Inclined Rotary Fin Principles, construction, geometry, design, and operation o f multistart, coned, helical-finned vacuum fractionator
for heat-sensitive compound ince many papers have been written on low-pressure
S drop-fractionation equipment, as though this were a
goal in itself, and without regard to requirements, the market for such equipment will be discussed first. Heat-sensitive and high-boiling materials that cannot be satisfactorily fractionated in conventional equipment occur in the manufacture of pharmaceuticals, perfumery and flavoring compounds (both natural and synthetic), vitamins, hormones, dyestuffs, and their intermediates. These industries, at the same time, are large users of conventional fractionation equipment which does not, in the majority of cases, give rise to trouble through thermal degradation. Although the specialized equipment described here can handle a number of products beyond the range of conventional equipment, it too will have its limitations. Its market may be considered as the continental shelf, largely uncharted, which borders the land masses that are the market for conventional equipment. Packed columns, commonly employed in the industries referred to, are straightforward in design and construction, and may be filled easily with any one of several standard types and sizes of packing, along with considerable choice of materials. Some recent innovations, such as metal Pall rings, are now successfully employed for fractional distillation at pressures of 20 torr and less, where they give markedly better results, both in terms of pressure drop and fractionating efficiency, than traditional stoneware Raschig rings. The versatility of the packed column ensures that it will be used to the limit before more sophisticated equipment is considered. The pressure drop through the fractionator is but one AUTHOR R a l p h W. K i n g is a Director of PenChem DeveloFments L t d . , 79 Grosvenor Place, London, S. W. 7 , England. Mr. K i n g thanks Technical Development Capital L t d . and the National Research Development Corp. for financial assistance and Roche Products L t d . f o r permission to publish the spectrog r a m s in Figure 76. 66
INDUSTRIAL A N D E N G I N E E R I N G CHEMISTRY
of several factors which affect thermal degradation of the product, and careful attention to the kettle or reboiler design may succeed in reducing the problem to manageable proportions while still allowing a packed column to be employed. Even so, some materials cannot be satisfactorily fractionated in conventional equipment because the pressure drop per theoretical plate is too high, leading to base pressures and temperatures unacceptable from the point of product degradation. Usually such materials are high in price, from $2.00 to $200 per pound. Examples found to date include triheterocyclic drug bases, whose hydrochlorides are used as tranquilizers and antidepressants; ethers of glycerol and various substituted phenols used as muscle relaxants in surgery; oxygenated sesquiterpene derivatives used as “base notes’’ in perfumery; and various unsaturated aldehydes containing 14 or more carbon atoms which are intermediates in the manufacture of synthetic vitamins and perfumery compounds. Such materials are often purified by a straight vacuum distillation (no fractionation), followed by crystallization from a solvent, centrifuging, and drying. This is expensive, both in capital and running costs, and losses of product (mainly in the mother liquor) are usually high. But, unless production is to be increased, or fresh production planned at a new site, the manufacturer is unlikely to invest in new and novel equipment to displace what has been bought and paid for, even though he recognizes that the new equipment gives a higher yield and/or a better product at a lower operating cost. These are some of the marketing problems to be faced for the more efficient equipment. Both the limitations of conventional packed and plate columns at low absolute pressures, as well as the special systems developed mainly in Europe for dealing with such problems, have been discussed at some length in earlier articles (4-6). Principle of PenChem Fractionator
Most very low pressure drop fractionators are based on the “wetted wall falling film principle,” the theory of
RALPH W. KING
Distillation Column
which has been dealt with in an earlier article (5). These include the Kloss system and the ACV trickling column. Apart from the difficulties of reflux distribution which these designs entail, the H E T P of these systems is rather high--i.e., from 2 to 4 ft. For a typical commercial column handling a modest throughput of a n expensive heat-sensitive compound, this leads to a tall and narrow column which often has to be thermally compensated and placed in a building to shield it from the elements. Even though the space within the column may be used efficiently, the whole thing becomes very cumbersome. T o reduce the H E T P of wetted wall columns and improve on the height-to-diameter ratio, Morton and others at Manchester University have studied the use of stationary helical coils, with the axis vertical, for fractionation at low absolute pressures (2, 7, 8). The reflux liquid running down the coil can cover only a fraction of the wall area (i.e., well under 5oy0),but measured vertically, such columns do give a low HETP, corresponding to two to three turns of the helix per theoretical plate, as well as a low pressure drop, over the very low range of throughputs which they can handle. The size of tube from which the helical coils are formed is limited to less than 1 in., and the free cross-sectional area of the passage, expressed as a percentage of the area of a cylinder containing the helical passage, is less than 2Y0. If the helical passage were formed from a finned tube inside a close fitting cylinder, the percentage free crosssectional area could be higher, and by employing several parallel passages by the use of interlayered fins, akin to a multistart screw, the percentage free cross-sectional area could be increased to some reasonable figure-Le., about 10%. This is illustrated in Figure 1, a-c. Apart from the constructional difficulties, this proposal has the serious disadvantage that reflux liquid applied to the top of a vertical wide-finned tube will take the shortest and steepest path down--i.e., it will cling to the inner tube and leave the greater part of the fin area unwetted.
T h e first aim of the PenChem fractionator was to wet positively the whole of the upper surface of a wide helical fin mounted on an inner cylinder by reflux liquid flowing from top to bottom. This is achieved by three features, taken in combination. First, the fin is made slightly coned, instead of sticking out a t right angles to the axis. T h e geometry and construction of such coned fins, as described later, have turned out to be simpler and easier than the “normal” helical fin. Second, the axis is inclined at such an angle to the vertical that the fin is roughly tangential to a horizontal plane in the direction to which it is tilted. When
Figure 7.
Cross sections of helical passages
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encased in a n outer cylinder, a stationary finned tube as described above has hollow troughs in the surface. O n pouring liquid onto the fin, pools extending the whole width of the fin form in the troughs. Third, by rotating the finned tube slowly about its axis, the pools of liquid can be screwed u p or down a t will, thereby wetting the entire upper surface of the fin once per revolution. Thus, the first objective was achieved (7). The second objective was to wet the undersurface of the helical fin in a similar way, so that virtually the entire wall of the helical passage was positively wetted. Attempts to do this by perforating the entire fin or making it porous were not successful. O n rotating such a perforated fin with the axis inclined as before, and with liquid supplied to the top, liquid was found to flow through the perforations close to the inner cylinder without spreading over the rest of the fin. T h e problem was solved more or less accidentally. When the unperforated fin was rotated slowly so as to screw the liquid down, the pools were displaced by the drag of the surface and retreated toward the inner cylinder. T o overcome this, the angle of tilt was increased slightly. When the liquid pools were screwed down again, a new phenomenon was observed (Figure 2 ) . From the inner cylinder, liquid streamed continuously outward over the fin, along the shallow trough formed by the moving surface. Before reaching the outside of the fin, the liquid was displaced and lifted by the drag of the surface to a point where the radial slope was reversed--i.e., where the outer edge was higher than the inner edge. The liquid could no longer flow outward any farther, and, after first being dragged still higher on the fin at the same radius, it flowed back over the sloping fin to the inner cylinder, and following this, down to the start of the trough again. Thus there was a continuous circulation of liquid over the surface, from the inside to the outside and back again. The extent of the outward flow of the liquid was governed by the total flow of liquid, the speed of rotation, and the angle of tilt. Rotational speeds were in all cases lower than those at which the liquid was thrown to the outside by centrifugal force. The whole surface was also covered effectively over a fairly wide speed range ( i . e . , + 3070 of optimum), with other factors constant. At this point, a row of small holes was drilled, closely spaced, and close to the edge of the fin. As expected, the liquid flowed outward over the moving surface until it reached the row of holes, flowed through them, and then back across the underside to the inner cylinder. The experiment was first carried out with a shallow truncated cone of thin gauge stainless steel, 10 in. 0.d. and 3 in. i.d. on a variable speed turntable, the angle of which could be adjusted. A stream of water was directed onto the upper surface near the inner radius. A trace of detergent in the water ensured that the liquid wetted the under surface of the cone and flowed across it without dropping off. I t has since been verified on a number of helical fin elements, both visually and in fractionating trials ( 3 ) . These two features of the PenChem fractionator assure 68
I N D U S T R I A L A N D ENGINEERING C H E M I S T R Y
complete and positive wetting of the walls of the helical fractionating passages even at the low liquid rates characteristic of low-pressure operation ( i . e . , under 5 torr). They also ensure good radial mixing and low axial mixing of the liquid. These features obviate the need for reflux distribution at the top, and periodic redistributors lower down, which are characteristic of packed columns. The surface of the fin, when smooth, tends to drain too completely between successive wettings when low viscosity fluids are distilled. (For a typical speed of 12 rpm, the surface is wetted once every 5 sec.) To obviate this, the surface is covered with shallow depressions which are rolled or pressed into the fin during fabrication. These dimples (on the upper side) and corresponding pimples (on the under side) increase liquid retention on both sides of the fin and improve performance. Vapor flows in helical passages have been discussed in detail in relation to mass transfer, pressure drop and transverse mixing by Morton and others in the references cited. Geometry
Since a knowledge of the geometry of coned helical fins was necessary for their development and construction, some of these features are described below. First, a coned helical fin or surface is a hybrid between a normal helical fin (which is at right angles to the axis) and a cone. An important difference between a cone and a normal helical fin is that the former can be developed from a flat sheet by cutting, bending, and joining edges, whereas the latter cannot. A normal helical fin can be formed only from flat sheet (or strip) by differential stretching. This raises the riddle, “When is a coned helical fin like a cone and when is it like a normal helical fin?” The answer, which follows from the geometry, is that it is usually like a cone, and all
Figure 2. Path of liquid on surface of rotating cone with inclined axis
T h e height y of the point P on a helix of radius r is plotted against w in Figure 4 for fixed values of p and a! and for varying values of r. The resulting curves may be described as “sine waves on a slant.’’ For values of r greater than a critical value rC, the curves show maxima and minima in y . At r = rc, the curves show horizontal “points of inflection,” while for r < ro the gradients of the curves are positive throughout. For r = rc, the horizontal points of inflection occur at values of w of 0, 27, and 4n-, and both dy/dw and d2y/dw2 are equal to zero a t these points. Differentiating y with respect to w,
9 --@ cos - r cos w sin do 27r a!
_ d2y - r sin w sin a! = 0 when w
CY
= 0, 28, etc.
do2
When dy/dw = 0 and w = 0, cos a = 1
P cot rc = -
a!
2T
Figure 3. Projection of helix with inclined axis on vertical plane
fins with which we are concerned here have the property, inherent in a cone, that they can be developed from flat sheet by cutting, bending, and joining edges. Consider first a helix formed around a n inclined axis OY‘ which is a t right angles to a plane with rectangular axes OZ, OX’ (Figure 3). The axis O Z is horizontal and a t right angles to the plane of the paper. Horizontal and vertical axes, O X and OY, lie in the vertical plane X’OY‘ such that the angles XOX’ and YOY’ are both equal to a!. The coordinates of a point P on the helix are ( x ’ , y ’, z), with reference to the axes OX‘, OY’, OZ, and ( x , y), with reference to the axes OX, O Y . Then x = x‘cos~~+y‘sina! y = y’ cos
a!
- x’
sin
Condition for pool formation. Figure 4 shows that helices for which r > re, the value of y is negative a t the minimum where dy/dw = 0, and that this minimum value of y corresponds to a positive value of w , designated am. There is also a maximum value of y at a negative value of w, where again dy/do = 0. The values of urnare given by
r cos urn=
P
-
2a
cot
a!
= ro
or
Figure 5 shows the projection of the minima and maxima in y onto the plane X‘OZ. The coordinates ( x ’ , z ) are clearly given by
a!
The position of P may also be defined in terms of the polar coordinates (r, w ) and P, the pitch of the helix, where r is the radius of the helix and w the angle around OY‘ which the point has traversed on the helix from the starting point where x , y , and w are all zero, and z = r. Then clearly w
Y‘
=gP
x ‘ = r sin w z
= r cos w
And referred to the axes OX, O Y , UP
y = - cos a 27r
- r sin a! sin w
x = r sin w cos
CY
+ 2-r sin UP
a!
Figure 4. Vertical and angular projections of points on plane helical surface with inclined axis V O L 61
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z
Thus the value of y at any point on the coned helical surface of radius r (> r c ) is given by:
= rc
x ’ = rc tan w,
= r sin w, =
d r 2
- r:
wr,
Positive values of x ’ and w m correspond to minimum values ofy, and negative values of X I and w m to maximum values of y. Again, Figure 4 shows that the surface on which helices of radius greater than rc lie would be one formed at right angles to the axis OY’, and that ymindecreases as the radius r increases. I n other words, when such a surface is bounded by concentric inner and outer cylinders, liquid poured onto the surface would initially form pools only at the periphery of the surface, in contact with the outer cylinder. I n order that pools should have the same maximum depth at any radius, the helical surface must be coned. This requires that the sine curves shown on Figure 4 for values of r greater than rc be raised by an amount (-ymin) so that y = 0 and dy/dw = 0 for any helix of radius r which lies on the surface (Figure 6). Referring to the original surface projected in Figure 4,the value of yminis given by : Ymin
Urn
(i
cos a)
- r sin w, sin a
rc : sin . w, = r
cos w, = -
li,r2
c0s-l r c / r . The “line of depression,” which is tangential to an imaginary inner helix of radius rc, is a generator for the surface. Maximum depth of pools. The surface is bounded by inner and outer concentric cylindrical planes of radii r f and ro formed about the axis OY. The height of the surface y has a maximum at any radius within these limits. The maximum occurs at the negative value of w given by : w =
where
- -sin a( l / r 2 -
-
rc2
r - rc cos-1 2)
r
Figure 5. Projection of line of maxima and minima on surface on plane perpendicular to axis INDUSTRIAL A N D ENGINEERING CHEMISTRY
-1
r cos-1 2
r
1
Figure 6 shows that the maximum value of y is lowest at the smallest radius rc, and that a pool of liquid on the surface will “spill over” at this point. The maximum pool depth is given by the value ofymanat ri
y,,
ym = -sin a (r sin w, - rewm)
70
A series of such curves for values of r/rc = 1, 2, and 4 is plotted in Figure 6. Both y = 0 and d y / d w = 0 at the “line of depression” on the surface for the positive value of w given by w =
- rc2
from which
- r sin w
= 2sin a
c
‘1
- rc cos-1 2 Ti
P
rc = - cot a 2n
The solid analytical geometry given above was worked out entirely from first principles, since no treatment of coned helical fins could be found in the literature. It forms the basis from which exact formulas were derived for cutting disks from which finned ele-
Figure 6. Vertical and angular projections of points on coned helical surface with inclined axis
ments of given dimensions can be constructed. I t was also invaluable for solving problems of scale-up. An approximate derivation of the free cross-sectional area of a fractionating element so constructed is given in Figure 7. Figure 7 represents a section through the axis of a multiple-start helical fin-fractionating element containing n starts, of pitch @, constructed of thin metal sheet, thickness t , and bounded by inner and outer cylinders of radii rl and ro; d is the gap between fins measured in a line parallel to the axis. A section through a helical passage is very nearly a parallelogram (slight curvature in the longer sides has been noted earlier). The shorter sides are given by: d = p/n
- t cos a
and the longer sides are approximately equal to :
T h e cross-sectional area of the passage, A , is given by :
A = d(r0
- rJ
Design and Construction of Fractionator
I n designing a fractionating device based on the principles discussed, choices had to be made on several questions : (1) Is the unit to be batch or continuous? ( 2 ) What are its intended capacity and fractionating efficiency? (3) What is the value of the cone angle a to be, and what should be the angle of inclination of the axis to the vertical? (4) What should be the relationship between the gap d and the radii of the inner and outer cylinders (r( and ro) bounding the fin?
Figure 7. Section through axis of four-start fractionating element
(5)
How many starts are to be employed? (6) How is the fin to be constructed and attached to the inner cylinder? (7) Is the outer cylinder bounding the helical passages to be stationary or attached to the moving element? (8) What speed of rotation is to be employed? I n deciding on answers to these questions, a great deal of useful experience had been gained through the design, construction, and operation of an earlier prototype unit. Where improvements on the prototype unit were indicated, these were attempted without hesitation in the new design. The answers then were as follows : (1) Batch or Continuous? The unit was designed in the first place for batch operation, with a view to possible conversion to continuous operation later. ( 2 ) Duty. T h e unit was intended to handle vapor rates u p to 35 kg/hr under various conditions. For this, several different fractionating elements were to be constructed, that designed for the lowest absolute pressure (i.e. 1 to 2 torr) to have a n efficiency of a t least 4 theoretical plates, while others suitable for higher pressure operation would have efficiencies u p to 30 theoretical plates. (3) Cone Angle “a+” Previous work had shown the desirability of keeping this angle as small as possible, so long as it does not restrict the overall cross-sectional area of vapor passage too severely. I t was also something which requires standardization, and it was finally decided to employ a n angle of 1 2 f.1 At the same time, the unit was designed so that the angle between the axis of the element and the vertical could be varied between 9’ and 20’ for experimental purposes. (4) Relationships between ri, ro, and d. Taking first the ratio ri/ro with other things constant, one would expect the vapor capacity to fall and the fractionating efficiency to rise as this ratio is increased, so that a compromise must be made. The value chosen was 0.45, based on experience from the prototype unit. For d/ro, values of 0.131 and 0.087 were chosen for different elements constructed. There is scope for further work in which these ratios are varied systematically over a wider range. (5) How Many Starts? Different elements have been designed and constructed with the number of starts varying from 1 to 4. As the number is increased, the total cross-sectional area for vapor flow increases while the length of passage, and hence the fractionating efficiency, decreases. The work a t low absolute pressures described in this paper was carried out entirely on a four-start element. (6) Fin Construction. T h e fins are constructed out of light gauge sheet stainless steel by a sequence of operations involving cutting, rolling, bending, and welding. They are then fitted in helical grooves turned on the inner core, and suitably fixed. A certain amount of trial and error was needed before completely satisfactory methods were evolved.
’.
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( 7 ) Element Casing. The rotating element and the stationary outer casing were so constructed that the fins make a very close running clearance within the shell. Other arrangements had previously been tried, with rather less satisfactory results. (8) Speed of Rotation. Since previous work on the prototype indicated that the optimum speed was likely to be between 12 and 20 rpm, a variable speed drive was used which could be adjusted to any value between 8 and 40 rpm. Description of Unit Used for Tests
The unit in which these tests were carried out was a preproduction model shown in Figures 8 and 9. The main features are as follows : (1) A 25-1. stirred stainless steel kettle with conical base, thermostated 3.5-kw heater, lower bush bearing, continuous sample removal and return points, thermopockets for liquid and vapor temperatures, manometer branch, lagging case, glass illumination and viewing ports, and other accessories (2) Flanged mild steel shell, accurately bored and buffed, and internally plated with a hard chrome finish. This not only houses the fractionating element, but also the condenser, reflux splitter, and product offtake (3) Four-start fractionating element, with dimensions given below, with flanged lower half shaft and kettle stirrer, and flanged upper half shaft with reflux splitter, drive plug, and antiswirl baffles (4) Top flange assembly with 0.5 m2 stainless steel condensing coil, cooled seal and bearing housing, fixed
reduction gear, and reflux ratio operating mechanism (5) Flameproof electrical heating system with controls including two 1-kw shell compensating heaters with energy regulator (6) Two-stage oil-sealed rotary piston vacuum pump, rating 150 l./min with flameproof motor and belt drive, and a variable speed in-line reduction gear driven by the same motor. This was linked by a flexible shaft to the fixed reduction gear on the top flange assembly driving the fractionating element, covering a speed range from 8 to 40 rpm. ( 7 ) Panel-mounted vacuum manifold with cold trap and such items as a self-actuating absolute pressure controller, vacuum reservoir, rapid evacuation line, and isolating valves (8) Framework and instrument panel, the latter including one 0" to 300°C indicating pyrometer and six-way switch, oil-filled pressure drop manometer, two 0 to 20 torr absolute pressure gauges, two compound vacuum gauges, and one 0 to 5 l./min cooling water rotameter with control valve Fractionating element. The fractionating element consisted of four coned helical fins mounted in grooves turned on a tubular core. The fins were made from flat annular disks of 26-gauge sheet stainless steel passed through special rollers to give 12 spoon-shaped depressions/sq in., of approximately 0.9 mm depth. This pattern was flattened over narrow annular bands bordering the inner and outer edges of the disks, and a series of equal slots was punched in the outer band a t an angle of 45' to a radial line through the center of the
PC.
Figure 8. Diagrammatic view of fractionator used for tests Pressure controller. PI. Pressure indicator. TC. Temperature controller. TI. Temperature indicator.
72
INDUSTRIAL A N D ENGINEERING CHEMISTRY
TIC. Temperature indicator controller
slot. T h e dimensions of the flat disks common to all elements of this series are given in Table I. The disks were cut radially and four sets welded together on a specially designed jig to form four flat, lefthand helical fins each containing 15-1/2 disks. Each set was then bent into conical form on a special jig, and finally was fixed under tension into each of four helical grooves, 2 mm deep, turned on a cylindrical former. The whole assembly gave a four-start helical fin shown in Figure 10 with the following dimensions : ri = 66.75 mm
d = 18 mm
ro = 1 3 4 m m
p
a = 12"
= :2mm
Other fractionating elements have been constructed with the same overall dimensions, but with different cross-sectional areas and lengths of passages. Although these are less tried or developed than that reported on here, the flexibility possible with this type of design is apparent. Choice of test mixture. The choice of a binary test mixture for evaluating the efficiency of fractionating columns a t pressures below 5 torr is not easy or straightforward, and, as recent studies a t Birmingham University have shown, most of the systems recommended and reported in the literature are unsatisfactory. As a result of some discussion it was decided to employ a mixture of dimethyl phthalate and diethyl phthalate, to meet the following requirements. (1) Owing to the close similarity in molecular structure of the components, the system may be treated as an ideal one, and reliable figures for the relative volatility calculated from the ratio of the vapor pressures. (2) The boiling points of the components fall in a suitable range (100" to 150°C) for test work a t these pressures. (3) T h e relative volatility (1.4 to 1.5) is suitable for evaluating columns with up to 1 2 theoretical plates.
The finned element rotated freely in the bored shell with a clearance of approximately 0.2 mm. The element and shell formed four helical passages, each of 1 5 turns, with a total cross-sectional area of 46 cm2, which amounted to 8.2y0 of the cross section of the cylindrical shell. T h e surface area of the fins, without allowing for indentations, was approximately 5 m2, and average length of each helical passage was 9.5 m.
Figure 9. Photograph of fractionator used for tests
Figure 70. Photograph of four-start fractionating element VOL. 6 1
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(4) Samples of the mixture can be speedily analyzed by refractive index with an accuracy of 1% using standard equipment. (5) Both materials are available in commercial quantities at satisfactory purity. Properties of test mixture. Dimethyl phthalate and diethyl phthalate were kindly supplied for test purposes by ABRAC, and samples were first examined for purity by GLC analysis by Yarsley Testing Laboratory, Ashtead. The results are given in Table 11. The total D M P DEP content of the test mixture was
+
thus nearly 99y0, the main impurity being identified with near certainty as methyl ethyl phthalate. T h e effect of this, small as it may be, on the measured fractionating efficiency will be to show a somewhat lower figure than the true one. The lower boiling unknown impurities 1 and 2 were removed by withdrawing and discarding a few per cent of the overhead product from the fractionator prior to commencement of tests. The higher boiling impurities 3, 4, 5, and 6 should remain in the kettle and should not affect the samples. The relevant physical properties of DMP and DEP are given in Table 111. Vapor pressures and "ideal" relative volatility are plotted against temperature in Figure 11. Analyses. Since there is a difference of 0.0132 in the refractive indices of D M P and DEP at 25"C, it should be possible to analyze mixtures of the two to an accuracy of 1% by refractive index if an accuracy of 0.0001 in refractive index measurement is achieved. Three intermediate sample mixtures of B.D.H. analytical reagent grade materials were made up and their refractive indices measured at 25°C. The results (Table IV) indicate that this assumption is justified when using a linear relationship bctwcen refractive index and composition. Sampling arrangements for tests at total reflux.
'The product cooler was removed and in its place a sampling system was installed as shown in Figure 12. A similar system was installed for sampling liquid passing from the lower end of the fractionating element to the kettle. Both systems allowed a small bleed of liquid to pass continuously through a small glass collector back into the fractionator. The bleed could be interrupted and the sample collectors drained and re-evacuated without upsetting the column conditions. Distillation tests, Five 1. of a test mixture (20y0 DMP, 80% DEP) were initially placrd in the kettle. As a result of sample removal, the kettle mixture became progressively depleted in DMP, which was periodically replaced. The kettle liquid thus varied in composition from 13 to 26y0DMP during the course of the tests. Four series of tests are reported in this paper. I n the first three, the speed of rotation of the element was varied during each series (from 8 to 25 rpm), while the other variables-angle of element axis to vertical, boilu p rat?, and column top pressure-were maintained at constant values. The fractionating efficiency (number of theoretical plates a t total reflux) and pressure drop were measured for each experiment in the series, so that the optimum speed of rotation could be found. I n the last series, the angle of the element axis to the vertical was varied from 10" to 19", while the other three variables were held constant, using the best speed of rotation found from the first three series. The vapor throughput was maintained sensibly constant during each series by setting the control temperature of the heater at a constant value (225" to 350°C). The electrical heat input to the kettle was measured by a kw-hr meter, and the heat removed from the con74
I N D U S T R I A L A N D E N G I N E E R I N G CHEMISTRY
Figure 7 7 . Vapor pressures and "ideal alpha" of dimethyl phthalate and diethyl phthalate as function of temperature
Figure 12. Sampling system f o r tests at total rejux
denser calculated from the water flow and temperature rise. Before commencing tests, the jacket compensating heater was adjusted until the rate of heat removal from the condenser came within 10% of the heat input to the kettle. Samples of top and bottom liquid were taken and analyzed at 30-min intervals for each set of conditions employed until constant analyses were obtained. Two causes of error, reduced as experience was gained and which gave efficiencies lower than the true ones, may be mentioned here. (a) T o maintain conditions as nearly as possible to those of total reflux, the flow of top sample through the collector must be very low. This was difficult to achieve with the diaphragm valves employed.
(b) Any disturbance in the controlled pressure would upset the concentration gradient established in the column, and some time was necessary to restore this. T h e column top pressure was measured by a 0 to 20 torr Edwards capsule dial gauge. The pressure drop was measured between the kettle and a point on the vacuum line above the condenser by an oil-filled U-tube manometer. The pressure drop over the condenser was also checked by the manometer during the first series, but as this was only a small part of the overall pressure drop, this measurement was discontinued. Results of tests. The results of the tests in series 1 to 4 are summarized in Table V. The vapor throughput quoted for each series was a n average value, calculated as the arithmetic mean of the heat generated in the kinetic heater and the heat removed by the condenser, using an average latent heat of 74 cal/g. T h e vapor rate of 25 kg/hr in series 2, 3, and 4 was close to the maximum possible with the existing kettle heater and control system. This was below the load point of the fractionator, even a t the lowest absolute pressure. The number of theoretical plates at total reflux was calculated from the analyses of top and bottom samples by the Fenske-Underwood equation. T h e figure quoted is for the fractionating element only and does not include the extra plate to be added for the kettle to give the overall figure for the fractionator. The number of theoretical plates determined for series 1 to 3 is plotted against rotational speed of the element in Figure 13. The number of theoretical plates found in series 4 is plotted against the angle of the axis to the vertical in Figure 14. T h e pressure drop per theoretical plate at the optimum speed is plotted against vapor throughput in Figure 15. Measurements of liquid holdup under fractionating conditions were not possible, but some measurements were made at atmospheric pressure by supplying a stream of soapy water at a known rate to the top of the rotating element, injecting a shot of dye at the top, and determining the time taken for it to emerge. From this, the dynamic holdup under fractionating conditions was deduced as 0.4 to 0.6 l., depending on throughput. The static holdup was extremely small. After a few minutes of continued rotation when the flow of liquid to the element had ceased, drainage was virtually complete leaving less than 50 ml of liquid adhering to the element.
Discussion
Number of theoretical plates at total reflux (NTP). The N T P depends on the speed of rotation and shows a flat maximum at 8 to 12 rpm (Figure 13), while falling off at higher speeds. This drop in N T P with increasing rpm is more pronounced a t lower throughputs. T h e NTP is not critically dependent on the angle of tilt, being almost constant from 13" to 19", and falling off below 13" (Figure 14). At lower vapor rates, the VOL. 6 1
NO. 9
SEPTEMBER 1969
75
Figure 73. Number of theoretical plates us. speed of rotation 0
-. - -.
7- Torr top pressure, 25 kg/hr 5-TOrr top pressure, 25 kg/hr
A - - - - -. 5-Torr top @ressure, 11.5 kg/hr
Figure 75. Pressure drop per theoretical plate us. vapor throughput
0.5-Torr top pressure, 25 kglhr
0. 7-Torr top pressure, 25 kg/hr
Figure 74. Number of theoretical plates us. angle of tilt 76
INDUSTRIAL AND E N G I N E E R I N G CHEMISTRY
range of angles over which the KTP is constant is somewhat narrower. These features are undoubtedly due to a common cause--i.e., a t higher rotational speeds and at lower angles, the liquid is no longer able to reach the peripheral holes and flow over the underside of the helical fin. This has been confirmed in both cases by visual observation with the element mounted outside the shell, and a stream of soapy water supplied to the top. Vapor throughput appears to have little influence on the fractionating efficiency at the optimum speed of
rotation (series 1 and 2). This was also found to be true in other tests, not reported here, which covered a wider range of throughputs, and other elements as well. The somewhat lower N T P at 5 torr top pressure (series 2) compared with 1 torr top pressure (series 1) found in these tests are not thought to be significant, since higher N T P (6 to 7 ) were found in other tests with the same mixture at 5 torr top pressure. I n general terms, one theoretical plate requires between 2-1/4 and 2-1/2 turns of each of the 4 helical passages of this element. This holds over a wide range of 8 to 12 rpm, and angles of 14" to 18" from the vertical. Almost identical fractionating efficiencies were found with two other test mixtures, o- and p-dichlorobenzene, and nerol geraniol. There appears to have been no noticeable adverse effect on the fractionating efficiency of this four-start element caused by liquid or vapor leakage through the narrow gap (under 0.5 mm) between the stationary shell and the outer edge of the fins. This is not surprising, since leakage past one fin of a four-start element bypasses only one quarter of a turn of a helical passage, or one tenth of a theoretical plate. The effect of such a leakage might be expected to be more serious on a singlestart element, and this indeed is the case. Some tests have been carried out both on single-start and two-start elements in the same shell. The single-start element required four to five turns of the helical passage per theoretical plate, whereas the two-start element required 2.7 to 3.3 turns. Pressure drop. T h e pressure drop through the element was virtually unaffected by changes in the rotational speed or the angle of the tilt. At the maximum vapor rate of 25 kg/hr, the average pressure drop per theoretical plate at speeds of 8 to 12 rpm was 0.42 m m Hg a t a top pressure of 5 torr and 0.63 mm H g at a top pressure of 1 torr. At a top pressure of 1 torr, and a vapor throughput of 18 70'% of the maximum-the average preskglhr-Le., sure drop per theoretical plate was 0.35 mm Hg, giving a kettle pressure below 4 torr in conjunction with a total fractionating efficiency of 8 theoretical plates. These conditions were successfully used for the trial fractionation of a number of heat-sensitive products which could not be handled in conventional packed columns. I t was surprising to find that the column did not flood at a vapor rate of 25 kg/hr and a top pressure of 1 torr, since the vapor velocities reached near the top of the helical passages must have been well in excess of 100 m/sec. T h e reason for absence of flooding may be that drops of liquid entrained by the vapor are thrown onto the outer wall by centrifugal force. Both the pressure drop per theoretical plate and the height of a theoretical plate for low-pressure distillation are markedly similar to those published for Sulzer BX packing (9) in spite of major differences in geometry and principle of operation. The pressure drops measured here included three
additional sources of loss as well as the pressure drop through the element. These were : (a) Pressure drop through the condenser (b) Pressure drop in the antiswirl baffles through which the vapor passed on leaving the top of the element. These were fitted to ensure correct operation of the reflux splitter (c) Pressure drop due to sudden acceleration of the vapor entering the element Of these, (b) and (c) are still significant factors, and improvements are clearly possible by streamlining the inlets and outlets of the passages. General. Some qualitative comparisons may be made with other equipment used for fractional distillation a t pressures below 20 torr. T h e most widely used appear to be stainless steel Pall rings, and the Sulzer BX packing. A direct comparison with 2-in. stainless steel Pall rings in the distillation of a heat-sensitive compound is shown in the gas chromatographic analyses of the same batch of material fractionated: (a) in a batch still with a column 18 in. in diameter packed with 2-in. Pall rings to a height of 20 ft and (b) in the present Penchem fractionator using the element employed in the tests reported in this paper. The GLC analyses are shown in Figure 16, a and b. The distilled product is in each case a heart cut of 97y0 purity, and the distillations were carried out a t the same reflux ratio taking the same time to complete. T h e column top pressure was held at 1 torr in both cases. Apart from small differences in fractionating efficiency, the main difference in conditions was that the kettle liquid temperature lay between 125" and 128°C in the PenChem fractionator, with a pressure of less than 5 torr,
Figure 16. Gas-liquid chromatograms of distilled heart cuts of heat sensiiive product under comparable conditions VOL. 6 1
NO. 9
SEPTEMBER 1969
77
while the kettle liquid temperature of the column packed with Pall rings lay between 138” and 145°C corresponding to a pressure of 1 5 to 20 torr. The GLC chart for the PenChem distilled product shows a lower boiling impurity A, the main component peak B, and a higher boiling impurity C. I n the chart for the product distilled in the column packed with Pall rings, only peaks A and B are present. T h e higher boiling impurity C had completely reacted a t the higher temperature in the kettle with some of the main product to form a nonvolatile tarry material, with consequent loss of product. Had the same distillation been carried out in a column packed with Sulzer BX gauze cartridge packing, the kettle temperature could have been closer to that found with the PenChem fractionator. I t is thus useful to compare the essential features of both: (1) The Sulzer packing is stationary, while the Penchem fractionating element rotates slowly. This is a disadvantage only in that it places a restriction on the height of a PenChem fractionator, and limits it to about 30 theoretical plates. Slow rotation is an advantage in that it prevents the element from becoming bound to the shell (by polymer). (2) Both require internally machined shells with a close tolerance on the internal diameter. (3) Heights of packed shell or fractionating element for the same duty are the same. The Sulzer BX column may have a somewhat smaller diameter. (4) The Sulzer BX packing must be cleaned by suitable solvents. I t is too fragile to be removed and cleaned mechanically. The PenChem fractionating eIement is mechanically much stronger, and can be removed completely from its shell for cleaning and inspection. T h e whole surface can be cleaned mechanically--e.g., by brushing, scraping, or light shot blast. (5) The Sulzer BX packing has a greater area of exposed metal surface per unit volume than the PenChem fractionator element, which relies for its efficiency on complete wetting and high “j” values. The Sulzer packing is formed from thin wire gauze, while the PenChem fractionator element is made from sheet metal. From this, one would expect the Sulzer packing more likely to cause catalytic degradation of the product and more subject to corrosive wastage than the PenChem fractionator element. (6) T h e liquid holdup for columns designed for identical duties appears to be higher for the Sulzer BX packing. ( 7 ) I t also appears that equilibrium is achieved more rapidly and reflux distribution more easily attained in the PenChem fractionator. From a study of these points, the following circumstances would tend to favor the use of the PenChem fractionator : (a) I n handling materials which tend to polymerize or foul the column internals (b) I n handling materials which may be catalytically decomposed by the metal of the packing or element (c) I n handling materials which may corrode the packing or element 78
INDUSTRIAL AND ENGINEERING CHEMISTRY
(d) I n distillations which require column base pressures of 5 torr or less to prevent thermal degradation ( e ) In preparations which require fewer than 20 theoretical plates Scale-up. An analysis of the theoretical factors has revealed no serious scale-up problems. Larger elements are scaled up geometrically, all dimensions being increased in proportion. The vapor capacity is expected to increase in proportion to a power of approximately 2 . 2 of the linear dimension. The rotational speed will be reduced somewhat as the diameter increases. T h e problem is one to be tackled in stages. From present experience, elements of twice the present diameter can confidently be designed and made and their performance reasonably predicted. This, however, needs to be carefully checked before still larger fractionators of this type are constructed. Prior to building the fractionator described here, an earlier 6-in. model had been built and tested, and nothing unexpected was revealed in going from the 6-in. to the 10-1/2-in. model. Fractionating elements of this type are also entirely suitable for continuous distillation and should be free of the problems of ensuring even feed distribution, problems associated with packed columns operating at low absolute pressures. Equipment operating on the same principle has several possible applications for other mass transfer operations, which may, if required, be combined with heat transfer.
Nomenclature A = cross-sectional area of helical passage d = gap between fins measured parallel to axis n = number of starts of multiple helical fin = pitch of helix or helical fin p = critical radius of helical surface rc = inner radius of helical surface or fin rz = outer radius of helical surface or fin ro x , y , z = Cartesian coordinates referred to rectangular axes (OY vertical) of a point P on a helical surface with axis in plane X O Y at angle a to vertical x ’ , y ’ = Cartesian coordinates referred to rectangular axes X‘OY in plane X O Y of a point P on a helical surface whose axis is 0 Y’ r , w = polar coordinates of a point P on a helical surface whose axis is OY’ Greek Letters
a
= angle YOY’
REFER ENC ES (1) Cooke, E. V., and King, R. W., I.C.I. Ltd., Brit. Patent 698,246, March 3, 1950. (2) Huggett, R., andKing, P. J., Trans. Inrt. Chem. Eng., 46, T.lO1 (1968). (3) King, R. W., PenChem. Developments Ltd., Ital. Patent 831,122, April 11, 1968. (4) King, R . W., Brit. Chem. Eng., 12 (4), 568 (1967). ( 5 ) King R W ibid. (5) 722 (1967). (6) King: R: W:: ibid.: (id), 1599 (1967). (7) Morton, F., King, P. J., and McLaughlin, A,, T r a n s . Inst. Chem. Eng., 42, T.285 I1 964). ~.~ - ,. (8) Morton, F., King, P. J., and McLaughlin, A., ibiu‘., T.296 (1964). (9) Sperandio, A,, Richard M., and Huber, M., Chem. Eng. Techno!., 37 (3), 322 (1965). (10) Jordan, T. E., “Vapour Pressures of Organic Compounds,” Interscience, New York, N.Y., 1954. (11) Yarsley Testing Laboratories Ltd., T h e Street, Surrey, Ashtead, England.
