High Temperature Reaction Rates in Hydrocarbon Combustion

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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

1634

Vol. 47, No. 8

T h e microscopic size analysis was made by G. C. Williams of thiE laboratory.

Langmuir, I., J. Am. Chem. Soc., 37, 1154 (1915). Letort, M., and Magrone, R., J. chim. p h y s . , 41, 576 (1950). Long, F. J., and Sykes, K. W., Proc. Roy. SOC.(London), A193,

LITERATURE CITED

Mellor, J. W., “Comprehensive Treatise on Inorganic and Theoretical Chemistry,” Longmans, Green, London, 1924. Natl. Bur. Standards, Washington, D. C., “Tables of Selected Values of Chemical Thermodynamic Properties,” 1949. Petersen. E. E.. Ph.D. thesis. Pennsvlvania State College. 1953. Petersen; E. E.’, Walker, P. L., Jr.,”and Wright, C. C.: A S T M

377 (1948).

Abbott, H. W., private communication, 1950. Barrett, E. P., Joyner, L. G., and Skold, R., J . Am. Chem. Soc., 73, 373 (1951).

Ronner, F., and Turkevich, J., Ibid., 73, 561 (1951). Bowring, J. R., and Crone, H. G., J. chim. p h y s . , 47, 576 (1950). Broom, W. E. J., and Travels, XI. W., Proc. Roy. SOC.(London), 135A, 5 1 2 (1932).

Gbdsby, J., Hinshelwood, C. N., and Sykes, K. W., Ibid., 187A, 129 (1946).

Gadsby, J.,’Long, F. J., Sleightholm, P., and Sykes, K. W., Ibid., 193A, 357 (1948). Goring, G E., Curran, G. P., Tarbox, R. P., and Gorin, E., IND,ENQ.CHEM.,44, 1057 (1952). Gulbransen, E. A., and Andrew, K. F., Ibid., 44, 1048 (1952). Hougen. 0. A., and Watson, K. SI., Ibid., 35, 529 (1943). Kelley. K. K., U. S. Bur. Mines, Bull. 371, 17 (1934). Kpttler, F., J. Franklin Inst., 250. 339 (1950). Ibid., 251, 499 (1951).

Bull., 183, 7 0 (1952).

Ritter, H. L., and Drake, L. C., IND.ENG.CHEM.,ANAL.ED., 17, 782 (1945).

Spencer, H. M., and Justice, J. L., J . Am. Chem. Soc., 56, 2311 (1934).

WAlkeryP. L., Jr., Foresti. R. J., and Wright, C. C., IND. ENQ. CHEM.,45, 1703 (1953). Washburn. E. W.. Proc. Natl. Acad. Sci., 7 , 115 (1921). Wheeler, A., Advances in Catalysis, 3 , 249-327 (1951). RECEIVED for review May 10, 1954. ACCEPTED February 28, 1955. Contribution 53-56, College of Mineral Industries, The Pennsylvania State University.

High Temperature Reaction Rates in Hydrocarbon Combustion J

JOHN P. LONGWELL AND MALCOLM A. WEISS Esso Research and Engineering Co., Linden, N . J.

I

N THE design of high output combustion chambers, par-

ticularly for supersonic missiles and aircraft, minimizing the combustion chamber volume is invariably a major goal. An ultimate minimum is set b y the over-all kinetics of chemical reaction between fuel and oxygen. It is more likely that this ultimate will be reached a t low combustion chamber pressures where the reaction rates are much slower (both absolutely and relative to mixing) and where large combustion volumes may thus be required. I n low output combustion equipment (furnaces, for example), combustion rates are usually limited by heat and mass transfer processes. These include the evaporation of fuel drops, mixing of air and fuel, heating the mixture to ignition temperature, etc. It is usual, and usually valid, to assume that the actual burning process (of chemical reaction) takes negligible time in comparison

E X I T HOLE 1/4” D. I ONE OF1 6 0 ?

R

\-

“\A,+,*\,

Figure 1. Typical spherical combustion reactor

with the time required for the transfer processes. However, as the transfer processes are made faster (or reaction rate slower), conditions can eventually arise such that the time required for chemical reaction becomes important and eventually controls. Clearly, this may occur in low pressure, high output equipment where the maximum possible space heating rates are desired Knowledge of the reaction times required by chemical kinetics limitation, with no mixing limitation, is then of the greatest significance. No improvement in equipment design can reduce the total reaction time to a value lower than that demanded by the over-all kinetics. Similarly, knowledge of kinetics-limiting time points out what theoretical gains in performance are possible in actual equipment. The work described in this article is part of a continuing study to determine experimentally, and interpret theoreticall), combustion rates limited only by chemical kinetics. Interest in combustion rates limited by kinetics has been expressed in many recent reports (1, 3, 8, 9, 1.6, 16, 17’). I n most of this work, the intent was t o use kinetics simply as a tool to describe qualitatively the particular aspect of combustion under study. Numerical values of the kinetic constants were thus simply assumed for illustrative purposes or approximated from other types of data b y making admittedly questionable major assumptions. These approximations include, for example, the extrapolation of low temperature hydrocarbon oxidation rates t o high temperature hydrocarbon combustion rates. Other schemes involve application of semiempirical laminar flame speed theories or of flow patterns in the wake of a bluff body immersed in an air stream. Workers at the Applied Physics Laboratory, The Johns Hopkins University, have developed very promising techniques for direct determination of high temperature kinetics. Temperature profiles through a laminar flame front have been determined by a particle track method ( 5 ) , and composition profiles have been determined by traversing with a very small probe leading directly

August 1955

INDUSTRIAL AND ENGINEERING CHEMISTRY

to a mass spectrometer ( I S ) . The combination of the two profiles yields direct reaction rate information. I n a similar type of work, Friedman and Burke ( 4 ) traversed a laminar flame front with a small thermocouple to obtain temperature profiles. By making assumptions about local compositions, which were not determineb, local space heat release rates were estimated. Obviously, such estimates can also be made from the data of Fristrom and coworkers ( 5 ) and Prescott and coworkers ( I S ) . All of these experiments require a high degree of precision, and only limited results are a t present available. Regardless of the success of such measurements, any other method for direct determination of the reaction rates in high temperature combustion would be desirable. The difficulties of direct determination of the rates are evident if the speed of the reaction and the extreme temperatures involved are considered.

Figure 2.

INSULATED SPHERICAL REACTOR USED TO DETERMINE COMBUSTION RATES

Tho chief experimental problem in this work was, of course, the construction of a n experimental device in which high temperature combustion rates could be obtained without mixing limitations. After 2 years of development, a suitable react,or design was achieved. A typical reactor of this design is sketched in Figure I, and the reactor, with one hemisphere removed, is shown in Figure 2. The design objective was the creation of a reactor to simulate a perfectly stirred flow system a t steady state. Ideally, a homogeneous mixture of air and fuel of known composition and temperature enters the reactor and is instantaneously and perfectly mixed with the burning combustion gases within the reaction volume. (The reactor contents are thus uniform in composition and temperature.) Simultaneoasly, products of combustion are withdrawn from the reaction volume; these products are identical in temperature and composition to the reactor contents, and the exiting mass flow is equal to the entering mass flow. In addition, the reactor is adiabatic, has minimum surface-tovolume ratio, and can withstand reaction temperatures exceeding 2000' K. under both oxidizing and reducing conditions. It will provide for average residence times as short as about second and can be operated over a wide range of internal pressures. The spherical reactor of Figures 1 and 2 necessarily is a compromise among the various characteristics of the ideal reactor described above.

1635

C o m b u s t i o n reactor w i t h o n e hemisphere removed

perature. Thus the problems of fuel evaporation and fuel-air mixing are eliminated. The homogeneous mixture, under pressure, enters both ends of the inlet tube (Figure 11 of the reactor. The mixture then flows to the center of the small central perforated ball and issues as small sonic jets from the holes in the ball into the surrounding combustion gases. These burning combustion gases fill the reaction volume--the volume bounded by the inner insulation surface and the inlet tubes and ball. The numerous small jets of inlet mixture stir the reactor contents vigorously; thus, the inlet fuel-air niixture rapidly (compared to the average residence time) loses its identity, and temperature and composition within the reactor approach homogeneity. Burned mixture exits through the many holes, arranged as symmetrically as possible, in the insulating shell. This shell is provided both to withstand combustion temperatures and to maintain the reaction zone as adiabatic as possible; it is fabricated in two hemispheres carved from high temperature insulating firebrick with a nominal service rating of about 1900" K. The outer thin nickel shell serves to retain the firebrick and to facilitate making connections to the shell. This shell also consists of two hemispheres and is assembled by bolts through rim lugs. The entire assembly is enclosed in pipe sections (provided with appropriate connections and sight ports), which are connected to a suitable exhaust system, allowing control of the reaction pressure.

Several different reactors of this general design were used in the tests, as described in Table I. The reactor shown in Figure 1 is designated as reactor 3b in Tables I and I1 and the reactor of Figure 2 is reactor 3e. I n order to ignite the reactor, fuel and air rates were first set approximately and the pressure was maintained near atmospheric. The hITdrogen ignit,ion jet of Figure 1 allowed a jet of burning hydrogen (ignited by the spark shown) to penetrate I n accessory equipment, air and fuel are separately metered through one of the exit holes into the reaction volume. The mixthrough calibrated orifice meters and rotameters, respectively, t ire ignited readily and, as soon as the reactor walls reached a mixed to homogeneity, and preheated t o the desired inlet temreasonably stable temperature, both hydrogen and spark were turned off. They remained off until reignition was required. T a b l e I. Spherical Reactors After burning began, inlet temperaNomiture, pressure, and equivalence ratio nal NomiInlet Tube Inside nal (fuel/air ratio divided by the stoichioReaction Inlet Holes Outsid? Diam. Outside Exit Volume metric fuel/air ratio) were adjusted ae Reao- (an), Diam., Size, Size, Diam., ( V ) ,Liters tor Inches Inches Number inch 'in,. Number inch Inch desired. I n the majority of runs, blow0 242-0 257 68 0,047 3/1b '/P 2,'s 3s. 3 5 60 outs were determined. That is, for a 0 230-0 249 68 0.047 '/< '/r 3,'s 3b 3 5 BO 0 239 68 0.047 5/16 3/s 3/s 3c 3 5 BO burning system, a sufficient increase in 0 236 3d 3 5 60 3/s '/? 102 0.047 '/1R 0 242 3/16 3e 3 5 60 '11 x/& 68 0.047 mass flow, decrease in pressure, or change 1 800 1 880 84 0 094 1-5/a 1/2 3/4 6a 6 8 84 in equivalence ratio away from stoichio(conical) metric will cause the flame suddenly

INDUSTRIAL AND ENGINEERING CHEMISTRY

1.636

~~

Table 11. Inlet Mixture Equiva- Pressure Air R a t e Temp. To lenee ( P ) Atm. ( N ) , G.- (To) Toa, Reactor Ratio (6) Abs. Mole/Seo. K'. K. 3b 3b 3b 3b 6a 3b 6a 3b 3b 6a 6a 6a 6a 6a 6a 3b 3b 6a 3b 6a 3b 3c 3b 30 3s. 3d 3h 6a 3d ~~

Cia

30 3d 3a 3d 3a 3b 3b

0.412 0.439 0.469 0.476 0.476 0.482 0.495 0.497 0.519 0.521 0.524 0.530 0.530 0.540 0.553 0.555 0.567 0.571 0.672 0.584 0.588 0.590 0.602 0.603 0.606 0.620 0.624 0.625 0.652 0.653 0.658 0.669 0.671 0.675 0.700 0.707 0.718

.

0.967 0.650 0.965 1.002 0.991 0.466 0.662 0.667 0.500 0.650 0.500 0.500 0.995 1.005 0.823 0.682 1,000 0.331 0.333 0.690 0.498 1.000 0.233 0.833 0.213 1.000 0.250 0.575 1.000 0.244 0.668 0.833 0.296 0.836 0.196 0.483 0.187

0.158 0.163 0.255 0.717 2.44 0.174 2.46 0.257 0.255 2.40 2.43 2.46 5.64 6.20 6.22 0.683 1.28 2.41 0.257 6.34 0.703 1.82 0.161 1.83 0.248 2.34 0 255 A 34

i.3.j

2.50 1,78 2.36 0.535 2.38 0.299 1.14 0.255

420 425 405 425 393 418 420 403 412 422 419 419 420 403 410 415 407 416 405 401 430 382 410 382 468 373 401 400 367 414 377 372 436 371 440 412 401

550 53" 34 11 11 50" 11 36 35 10 I1 11 4 4" 4" 13 9 12 33 4a 13 7 570 7 42 6 43 4a

6 11 7 5 26 6 46 lla

47

1.0

Corrected t o TOE __ = 400° K. N/VPz N / V P l . @ 0.974 2.05 1.57 2.73 1.61 4.32 2.91 3.20 5.09 2.93 5.0s 5.15 2.78 3.39 4.86 6.22 5.02 11.6 11.5 7.46 10.8 8.90 16.3 12.8 18.8 12.1 21.2 10.8 12.5 22.4 19.8 17.2 22.6 17.9 33.6 20.2 38.9

O 0 7

m

a

0

5 0.6

0.968 1.88 1.56 2.73 1.61 3.71 2.68 2.95 4.43 2.69 4.40 4.49 2.78 3.39 4.68 5.76 5.02 9.30 9.23 6.92 9,40 8.90 12.2 12.3 13.8 12.1 16.1 9.67 12.5 16.9 18.3 16.6 17.7 17.3 24.3 17.5 28.4

'\ HIGHEST

a

00

MASS

0

0

I

3

0

ob

0

t-

w

1-

00

Inlet Mixture Equiva- Pressure Air Rate Temp. Tolenee ( P ) Atm. ( N ) , G.- CZo), Toe, Reactor Ratio (6) Abs. Mole/Sec. K. K. 3e 3d

3% 30

3a 3e 3d 3d 3e 3a 3d 3a 3d 3a 3E

3d 3a 3d 3a 3a 3s 3a 3a 3a 3e 3e 3a 3b 3b 3b 3b 3b 3e 3c a

0.721 0,732 0.740 0.749 0.773 0.793 0,807 0.840 0.868 0.869 0.869 0.913 0.942 0.980 0.986 0.989 1.06 1.09 1.12 1.25 1.31 1.37 1.45 1.60 1.61 1.67 1.68 1.70 1.73 1.74 1.74 1.82 1.83 1.96

1.02 2.36 0.243 1.80 0.534 1.02

ni

411

1 420 __..

2.07 1.04 0.537 1.02 0.252 1.02 0.531 0.255 1.03 0.254 0.989 0.259 0.258 0.255 0.482 0.482 0.484 1.32 0.775 0.484 0.418 0.162 0.316 0.410 0.324 0.402 0.271

13 5 53 7 29 12 12 8

422

14 30 13 65 _12 . 33 66 12 65 11 78 74 68 29 31 30 10 17

421 439

28 46 300 25 30a 30 40

460

41 6 ~

445 462 418 450 41 9 447

455

27 -.

Corrected to

= 400' K. _______

N/VP'

h'/VP"'

25.5 27.5

21.4 25.4

32.9 40.6 36.9 42.1 46.1 48.1 60.3 51.2 61.5 65.6 56.2 68.6 59.5 46.2 34.2 30.5 23.9 16.9 6.89 5.79 3.49 3.04 3.78 5.91 2.59 1.88 1.31 2.06 1.34

26.8 32.4 32.4 33.5 36.0 37.7 40.0 39.9 37.2 43.2 43.4 45.1 45.5 31.9 24.1 21.7 18.3 13.4 6.00 5.78 3.45 2.87 3.49 4.79 2.39 1.86 1.31 2.03 1 32

4:3:6 . 9

E:; 27.7

Estimated.

BLOWOUT DATA TAKEN OVER WIDE RANGE OF VARIABLES

Two sphere sizes and two similar fuels were used in these tests. The spheres were nominally 6 and 3 inches in inside diametergiving an 8 t o 1 reaction volume range. I n the 6-inch sphere, the fuel was solvent naphtha-a commercial petroleum fraction with physical properties as previously reported (IO) and average chemical composition close to heptane. I n all 3-inch sphere tests, the fuel was reference quality iso-octane (99+ 2,2,4-trimethylpentane). Extensive data from an earlier type of spherical combustion reactor, previously illustrated (8), showed no difference in performance between these two fuels when they were mixed homogeneously with air; this conclusion has been confirmed by all experience in this laboratory with various burner types. Burning pressures were varied between about 1 and '/a a t mosphere absolute. The upper limit was dictated by the fact that reaction rate increases much more rapidly with pressure than does mixing. It j s easier to avoid mixing limitations by remaining in a pressure regime where reaction rates are relatively slow, Thus, an arbitrary upper limit of 1 atmosphere was set. Mass flow limitations prevented the use of pressures below */( atmosphere.

00

0

0.6

Vol. 41, No. 8

~

Spherical Reactor Blowout Data

to be extinguished. I n practice, blowout points were determined a t constant air rate by either changing fuel rate or decreasing pressure in very small increments. Sufficient time was allowed a t each step for steady state to be established. Finally, any additional small change in the same direction would extinguish the flame. This was taken as the blowout point. For lean mixtures, blowouts were very sharp and reproducible. For rich mixtures, they were much less so. Rlon~outswere denoted by a sudden cooling of the hot internal firebrick wall-observed by looking a t the wall through a sight port and gas exit hole. If the pressure was then rapidly increased, or equivalence ratio quickly changed toward stoichiometric, reignition could usually be accomplished by the hot brick wall without requiring use of spark and hydrogen.

vj

~~

0

'\

S,LVER

STAINLESS STEEL HYPODERMIC T U B I N G

A\

SOLDER

COOLANT EXIT

P.NNULUS

c

13 0

30

00

0 -J

m

0 - BLOWOUT EQUIVALENCE RATIO

Figure 3.

Range of blowout data

~OOLANT FLOW

Figure 4 .

COOLANT/ I N L E T ANNULUS

Cross section of sampling probe tip

August 1955

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

1637

crease in inlet temperature. There is a slight error caused by Early work suggested that it would be possible to correlate preheating of the inlet mixture in the section of inlet tube the blowout data by plotting the blowout equivalence ratio actually passing through the hot firebrick wall. However, this against a correlating group of the type N/VP",where N is the error has been neglected. mass flow, V the reaction volume, P the pressure, and n some Effective Inlet Mixture Temperature. Despite the presence unknown pressure exponent. T o determine this exponent i t is of the insulating shell, there is some heat loss by conduction desirable, for a given value of NIVP" (and thus presumably for through the shell. The reactor is nonadiabatic to the extent of a given value of equivalence ratio), t o vary N , V , and P indethis loss. By means of a n optical pyrometer, used by sighting pendently as widely as possible. Minimum values of V are set through sight port and exit hole, both interior ceramic and b y the increasingly large surface-to-volume ratios of smaller exterior nickel average wall temperatures were measured. The spheres. The results would be undesirable increased surface effects and heat losses. The upper limit to V and sphere size is observed temperature differential varied from about 250" to 700" K. with absolute interior temperatures to about 2000" K. set b y laboratory limitations of air supply and exhaust capacity. The range of blowout pressures for the data of Table I1 is Heat loss through the brick shell was computed from this differential and the brick thermal conductivity. Values of the latter, shown in Figure 3. This range is, of course, conditioned by mass supplied b y the manufacturer, averaged about 0.002 gram-caloflow limitations in order to maintain a given NIVP" for a given ries per (second) (cm.) (' K.) in the temperature range of interest. V . Thus the low pressure limitation is really a minimum mass I n the equation for spherical conduction, the presence of the exit flow limitation. At lower mass flows, heat losses from the holes was ignored. These holes occupy between about 10 and sphere become progressively more important and cause unde25% of the interior surface area and less than 10% of the exterior sirably large corrections. Where 1 atmosphere is not the upper surface area. Although some error is thus introduced, i t tends pressure limit, this limit is set by the upper limiting mass flow. Beyond this mass flow, damage t o the reactor is so rapid and to be offset by radiation from the hot walls of the exit holes. severe that there is not enough time to obtain useful data. These There is negligible direct radiation loss from the "bottom" of considerations apply for lean mixtures. Because of poor reprothe hole because, looking along the hole axis, one sees only the ducibility and reactor damage, no attempt was made to investirelatively cool central injection ball. Once the heat loss has gate thoroughly the rich range. been computed, i t is most conveniently expressed as an effective For these reasons, mass flows in the small reactors were varied lowering of the inlet temperature; reactor behavior is not infrom about 0.16 to 2.4 gram-moles of air per sec6nd; in the large fluenced by the fact that a given amount of heat is lost by conreactor, extreme values were not attempted and mass flow rates duction through the walls rather than by being absent from the varied from about 2.4 t o 6.3 gram-moles per second. I n all inlet mixture. Lowering of the effective inlet temperature is cases the inlet fuel-air mixture entered at about 400' K. The directly computed b y dividing the heat loss by the product of over-all equivalence ratio range extended from about 0.4 t o 2.0. entering total mass flow and its specific heat. Equivalace Ratio. All blowout data are tabulated in Table The variations in reaction volume for a given reactor type I1 in order of increasing equivalence ratio. During all runs (Table I ) were chiefly due to small differences in the diameters of with lean mixtures a small water-cooled probe (Figure 4 ) made the many firebrick shells used. For a given inlet assembly and from three concentric lengthp of stainless steel hypodermic tubing nickel shell, a dozen or more firebrick liners may have been used. was kept in position in the sphere, as shown in Figures 1 and 2. The life of a liner is critically dependent on the conditions of opThrough this probe, a sample of the combustion gases was eration. A single liner might last for only one or two runs at high withdrawn continuously, compressed by a diaphragm pump, mass flows and near-stoichiometric equivalence ratios. For passed through a packed furnace t o complete combustion, and very lean or very rich mixtures a t low mass flows, a liner would then led to a n oxygen meter (paramagnetic type). By simple remain intact (without fusing or crumbling) for ten or more runs. stoichiometry, the reqidual oxygen in the completely burned Table I1 also includes the groups correlating mass flow, presgases could be translated to equivalence ratio, 6. The + so sure, and reaction volume corrected to a base of 400' K. inlet obtained is the value tabulated in Table 11. For rich mixtures, temperature. The method of correction from TOE(the effective this technique is not applicable without modification, and the + given in Table I1 was computed from the individual air and inlet temperature, the actual inlet temperature minus the effecfuel rates. T h e tabulated air rates were calculated from these equivalence ratios and the known fuel rates; the combined 20 accuracy of oxygen meter and fuel rate RESSURE OISSOCIATION was believed t o be greater than that of the orifice. However, in all cases the 0 air rate so calculated agreed with the * a air rate indicated by the orifice meter W I O w z within a few per cent. The static pres20 8 sure, P, was t h a t obtained b y the ? pressure tap shown in Figure 1. SO6 The inlet mixture temperature, TO, was indicated b y thermocouples lying 04 inside the inlet tubes and with junctions a t the points where the inlet tubes enter the insulating shell. It should be noted 08 I 40 60 N / V P " - G R d M O L E S 2iR.&o,o)'?~,i,R) biMl8j O ' that there is no error due to preheating Figure 5. Blowout data correlation for 1.8-order reaction (n = 1.8,s = 0.8), of the inlet mivture within the inlet corrected to 400' K. inlet temperature tubes and ball (if the latter are within Equivalence Ratio. + the reaction volume). This heat is 0.4 0.5 0.6 0.7 0.8 0.9 1.0 1.0 1.1 1.2 1.5 2.0 simply recycled through the reaction 9.13 1.57 35.6 26.8 8.92 19.3 32.6 43.7 4,8.0 48.0 NIVP1.S 0.539 2.82 zone-the apparent loss in flame tem0.75 0.76 0.77 0.78 0.81 t 0.94 0.93 0.87 0.83 0.92 0.91 0.89 2010 2035 1995 1940 1890 1715 1470 T 1315 1490 1655 1800 1915 perature recovered as a n equivalent in-

1638

tive temperature lowering due to heat losses) t o 400" K. will be described in detail. -4s expected, the blowouts could be correlated by plotting equivalence ratio against NIVP". The pressure exponent, n, appeared to have a value between about 1.5 and 2.0. Six separate large scale plots were thus made with values of n in 0.1 increments in this range. The data were examined for absolute pressure trends in the direction of deviation, I n addition, the best smooth curve was drawn for each plot, and the equivalence ratio root mean 100

# I

80

0

1I

AIR RATE (G;MOLS/SEC)

PRESSURE LATM. ABS)

RAOIUS

It / 3

30°

Figure 6 .

e

I/ 3

a modified radius cubed abscissa. This abscissa was chosen to show directly the fraction of reaction volume in which the indicated oxygen efficiencies were pertinent. At least for this time-average type of measurement, the composition is uniform (within experimental error) through more than 90% of the reaction volume. Tests similar to these were made in the 6-inch reactor by moving the probe in 0.2-inch increments. Similar results were obtained, again showing very good time-average homogeneity. I n another series of runs, the probe was fixed in position in the reactor a t constant air mass flow and inlet temperature. For three pressure levels (1, 2/3, and atmosphere absolute), combustion gas compositions were studied over a range of lean equivalence ratios. The gas samples were analyzed by a combination of oxygen meter (for oxygen), Orsat (for oxygen, carbon dioxide, and carbon monoxide), and mass spectrometer (chiefly for hydrogen but determining other components as well, except carbon monoxide). The results for oxygen, carbon dioxide, carbon monoxide, and hydrogen are presented in Figure 7. As pressure decreases for a given equivalence ratio, oxygen

cop

-

-

P=i 2/3 1/3 -

C O D

v

H o O

*>

I

v

-

90'

1

I

I

0.25

0.50

0.75

C(r - r,N J)/ ( b- r, NJ)I3

Combustion efficiency traverses in a 3-inch reactor

square deviations for each curve were computed. At a pressure exponent, n, of 1.8, no consistent trend of deviation with pressure could be detected. Furthermore, this exponent gave the smallest root-mean-square deviation (3.4% in equivalence ratio compared, for example, to 4.6% for exponents of 1.5 and 2.0). A similar difference in root-mean-square deviation was observed in the restricted range of 4 between about 0.6 and 0.8 where the largest variations in P for a given + were obtained. These computations were performed only for the 53 lean blowout points. Rich blowouts were not considered. The data have thus been plotted in Figure 5 as + versus N/VPl.8. The curve shape is typical in t h a t i t shows a maximum near stoichiometric and is roughly symmetrical about the stoichiometric axis. PROBE WITHIN REACTOR DETERMINES COMBUSTION GAS COMPO SITION S

,

Vol. 47, No. 8

INDUSTRIAL AND ENGINEERING CHEMISTRY

I n addition t o runs t o obtain blowout data, several runs were made a t steady state during which the gas sampling probe was used to study composition of the burning gases. I n one series of such runs, traverses were made along radii of the 3-inch reactor, designated reactor 3b, by moving the probe in 0.1-inch increments from the injector ball surface to the insulating wall. These traverses were made under two sets of conditions and on two different radii in the sphere-one a t about 30" to the inlet tube axis and one a t 90". The object was to assess the homogeneity of composition within the reaction volume. This was done by measuring the oxygen combustion efficiency of the gas samples. The oxygen content of the direct sample was compared with the oxygen content of the sample after completing combustion in the furnace used for measuring equivalence ratio. Some results are given in Figure 6, where the local oxygen consumption efficiency (the oxygen consumed divided by the oxygen i t would be possible to consume if combustion were complete) is plotted against

.4

0.6

0.8

EQUIVALENCE

Figure 7 .

IO

I

RATIO

Gas compositions in a 3-inch reactor, N = 0.3

content increases-Le., combustion efficiency decreases-as do carbon monoxide and hydrogen. Simultaneously, carbon dioxide decreases. Of particular interest are the large amounts of carbon monoxide formed and the relatively small (compared to the carbon monoxide) amounts of hydrogen. The water gas equilibrium ratios calculated from these data (using water fractions computed by material balances) correspond to equilibrium temperatures far higher than could be achieved under the conditions of operation. This implies, of course, that the water gas equilibrium reaction is too slow to have been achieved in the reactor. Average residence times in the reactor for the blowout tests ranged from about 0.015 to 0.00035 second. For the range of equivalence ratios of Figure 7, the mass spectrometer analysis showed only traces (below the limits of instrument accuracy) of other combustible components. Within experimental error, the carbon monoxide and hydrogen present can consume all of the oxygen present which would be consumed if combustion were complete. A few scattered runs of this type were also made for rich mixtures. The data were inconclusive but did show that for fairly rich mixtures hydrocarbons began to appear in the sample. Thus, for example, for the same mass flow as for Figure 7 and a t 2/8

INDUSTRIAL AND ENGINEERING CHEMISTRY

August 1955

atmosphere, the data of Table 111 were obtained. The mass spectrometer was not calibrated for acetylene, and it is not reported. I t s concentration is probably small, judging from the good material balances obtained when equivalence ratio and water concentration were computed in the several ways shown in Table 111. Two other series of runs were made in which oxygen consumption efficiencies were measured as described for Figure 6. I n all these runs the equivalence ratio was held constant a t 0.70 with inlet temperatures about 400" K. The fuel was iso-octane hurned in 3-inch reactors-thus giving approxiniately constant reaction volume for all runs. However, for one series of runs, the mass flow was held constant a t 0.3 mole per second while pressure was varied. I n the other series, the pressure was held constant a t 0.5 atmosphere while mass flow was varied. The results are shown in Figure 8. For a given value of the correlating group N/VP1.a (corrected to 400" K. inlet temperature), the oxygen efficiency is a constant regardless of the individual values of N and P. There is some departure as the curves approach blowout, but this is a reasonable amount of deviation for this type data. DATA INTERPRETED BY REACTION RATE THEORY

I t has been shown previously (9) that the phenomenon of blowout is t o be expected for an exothermic flow reaction in a perfectly stirred reactor. This conclusion follows from considerations of elementary chemical kinetics of the over-all reaction. Furthermore, the over-all kinetic constants can be obtained from the blowout data. Therefore, it is useful to interpret the present data hy a reaction rate approach.

Table 111.

1639 I \EQ.

I

40

LL

0.9

10,.

v,

z

0.N=0.300t.002, V= 0.243 O-P=0.500+.005. Vz0.239 OR 0.227 BLOWOUT' 0=0700+.008

-

f X 0

I 4

2

l

l 20

6 8 1 0 N/VP1.8

40

C o m b u s t i o n efficiency dependence on N/VP'.* in a 3-inch reactor

Figure 8.

Conatants for Equation 4s curve as shown in Figure 5

A quantitative description of this reaction-rate-induced blowout is straightforward. Let N moles of air enter the reactor per second containing m moles of inerts per mole of oxygen; define oxygen consumption efficiency between inlet and exit as E (again, the oxygen consumed divided by the oxygen which would be consumed were combustion complete). Then the rate a t which oxygen disappears between inlet and exit streams is N d e / ( 1 m ) for 5, < 1 and N e / ( 1 m ) for 5, > 1. This rate can be equated t.0 the rate a t which oxygen is consumed within the reactor due to chemical reaction. Thus

+

+

Spectrometer Analysis of C o m b u s t i o n Gases

+64 (from fuel and air rates) (from C C/N / N ratio)

HzO Hi0 (from N/O ratio), vol./ 100 vol. dry gas HsO (from C/H ratio), vol./ 100 vol. dry gas

1.70 1.71 13.1 12.5 Dry Gas Compn., Vol.

%

Hydrogen Oxygen Carbon monoxide Carbon dioxide Methane Ethylene Ethane Nitrogen (balance)

8.5 0.2 14.2 5.0

0.7 0.3

0.1 71 .O

1oo.o

To account for blowouts qualitatively, consider combustion to be a chemical reaction between fuel and oxygen whose rate depends on temperature and fuel and oxygen concentrations. For combustion approaching completion-that is, after a relatively long combustion time-the temperature is near a maximum, thus favoring a rapid reaction. However, the reactants have largely been consumed and their concentrations are small, thus favoring a slow reaction. If less time is allowed, the reaction will not have proceeded so far toward completion. The temperature will be lower but, more important, the reactant concentrations will be higher; the net result will be a faster over-all rate. As reaction time is still further reduced, the increase of reactant concentrations continues to more than offset temperature decrease and the over-all reaction rate still increases. However, a point is eventually reached such that an additional reduction in reaction time causes a temperature lowering with greater effect than the rorresponding concentration increase. This point represents the shortest possible reaction time and the fastest reaction rate. Further reduction in reaction time makes continued reaction impossible; the reaction thus stops and in effect, the flame has blown out.

according to collision theory ( 6 ) , using an Arrhenius exponential for the temperature effect and incorporating a term presumably to make the rate constant, k , temperature independent. The over-all reaction order is n, and zed,. and zoZ are the mole fractions in the combustion gases of the combustible component and oxygen, respectively. The composition data of Figure 7 suggest that the unburned material is not original fuel but rather chiefly carbon monoxide. Therefore, if the original fuel is C I H I ~but the remaining combustible is assumed to be only carbon monoxide (hydrogen is neglected), the reaction, for 4 moles of CsH18,is

dF

(1

-

GI/)

O2+ 25 (5, Con

- cy) GO + (25ey

-

176)

25 + 9+ HzO + 3.76 X 3 Nt

(2)

from which =

4.76

%d - ey) 1.36 5, -

+

1

E?/

202

= 4.76

-

ey

+ 1.36 4 - e~

(3)

By substituting Equation 3 in Equation 1and rearranging, letting m = 3.76 for air,

N 4.76k e - E / R T [25,(1 - €)If [I - +€In-' VP" = R* T n - 1 / 2 + ~ [ 4 . 7 6 5, (1.36 - e ) F N = 4.76k e--E / R T [2(+ - €)If [l - €ln-f Rich: __ VPn R" Tn-112 e[4.76 1.366 - e]"

Lean:

~

~

~

+

(4a)

+

(4b)

For rich mixtures, it seems a poor approximation to assume that carbon monoxide is the only combustible reactant. The illustrative data for 4 = 1.70, which show a larg,e quantity of hydrogen and traces of hydrocarbons, mggest this. Because the lack of

INDUSTRIAL AND ENGINEERING CHEMISTRY

1640

Vol. 47, No. 8

tional oxygen consumption, E . T was computed by a n enthalpy balance-summing the enthalpies of the reaction products (defined by Equation 2 or 5, as appropriate) and comparing this enthalpy with the heat of combustion of the fuel burned (correctFractions E of the original oxygen and original CsH18 are coning for carbon monoxide and hydrogen formation where necessumed to give new chemical species. sary). Values of T were tried until a balance was obtained or the T h e consumed material forms a mixture of carbon monoxide, carbon dioxide, hydrogen, and water in water gas equilibrium at correct value bracketed, allowing interpolation. These temperathe reaction temperature, 2'. tures were computed for homogeneous inlet mixture temperaFractions (1 - E ) of the original oxygen and CsHls remain and tures, TO,of 300°, 400°, and 500" K., for CsHla and dry air a t are regarded as the only reactants. atmospheric pressure. Component enthalpies were obtained from For lean mixtures the fuel consumed forms only carbon dioxide Huff and Gordon ( 7 ) and Rorssini and coworkers ( 1 5 ) . The comand water. Therefore, the reaction, for + moles of C B H I ~ is, putations above do not take dissociation into account and a p proximations for this effect were estimated from the work of Mulready ( I d ) . Therefore, any temperature exceeding 2100" to 2200" K. may be questionable. Fortunately, in this work temperatures were low enough and pressures high enough to make dissociation of importance over only a small fraction of the data 25 9+e (Hz +HzO) 3.76 X -5- Nz (5) range. Results of the computations are given in Table TV for the range of equivalence ratios. Flame temperatures calculated by where the individual quantities of carbon monoxide, carhon diassuming the reaction of Equation 5 and the reaction of Equation 2 oxide, hydrogen, and water for rich mixtures can be obtained by are given. For lean mixtures with e < 1.0, Equation 5 results in a carbon, hydrogen, and oxygen balances and noting that X C O X H ~ O / higher reaction temperature for a given oxygen consumption, e, X C O ~ X H= ~ K (the water gas composition equilibrium constant at than Equation 2. This occurs because the consumption of a T). From Equation 5, given quantity of oxygen in burning fuel to carbon dioxide rather than carbon monoxide liberates much more heat. 0 08(b(1 - E ) For lean mixtures, Equation 2 sets a lower limit of 0.68 for e. ZC8Hn = 4.76 0.08(b ~(1.28(b - y) A t this value, all of the carbon in the system appears as carbon 1 - Ey 'On (6) monoxide and no carbon dioxide remains. This limiting value of 4.76 0.08 (b ~(1.28(b- y ) E is >0.68 for + > 1, a p is evident from Equation 2. This beSubstituting Equation 6 in Equation 1 gives the alternate equahavior is an inronvenience in extrapolating reaction rates t o tions with original fuel rather than carbon monoxide as the comlower values of efficiency; however, there is no such limitation for bustible reactant. Thus Equation 5. The over-all reaction order, n, is 1.8 as previously shown by empiriral correlation of the data. T h e individual combustible and oxygen eyponents, f and ( n - f ) , cannot be determined (7a) directlg. There is, however, a clue to their approximate values. N 4.76k e - E / R T (O.O8+)f By assuming a trial value of E-say, 40,000 calories per moleRich: = -Rn7 close to the evpected final value, the influence of .f on the shape 1--s of the theoretical hlowout curve can be ohserved. The blowout (7b) [4.76 - E 0.08+(1 16e) curve is computed by selecting values of E for any particular 4. I n order to apply Equation 4 or 7 to the blowout data, e and T The correspoirding T from Table I V is used and these values are substituted i r i Equation8 4a and 4b. It is desirable to use both must be related and the constants, n,f, k , and E, must be known. T is the adiabatic reaction temperature corresponding to the fracparts of Equation 4 (rather than a combination of Equations 4 and 7 ) in order to have a consistent assumed reaction mechanism near d = 1. for both lean and rich mixtures. By Table IV. Flame Temperatures of Iso-Octane and Dry Air at Atmospheric Pressure carrying /c along as an unknown Flame Temperature, K. constant, a particular value K. E-= 0.8 a t To, e K. c 0.9 a t To, e = 0.6 a t To, K. c 0.7 a t To, K. e = 1.0 a t To, a K. ___of E will be foxind (ordinarily 300 400 500- ~ 0 0 4 6 ~ 0 03 - 0 0 ~ 4 0 0 ~ 0 0 ~ 300 ~ 0 400 ~ 45000 ~ 9 >0.7) such that NIVPl.8 is a Gas Composition Given by Equation 5 maximum. This is repeated 950 1039 862 937 1025 1112 1011 1098 11x4 I 084 1170 1256 875 965 0.3 786 for other equivalence ratios t o 907 1014 1101 1024 1109 1195 1119 1203 1288 1213 1296 1381 1307 1389 1473 0.4 1059 1144 1229 1175 1268 1342 1290 1372 1455 1403 1484 1566 1515 Id95 1677 0.5 give a blowo>it curve of + I184 1266 1349 1318 1400 1182 1451 1531 1612 1.582 I661 1741 1712 1791 1871 0.6 1302 1382 1464 1453 I533 1613 1603 I681 1761 1762 I829 1909 1899 1976 2055 0.7 versus (N/VP1.8)msx.. This 1414 1403 1573 1582 I659 1739 1748 1825 1904 I913 1989 2067 2077 2155 2210 0.8 curve can be plotted for 1320 1.598 1677 1704 1780 1858 1886 1962 2039 2067 2142 2205 2225 2275 2330 0.9 1.0 1622 1698 1776 1821 1896 1973 2017 2092 2165 2200 2250 2295 2315 2355 2390 various valueii of .f in Equation 1.1 1564 1641 1719 1754 1830 1907 1941 2017 2094 2130 2200 2260 2285 2345 2395 1.510 1.586 1664 I690 I766 1343 1870 1945 2022 2048 21 24 2200 2225 2295 2355 1.2 4, and i t is found that f affectb 1359 148li 1515 1515 1592 1670 1670 1746 1825 1826 1902 1980 1980 2058 2135 1.5 the eqwvalence ratio at the 2.0 I139 1218 1300 1260 1339 1419 1381 1460 1540 1501 1;781 1662 If322 1703 1783 944 102.5 1108 1036 1118 I201 1127 1211 1294 1219 1302 1886 1310 1344 1478 2.5 maximum value of 894 981 1069 935 828' 915 1002 959 1047 1135 1024 1113 1202 3.0 76 1 847 (N/VP1.8)max..For f > 1.0, Gas Composition Given by Equation 2 the maximum occurs at 6 > 9.52 970 1057 0.3 775 863 882 987 1073 1159 1084 1170 1256 1.1; a n d f o r f < 0.7, a t 4 < 1.0. 999 1087 1046 1133 1218 1178 1263 1348 1307 1389 1473 913 0.4 0.5 1042 1127 1214 1203 1288 1371 1360 1444 1527 1515 1595 1677 Because the data maximize at 0.6 1173 1219 1335 1351 1435 1517 1533 I616 1898 1712 1791 1871 0.7 1279 1366 1450 1492 1575 1656 1699 1779 1860 1899 1976 2055 + between 1.0 and 1.1, f is 0.8 1395 1478 1560 I627 1709 1789 1855 1933 2013 2077 2155 2210 in the 0.7 t o 1.0 range. An 1502 1584 1665 1767 1837 1915 2004 2079 2155 2225 2275 2330 0.9 1606 1686 1767 1879 1958 2036 2145 2215 2270 2315 23.55 2390 1.0 arbitrary choice within this 1783 1864 1943 20.53 2129 2195 228.3 2346 2400 1.1 1957 2035 2120 2215 2285 2350 range 1.2 ... ... ... - was made to fit the lean data most closely.

data permits no choice of specific reactants i t will be assumed, as previously ( 8 ) , that the original fuel is the combustible. For rich mixtures the following assumpttions are then made:

+

+

+

+

v&

+

-Tzv~

+

O

+

1"

-

...

O

... ...

... ...

-

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

August 1955

1641

heat losses affect the blowout rates in reducing total enthalpy; the enthalpy loss is accounted for by an equivalent decrease in inlet temperature. Equation 4a was also used, with these same constants, to predict the dot-dash curve of Figure 8. Values of from about 1.0 to 0.9 were simply substituted with the corresponding TI to obtain the curve. Although the data points fall below most of the curve, the efficiencies a t blowout, are in reasonable agreement. I n any case, the correspondence of e for given NIVP1.8, regardless of individual values of N and P, lends credence to the reaction rate concept. The exact level of predicted efficiency depends on the reactants assumed. For example, if oxygen were reacting with CsHIR rather than cakbon monoxide-thus requiring the use of Equation 7a rather than Equation 4a-agreement between observed and predicted efficiencies n-ould be excellent. 4001,

Figure 9.

OK.

Calculated blowout reaction rates for various effective inlet temperatures

T h e maximum value of NIVPl.8 of Figure 5 (dissociation corrected) is about 48. Equation 4a predicts that the combustion efficiency is 0.83. This rate of reaction corresponds t o a space heat release rate of about 3.0 X 1 0 8 B.t.u. per (hour) (cubic foot) (atm.l.8). Such a rate is an order of magnitude greater than the highest published esperimental value (11) for a n appreciable volume known to the authors. Lack of other data makes direct comparison of these results impossible. This rate will not necessarily correspond to the previously discussed highest local rate appearing in laminar flames. I n a laminar flame, it is certainly possible that a happy combination of heat and mass diffusions may give a local combination of temperature and reactant concentrations that results in a rate higher than observed here. I n previous theoretical treatments of homogeneous combustion ( I , 3, 9,14, 16), it was assumed that the over-all reaction order was 2.0 with equal combustible and oxygen exponenta. If such assumptions are made in this case, the data can be plotted as shown in Figure 10. However, there is a pressure trend in the lean data points-lower pressure points fall below and to the right. The solid curve best fits the lean data with Equation 4a using E = 42,000 and the rich data with Equation 7b and the same E. The most significant deviations again occur for very rich mixtures where the data are most questionable. If the further usual assumption is made that the combustible material is original fuel (rather than carbon monoxide), Equation 7a is used for the lean data. Thus the dotted curve of Figure 10 was calculated with a best-fitting E of 40,000 and this same E in Equation 7b for rich mixtures. The fit here is poorer, particularly near + = 1. For both curves, k was chosen for the rich mixture equation so as to make N I V P a t + = 1 coincide for both lean and rich branches. All of the data of Figure 10 were corrected to a 400" K. inlet temperature reference by using Figure 9. Figure 9 is negligibly in error for small variations in E,n, or f. Although use of N / V P * withf = 1 is a slightly more convenient approximation, it predicts somewhat optimistic heat releases (at pressures higher than the pressures used). Thus, the maximum observed N/VP2 is 74 (corrected for low pressure dissociation), and the maximum values predicted by the two curves are 68.5 and

Relative to rates at TOE= 400' K. Equations a n d constants as shown i n Figure 5

By using f = 0.8 in Equation 4a, the lean data were best correlated with an E of 42,000 calories per mole. The same .f and E were used in Equation 7b to fit the rich data of Figure 5. Computed values of N/VPl.*, e (within 0.01), and corresponding T a r e tabulated in the caption of Figure 5. Differences in k result from the different assumptions as t o the nature of the combustible and the combustion gas composition for the lean and rich mixtures. The semitheoretical curves computed from these constants are shown in Figure 5. The data points near + = 1 in Figure 5 (and Figure 10) were obtained a t relatively high temperature and low pressure. Dissociation corrections t o reaction temperature a t these low pressures are greater than those incorporated in the atmospheric pressure temperatures of Table IV. Thus the temperatures are lower and reaction rates correspondingly lower. The semitheoretical curve reflects the data corrected to atmospheric pressure reaction temperature, thus eliminating this extra dissociation effect a t low pressure. This correction did not exceed 8% In N/VP1.8 a t any point. Reduction in flame temperature a t low pressures is illustrated by the following Bureau of Mines calculations ( 9 ) for CloHzofuel a t 2'0 = 400" K. and e = 1.0: Flame Temperatures,

P 1.000 0 398 0.158 0 100

.

0.667 1935 1933 1929 1926

0.818 2177 2163 2146 2136

K.. at Equivalence Ratio, 6 1.222 1.500 2119 2360 2313

1.000 2317 2282 2263

2297 2275 2263

2115 2110 2107

Once the values of the kinetic constants in Equations 4a and 7b have been established, the blowout values of N/VP1.8 and 4 can be recomputed for various inlet temperatures. An inlet temperature other than 400' K. implies only a different T for a given e in the equations. The blowout NIVP1.8 for a given + can then be compared, for various to^, to the blowout value €or the same + a t TOE = 409' K. This has been done in Figure 9, using the kinetic constants of the curves of Figure 5. Figure 9 was used in Table I1 to correct the blowout rates to 400' K. for effective inlet temperatures other than 400" K. It is clear from this figure how

OBSERVED SPACE HEAT RELEASE RATES ARE VERY HIGH

1642

INDUSTRIAL AND ENGINEERING CHEMISTRY

Vol. 47. No. 8

tions. Numerical constants are thus availrtble for the theoretical ram-jet combustor analyses of Avery and H a r t ( 1 ) and Rosen and Hart ( 1 4 ) . Where combustion is the usual complex of mixing and kinetics, information on the latter process can be obtained from the equations. Applications to flame spreading from an ignition source in a stream and t o flame stabilization by bluff bodies or cans are possible. Where mixing patterns are uncertain, these patterns and mixing rates a t least can be hypothesized to give idealized systems. Such systems, combining optimum mixing and reaction rates, present specific goals which rational design methods for practical Figure 10. Blowout data correlation for second-order reaction ( n = 2.0, f = l.O), Corrected to 400' K. inlet temperature equipment can try to approach. The validity of these conclusions deEquivalence Ratio, $I pends chiefly on how well the experi0.4 0.5 0.6 '0.7 0.8 0.9 1.0 1.0 1.1 1.2 1.5 2.0 mental spherical reactor simulates the Equations 4a and 7b, E = 42.000 ideal perfectly stirred system. The timeaverage traverses of Figure 6 show no 0.81 38.0 14.1 68.5 51.3 41.5 58.5 68.5 10.4 23.4 0.575 3.30 Nl VP' e 0.92 0.91 0.90 0.89 0.87 0.85 0.81 0.74 0.75 0.75 0.76 0.78 gross nonhomogeneity, but small scale T 1290 1460 1615 1760 1870 1960 1985 1975 1925 1855 1685 1435 rapid heterogeneities are not disproved. Equations 7a and 7b, E = 40,000 However, the gas s:tmples indicate that 18.3 2.62 47.1 62.5 82.0 82.0 22.1 40.0 60.8 9.92 residual carbon monoxide and hydrogen ,V/VPp 8.624 3.18 0.76 0.77 0.74 0.74 0.73 0.73 0.77 0.83 0.81 0.85 0.88 0.87 account for all unburned combustible. 1905 1835 1685 1125 1953 1840 1905 1955 T 1280 1450 1595 1525 This implies that the pas sample drawn into the probe is nearly homogeneous in composition. If the gases consisted of alternating eddies of 82.0. These numbers correspond to heat releases of about 4.4, burned and unburned material, one would expect to find at least 4.1, and 4.9 X 10s B.t.u. per (hour) (cubic foot) (atm.2), respecsome combustible material as original fuel or as cracked hydrotively. carbons from the fuel. I n general, the available evidence does not prove that the reactor is homogeneous, but no evidence for CONCLUSIONS nonhomogeneities has been discovered. The reactor itself can be Lean mixture blowout data from insulated spherical reactors used as a tool for study of other homogeneous reactions-either suggest that the high temperature combustion of hydrocarbons highly exothermic reactions by using the blow-out technique, or in air is an over-all reaction of 1.8 order with activation energy nonexot,hermic reactions by measurement of temperature, time, of 42,000 calories per mole. No hydrocarbons are found in the and composition in the usual ways for a flon system. combustion gases of incompletely burned mixtures. However, Therc m:t.v be some question as to the efficienw of the gas large amounts of carbon monoxide are found and the carbon sampling probe. If the gas temperature is not reduced instantly monoxide present, together with the small quantities of hydrogen (compared to the reaction rate), reaction continues in the probe observed, can. account for all the combustible material that has and the observed compositions nil1 correspond to some reaction not yet burned. Therefore, the slow and controlling step in time and temperature other than those in the reactor. Rough kinetics-limited high temperature rombustion seems to be the heat transfer calculations predict that the quenching rate is rapid oxidation of carbon monoxide. These considerations apply to enough to stop the reaction. Honever, an ultimate test of the lean mixtures. Although the rate equations appear to describe probe is to compare the results obtained with those from a probing adequately the blowouts of rich mixtures, the nature of the reacdevice, as described by Prescott and coworkers ( I Y ) , for example, tants has not yet been established. It must be recognized that in which there is no doubt ahout the quenching action. This type combustion of a hydrocarbon is a complex process of sucressive test is planned. and competing reactions; the intent of the equations developed Comparison of data from 3- and 6-inch spheres shows no surin this article is only to describe the net over-all rates. face effert of the hot walls for this 2 to 1 change in surface-toThe limiting space heat, release is about 3 X 108 B.t.u. per volume ratio. Tests on spherical reactors of the type previously (hour)(cubic foot)(atm.l.*) for stoichiometric mixtures. This illustrated (8) similarly showed no surface effects for reactors limiting rate is reduced for nonstoichiometric mixtures. At C$ = ranging from 1.5 to 6 inches in diameter and using various types 0.7, for example, i t is about 9 X lo7 and, at C$ = 0.5, about 1 X of firebrick walls. 10'. The work described is being continued in an effort to obtain Heat losses from the combustion zone reduce these limiting more rich mixture data, to elucidate the reaction mechanism b y space heat releases b y reducing the reaction temperature and thus detailed gas composition studies, and to investigate other varithe reaction rate. For example, a heat loss equal t o 5 % of the ables-e.g., inlet temperature and type of fuel. maximum adiabatic heat liberated reduces the maximum heat release rate by more than 30%, while a 20y0 heat loss reduces ACKNOWLEDGMENT the rate about 85y0. These figures do not vary greatly with This work was performed for the Bureau of Ordnance, Departequivalence ratio. The heat loss effect may thus be of great ment of Navy, in cooperation with the Applied Physics Laboraimportance in practical combufition devices. I n general, these tory, The Johns Hopkins University. The authors are grateful to losses have usually not been considered in design or have not been the Bureau of Ordnance and to the Esso Research and Engineering subject to much control by the designer. Co. for permission to publish these reports. The rate equations used in this discussion have other applica-

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

August 1955

NOMENCLATURE

E

1643

DeZubay, E. A., Paper 54-SA-27, presented at Semiannual Meeting, Am. SOC.Mech. Engrs., Pittsburgh, Pa., June 1954. (4) Friedman, R., and Burke, E., J. Chem. Phys., 2 2 , 8 2 4 (1954). (5) Fristrom, R. hf., Avery, W. H., Prescott, R., and Mattuck, A., (3)

activation energy, gram-cal./gram-mole base of natural logarithms e f combustible concentration exponent K = water gas equilibrium composition ratio a t T = rate constant, (liters)*-l/( O K.)1’2(.gram-moles)”-’ (see.) k m = moles inert gas/mole of oxygen in air Iv = air rate, gram-moles/sec. n = over-all reaction order P = pressure, atm. abs. R = gas constant, 1.987 gram-cal./(gram-mole) ( ” K.) or 0.08206 (liter) (atm.)/(gram-mole) (’ K.) r, f i n , . , 7 0 = radius at’probe ti6, injector sphere, and reactor wall, respectively, inoches T = reaction temperature, K. To = actual inlet mixture temperature, O K. TO^ = effective inlet mixture temperature (TOcorrected for heat loss), K. V = reaction volume, liters = mole fraction in combustion gases 5 y =&for+l = fractional oxygen consumption efficiency e + = equivalence ratio = = =

O

LITERATURE CITED (1)

Avery, W. H., and Hart, R. W., IND.ENG.CHEM.,45, 1634

(2)

Brinkley, 5. R., Jr., Smith, R. W., Jr., and Sapsara, 0..U. S. Bur. Mines, Rept. PX3-107/3, 1952.

(1953).

Ibid., p. 106.

Glasstone, S., Laidler, K. J., and Eyring, H., “Theory of Rate Processes,” p. 5, McGraw-Hill, New York, 1941. (7) Huff, V. N., and Gordon, S., Katl. Advisory Comm. Aeronaut., Tech. Note 2161, 1950. (8) Longwell, J. P., in “Selected Combustion Problems, Fundamentals and Aeronautical Applications,” AGARD NATO Combustion Colloauium. Deo. 7-11. 1953. -DD. - 508-‘10. Butterworth, London, i 9 5 4 . (9) Longwell, J. P., Frost, E. E., and Weiss, M. A., IND. ENO. CHEM.,45, 1629 (1953). (10) Longwell, J. P., and Weiss, M. A., Ibid., p. 667. (11) Mullen, J. W., 11, Fenn, J. B., Garmon, R. C., Ibid., 43, 195 (6)

(1951).

Mulready, R. C., United Aircraft Corp., East Hartford, Conn., Meteor Rept. UAC-9, 1947. (13) Prescott, R., Hudson, R. L., Foner, S . N., and Avery, W. H.,

(12)

J . Chem. Phys., 22, 145 (1954).

Rosen, P., and Hart, R. W., J . Aeronaut. Sci., 20, 549 (1953). Rossini, F. D., et al., “Selected Values of Physical and Thermodynamic Properties of Hydrocarbons and Related Compounds,” Carnegie Press, Pittsburgh, 1953. (16) Spalding, D. B., Aircraft Eng.,25, 264 (1953). (17) Wohlenberg, W. J., Paper 53-A-103 presented at Annual Meeting, Am. SOC.Mech. Emgrs., New York, N. Y., December (14) (15)

1953. RECEIVED for

review October 14, 1954.

ACCEPTED

February 9, 1955

Drying Oil Oxidation Mechanism, Film Formation, and Degradation AND A. L. ALEXANDER Naval Research Laboratory, Washington 25, D . C .

S. B. CRECELIUS, R. E. ICAGARISE,

A

MOKG the more basic phenomena on which present

protective coating technology is based is the oxidation of drying oils. One would expect a process so basic to the coatings industry to be the subject of intensive research, a8 in fact it has been for the past several decades. Nevertheless, as Powers (18) has stated in a recent review, the oxidation mechanism of unYaturated oils is still obscure due to the complexity of the oxidation products and the uncertainty of their structure. Progress toward an understanding of the oxidation mechanism has been hampered in the past by the inadequacy of the methods of direct chemical analysis. With recent developments of physical methods, however, particularly the techniques of ultraviolet and infrared spectroscopy, the scope of the analytical attack has been broadened considerably; it is by these means, coupled with the information already obtained on chemical structure, that the hope of obtaining more complete information of the autoxidation mechanism seems closer at hand.

radical chain mechanism which results in the formation of a hydroperoxide group (-OOH) a t the a-methylenic group adjacent to t,he double bond carbon atom. He suggests that the oxygen adds first a t the double bond to form a biradical

H H H

-C-C==C H

+

H H -CHzC=C-

+

bH + I

0 2

-

b

H -0

I

I

HC or I>=O

HC

I

has been chalIenged by the hydroperoxide theory of Farmer ( 8 , 9) who postulates that the autoxidation of drving oils proceeds by a

CHI

8-C H H

0 +-

H

H

I n recent years the older theory ( 7 , 14, 27) that the perosidation mechanism ir due to the formation of a peroxide ring

+

CHz6-GH H H -&C=C-

PREVIOUS WORK

H -0

0 2

This biradical then combines with another unsaturated group t o yield two radicals

*

AH

O* 0

n -

I€

+

* -CH2-C-C--

0

H H

These radicals continue the chain reaction and lead t o the formation of a-methylenic hydroperoxides. As Powers (18) has commented, this explanation does not account for the uniform drop in iodine value during oxidation as found by previous workers ( 2 , 6 ) except for the possible formation of conjugated systems which would give a reduced iodine value. However, it is difficult to explain by this means alone the extensive and uniform drop in iodine value which has been observed.