in conventional distillation carried out by Septilveda (1989)and concluded that energy savings ranged from 38 to 52%. Other results computed from experimental data are the following: overall plate efficiency averaged 45%, the vaporized fraction after the expansion valve VC-7 was about 11%, and thermodynamic efficiency (second-lawanalysis) was of the order of 16% in comparison with 6.5% for conventional distillation. The effect of different operation conditions of the reboiler-condenser upon column performance can be determined by manipulating the expansion of the compressed condensate in valve VC-7. For a valve opening the liquid flow rate increased and the liquid level within the tubes decreased, leaving more uncovered heat-transfer surface for vapor condensation; vapor temperature and pressure both fell, conserving thermodynamic p T equilibrium. For
Table VI. Heat Balances in Steady-State Vapor Recompression Distillation expt A at % expt B at % lift VC-7 lift VC-7 63 80 80 60 Qn kW 11.16 11.28 11.16 11.25 2.92 3.14 2.95 Wc,kW 3.15 Q,, kW 0 0 0 0 PIC, P S k 46 47 35 34 110 108 TI,, "C 118 118.5 17.98 17.77 Qrc, kW 15.31 14.73 15.4 14.6 DQ, % 15.3 14.8 51.5 50.5 45.3 46 energy saving, % 15.7 14.9 27.7 AT,~:-OC 25.6 1145 1193 UA. W/K 598 532 63 62 53 285.4 286.6 218 Table VII. Steady-State Gains for Vapor Recompression Distillation gains variable expt A expt B units D 0.571 0.53 (kg/h)/% lift B -0.582 -0.51 (kg/h)/% lift P 0.059 0.060 psi/% lift PIC -0.647 -0.65 psi/% lift TB 0.118 0.115 "C/o/, lift Tfi 0.065 0.06 "C/o/, lift XB -0.465 -0.475 wt % / % lift XD -0.212 -0.18 wt % J % lift
pressures below 28 psig no changes on column responses were observed after these manipulations, because the transferred heat Q,, showed little change (see Table IV). Jimhez (1990)conducted two carefully done and different experiments at higher pressures, a valve opening from 63 to 80% valve lift (experiment A) and a valve closing from 80 to 60% lift (experiment B). The experimental conditions and results are presented in Tables V and VI. Examining the numbers, it can be seen that column operation was affected by the liquid expansion. Comparing the experiments, it is also noted that column responses changed almost in symmetrical amounts, which is better confiied observing the steady-state gains for both experiments in Table VII. This can be explained because the heat duty Q, experienced a change of about 12% (see Table VI),the total heat input Q,+ Q , changed by lo%, and the process evolved within its linearized region of operation. In fact the process was subjected to a square pulse i n p u t 4 3 to 80 to 60% valve lift-and it reached two successive steady states. Static linearity cannot be inferred at all, since a third symmetrical valve movement below 63% lift would have been needed to prove such an assertion. The overall mass balance is satisfied with little error. In Table VII it can be noted that the sum of the gains for D and B almost cancelled, a consistent result since the feed F is kept constant in each experiment. Also, the sum D + Lofrom Table V is very like the figure V,of Table VI, because all top vapors were compressed (cooling water was not used). An interesting result is that the heat transferred at the reboiler-condensr increased after valve opening. The heat-transfer surface practically doubled but the temperature difference changed by less than a factor of 2, and the heat duty augmented. This is in accordance with the rise in vapor flow rate and the increase in vapor latent heat of condensation. B.2. Dynamic Results. The transient responses for experiments A and B were characterized as first- and second-order transfer functions. Figure 11 illustrates the pressure evolution of compressed vapors at the reboilel-
2554 Ind. Eng. Chem. Res., Vol. 31, No. 11, 1992 46
a .? W W 2
5 m:
0
. 43
I"
,
a
LLm O W 40-
32
I n 0
m u W a a
37-
3L -F 0
I
500
2030
laxl
TIME,
3000
seconds
Figure 11. Reboiler-condenser pressure response for step closing of the expansion valve (80 to 60% lift). Table VIII. Time Constants for Vapor Recompression Distillation (rnin) expt A expt B variable 7, T2 0 71 72 6 D 9.57 1.33 10 1.66 B 10.23 4.72 12 5.33 TB 4.5 3.82 3.66 6.66 6.5 2.83 T6 5.0 4.66 3.0 6.33 5 4.16 P 4.6 4 4 6.3 5 4 prc 9.15 2.8 10.83 2.5
condenser after valve VC-7 was closed. Table VI11 shows the resulting time lags 71 and 72and time delays 6'. It can be seen that time constants for experiment B are consistently higher than those for experiment A. The differences are small and may be explained because it was faster to lose pressure opening the valve than to build up pressure when the valve was closed. Response times for both experiments are slightly different, and a quasi-linear dynamic behavior was obtained as a result, which again may be explained because of a small departure in heat duty at the reboiler condenser. Time constants in Table VI11 for top and bottom temperature responses are higher than those corresponding to atmospheric conventional distillation (see effects of P, on T6and TBin eq 1). The mutual interaction between the tower and the reboiler-condenser enlarged the inertia of the whole process. The actual liquid inventory in the reboiler-condenser during operation is about 90 L against 45 L as total hold-up of the tower. Response times estimated from Table VI11 indicate that all responses had to adapt to the dominant inertia of the reboiler-condenser. An exception was the smaller time delay of the distillate product because the flashed mixture from VC-7 discharged directly in the reflux drum, whose level was controlled precisely acting on the distillate flowrate. Interactions can be exemplified considering an opening of valve VC-7: pressure at the reboiler-condenser decreases but column pressure rises due to the increase in heat duty. Discharge pressure at the compressor consequently increases, so attenuating the pressure decrease at the reboiler-condenaer. This effect of column pressure on reboiler-condenser operation is also found if the cooling water flow rate is altered. In an experiment carried out by Barrientos (1991))valve VC-6 was opened while valve VC-7 was kept unmoved. Total heat at the bottom of the column was kept constant (decreasingthe heat duty Q,to compensate for reboiler-condenser effects). In this way the influence of varying cooling water flow rate was isolated from heat inputs but not from process interactions. The resulting gain for column pressure was -0.025 psi/(kg/h) which compares well with the gain of the first effect in eq 2, but the time lag increases to 11 min instead of 3 min in eq 2. The decrease in column pressure was now followed
by a reduction in pressure at the reboiler-condenser causing the heat rate of condensation to get higher. The vapor flow rate through the compressor remained fixed as the condenser took account of the excess vapors from the top of the column. This experiment helped further understanding of the occurring phenomena in vapor recompression distillation. Control of the pilot distillation plant under vapor recompression has not been attempted yet, but it is considered as future work in a new stainless-steel packed distillation tower which is under installation. Compressor control configurations in vapor recompression distillation have been studied by Miihrer et al. (1990). Of all four alternatives discussed in their paper (variable reboilercondenser area, compressor variable speed, compressor bypassing, and compressor suction throttling), variable condenser area can be easily implemented in the pilot plant, closing the loop between valve VC-7 and the bottom temperature controller programmed in the OPTOMUX A/D/A process-computer interface. Variable-speed compressor control was disregarded, even when an ac variable-frequency driver was available, because the compressor is working at its capacity limit and any substantial increase in speed may result in severe mechanical damage. Compressor control through bypassing or suction throttling can also be accomplished reinstalling control valve VC-7 in the appropriate line. These last options should be first studied in open-loop control in order to characterize plant dynamic responses for control purposes. A final comment refers to the reciprocating compressor, whose characteristic line head v8 flow is a very sharp line which moves to the right with increasing speed, but the analysis of the operation point follows a similar treatment to that discussed in the referred paper, in which a centrifugal compressor was considered. Conclusions The vapor recompression pilot distillationplant has been successfully operated, providing useful steady-state and dynamic data for future control work. The heat and maw balances are satisfied within normal margins. Despite the errors asaociated with heat and mass losses in the column and in the compressor crankcase, it is concluded that the pilot plant properly reached steady-state conditions. The most remarkable result is the reduction in energy consumption as compared to conventinal distillation,which ranged from 45 to 56%. Thermodynamic efficiency was higher as well, 16% against 6.5%) which also indicates better energy utilization. Heat interactions at the reboiler-condenser slowed down dynamic responses. The process showed steady-state consistency for varying position of the expansion control valve, and the transient behavior exhibited small departure from linearity, as vapor decompression resulted a little faster than compression. Manipulation of reboiler-condenser area has demonstrated to be a feasible control agent of the pilot plant, if the compressed vapors pressure exceeds 28 psig. Acknowledgment We acknowledge the National Fund for Scientific and Technological Development (FONDECYT),Chile, for financial support through Project 89-0722. Literature Cited Acuiia, 0. A. Experimental Dynamic Identification of Vapor Recompression Distillation. Chem. Eng. Thesis, University of Concepciijn, 1989.
Ind. Eng. Chem. Res., Vol. 31, No. 11, 1992 2666 Barrientos, R. Dynamica Study of Vapor Recompreeaion Distillation. Chem. Eng. Thesis, University of Concepci6n, 1991. Bjam, I,; G r h , U.; Stram, K. A Study of a Heat Pump Distillation Column System. Chem. Eng. Process. 1991,29,185-191. Brousee, E.; Claudel, B.; Jalut, C. Modeling and Optimization of the Steady State Operation of a Vapor Recompression Distillation Column. Chem. Eng. Sci. 1985,40,2073-2078. Carte, R.; Kovacic, A; Tola, G. Separation of Ethyl benzene-Xylenes Mixtures by Distillation: A Possible Application of Vapour Recompression Concept. Chem. Eng. Commun. 1982,19,157-165. Collura, M. A.; Luyben, W. L. Energy-SavingDistillation Designa in Ethanol Production. Znd. Eng. Chem. Res. 1988,27,1686-1696. Danziger, R. Distillation Columns with Vapor Recompression. Chem. Eng. Prog. 1979, 75 (9),58-64. De la Fuente, G. E. Dynamic Identification and Control of a Conventional Distillation Column. Chem. Eng. Thesis, University of Concepcibn, 1991. Ferre, J. A.; Castella, F.; Flores, J. Optimization of a Distillation Column with a Direct Vapor Recompression Heat Pump. Znd. Eng. Chem. Process Des. Dev. 1985,24,126-132. Finelt, S. Better C3 Distillation Pressure. Hydrocarbon Process. 1979,58 (2),95-98. Flores, J.; Castella, F.; Ferre, J. A. Recompression Saves Energy. Hydrocarbon Process. 1984,63,59-62. Gondlez, F. E. Steady-State Simulation and Start-up of a Vapor Recompression Distillation Pilot Plant. Chem. Eng. Thesis, University of Concepcibn, 1988. Gopichand, S.; Omiyedi, T. 0.; Kasprzycki, J.; Devotta, S. The Economics of Heat Pump Assisted Distillation Systems-11. Analysis of Ethanol-Water Mixtures. J. Heat Recovery Syst. 1984,4 (a),271-280. Henley, E.; Seader, J. D. Equilibrium-Stage Separation Operations in Chemical Engineering, 1st ed.; Wiley New York, 1981;p 510. Jimhez, A. A. Experimental Dynamic Identification of Vapor Recompression Distillation. Chem. Eng. Thesis, University of Concepcih, 1990. Leiva, J. C. Digital Control h a y s in a Distillation Column. Chem. Eng. Thesis, University of Concepci6n, 1987. Loken, Per A. Process Intagration of Heat Pumps. J. Heat Recovery Syst. 1985,5(l), 39-49. Lynd, L. R.; Grethlein, H. E. Distillation with Intermediate Heat Pumps and Optimal Sidestream Return. AZChE J. 1986,32 (a), 1347. Meili, A. Heat Pumps for Distillation Columns. Chem. Eng. Prog. 1990,60-65. Meili, A. Stuecheli, A. Distillation Column with Vapor Recompression. Chem. Eng. 1987,94,133-143. Menzies, M. A.; Johnson, A. I. Steady-State Modelling and Parametric Study of a Vapor Recompression Unit. Can. J. Chem. Eng. 1981,59,487-491. Mostafa, H. A. Thermodynamic Availability Analysis of Fractional ,
Distillation with Vapor Recompression. Can. J. Chem. Eng. 1981, 59,487-491. Mtihrer, C.; Collura, M. A.; Luyben, W. L. Control of Vapor Recompression Distillation Columns. Znd. Eng. Chem. Res. 1990,29, 59-7 1. Null, H. R. Heat Pumps in Distillation. Chem. Eng. Prog. 1976,73, 58-64. Omiyedi, T. 0.;Kasprzcki, J.; Watson, F. A. The Economics of Heat Pump Assisted Distillation Systems-I. A Design and Economic Model. J. Heat Recovery Syst. 1984a,4 (3),187-200. Omiyedi, T. 0.; Parande, M. G.; Kasprzycki, J.; Devotta, S. The Economics of Heat Pump Assisted Distillation System-111. A Comparative Analysis on Three Alcohol Mixtures. J. Heat Recovery Syst. 1984b,4 (4), 281-286. Omiyedi, T. 0.; Parande, M. G.; Supranto, S.; Kasprzycki, J.; Devotta, s. The Economics of Heat Pump Assisted Distillation Systems-Part 4. Experimental Assessment with MethanolWater Mixtures. J. Heat Recovery Syst. 1985,5 (6),511-518. Ogunnaire, B. A.; Lemaire, J. P.; Morari, M. Advanced Multivariable Control of Pilot-Plant Distillation Columns. AZChE J. 1983,29 (4),632. Oyarzfn, N. Dynamic Simulation and Control of a Binary Distillation Column. Electr. Eng.Thesis, University of Concepcibn, 1991. Papastathopoulou, H. S.;Luyben, W. L. Control of a Binary Sidestream Distillation Column. Znd. Eng. Chem. Res. 1991, 30, 705-713. Quadri, G. P. Use of Heat Pump in P-P Splitter, Part 1: Process Design. Hydrocarbon Process. 1981a,60, 119-126. Quadri, G. P. Use of Heat Pump in P-P Splitter, Part 2: Optimization. Hydrocarbon Process. 1981b,60, 147-151. Rodriquez, L. Dynamic Simulation of a Distillation Column. Chem. Eng. Thesis, University of Concepcibn, 1985. Roffel, B.; Fontein, H. J. Constraint Control of Distillation Processes. Chem. Eng. Sci. 1979,34,1007-1018. Shchez, A. Dynamic Simulation and Control of Lumped Parameters Systems. Chem. Eng. Thesis, University of Concepcibn, 1988. Septilveda, L. A. Dynamic Behavior of a Conventional Distillation Column, M.Sc. Thesis, University of Concepcibn, 1989. Septilveda, R. M. Experimental Study of Bivariable Pressure and Temperature Control in a Distillation Column. Chem. Eng. Thesis, University of Concepcibn, 1991. Sundaresan, K. R.; Krishnaswamy, P. R.Estimation of Time Delay Time Constant Parameters in Time Frequency, and Laplace Domain. Can. J. Chem. Eng. 1978,56,257-262. Sundaresan, K. R.; Chandra Prasad, C.; Krishnaswamy, P. R. Evaluating Parameters from Proceas Transients. Znd. Eng. Chem. Process. Des. Dev. 1978,17,237-241.
Received for review December 16,1991 Revised manuscript received June 12,1992 Accepted July 1, 1992
