Measurement and CornDarison of Pumping Speeds Benjamin B. Dayton DISTILLATION PRODUCTS, INC., ROCHESTER 13, N. Y.
A study has been made of the effect on measured pump speeds of varying the position of the mouth of the gage, the position of the a h inlet, the shape and size of the testing dome or head, and the design of the buret, ion gage, and other measuring instruments. Various techniques of speed measurement are described and the common errors pointed out. The formulas for pumping speed are reviewed and the relation 1/S = 1/So 1 / U for computing net pumping speed, S, at the end of a pipe of conductance, U , is shown to be incorrect f6r short lengths of pipe when the pump speed, So, is measured with a test dome in the usual way. The need for improved standards is shown by comparison of speed curves and other data obtained by different 1Fboratories for the same pump and by a review of the literature on measurement techniques.
+
air is then introduced from a point high on the side of the dome, so that the molecules are scattered into a random distribution about the top of the dome. The molecules can then diffuse through the mouth of the pump from all directions in a way which approximates the actual flow in most applications. Data obtained in the Distillation Products laboratories and also in the university laboratory mentioned show that the speed as measured with a dome may be 50% of the pseudo-speed as measured with a blank-off plate with air beamed directly into the jet of a n unbaffled pump, and even with a baffle located in the top of t h e pump casing $he speed with a dome may be 25y0 less than the value obtained with a plate. //
A
*
T T H E present time the comparison of results from different pump testing laboratories is difficult because of the lack of standard test procedures and nomenclature. I n particular there is a great need for standard techniques in the measurement of the speed of diffusion pumps. I n the course of the past few years the author has had the opportunity t o observe several important sources of error in these measurements which have not been adequately stressed in the literature. During the recent war laboratories had several opportunities t o compare notes on the same pump. The wide discrepancy in the data obtained by independent groups showed very clearly the need fo? a s h n d a r d technique.
BEAM O f AIR MOLECULES
Directional Effects in the Test Dome
\ 1
/I
On visit t o one of the university laboratories in 1943 the author found one of the large Distillation Products pumps being tested by placing a “blank-off” plate over the mouth of the pump with air from the measuring buret streaming from a small hole directly down into the pump. This technique, which is illustrated in Figure 1, has been used in many laboratories. For example, diagrams of systems for measuring the speed of diffusion pumps in which this technique was employed are found in articles by Alexander (I), Wachter and van der Scheer (30),and Westin and Ramm (31). It can be shown that this procedure gives abnormally high values, or pseudo-speeds. I n the normal use of the pump the air enters the mouth of the pump from all angles according t o the cosine law, and some of the molecules strike the side walls or top cap on the jet assembly and then are scattered back out of the pump. However, when the air issues from a small hole in a flat plate over the pump, the molecules may be “beamed” directly into the vapor jet or otherwise aided in reaching the pumping area without the normal amount of diffusion and scattering in the direction of the mouth of the vacuum gage. The vacuum gage thus gives lower readings and the speed corresponding t o a given air leak will be higher than normal. I n order to simulate the conditions under which the pump will normally be used, the proper procedure is t o build a test dome of diameter equal to, or preferably larger than, t h a t of the pump casing and height at least equal t o the diameter. The
DIFFUSION
IR
PUMP
Figure 1. Incorrect Method of Measuring Pump Speed
The position of the mouth of the vacuum gage within the dome has also been found to influence the speed readings, as shown by Figure 2. The number of molecules entering the mouth of the vacuum gage depends on the flux of molecules out of each volume element at the distance of one mean-free-path in anx given direction and on the cosine of the angle between that direction and the normal t o the mouth of the gage. As the
0
1 0 L
-
795
i
-
796
Vol. 40, No. 5
INDUSTRIAL AND ENGINEERING CHEMISTRY
peak speed of large diffusion pumps occurs in the neighborhood of mm. of mercury, the mean free path near the peak speed is between 50 and 100 cm., which is larger than the diameter and length of most test domes. Thus if the mouth of the gage is faced toward the pump, a low reading will be obtained, because the distribution of air molecules near the pumping area is such t h a t not as many molecules are directed back out of the pump as enter the pump from the dome. If the mouth of the gage is faced away from the pump and toward the top of the dome, a higher pressure will be recorded than in the previous case because the molecular flux from this region is much higher. These effects were studied experimentally at Distillation Products, Inc., in 1940 with a 4-inch pump, and the wide differences i n results obtained are shown in Figure 2. A standard test dome was therefore adopted with the leak placed high on the side of the dome and directed across to the opposite side wall, or dispersed by a screen or base over the top of the dome; the vacuum gages were placed a t right angles t o the leak low on the side wall near the mouth of the pump lvith the tubulation extending inside the dome t o about one third of the radius. The mouth of the ion gage has always been pointed radially across the dome in tests conducted by the author, but other laboratory groups have been observed using a dome in which the mouth has been tilted toward the pump and correspondingly higher speeds have been obtained. That directional effects have been frequently overlooked is evident from the diagrams of pump speed measurement appearing in journal articles. Figure 3 shows the test chamber used by Copley, Simpson, Tenney, and Phipps (4) in their study of the speed of divergent nozzle pumps published in 1935. They introduced the air at L and measured the pressure at both B and C with McLeod gages. If one defines the true speed of a pump i n terms of the speed of exhaust of a large chamber attached directly to the mouth of the pump, and point B i n Figure 3 is considered as the mouth of t h e pump, then the writer claims t h a t the results obtained by these authors for the speed at B will be higher than the true speed because the opening t o the gage is pointed into the mouth of the pump. Examination of Table I in their article shows that they obtained a speed factor, or Ho coefficient, of 50 t o 60% for some of the pumps studied, They define the speed factor as the ratio of the speed measured a t the inlet t o the nozzle chamber t o the ideal speed as calculated by the kinetic theory for the annular area between the mouth of the nozzle and the pump wall. The ideal speed is assumed to be 11.7 liters per second per square centimeter for air a t 25 C.
in practice and regards with suspicion any claims t o the contrary. Investigation of such claims has always revealed t h a t either the mouth of the gage is pointed into the pump or the air has been beamed j n t o the pump from a narrow opening. Hundreds of different nozzle designs have been tested in the Distillation Products laboratories, and in spite of every effort t o improve the Ho coefficient no one has ever obtained a reproducible speed corresponding to a speed factor over 45VGwhen using a properly designed test dome. Recently a n independent group a t the Distillation Products laboratories reported speeds of 1000 t o 1200 liters per second for a new 6-inch pump, corresponding t o a speed factor of about 607a, This group used a large test dome in which the air was admitted properly by diffusion over a broad area before entering the pump, but the tubulation of the ionization gage extended into the dome and was bent towards the mouth of the pump, so t h a t the opening t o the gage faced directly into the jet. When it was pointed out that the only molecules which could enter the gage were those scattered back from the jet and the upper part of the pump casing and that these could not possibly equal the flux of molecules into the pump unless the pumping speed were zero, the tubulation was cut off so that the plane of the mouth of the gage was perpendicular t o the plane of the mouth of the pump. The speed then dropped to a maximum of about 750 liters per second, corresponding to an Ho coefficient of about 40’%’,. The question then arose as t o whether the normal t o the mouth of the gage should point back into the test dome in a direction parallel t o the normal to the mouth of the pump as shown at A on Figure 4, rather than being perpendicular as shown at B. At first sight this might seem the logical position, as the aspects of the mouth of the gage and of the pump are then identical and the same number of molecules from the dome should pass through the openings per square centimeter of cross section. However, if one refers t o the exact definition of pressure as given, for example, by Kennard (19, p. 6), it becomes evident t h a t the pressure recorded in this case will be too high. If an imaginary plane surface is drawn in the gas, momentum is continually being transmitted across the surface in both directions by molecules which themselves actually cross the surface. The pressure acting across the surface can be defined as the net rate a t which momentum normal t o it is being transmitted across
The writer is of the opinion that speed factors of over 5070, while theoretically possible, have probably never been realized TEST
DOME
...
Figure 3.
Testing Chamber Used by Copley and Co-workers
Figure 4. Test Dome for Measuring Pump Speeds
May 1948
.
797
INDUSTRIAL AND ENGINEERING CHEMISTRY
it per unit area in the positive direction, momentum transmitted III the opposite direction being counted as negative. Assuming that this definition can be applied to the present case in which the gas is not a t rest, the pressure a t the mouth of a diffusion pump should be obtained by computing the rate of transfer of momentum due to gas molecules coming from the dome and averaging with this the rate of transfer of momentum due to gas molecules scattered back out of the mouth of the pump into the dome. These latter molecules carry negative momentum, but they are traveling in the negative direction; hence the rate of transfer of momentum is positive. Because there is a net flow of gas into the pump, the transfer9 of momentum by molecules passing out through the mouth will be less than the transfer of momentum in the other direction. Data on several standard pumps reveal t h a t the Ho coefficient or speed factor based on the mouth area rather than the jet clearance area is usually about 30%; therefore we may assume that the flux out of a pump is about 70% of the flux into the pump. The total transfer of momentum is thus about 1.7 times the transfer from the dome into the pump. However, in a stationary gas the total flux should be two times the flux in one direction, since just as many molecules would travel u.p through the imaginary plane as pass down from the other side. This would be the condition prevailing in the mouth of% vacuum gage having the tubulation bent in the form of a J, so t h a t the mouth points back towards the dome. Thus the pressure recorded by the gage in this case should be higher than the true pressure in the ratio of 2 over 1.7. It was for this reason t h a t the author had adopted the practice of placing the plane of the mouth of the gage perpendicular t o t h a t of the pump, so that a n intermediate pressure could be recorded. The perpendicular position also conforms to the ordinary method of measuring the static pressure in a moving fluid. On the other hand, it may be shown t h a t the J tubulation with the mouth facing upstream gives the true dynamic.or impact pressure in the gas flowing into the mouth of the pump and this should be equal to the true static pressure at the distance of several mean-free-paths within a sufficiently large dome where the stream velocity is negligible. This point is left open for discussion. It has been suggested that directional effects on ionization gages could be eliminated by removing the electrodes from the glass envelope and mounting them directly inside the test dome as shown at C. This technique was used by Blears (2) in measuring the ultimate pressure of oil diffusion pumps. This suggestion is also left open for discussion. If the air is beamed away from the mouth of the gage, the pressure recorded will be lower than normal; on the other hand, if the mouth of the gage is located directly opposite the air inlet so that the air is beamed directly into the mouth of the gage, the pressure, reading will be higher than normal, and the measured pump speed will be less than the trGe speed. This latter condition is illustrated by Figure 5, which is taken from a diagram of the apparatus used by Matricon (23)for the measurement of pump speeds.
/I
MsLEOD OAUGE
Figure 5. Method of Speed Measurement Used by Matricon will be beamed into the gages, which then record pressure readings greater than the true static pressure in. the two chambers V and B. This effect has been verified by experiments in the author’s laboratories in which the measured speed was nearly one half the true speed. The importance of proper location of the mouth of the vacuum gage and air inlet was first pointed out by Ho (f5)in 1932. H e did not, however, stress the importance of the direction of the mouth of the gage and of the air inlet. The amount of beaming which occurs when air enters a chamber from a narrow tube under conditions of molecular flow was calculated by CIausing (3) in 1930. Ellett (IO) checked Clausing’s results experimentally in 1931. Korsunsky and Vekshinsky (20) also studied the formation of molecular beams in 1945. Figure 6 is a reproduction of a diagram prepared by Clausing (3,38) which illustrates the beaming eEect of the entrance tube. The dotted line shows the cosine law distribution t o be expected from molecular flow through a hole in a thin plate, while the solid line indicates the distribution calculated by Clausing for a short tube with length
Air enters the high vacuum side a t C and flows out through the connection a t D into the fore pump when cutoff R is open. Bulb V and test head B are first well evacuated by operating both pumps. Then a gas t o be studied is admitted to V and R is closed when a suitable pressure is reached. The gas circulates around path ABCD through the capillary at A . The conductance, U,of the capillary is computed from Knudsen’s equation and the speed of the pump is then given by
where p 2is the high vacuum and PI the forepressure.
.
The principal error in the arrangement shown in Figure 5 is the location of the openings t o the McLeod gages directly opposite the exit of narrow tubes, through which gas is flowing. The gas
Figure 6. Beaming Effect of Entrance Tube
INDUSTRIAL AND ENGINEERING CHEMISTRY
798
equal to its diameter. Korsunsky and Vekshinsky calculated the distribution to be expected when silve is kept less than 1.05; because the pressure in the test dome will fall i n the same ratio from pl t o p2, the interval p 1 - pz will be so small that the pumping speed can be considered constant. Thr speed of the pump, S, is then given by
Because there is always the danger of contaminating the pump by accidentally sucking the buret fluid through the needle valve, such fluids as alcohol or mercury have been avoided a t Distillation Products and the custom has been adopted of using the same pump fluid in the buret as is present in the pump under test. The viscosity of organic pump fluids is not as low as that of alcohol; but as it can be shown that the dynamic head is still negligible for these oils, there is no reason why they should not, be used. However, care must be taken to allow the oil t o drain thoroughly from the buret before the next reading is taken. .4thin film of oils always remains in the buret, but in general the slight decrease in cross section is negligible compared to t h r errors involved in measuring the pressure in the test dome. It is advisable t o place only just enough oil in the reservoir t o permit the oil to flow a little beyond the mark a t which the stop ~ a t c h is to be stopped before bubbles of air enter the tube. This avoids to some extent the danger of sucking oil through the needle valve; in wide burets the bubbles rise through the oil, which then remains stationary.
It is possible to keep the factor TTP8log, PljPzequal to a coilstant, C, by always using the same two points, P , and P,, on the manometer and adjusting tube D to keep P3constant according to the prevailing barometric pressure. The speed equation then becomes S = C / p t . I t is also convenient t o choose Pz = PI. Every reading can be repeated as often as desired by rotating stopcock E through 360". Alexander used a reservoir systeni having V = 1000 cc. Since time intervals as short as 5 seconds can be measured accurately with a stop watch, this system could easily measure rates of flow up to 4000 micron-liters per second or a little over 5 cc. per second of air a t atmospheric pressure. For larger rates of flow V should be greater than 1000 cc. Alexander suggests the use of electrical contacts i n the manometer arm to signal points PI and P, and recommends luhricat,ing the meniscus with a few drops of benzene.
Figure 9.
Alexander's Apparatus for Metering Air Flow
Alexander ( 1 ) has described a novel method of metering the air flow. He points out that the mercury pellet and calibrated capillary method is not suitable for high capacity pumps, but does not mention the possibility of using oil burets of large diameter. Nevertheless his apparatus (Figure 9) appears to be convenient for measuring the speed of large pumps and may have some advantages over oil burets. Alarge glass reservoir, A , is connected to a n open-end manometer, B , a flow regulator, CD, and a tube leading to the pump through a two-way stopcock, E , and needle valve, F. Mercury is placed i n both A and C. Tube D is adjustable, so t h a t the length of the submerged part can be altered. The volume of air, V, trapped i n the system, AC, up to the stopcock, E, is accurately calibrated first. This volume, V, remains approximately constant throughout the measurements. Air is then sucked through valves E and F until a steady rate of flow is reached with air bubbling into C from the lower end of D.
s = 1'P pt 3-
log,
P -' pz
A few other methods of metering the air flow have been tried but have not come into general use. A series of calibrated fixcvi leaks can be used to obt'ain several points on the speed curve, but such leaks do not hold their calibration very well unless thr air is thoroughly filt,ered. Sivertsen (28) has designed calibrated leaks consisting of glass capillaries with an inner member which gives a ring-shaped cross section. It is claimed t h a t these leaks are free from clogging by dust particles, because i t is not possible for the whole cross section to be blocked a t once as in thermometer capillaries. The calibration changes linearly with temperature and thus can be calculated for any temperature. Sivertsen points out that' by using various series-parallel combinations almost any desired leak rate can be approximat'ed, and the combined resistance tBogas flow is calculated by formulas resembling those for analyzing electrical networks. Calibrated fixed leaks can also be constructed by flat'tening a length of copper or aluminum tubing, or by sealing porous plugs of fritted glass, porcelain, or other sintered ceramic powder in glass tubes. However, the general opinion of those who have used such fixed leaks is that, the calibration must be checked every few weeks. One method which has been frequently employed is the cir,culation of the gas through the pump and back into the high vacuum side through a capillary t,ube as shown in Figurc, 5 . The pressures involved are such that the conductance of thta capillary can be calculated from Knudsen's equations. This method is described by Dushman (8), Matricon (25), and Hickman and Sanford (14). It is not, suited for obtaining the apcwlpressure characteristics of diffusion pumps under normal opcrating condit'ions, since the forepressure depends on the capillary and quant,ity of gas circulating rat,her than the capacity of a chosen forepump, and the part of the speed curve near the ultimate vacuum cannot be measured accurately because of the partial pressure of organic vapors which circulate with the gas t,hrough the capillary. A unique method of determining the leak rate through a variable leak or needle valve involves measuring the rate of rise of pressure in a large chamber connected t o the pump through a low-impedance vacuum valve. The valve is first opened and the leak adjusted t o give a convenient pressure. The valve id then closed and the rate of rise of pressure measured. The process is repeated for several settings of the leak and a graph of rate of rise us. equilibrium pressure is plotted. The slope of
May 1948
.
INDUSTRIAL AND ENGINEERING CHEMISTRY
the curve a t any pressure times the volume of the chamber gives the speed a t this pressure. The great importance of this method is that errors due to incorrect calibration of the vacuum gage cancel out of numerator and denominator in computing the slope when the same gage is used for determining rate of rise and equilibrium pressures. The volume of the chamber in liters should be approximately equal to the speed of the pump in liters per second for greatest accuracy and convenience in timing. The constant volume method (24, 32) of measuring pump speed is not easily adapted t o obtaining performance curves for high speed pumps and is therefore seldom used for modern diffusion pumps.
Errors Associated with Calibration of Ionization Gages for Air and Pump Fluid’Vapors Ionization gages and Pirani gages have been used almost exclusively a t Distillation Products for measuring pump speeds. Occasionally a system is checked with a McLeod gage and the higher pressures are checked with an oil manometer. The ionization gage is particularly subject to misinterpretation because of the effects of oxygen and organic vapors on the filament, errors due to incorrect calibration for air and pump fluid vapor, and errors due to outgassing and electrical cleanup which depend on the length and diameter of the tubulation. It has been repeatedly observed that high speed values are obtained with an ionization gage (which has been thoroughly outgassed) during the first few hours of operation. This might possibly be due t o removal of oxygen from the air by the clean tungsten filament and would presumably be inhibited as the filament ages by the formation of tungsten carbide from the action of organic vapors on the surface of the filament. It may also be associated with electrical cleanup by driving of positive ions into the collector as described by Schwarz (26). The author has been considering the advisability of running speed curves with pure nitrogen to eliminate any variable effects due t o oxygen in the air. When two or.more ion gages of the same type but with tubulations of different length and diameter are placed on the test dome, the pressure indications are seldom the same and consequently the measured speed is different for each gage. Usually the pressure is lowest in the gage whose tubulation has the greatest impedance t o gas flow because of cleanup in the bulb; however, if the gages are still outgassing, the pressure may be higher in the gage with the higher-impedance tubulation. T h e effects of variations in the size of the tubulation have been studied by Blears ( 2 ) and the author has carried out many experiments a t Distillation Products along the same lines. The errors arising from incorrect calibration of the ion gages for air have nothing to do with the technique of speed measurement (except as they are eliminated by the rate of rise method), but unfortunately have usually been present because of problems involved in calibrating ion gages against McLeod gages. A comparative study of several commercial ionization gages has revealed a wide discrepancy in the rated sensitivity values, and i n general before comparing speed curves it is necessary to know what ion gage was used and what sensitivity was assumed. The shape of the speed curve in the ultimate vacuum region depends on whether or not the partial pressure of the pump fluid vapors is included in the measured pressure. If a McLeod gage or trapped ion gage is used, the measured presshre involves only the partial pressure of air. If the pump fluid vapor is allowed t o enter the ion gage, the form of the speed curve a t the lower pressures depends on the pump fluid used and whether the plate current readings were converted into pressure units by using the sensitivity factor for air or a factor corrected for the presence of the pump fluid vapor. This part of the speed curve also depends on whether or not the ultimate vacuum was subtracted from the pressure reading before the speed was computed and plotted. Before two different speed curves are compared on the same
801
pump, or similar pumps, all these factors must be taken into consideration. It is possible t o evaluate some of these factors and arrive at a common basis for comparison of pumps by using certain formulas presented below. Derivation of these formulas is omitted but should be clear to those familiar with the usual equations in the books by Dushman (8),Strong (29), Dunoyer ( 7 ) ,Monch (25),Yarwood (Sd), et al. Formulas. The formula which is used for computing speed at Distillation Products is
,
s = 750 Pt
where S is defined as the speed of exhaust ( I S , 21) of the test dome in liters per second, when p is the total pressure in microns as indicated by an untrapped ion or Pirani gage using the calibration factor for air, and t is the time in seconds for I cc. of air a t atmospheric pressure (assumed to be 750 mm. of mercury) to enter from the buret, Some-laboratories prefer t o compute the so-called ‘’true speed’’ (8, 21) or admittance A from the formula
where po is the ultimate vacuum (pressure corresponding io t = m and S = 0 in the first equation). A is theoretically independent of p when the fqrepressure is sufficiently low. The speed of exhaust is related t o the admittance by the equation
(3)
A graph of this formula shows that S is approximately constant and equal to A until p is 20 times po and then decreases rapidly t o zero as p approaches PO. The true ultimate pressure, pol is difficult to measure because of the presence of pump fluid vapors for which the ion gage calibration factor is uncertain. Some laboratories prefer to use trapped gages or McLeod gages and ignore the partial pressure of the vapors. This is the best procedure for determining the admittance by Formula 2, because the presence of vapors makes the conversion of ion gage readings into the true pressures, p and PO, very difficult. I n this case p and po should be replaced by different symbols, p’ and p i , indicating partial pressures of air rather than total pressures. However, a graph of A us. p ’ does not represent the true performance of a pump unless a large cold trap is added, and the latter introduces an impedance which reduces the effective value of A. Because one of the outstanding features of the pumps developed a t Distillation Products is the ability to attain a low ultimate pressure without the use of cold traps, and the ultimate vacuum of untrapped fractionating oil pumps has been shown to be determined primarily by the vapor pressure a t room temperature of the pump fluid which collects on the walls back of the top jet (or high vacuum jet) due to backstreaming, the Distillation Products laboratories prefer, t o use untrapped ion gages and compute the speed from Formula 1. One difficulty with this procedure arises from the uncertainty as t o the calibration factor of the ionization gage for different pump fluid vapors (17). Rough measurements indicate that the sensitivity of an ion gage for Octoil vapor is 3 to 5 times the sensitivity for air, but the sensitivity depends in such a complicated way on room temperature, drafts on the tube, and the tubulation ( 2 ) t h a t no attempt is made t o correct the readings for the vapor calibration factor. The plate current in microamperes is converted directly into pressure units by using the calibration factor for air and this value of p is substituted in Equation 1. Although this admittedly introduces some error in the performance curve near the ultimate vacuum, it a t least avoids any misrepresentation by displaying a falsely low ultimate pressure and high speed in this region. If p is replaced by the quantity (p’ k p ” ) where p ‘ is the partial pressure of air, IC is the ratio of the sensitivity of the
+
,
a
802
*
s
INDUSTRIAL AND ENGINEERING CHEMISTRY
’
Vol. 40, No. 5
tion of the jets, or changes in the partial pressure of the vapors during the time involved in obtaining the original speed curves. Comparison of Speed Curves i n Forepressure, Breakdown Region
I
PO
10-5
PB I0-5
10-4
10-2
10.3
IO0
10-1
P Figure 10. Typical Speed Curve of Oil Diffusion Pump ion gage for air to the sensitivity for pump fluid vapor, and p‘’ is the partial pressure of pump fluid vapor using the air calibration factor, then the computed speed would be higher and the curve would show a better ultimate vacuum, p o = p6 kp;, since IC is less then unity. If.for some reason \ye wigh to compare jets on the basis o f the assumed theoretical admittance, the relation
+
A =
S P
A
- Po
+
may be used, providing the ultimate pressure, p, = p; kp:, is measured within a few minutes of the recording of the total kp“. Under these conditions the partial pressure, p = p’ pressure of pump fluid vapors in the dome (and more particularly in the ion gage) will be constant, or pg = p”. Thus Equation 4 reduces t o
+
(5)
which is ip agreemcnt with the admittance as computed by those laboratories which use trapped ion gages or McLeod gages. However, the ultimate pressure, pa, as measured by a n untrapped ion gage changes rapidly with time because of the evolution and adsorption of vapors by the wall of the tube. The wall is heated when the filament wattage is increased t o maintain emission at the higher pressures and cools again when the filament temperature is lowered as the pressure falls. Therefore, we frequently prefer t o compute the admittance A from the formula
where SI, p l and S2,p~ are coordinates of two points closely spaced on the “smoothed” speed curve through experimental points obtained from Equation 1. The highest value of p2 should not exceed 15 po. If the pump operates over a wide pressure range, so that a plateau of constant speed is reached for pressures above 20 p o , it will usually be found that the value of A computed from Equation 6 is equal t o this maximum, or plateau speed. However, if the pump begins t o break down before the pressure exceeds 20 po, then A will probably be greater than the peak of the speed curve (S vs. p ) and represents the peak speed which could in general be obtained if a larger forepump, or a n oil of lower vapor pressure were t o be used. It is sometimes very instructive t o plot values of A obtained from Equation 6 for a series of successive points on thg. smoothed speed curve between p = po and p = 15 p , against the pressure midway between two successive points, and to compute values of p o from
(7) ’
where A and p are coordinates of points on this new curve. If the values of A and pa thus obtained are not constant, investigation usually reveals errors in ion gage readings, abnormal opera-
The speed of diffusion pumps is practicallj independent of thcs foiepressure when the latter is less than a certain critical value known as the breakdown forepressure or limiting forepressure, denoted by FB. As shown by a typical speed curve, Figure 10, the speed of pumping decreases rapidly to that determined by the forepump as the pressure, p, on the high vacuum side of the diffusion pump rises above the critical value, p ~ at, which the forepressure, F, equals F B . If S B represents the maximum speed of the pump (which usually occurs at pressure p ~ and ) C B is the capacity in liters per second of the forepump a t pressure F B , then the equation of continuity requires that ;DB = F B C B / S B ( 8) The limiting forepressure, F B , depends on the jet design, pump fluid, gas load, heater input, and cooling. Sometimes F B is distinguished as ‘[static” or “dynamic,” depending on the method of measurement. The static value is obtained by throttling the forepump with a valve or by introducing a leak in the fore vacuum line. The dynamic value is the forepressure reading when the jet breaks as the result of increasing the leak rate on the high vacuum side when using the full capacity of the forepump. 811 these factors, as related by Equation 8, must bv taken into consideration before comparing speed curves in the forepressure breakdown region. Equation 8 indicates that the maximum stable fine pressure or “head” pressure, p a , can be increased by increasing the capacity, CB,of the forepump whenthe limiting forepressure, FB,remains relatively constant. However, for multistage pumps there is ail absolute upper limit for p~ which depends only on the design and operation of the diffusion pump, and above this the dynamic value of F B decreases as C B is increascd. This limit depends 011 the forepressure breakdown of the weakest stage a t the mat;imum gas load, Therefore, unless the capacity of the forepump is so large that this upper limit of p~ has been reached a t the given heater input, the exact pumping speed of the forepump system a t all forepressures must be known before the performance of diffusion pumps is compared in the breakdown region. Also, since F B depends on the boiler pressure, a true comparison of jclt designs cannot be made in this region unless the same oil and the same boiler temperature are used, or for over-all comparison, the same heater input and rate of cooling must be applied t o the pump casings. If the pump fluid is a hydrocarbon mixtuie, the length of time during which the oil has been boiled undci vacuum must be taken into account, as most hydrocarbon oils tend to become heavier by loss of volatile fractions (and perhap& by cracking and polymerization) during the first few days o f operation. The measurement of speed in the forepressure breakdo\\ 11 region is complicated by the unsteady operation of the jets The vacuum oscillates between rather \Tide limits, as indicated by the swinging of the needle on the pressure meter of a Pirani gage while the individual jets of a multistage pump beconic, overloaded and then recover again. This action is usual13 accompanied by oscillations in the forepressure, F , which increases as the head pressure, p , decreases, and vice versa. The speed, S,and the ratio, F / p , thus rise and fall together as thc steady rate of leakage of air through the needle valve is broken into surges of air from the test dome into the fore vacuum line. The period of oscillation probably depends on the volume of the fore vacuum line. The effect is also usually associated with turbulence in the vapor jet and fluctuations in boiler temperature which occur as the amount of oil circulated begins t o vary. The discharge of vapor from the nozzles is constant as long as the forepressure, F , is less than about 0.6 times the boiler pressure,
May 1948
INDUSTRIAL AND ENGINEERING CHEMISTRY
P, but falls rapidly when the forepressure exceeds this value. As t h e discharge rate decreases the boiler temperature rises until the rate begins to increase and normal jet action is temporarily restored. If one records the upper and lower limits of oscillation on the Pirani gage for a given leak rate and plots both values, as well as the mean, on the graph of speed vs. high vacuum, it will usually be found that all of the points fall nearly on the same smooth curve regardless of what pressure is chosen within the range of oscillation on the gage. This is merely a coincidence arising from the fact t h i t the speed is computed,from the formula
where t is constant during the measurement while the actual variation of speed with pressure in this region follows the law
FC
#=P where FC is nearly constant over a range of p corresponding to‘ the range of the oscillations.
Literature Cited Alexander, P., J . Sci. Instruments, 23,11-16 (1946);21,216-18 (1944). Blears, J., Nature, 154,20 (1944); Proc. Roy. S O L ,188, 62-76 (1946). Clausing, P., 2. Physik, 66,471 (1930). Copley, M., Simpson, O., Tenney, H., and Phipps, T., Rev. Sci. Instruments, 6, 265-7 (1935). Czerny, M., and Murmann, H., Physik. Z . , 30,462-3 (1929). Downing, J. R., private communication from National Research Corp. Dunoyer, L., “Vacuum Practice,” pp. 1-4, New York, D. Van Nostrand Co., 1926.
803
(8) Dushman, S., “High Vacuum,” pp. 38-40, Schenectady, N. Y., General Electric Co., 1922; J . Franklin Inst., 211, 691-6
(1931). (9) Ebert, H., 2. Instrumentenk., 51, 337 (1931). (10) Ellett, A , , Phys. Rev., [2]37, 1699 (1931). (11) Eltenton, G.C., J. Sci. Instruments, 15,415 (1938). (12) Farkas, A.,and Melville, H. W., “Experimental Methods in Gas Reactions,” London, Macmillan Co., 1939. (13) Gaede, W., Ann. Phys., 41,337 (1913). (14) Hickman, K. C.D., and Sanford, C. R., Rev. Sci. Instruments, 1, 154-9 (1930). (15) Ho, T.L.,Ibid., 3,133 (1932);Physics, 2,386-95 (1932). (16)Howard, H.C., Rev. Sci. Instruments, 6,327 (1935). (17) Kapff, S. F.,and Jacobs, R. B., Ibid., 18,581-4 (1947). (18) Kaye, G. W.C., “High Vacua,” p. 162,New York, Longmans, Green and Co.,1927. (19) Kennard, E. A., “Kinetic Theory of Gases,” pp. 6-9, 306-8, New York, McGraw-Hill Book Co., 1938. (20) Korsunsky, M.,and Vekshinsky, S., J . Phys., U.S.S.R., 9,399404 (1945). (21) Langmuir, I.,Gen. Elect. Rev., 19, 1062-6 (1916). (22) Loeb, L. B., “Kinetic Theory of Gases,” 2nd ed., pp. 301-10, New York, McGraw-Hill Book Co.,1934. (23) Matricon, M., J . phys. radium, 3, 127-44 (1932). (24) Mills, P. J., Rev. Sci. Instruments, 3,309 (1932). (25) Monch, G., “Vakuumtechnik im Laboratorium,” Ann Arbor, . Mich., Edwards Bros., Lithoprint, 1944. (26)Sohwarz, H., 2. Physik, 122,437-50 (1944). (27) Sinel’nikov, K. D., Val’ter, A. K., Gumenyuk, V. S., and Taranov, A. Ya.,J. Tech. Phys. (U.S.S.R.), 8,1908-22 (1938). (28) Sivertsen, J., Instruments, 20, 333-4 (1947). (29)Strong, J., “Procedures in Experimental Physics,” pp. 97-101, Ne* York, Prentice Hall, 1938. (30) Wachter, H., and Scheer, J. van der, 2. tech. Physik, 24, 287-91 (1943). (31) Westin, S.,and Ramm, W., Kgl. Norske Videnskab. Selskabs, Skrifter, 1936, No. 9. (32)Yarwood, J., “High Vacuum Technique,” Chap. 3, hfew York, John Wiley & Sons, 1945. RECEIVEDNovember 3, 1947. Communication No. 128 from the Laborstories of Distillation Products. Ino.
Dehydration of Orange Juice A. L.
Schroeder
NATIONAL RESEARCH CORPORATION, CAMBRIDGE, MASS.
R. H. Cotton‘ NATIONAL RESEARCH CORPORATION, PLYMOUTH, FLA.
-4 study of the drying of orange juice is presented. The general approach and technique are applicable to the drying of most dilute solutions. Data are presented regarding the variation of drying rates under different conditions of temperature, pressure, and charge in the dryer. The method of drying from viscous thin films is outlined by which maximum production rates and product quality may be obtained. The effect of processing temperature on production rates is discussed, and a brief rBsum4 of the storage stability of orange juice powder is given.
T
HE dehydration of orange juice has been of great interest for many years. The economic advantages are obvious. Suffice it t o say that the citrus crop is growing larger every year, particularly in Florida, and the citrus industry may be confronted with a serious distribution problem. Development of processes and techniques for handling t h e crop willopen the road for stabilizing t h e industry. The large volume of canned citrus juice and the relatively new concentrate illustrate the current trend. The production of a satisfactory citrus powder, such as orange juice pow1
Present address, Holly Sugar Corporation, Colorado Springs, Colo.
der, will add another means for obtaining a broader market domestically and eventually throughout the world. Numerous investigators have been interested in drying orange juice. Some excellent work was done at the Massachusetts Institute of Technology during the recent war by t h e staff of the D e partment of Food Technology (I).‘ Flosdorf has described equipment and summarized a drying technique (2, 3)and Moore et al. (6) have reviewed t h e general developments of drying and concentrating in practice u p t o 1945. During t h e past seven years, this laboratory has been working on t h e general problem of drying heat-sensitive liquid materials, of which orange juice offers a good example. The work was instituted in Boston on a small scale and in 1944 was transferred t o a Citrus Regearoh Laboratory at Plymouth, Fla. Hayes, Cotton, and Roy (4) reviewed some of t h e problems and results in 1946. The present paper presents additional data regarding the drying of orange juice: Initial data comparing the drying of whole orange juice and a concentrate of orange juice. Results of an investigation of temperature ?nd moisture conditions existing within a drying mass of orange juice. A curve t o illustrate the “falling rate” period of drying. The effect of temperature on production rate for different final moistures.
