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Ind. Eng. Chem. Process Des. Dev., Vol. 18, No. 1, 1979
Mathematical Modeling of THPOH Ammoniation Julius P. Neumeyer,’ Esmond J. Keating, and Nestor
B. Knoepfler
‘
Southern Regional Research Center, New Orleans, Louisiana 70 179
A mathematical model of the ammoniation step in the THPOH-NH, flame retarding finishing process for cellulosic textiles has been developed. The basic features of commercial ammoniation systems are considered in describing physical phenomena that occur during the curing operation. The model demonstrates that thermodynamics of ammonia-water vapor-liquid equilibrium control the operating characteristics of the chamber. Increased flame-retarding polymer formation is predicted by the model, with (1) increases in ammonia flow rate, (2) decreases in chamber operating temperatures, and (3) cooling the fabric prior to curing. These predictions agree with experimental observations. The role of moisture in controlling the rate of conversion is further investigated. A procedure for investigating effects of different fabric constructions is suggested.
Introduction Although several different flame-retardant finishes for cotton textiles have been available, one has continually maintained keen industrial interest. In the process, a tetrakis(hydroxymethy1)phosphonium salt is partially neutralized with sodium hydroxide solution. The mixture is applied to fabric, which is then dried to a specified moisture content. The dried fabric is exposed to gaseous ammonia in a suitable enclosure to form an insoluble polymer on and within the cotton fibers. Ammoniation is followed by a hydrogen peroxide oxidative wash to increase durability and minimize odor of the flame retardant finish by essentially assuring that the trivalent phosphorus contained in the polymer is converted to the pentavalent state. A slightly different procedure, also referred to as the THPOH-NH3 process, is to pad the fabric with a condensation product of the phosphonium salt and a nitrogen compound. The drying, curing, and oxidation procedures remain essentially unchanged. A harsh fabric hand can readily be remedied with sanforization. Ammoniation of the fabric, often referred to as the curing operation, is often cited as the most difficult step of the process. The effectiveness of curing depends on numerous factors, including ammonia concentration and temperature in the chamber, construction, moisture content, and temperature of the fabric, and length of time the fabric is exposed to the curing atmosphere. Drake et al. (1969) described the usage of a modified tensionless open-width washer as pilot plant ammonia cure equipment. An enclosure was built over the washer to form the reaction chamber. Two pairs of slotted pipes were located inside the chamber so that the slots are in intimate contact with the fabric as it proceeds through the reactor. Ammonia vapor is forced out of the pipes and through the fabric. Two oscillating fans in the enclosure provide circulation of the chamber atmosphere. An exhaust port is located a t the top of the chamber to remove as much excess ammonia and reaction byproducts as possible. Perkins et al. (1974) later described a vertical reaction chamber wherein fabric enters at the bottom and travels upward, passing against two pairs of opposing slotted pipes before exiting at the top. Ammonia is forced through the fabric. Steam coils in the reactor walls maintain the desired chamber temperature and reduce condensation on the walls. Corrugated walls help to create turbulence in ‘One of the facilities of the Southern Region, Science and Education Administration, U.S. Department of Agriculture.
the chamber atmosphere. Excess ammonia and the reaction byproducts exit from the top of the chamber. Wagner (1974, 1976) described a curing apparatus consisting of upper and lower portions contained in a single housing. Ammonia gas is fed through a perforated plate located in the upper portion (referred to as gas treatment chamber). The fabric enters and leaves through the lower chamber. A partition between the upper and lower chambers is designed to permit the fabric to travel up into the treatment chamber, then back down into the lower chamber before exiting the housing. This configuration minimizes leakage of ammonia into the surrounding atmosphere. An exhaust outlet is located in the lower portion. It is further suggested that the housing might be provided with a cooling system to regulate the temperature of the chamber atmosphere. A liquid discharge port is positioned in a convenient place to remove accumulated condensate. Getchell et al. (1976) described an ammoniation apparatus comprised of two vertical reaction chambers arranged within a larger enclosure designed to control any noxious fume leakage from the reactors. The padded, dried fabric enters the enclosure and travels through the first reactor, wherein ammonia gas is forced through two pairs of perforated pipes, arranged so that gas flow impinges onto the moving fabric. The fabric exits the first reactor and then passes directly into an identical second reactor, where ammonia gas is again directed upon the fabric. The reactors are each equipped with evacuation means at both ends to collect excess ammonia and reaction byproducts. The apparatus is also designed to permit an area for holding the fabric within the enclosure for a desired time before it leaves the system. Heaters are located at selected sites within the enclosure to minimize moisture condensation. A drainage port permits removal of accumulated condensate. Carter (1977) described an apparatus for ammonia curing wherein the fabric is passed around the perimeter of a perforated rotating drum. The flow of gas is controlled so that substantially all of the flow passes directly through the fabric. Undue accumulation of condensate is avoided by an exhaust system, which tends to draw water vapor from the reaction site. A slanted ceiling configuration promotes condensate runoff and prevents dripping onto the fabric. The ammonia cure systems described by Perkins and Getchell are similar in that both are vertically orientated, plug-flow reactors wherein ammonia is forced directly upon the fabric from slotted or perforated pipes. Drake and Carter also utilized “forced diffusion” in their chamber
This article not subject to U.S. Copyright. Published 1978 by the American Chemical Society
Knd. Eng. Chem. Process Des. Dev., Vol. 18, No. 1, 1979
\:-
109
H O
"3
IENDO)
POLYMER
POLYMER
"'.\ /
2 do" Or
-?I-
t
CONDENSATE
Figure 1. Hypothetical ammonia cure apparatus.
designs. The system described by Wagner incorporated flooding of the treating chamber, with less forcing action, thus permitting "ordinary diffusion" to provide the primary driving force for transferring ammonia to the fabric surface. Lambert (1976) modified Perkins' vertical ammoniator to provide multiple fabric passage through the gaseous atmosphere. He then flooded the chamber with ammonia, rather than forcing the gas through the fabric, and further verified that ordinary diffusion alone can provide adequate transfer of ammonia to the fabric. Although research on the finish and its commercial application have been extensive, the design of the ammoniation chambers built and the processing methods used have been empirical. T o aid further investigation and development of the ammoniation process, the basic features of the systems previously described have been incorporated into a mathematical model describing physical phenomena that occur during the cure step. Solving the mathematics of the system allows investigation of the interplay of important process parameters. Such information often lessens the number and extent of experimental determinations needed for process optimization.
Model Development The system used in defining the mathematical model is illustrated in Figure 1. A controlled flow of ammonia a t a specified pressure and temperature enters the reaction chamber via the feed stream. The reaction byproducts and any excess ammonia are removed from the chamber as vapor exhaust and/or liquid condensate. Specifying the chamber temperature, along with the assumption of vapor-liquid thermodynamic equilibrium, fixes the ammonia-water concentrations of these exiting streams. Heat losses through the chamber walls are considered with the Q, term. Additional heat, Q,, can be transferred to or from the chamber via the internal heat exchanger. The fabric enters the chamber containing phosphorus, water, and other solids remaining from padding and drying operations. What occurs once the fabric is exposed to the chamber atmosphere is shown in Figure 2. Ammonia is transferred via gaseous diffusion to the fabric surface. As contact occurs, the ammonia is dissolved by water remaining in the fabric. This is an exothermic process and the energy released raises the fabric temperature. Once in the fabric substrate, ammonia reacts with the phosphorus compound to form the flame-retarding polymer. The net thermal effect of the polymerization reaction is exothermic, and the heat released further contributes to heating the fabric. This thermal energy is removed from the fabric by providing heat for evaporating moisture from the substrate and, to a lesser extent, via convective transfer to the chamber atmosphere. Excess moisture is available for evaporation as it is a byproduct of the polymerization reaction. The fabric exits the chamber containing polymer, unreacted phosphorus, ammonia, water, and the entering
H20
FABRIC
SOLIDS
1
HEAT
Figure 2. Scenario of physical and chemical phenomena occurring on fabric during curing.
solids. These variables, as well as the fabric temperature, are dependent upon exposure time in the chamber. The exposure time is directly proportional to the distance traveled by the fabric in the chamber and can be calculated for any given production rate. Equations 1, 2, and 3 present the mass balances ap-
(3)
plicable to the fabric. The change in water content, eq 1 over a differential length of fabric, d(H,O)/dl, is affected by both reaction rate and mass-transfer phenomena. The water content will increase with increased reaction rate, since 3 moles of water are theoretically produced for each mole of phosphonium compound reacted. It will decrease with a corresponding increase in the mass-transfer rate a t which water leaves the fabric via evaporation. I t is apparent in the mass-transfer term of eq 1that an increase in the rate of evaporation can result from (1)a decrease in the partial pressure of moisture inside the chamber, ( 2 ) an increase in the vapor pressure of water on the fabric, (3) an increase in the mass-transfer coefficient, and (4)an increased interfacial area. In contrast, the fabric's ammonia content, eq 2 , decreases with increased reaction rate and increases with increased mass transfer. The change in phosphorus content over a differential length of fabric, defined by eq 3, is equal to the rate of disappearance of phosphorus, as is defined by the reaction rate, eq 4. The rate of reaction is assumed to be pro-
R = hr(P)("J
(4)
portional to the concentrations of phosphorus and ammonia present in the fabric substrate. The reaction rate constant is assumed to be independent of temperature and is assigned a value that assures the reaction is not the rate-determining step of the process. This gross oversimplification of the polymerization reaction is permitted by assuming the reaction to be very rapid with respect to the other transport phenomena occurring during the overall curing step. If the polymerization reaction were the rate-determining step during the cure process, a more adequate reaction expression would be required. The energy balance defining the fabric temperature is presented by eq 5. This equation defines how increased
110
Ind. Eng. Chem. Process Des. Dev., Vol. 18, No. 1, 1979
Table I. Variables Considered by Mathematical Model fabric variables speed weight width temperature interfacial area phosphorus content moisture content other solids
a
chamber variables length of fabric path coollheat load heat loss thru walls feed variables rate concentration temperature physical constants heat of reaction reaction rate constant
60 yd/min 8.8 o z l y d 2 1 Ydo 1 2 2 Fa 1.0y d z / y d 2yda 0.03 lb/lb of fabric 0.15 lb/lb of fabrica 0.22 lb/lb of fabric
10.9 yd varied 0 varied 100% NH, 1 0 5 "F 20 350 cal/g-mol 1.5 X lo-' g-mol/cm' cm s
Otherwise varied.
ammonia transfer and reaction rate both contribute to increased fabric temperature. It also defines how cooling of the fabric is dependent upon increased rates of moisture evaporation and convective heat transfer. The heat- and mass-transfer coefficients of eq 1, 2 , and 5 significantly affect the rate of transport between the gaseous atmosphere and the fabric surface. These coefficients normally require experimental determination for the specified system. However, as a first approximation, the transfer coefficients used in this simulation are derived from an empirical correlation of experimental data relating transfer coefficients for a smooth, flat plate in tangential flow (Schlichting, 1955), by means of the Chilton-Colburn analogies (1934) relating mass, heat, and momemtum transport. The heat of reaction, HR,used in eq 5 is the average of the net heat effect of the overall polymerization reaction as reported by Hendrix (19721,Trecek (19761, and Lambert (1976). The fabric materials and energy balances are coupled to the chamber materials and energy balances through transfer of moisture, ammonia, and heat between the fabric substrate and the chamber atmosphere. Inside the reaction chamber, a mass balance for the ammonia component, eq 6, states that under steady-state operating 1
F
- V y - LX =
(MW)Wf& k a ( p - P*)",
dl
(6)
conditions the ammonia feed input minus the ammonia output through the vapor exhaust and liquid condensate streams must equal the amount of ammonia transferred to the fabric. Similarly, a steady-state material balance on the water component, eq 7 , states that the water leaving through the
exhaust and condensate streams must equal the amount of water evaporated from the fabric. Equation 8 reminds y,x = f(T,)
(8)
us that vapor-liquid thermodynamic equilibrium is assumed, thus defining ammonia concentrations of the vapor and liquid phases as a function of chamber temperature. The chamber temperature is defined by eq 9, a steady-state
FhF - Vhv - LhL (MmiyWfJ1(ka(p
-
P*)NH3")
(MW)HW,S0 ' ( k a ( p * - P)H&H) WfL1hu(Tf- T,) dl = Qr
dl + dl
+
+ Qw = Q
(9)
CHAYBER TEYPLRbTWIL-'F
Figure 3. Steady-state operating lines at various ammonia/phosphorus molar ratios.
energy balance around the chamber. This relation establishes that the net thermal effect associated with all material streams entering and leaving the chamber is balanced by the heat transferred via the internal heat exchanger, in combination with heat loss through the chamber walls. Solving eq 1-9 permits defining how variations in certain parameters can affect operating characteristics of the real process. The causes and effects of observed phenomena can more readily be identified by correlation with mathematical predictions. A few of the many variables capable of being studied with the model are listed in Table I. Note that several variables were constrained a t fixed values to facilitate clarity of this presentation.
Results and Conclusions A large number of computed solutions produced calculated data that are graphically presented in Figure 3 to illustrate steady-state operating lines for chamber operation under various ammonia/phosphorus molar ratios. The ordinate indicates net heat added to or removed from the chamber. For a given heat load and a specified NH3/P molar ratio, the chamber attains a steady-state operating temperature defined by the abscissas. Each operating line exhibits discontinuity a t two temperatures. (The lower temperature discontinuity point is off the scale for ratios greater than 2.5:l.) Between these temperatures the chamber operates under equilibrium conditions, with both vapor exhaust and liquid condensate exit streams. A t temperatures above the range of equilibrium operation (above the upper discontinuity temperature), condensation cannot occur and vapor alone is exhausted from the chamber. Operating below this range permits only liquid condensate to exit the chamber. For a specified NH,/P molar ratio, the chamber conditions can be controlled
Ind. Eng. Chem. Process Des. Dev., Vol. 18, No. 1, 1979
Table 11. Steady-State Chamber Temperatures at Zero Cooling NHJP molar ratio
-
111
temp, " F
12 8 4 2 1.75
_ . _ _ _ _ ~
133 143 159 182 188 COOLING D-R 8 TOMS 5 7 TONS
CHAMBER TEMPERATURE 167.F 131 T
NH, P MOLAR RATIO.2 5 1
RESIDENCE TIME (SEC.1
Figure 5 . Effect of cooling on phosphorus conversion at 2.5:l ammonia/phosphorus molar ratio. 143 *F
182.F
RESIDENCE TIME (SEC)
Figure 4. Effect of amrnonia/phosphorus molar ratio on phosphorus conversion.
anywhere along the operating line by varying the cooling capacity of the internal heat exchanger. For example, at 8:l molar ratio with 4 tons of cooling, a steady-state chamber temperature of about 122 O F is attained. The process operates under vapor-liquid equilibrium conditions, and both vapor exhaust and liquid condensate occur. As the cooling is reduced down to zero the chamber temperature rises along the 8:l operating line, eventually reaching another steady-state value at 143 OF. This operating point is still within the vapor-liquid equilibrium region; however, only a small amount of condensation occurs. With additional heat the system soon begins operating under superheated conditions, and only vapor exhaust exits the chamber. If the heat load is specified equal to zero the chamber operates along the abscissas of Figure 3. Steady-state temperatures under these conditions are presented in Table 11. Considerably higher chamber temperatures result as the NH3/P molar ratio decreases. consistent with observations reported by Herbes et al. (1976). The higher temperatures are detrimental to the process because they decrease the extent of the polymerization reaction. This effect is illustrated in Figure 4, where the phosphorus conversion, defined as the amount of phosphorus reacted divided by the amount that entered on the fabric, is shown to be dependent on the NH3/P ratio over the entire residence time of exposure. It is apparent that conversion can be increased at a given exposure time by increasing the N H J P molar ratio. The benefit realized is due to decreased chamber temperature, as determined by the thermodynamics of the system. Another alternative for decreasing the chamber temperature is to introduce cooling of the chamber via the internal heat exchanger. This approach permits a decreased ammonia flow at the cost of additional cooling. Wagner (1974, 1976) suggested cooling for regulating the temperature of the chamber atmosphere. Herbes (1976) suggested that cvoling coils could lower fabric temperatures via control of moisture
TEMPERATURE 'F
Figure 6. Ammonia-water vapordiquid equilibrium data.
transfer between the fabric substrate and the chamber atmosphere. The effect of cooling on conversion for an NH3/P molar ratio of 2.5:l is shown in Figure 5. Again note that higher conversion results at the lower chamber temperature. An explanation for these observations is found in ammonia-water equilibrium data (Perry, 1963) shown in Figure 6. Considering the saturated liquid curve as the liquid contained in the fabric substrate, it is obvious that higher fabric temperatures, due to elevated chamber temperatures, result in less ammonia present in the moisture contained on the fabric substrate. In addition, the rate of ammonia mass transfer to the fabric is reduced because of increased ammonia vapor pressure, p*NH3. Equation 4 indicated how a lower ammonia concentration would undoubtedly result in a lower rate of polymerization. The data of Figure 6 also suggests that the entering temperature of the fabric can affect the extent of conversion. Eggenweiler (1973) suggested that it is essential that the fabric be cooled prior to entering the curing apparatus. Computed data comparing conversion for different entering fabric temperatures, with zero internal cooling, are presented in Figure 7 . At a 5-s exposure, the conversion increased from 74% to 89% for the cooler fabric. For a 122 O F entering fabric temperature, the conversion at 5 s was 82% (not shown in Figure 7). These results are a consequence of increased ammonia content
Ind. Eng. Chem. Process Des. Dev., Vol. 18, No. 1, 1979
112
loor l 90 o0l
I
INTERFACIAL AREA PER UNIT FAERIC AREA
,
d 2
3
4 5 6 7 RESIDENCE TIME ( w c
8
9
10
I1
T w
20
0/2 0/4
N%:P MOLAR RATIO
8 I
Ib
Figure 7. Effect of entering temperature of fabric on phosphorus conversion at 8:1 ammonia/phosphoruu molar ratio.
ENTERING MOISTURE GMI+O/GM FABRIC
CUAMIER TEMPERATURE AT ZERO COOLING 140 'F 122 104
ChAMBER TEMPERATURE AT ZERO COOLING 148 'F 145 140
131 NH, P MOLAR RAl 10 . 8 I
RESIDENCE TIME Ifec
Figure 8. Effect of fabric moisture content on phosphorus conversion at 8:1 ammonia/phosphorus molar ratio.
on the fabric at the cooler temperatures. Considerable attention has been focused on the effect of fabric moisture content on the curing operation (Wagner, 1974, 1976; Hooper et al., 1974; Raitinger, 1975; Herbes et al., 1976). Too little moisture on the entering fabric can result in insufficient polymerization. Too much moisture results in a nondurable finish, In a series of model solutions, the moisture content (dry basis) of the fabric was varied from 20% down to 1%. The results are shown in Figure 8. A significant drop in conversion is noted when the moisture content decreases below 5%. These calculated results are consistent with experimental data and are a consequence of the model's assumption that moisture must be present on the fabric before ammonia transfer can commence. However, even at the very low moisture content of 0.01, the reaction would still go to completion if the exposure time were sufficiently long. The model indicates that conversion is much more rapid a t higher moisture contents. It is cautioned, however, that excessive moisture can result in nondurable finishes. Fabrics of different constructions cure differently under similar process conditions (Calamari et al., 1975; Smith, 1976). The model can be used to simulate construction effects by varying the interfacial area (total surface area/apparent surface area). This parameter affects the rate of heat and mass transfer between a fabric and its surroundings. Calculated conversions for different a values
RESIDENCE TIME (SEC
Figure 9. Effect of interfacial area ammonia/phosphorus molar ratio.
1
I
or1 phosporus
conversion at 8:1
are presented in Figure 9. A reduction in conversion from 83% down to 38% resulted at a 5-s exposure time when a was decreased by a factor of 4. With good a values the model might be used to suggest how process operating conditions could be altered to account for fabric construction in obtaining a required conversion. Adjusting the a value should also make it possible to simulate curing cotton-synthetic blend fabrics. Correlation of construction with interfacial area will, however, require extensive experimental data. Summary Mathematical definition of the physical phenomena governing ammonia curing has permitted computed predictions of process behavior that agree with experimental observations, The model assumes that polymerization occurs between the phosphorus compound and ammonia dissolved in moisture on the fabric substrate. The polymerization reaction is further assumed to be very fast with respect to the other transport phenomena that occur during the cure process. The effectiveness of the ammoniation process was measured by the extent of the polymerization reaction. The parameters considered in this investigation included variation in ammonia/phosphorus molar ratios, different cooling loads on the chamber condenser, different chamber temperatures, and variations in fabric temperature, moisture content, and construction. Polymerization increased with (1) an increase in amnionia/phosphorus molar ratio, ( 2 ) increased cooling of the chamber, (3) increased moisture content of the entering fabric, and (4) decreased entering temperature of the fabric. Effects of ammonia feed concentrations and temperatures, production speeds, and chamber residence times will be evaluated in a future study. An ammonia cure reactor designed for a wide range of operational conditions is presently being constructed for experimental pilot-plant investigations. The experimental data obtained will be used for further model development and process optimization. Nomenclature a, interfacial area per unit volume of fabric, yd2/yd2yd C,, heat capacity of fabric, Btu/lb O F C H 2 0 , heat capacity of liquid water, Btu/lb O F CNH3,heat capacity of liquid ammonia, Btu/lb O F C,, heat capacity of solids, Btu/lb O F F, feed rate, lbimin HH,enthalpy of water vapor, Btu/lb H K , enthalpy of ammonia gas, Btu/lb
Ind. Eng. Chem. Process Des. Dev., Vol. 18, No. 1, 1979
HR, net heat of polymerization reaction, Btu/lb-mol H,, heat of solution of ammonia in water, Btu/lb-mol H,, heat of vaporization of water, Btu/lb-mol h, heat transfer coefficient, Btu/yd2 min O F h f ,enthalpy of the feed, Btu/lb h L , enthalpy of condensate, Btu/lb h,, enthalpy of exhaust, Btu/lb k , mass-transfer coefficient, lb-mol/yd2 min mmHg k,, reaction rate constant, lb-mol/yd2 yd min L , liquid condensate stream, lb/min 2, fabric path length in chamber, yd M W , molecular weight, lb/lb-mol p , partial pressure in chamber atmosphere, mmHg p * , vapor pressure at fabric surface, mmHg (H,O), moisture content of fabric, lb of H,O/lb of fabric (NH,), ammonia content of fabric, lb of NH3/lb of fabric (P), phosphorus content of fabric, lb of P/lb of fabric Qr, heat transferred via internal heat exchanger, Btu/min Qu;, heat losses through walls, Btu/min R , rate of polymerization reaction, lb-mol/yd2 yd min T f ,temperature of fabric, O F T , temperature of chamber atmosphere, O F vapor exhaust stream, lb/min V,, velocity of fabric, yd/min Wf, fabric width, yd x , weight fraction ammonia in liquid phase y, weight fraction ammonia in vapor phase pf, density of fabric, lb/yd2 yd ~
113
Literature Cited Baitinger, W. F., "Proceedings, 1975 Symposium on Textile Flammability", pp 280-296, LeBlanc Research Corp., East Greenwich, R.I., 1975. Calamari, T. A., Harper, R. J., Beninate, J. V.. Trask, B. J., JFFIFire Retardant Chem., 2, 121-131 (1975). Carter, W., US. Patent 4 009 002 (Feb 22, 1977). Chilton, T. H.. Colburn, A. P.. Ind. Eng. Chem., 26, 1183 (1934). Drake, G. L.. Beninate, J. V., Cooper, A. S.,Walker, A. M., Reeves, W. A,, hoceedings, 1st International Symposium Textile Research Conference, Paris, April 22-25, 1969. Eggenweiler, R. B., U S . Patent 3 775 155 (Nov 27, 1973). Getchell, N. F., Hollies, N. R., Stanton, S . S., U S . Patent 3982410 (Sept 28, 1976). Hendrix, E. C., unpublished data, Southern Regional Research Center, 1972. Herbes, W. F., American Cyanimid Co., personal communication, 1976. Herbes, W. F., Remley, K. H., Trecek. J. B., Am. Dyest. Rep., 72, 73, 75, 76 (Sept 1976). Hooper, G..Nakajima, W. N., Herbes, W. F., "Proceedings, 1974 Symposium on Textile Flammability", pp 30-46, LeBlanc Research Corp., East Greenwich, R.I., 1974. Lambert, A. H.. unpublished data, Southern Regional Research Center, 1976. Perkins, R. M., Calamari, T. A., Moore, H. B., Schreiber, S. P., Cooper, A. S., Am. Dyest. Rep., 63,5 (May 1974). Peny, J. H., "Chemical Engineers' Handbook", 4th ed,pp 3-65-3-67, M&aw-Hill, New York, N.Y., 1963. Schlichting, H., "Boundary Layer Theory", Pergamon Press, New York, N.Y., 1955. Smith, L., Cotton Incorporated, personal communication, 1976. Trecek, J. B.. American Cyanamid Co., personal communication, 1976. Wagner, G. M., U S . Patent, 3846 155 (Nov 5, 1974). Wagner, G. M., U.S. Patent 3933 122 (Jan 20, 1976).
Received for revieu November 7, 1977 Accepted August 14, 1978
Development of Feed Changeover Policies for Refinery Distillation Units R. W. H. Sargent * and G. R. Sullivan Department of Chemical Engineering and Chemical Technology, Imperial College of Science and Technology, London, SW7, England
A two-stage approach was adopted for the control system design of a binary distillation column and multiproduct crude unit subject to feed changeovers. Firstly, off-line open loop optimal control studies were performed which provided the best possible control trajectories. Then sub-optimal control laws were proposed which retained the dominant features of the optimal responses yet allowed for the practical aspects of implementation. Fast feed changeovers were found to be effected by causing the entire contents of a column to become either heavier or lighter together. Implementable sub-optimal policies based upon this observation took the form of feedforward control levels with switching between the levels governed by simple feedback laws. Simulation of these policies revealed close agreement with the optimal responses.
Introduction Continuous processing plants normally operate in a steady state, and in refinery operations it is common practice to determine optimum conditions for a given crude in the prevailing market situation. When the crude feedstock changes, the balance of products also alters and the whole refinery operation must be changed to the new optimum conditions. The problem then exists of establishing control policies for the transition period. As feedstock costs increase, changeover operating policies become more important and with oil reserves expected to dwindle in the future, the frequency of changeovers can only be expected to increase. There are many methods available for designing the control system, but before one is chosen the nature of the 0019-7882/79/1118-0113$01.00/0
processes and disturbances should be considered. Vast operating regions are spanned during typical crude changes and since the distillation processes of interest are known to be nonlinear, any design based on linear control theory would be inadequate. The nature and timing of the changes are known in advance, so that any proposed control system should utilize this information through feedforward elements. Hence, the changeover problem appears to be an ideal candidate for application of the optimal control design technique. With this approach, nonlinearities and interactions are dealt with naturally and any special feedforward structure can be deduced from the optimal response curves. Recently, the optimal control philosophy has been at the center of some controversy. Foss (1973) heavily criticized @ 1978 American Chemical Society
