596
Ind. Eng. Chem. Prod. Res. Dev.
1988,25, 596-603
Kinetic Modei for High-Temperature Oxidation of Lubricants Satlsh K. Naldu,' Elmer E. Klaus, and J. Larry Duda Department of Chemical Engineerlng, The Pennsylvania State Universiv, University Park, Pennsylvania 16802
A kinetic model has been proposed for high-temperature oxidative degradation of lubricants under boundary lubrication conditions. This model incorporates primary oxidation reactions as well as subsequent condensation polymerization reactions that result in viscosity increase and sludge formation. Evaporation of the oil and its volatile primary oxidation products has also been considered. The model has been examined by conducting oxidation and evaporation studies on a commonly used neopentyl poly01 ester, trimethylolpropane triheptanoate, in the Penn State microoxidation test. This thin-film laboratory test simulates lubricant behavior in bearings under elastohydrodynamic and boundary lubrication conditions. The kinetic rate constants for the primary oxidation and the subsequent polymerization reactions have been determined. The rates of reaction appear to be first order. This information coupled with evaporation rates has been used to examine the ability of the model to correctly predict the oxidative behavior of the lubricant.
Introduction The principal process involved in lubricant degradation is oxidation. It appears that a complex oxidation process involving the bearing surface as well as the lubricant and dissolved oxygen is a key process in boundary lubrication. An analysis of most lubrication systems suggests that in the bulk condition lubricant temperatures are designed to be low enough so that the oxidation inhibitors in the lubricant are effective. In addition, lubricant additions or changes are generally frequent enough to maintain oxidation inhibitor in the system at all times. However, in spark ignition, diesel engines, and aircraft gas turbine engines significant oxidative degradation of the lubricant has been observed. One explanation for lubricant oxidation despite the presence of inhibitor is the presence of hot spots in the lubrication system where the temperature exceeds the capability of the inhibitor to prevent oxidation. In the engines mentioned, these hot spots occur in bearing areas and on metal surfaces where heat from the combustion system and thin films of lubricant coexist. These latter areas are illustrated by the piston ring-cylinder wall area of the internal combustion engines and the wall of the rear turbine-bearing compartment in the gas turbine. If the oil from the bearings and the hot surfaces circulates freely back to the reservoir and, therefore, spends most of the operating time in the reservoir, the oil in the bearings and on the hot surfaces can be considered to be saturated with oxygen from air solubility. The lubricant in the bearing has a short residence time, and the thin film of the lubricant on the hot nonbearing surfaces is generally in contact with an environment containing oxygen. In both cases, oxidation can take place with minimum or no limitations from oxygen diffusion into the liquid system. The oxidative reactions in these hot thin film zones would be expected to follow chemical reaction kinetics uncomplicated by diffusion effects. The similarity of the oxidation products and the apparent first-order reaction rates measured in thin-film laboratory systems with organic esters, synthetic hydrocarbons, mineral oils, and phosphate esters (Lockwood and Klaus, 1981; Bailey, 1983; Cho and Klaus, 1981, 1983) suggests that the primary reaction is associated with the C-C chain and not the heteroatoms in these molecules. This study treats the oxidation of a polyolester, trimethylolpropane triheptanoate (TMPTH), as a simple reaction with no oxygen diffusion limitations. The basic data are used to model the primary oxidation reaction and subsequent reactions, thereby providing a relationship that can be used to predict the behavior of lubricants given the
environments in which they function.
Experimental Met hods The Perm State microoxidation test was used to conduct all oxidation and evaporation studies. The test apparatus and procedure have been described in detail by Cvitkovic and Klaus (1979). A schematic of the apparatus is shown in Figure 1. The system consists of a glass tube with a flat bottom and a glass top with gas inlet and exit tubes. To start an oxidation test, a metal cup is placed at the bottom of the reactor, and the entire unit is immersed in a constant-temperature bath. The system is equilibrated by circulating nitrogen at 20 mL/min for 30 min. Dry air is then circulated for 10 more minutes. A t this point, the oil sample is injected onto the metal cup, and the air flow is continued for the chosen test time. For evaporation tests, no air is circulated in the system. Instead, nitrogen flow is continued throughout the test period. The metal cup is designed so that the oil forms a thin film (0.050.5-mm thickness) on the metal surface. This thin film is necessary to eliminate oxygen diffusion limitations in the liquid film. Low-carbon-steel (LCS) metal cups were used in all tests conducted in this study. The metal plays an important role in oxidation-polymerization reactions. Some metals act as catalysts, whereas others are inhibitors (Lahijani et al., 1982; Ugwuzor, 1982; Klaus et al., 1985). The organometallic corrosion products soluble in the liquid lubricant can influence the reaction mechanisms. LCS has been determined to be a catalyst in promoting both the oxidation and the subsequent polymerization reactions (Lahijani et al., 1982; Ugwuzor, 1982; Klaus et al., 1985). At the termination of the test, the reactor is removed from the hot bath and rapidly cooled by placing it in a stream of cold air. The products are then diluted with tetrahydrofuran and analyzed by gel permeation chromatography (GPC). The GPC analysis provides information on the molecular weight/size distribution of the products. The GPC columns consisted of 8 ft of 60-A Poragel and 4 f t of 100-A Styragel packing. The columns were calibrated by determining the elution time for polystyrene molecules of known molecular weights. The relative product molecular weights are determined relative to the elution time for a polystyrene of equivalent molecular size. The GPC unit used was equipped with an ultraviolet detector and a differential refractometer. The refractive index trace was used to determine quantitatively the amount of original oil remaining and also products of comparatively higher and lower molecular weights. A
0 196-4321/86/1225-0596$01.50/0 62 1986 American Chemical Society
Ind. Eng. Chem. Prod. Res. Dev., Vol. 25,
OUTLET TUBE METAL HOLDER
24/40 GROUND 01.ASS
1' ' 1
1
'
095 t 0.35 mm
0.5-
METAL CATALYST TEST CUP (DETAIL)
Figure 1. Penn State microoxidation test apparatus.
--a
ORIGINAL OXIDIZED
SAME MOLECULAR WEIGHT HIGH MOLECULAR WEIGHT LOW MOLECULAR WEIGHT
I
1
los
IO*
I
I 0'
I\ I
400
I
200
I
No. 4, 1986 597
basestocks and additives that would minimize oxidation problems in the engine. It is generally agreed that the first step in oil oxidation involves a free-radical chain reaction that results in formation of hydroperoxides (Zuidema, 1952; Emanuel, 1965; Reich and Stivala, 1969; Jensen et al., 1979). At the temperatures used in this study, hydroperoxides are unstable and are rapidly converted to acids and ketones (Hamilton et al., 1980; Jensen et al., 1981). In ester oxidation, these primary oxidation products are of lower molecular weight (LMW) than the original ester (Czarnecki, 1971). Czarnecki (1971), in oxidation studies of trimethylolpropane triheptanoate (TMPTH), showed that the primary oxidation products included heptanoic acid and more than 10 individual compounds or isomers in the general boiling range of the diheptanoate of trimethylolpropane. These individual compounds or isomers were identified from a temperature-programmed gas chromatographic (TPGC) analysis. The large number of intermediate products appears to be the result of the molecular rearrangement of the trimethylolpropane after losing heptanoic acid. In this study (Czarnecki, 1971),only a small amount of these lower molecular weight products were formed, and all of the remaining products were of a molecular weight higher than that of TMPTH. These materials were too high in boiling point to be eluted in the TPGC. This behavior has been confirmed by GPC analysis in previous studies (Ali et al., 1979). From microoxidation tests in the current study using GPC analysis, it is clear that at higher conversion rates the low molecular weight (LMW) primary oxidation products polymerize to form higher molecular weight (HMW) secondary products. Spectroscopic studies have shown that these polymers contain unsaturated ketones (-C=CC 4 ) as functional repeating groups (Naidu et al., 1984). At longer test times these HMW products become insoluble and form sludge and varnish type deposits. A. Oxidation Model. Analysis of the complex reactions involved in lubricant oxidation can be simplified by assuming a simple first-order kinetic model such as
IO0
MOLECULAR WEIGHT
Figure 2. Gel permeation chromatogram (refractiveindex response) of unoxidized and oxidized ester.
typical refractive index trace is shown in Figure 2. Previous studies have shown that the average deviation of the data obtained by one operator with GPC analysis of the microoxidation test products is within f5% for each individual molecular weight fraction (Lockwood and Klaus, 1981). However, comparisons of data between two or more operators agree to within &lo%. The response of the refractive index detector on the GPC is based on the difference in the refractive index between the solvent and the component producing the response. Since oxidation products do not have the same refractive index as the original ester, correction factors were used to account for this difference. The refractive index correction factors for the low molecular weight (LMW) and high molecular weight (HMW) fractions of trimethylolpropane triheptanoate oxidation products are 1.1and 0.51, respectively, relative to the original ester (Colaianne, 1978).
Results and Discussion Lubricant oxidation is the major cause of oil thickening and sludge and varnish formation. Accumulation of sludge and varnish deposits on engine parts can cause poor lubrication and increased engine wear. An understanding of the chemistry of oil oxidation would help in developing
A -B-P hr
k2
7
D 6
where A represents the original oil, E the evaporated original oil (A) in the vapor phase, B the LMW liquid-phase oxidation products, F the evaporated LMW products (B) in the vapor phase, P the HMW liquid-phase condensation polymerization products, and D sludge and varnish deposits. The constants k,, Itz, and k5 are composite reaction rate constants. The constants k3 and k4 are composite evaporation rate constants. A diagram of the reaction scheme is shown in Figure 3. In the development of this model, it is assumed that there are no oxygen diffusion limitations in the bulk oil. Thus, oxygen can be considered to be in excess and its concentration constant. Previous investigations have shown oxygen diffusion limitations to be negligible for oil sample sizes of up to 60 pl(0.3-mm-film thickness) in the microoxidation test (Naidu, 1981). A mass balance on the mass of the original oil MA remaining in the cup is dMA/dt = [rate of reaction of A] + [rate of evaporation of A] The rate of reaction of A to B can be represented as follows: (rate of reaction of A) = -klMA
Ind. Eng. Chem. Prod. Res. Dev., Vol. 25, No. 4, 1986
598
c
EVAPORATED A AND S I N AIR
In this study, for the oxidation of TMPTH, under the test conditions used, the rate of deposit formation was negligible; therefore, the reaction step
P-D
-REACTING SAMPLE I N LIQUID PHASE +-
A
ORIGINAL O I L L M W LlOUlD PHASE OXIDATION PRODUCTS HMW L I Q U I D PHASE CONDENSATION POLYMERIZATION
B P
D E F
METAL CATALYST
'
PRODUCTS SLUDGE AND VARNISH DEPOSITS EVAPORATED ORIGINAL OIL ( A ) I N THE VAPOR PHASE
k5
in the proposed model can be neglected. It is convenient to examine the GPC data obtained as a mass fraction of the original oil present initially at the start of the test. This can be labeled as M b Thus, the following set of new variables can be defined: M A = MA/Mb M B = MB/Mb Mp = Mp/MA,
EVAPORATED LMW PRODUCTS ( 6 ) I N THE VAPOR PHASE
Figure 3. Reaction scheme for oxidation-polymerization reaction sequence.
where kl is the fust-order rate constant and MA is the mass of A at any time t in the liquid. Similarly, the rate of formation and depletion of the LMW product (B)and the rate of formation of the HMW product (P)are (rate of formation of B) = klMA (rate of reaction of B) = -k2MB (rate of formation of P) = k2MB where k2 is the first-order rate constant and MB is the mass of B at any time t. The evaporation rate of the original oil A is (rate of evaporation of A) = -KA(AR)(PA - PA")
PA", the average partial pressure of A in vapor, can be assumed to be zero. PA, the partial pressure of A in the liquid, can be written as PA = [mole fraction of A in liquid]PA" where PAo is the vapor pressure of A at the operating temperature. When there is no reaction, the liquid is pure A and the mole fraction is 1. When there is reaction, the stable reaction products in the liquid are polymers and the mole fraction of A in the liquid can still be considered to be close to 1. The Flory-Huggins equation predicts (Reid et al., 1977) that under the conditions used in this study the presence of the polymer does not have any substantial effect on the vapor pressure PAoof the solvent. Thus, the rate of evaporation of A can be considered a constant: (rate of evaporation of A) = -KA(AR)(PA") = constant = -k, The rate of evaporation for the LMW product (B) is (rate of evaporation of B) = -KA(AR)(PB - PBm) PB", the average partial pressure of B in vapor, can be assumed to be zero. PB,the partial pressure of B in the liquid, can be written as PB= [mole fraction of B in liquid]PB" where PBo is the vapor pressure of B at the operating temperature. If it is assumed that the LMW product (B) either reacts very fast or evaporates while it is being formed, then the concentration of B in the liquid can be considered a constant. This is a reasonable assumption since GPC analysis of TMPTH oxidation products usually showed no appreciable amount of the LMW product B. Thus, the rate of evaporation of the LMW product B is also a constant: (rate of evaporation of B) = -KB(AR) [mole fraction of B in liquid]PBo = constant = -12,
ME = ME/MAo
M F = MF/MA,
k3* = k s / M b
k4* = k 4 / M b
In terms of the above-defined variables, the proposed model leads to the following set of differential equations: dMA/dt = -k3* - klMA (1)
- k4*
d.d?B/dt = k1MA - k&B
(2)
dMp/dt = k 2 M ~
(3)
dME/dt = k3*
(4)
m F / d t = h4*
(5)
At time t = 0, the initial conditions are MB = MP= ME = MF = 0 MA = 1 The solutions to this set of differential equations are as follows:
M B
=
Mp = [k3* + k 4 * ] [
1 - e-&
]
7- t +
ME = k3*t M F = k4*t
(9)
(10)
From GPC analysis of the oxidation products, the mass fraction of original oil remaining (MA) and the mass fraction of HMW product produced (RP) are determined. The concentration of LMW product remaining (I&,) was too small to be quantified. The amount of original oil evaporated (ME)was determined by conducting an evaporation test under a nitrogen blanket and subsequent GPC analysis. Under a nitrogen blanket, the rate constants kl, k2, and k,* are equal to zero, and the evaporation rate constant k3* can be determined from eq 9. It is assumed that the rate of evaporation of original oil A under a nitrogen blanket (when no oxidation takes place) is equivalent to the rate of evaporation of A in the presence of air (when oxidation reactions also occur simultaneously). is then The amount of LMW product evaporated (i@~) determined by an overall mass balance: 1-
(MA + a
p ) -ME
= &fF
Ind. Eng. Chem. Prod. Res. Dev., Vol. 25, No. 4, 1986 599 ("C)
TEMPERATURE
Table I. Arrhenius Constants and Activation Energies for Reaction and Evaporation of TMPTH" activation temp Arrhenius constant, energy, EA, rate constant range, OC In [a1 kcal/mol kl 150-245 14.2 17.8 k3* 150-245 17.2 22.6 k3* 150-185 12.8 18.7 k4* 150-245 15.5 19.8 k4* 150-185 11.9 16.7
8 245 225
200 185
k'J
A
0 k:
"Test conditions: 40-pL sample size; gas circulation, 20 mL/ min; metal, low-carbon steel.
1
1
10-4k
1
0.2
1
c I9
I 20
I
2I
I 2.2
IITEMPERATURE
1 23 x lo3
I
2.4
1 2.5
A
0.8 2.6
(OK-')
4
Figure 4. Arrhenius plots for reaction and evaporation rate constants for TMPTH. Test conditions: sample size, 40 pL; gas circulation, 20 mL/min; metal, low-carbon steel.
The LMW product evaporation rate constant (k4*)can be determined from eq 10. The reaction rate constant for primary oxidation (k,) and subsequent polymerization (k,) are determined by regression analysis of eq 6 and 8. It should be noted that in all cases k2 was at least 100 times larger than kl. This implies that the condensation polymerization reactions are very fast compared with the initial oxidation. When k2 is significantly greater than kl, eq 7 and 8 for MBand M p can be simplified: MB
Mp
=
(k3
= approximately equals O
( ")
+ k4*)(-t) + k3*l (1- exp(k,t))
B. Correlations and Predictions. Oxidation and evaporation tests were conducted at 150,185,200,225, and 245 " C with a 40-pL sample of TMPTH on a lowcarbon-steel cup. On the basis of the kinetic rate model, rate constants for both reaction and evaporation were determined. The reaction rate constants, kl and k2, are inherent properties of the oil and the oxygen concentration. They vary with temperature according to the Arrhenius equation: k = (a) exp(-E,/RT) Arrhenius plots for k,, k3*, and k4* are shown in Figure 4. The evaporation rate constants k3* and k4*are functions of the mass-transfer coefficient and vapor pressure of the evaporating species. The mass-transfer coefficient depends on the air velocity, which is kept constant in these experiments. The activation energy for evaporation may be determined from a semilog plot of the evaporation rate constant vs. temperature. Semilog plots of k3* vs. T and
0
180
360 TIME
540
720
(min)
Figure 5. Comparisons of model correlations with experimental data. Test conditions: sample size, 40 pL; air circulation, 20 mL/ min; metal, low-carbon steel; temperature, 150 " C .
k4*vs. Tare shown in Figure 4. It is observed from Figure 4 that the activation energy for evaporation of TMPTH and its low molecular weight volatile products increases with increasing temperature. This is attributed to thermal decomposition at higher temperatures. The activation energy based on data obtained at 150 and 185 " C can be considered heats of vaporization. Activation energies and heats of vaporization for TMPTH and its low molecular weight volatile products are shown in Table I. The heat of vaporization for TMPTH based on this analysis is 18.7 kcal/mol. This compares well with an estimation of 15.7 kcaljmol based on the Kistiakowsky equation (Reid et al., 1977). TMPTH is thermally unstable in the thin-film microoxidation test at temperatures above 185 "C. The evaporation process is no longer purely physical but is coupled with thermal degradation. The increase in activation energy for k3* and k4*at higher temperatures is characteristic of the effect of a chemical reaction (thermal degradation). Both evaporation and thermal degradation have been shown (Ugwuzor, 1982) to follow essentially a linear relationship with respect to time. Therefore, at any temperature the rate constant k3* could be considered the sum of k3*' due to evaporation and k3*" due to thermal deg-
600
Ind. Eng. Chem. Prod. Res. Dev., Vol. 25, No. 4, 1986
-
0.IO
0.2
ME 0.3
0.1
180
0
360
540
720
c
0
900
180
120
60
T I M E (min)
TIME (min)
0.3
-
0.2
MF
0. I
0 180
0
360
540
720
900 T I M E (rnin)
T I M E (rnin)
Figure 6. Comparisons of model correlations with experimental data for TMPTH evaporation. Test conditions: sample size, 40 pL; gas circulation, 20 mL/min; metal, low-carbon steel; temperature, 150 "C.
-
Figure 8. Comparisons of model correlations with experimental data for TMPTH evaporation. Test conditions: sample size, 40 pL; gas circulation, 20 mL/min; metal, low-carbon steel; temperature, 185 "C.
- 1t MA
04 L
i
1
0.4
I
1
06
I I
i 0 L
I I
I I
,
I
0.21
I
I
1l
0
I
1
0
1 I
-
i
-MP 0 4 L
-MP 0.04
i
r
i
t-
~
1
0.02
I
-.
00 I
. 0
30
60
90
120
150
1
180
0
210
T I M E (min)
Figure 7. Comparisons of model correlations with experimental data for TMPTH oxidation. Test conditions: sample size, 40 pL; air circulation,20 mL/min; metal, low-carbon steel;temperature, 185
I
I
I
I
I
10
20
30
40
I
T I M E (min)
Figure 9. Comparisons of model correlations with experimental data for TMPTH oxidation. Test conditions: sample size, 40 pL; air circulation,20 mL/min; metal, low carbon steel; temperature, 200 OC.
OC.
radation. Thermal degradation of TMPTH becomes a factor above 185 "C (Naidu, 1985). Similarly, k4*,the rate constant for "evaporation" of low molecular weight volatile products, can be considered the sum of k4*' due to evaporation and k4*" due to thermal degradation. The rate constants kl,k3*, and k4*,obtained from this model can be used in eq 6-10 to correlate with the experimental data at various temperatures. The correlations of the experimental data with the model are shown in
Figures 5-14. It appears that for the most part the data expected by the model agrees reasonably well with the experimental data. The largest deviations are for Mp at 150 and 200 "C. There is no apparent trend to these deviations. All the correlations shown in Figures 5-14 were based on a 40-kL initial sample size (0.2-mm initial film thickness). The ability of the model to predict oxidation and polymerization rates at a different film thickness was examined by using experimental data for the 4 0 - ~ Linitial
Ind. Eng. Chem. Prod. Res. Dev., Vol. 25, No. 4, 1986 601
0.06
-
7
0.15
1
0.04
ME
0.02 0 5
0 T I M E (min) I
I
I
I
I
I
15
IO
20
25
TIME (rnin)
I
I
0.3 -
I
I
0.3
-
-
7
4
o,2
-
-
0.0I
MF
5
0
10
20
15
25
T I M E (rnin)
Figure 12. Comparisons of model correlations with experimental data for TMF'TH evaporation. Test conditions: sample size, 40 pL; gas circulation, 20 mL/min; metal, low-carbon steel; temperature, 225 "C.
0.6
0.2
i
0.I 0
0.00 0.06
MP
-
0.04 4
L
MP
0.02 O.O2l 0.01
i 0
I 5
I IO
I
I
15
20
T I M E (min)
O'O./ 0.0I
I 0
I
I
I
3 TIME
I 6
I
I 9
1
(min)
Figure 11. Comparisons of model correlations with experimental data for TMPTH oxidation. Test conditions: sample size, 40 pL; air circulation, 20 mL/miq metal, low-carbon steel; temperature, 225 OC.
Figure 13. Comparisons of model correlations with experimental data for TMPTH oxidation. Test conditions: sample size, 40 pL; air circulation,20 mL/min; metal, low-carbon steel;temperature, 245 "C.
sample size to predict oxidation rates for the 60-pL initial sample size (0.3-mm initial film thickness). These comparisons were made for tests run at 200 "C. The experimental data are compared to the predictions from the model in Figures 15 and 16. It is observed that the data predicted by the model agree moderately well with experimental values. The initial oxidation and the subsequent polymerization reaction rates are somewhat lower than those predicted by the model. This could be due to the effect of soluble metal concentration in the oil. The
oxidation-polymerization reactions considered in these predictions were conducted on an iron surface. The iron surface acts as a source of soluble organometallic corrosion product. The exposed metal surface area at the interface between the oil and the metal does not change when the initial sample size is increased from 40 to 60 pL. Since the exposed metal surface area is the same, the amount of soluble iron corrosion product will remain the same, but the concentration of soluble iron will decrease when the oil sample size is increased from 40 to 60 pL. This decrease
Ind. Eng. Chem. Prod. Res. Dev., Vol. 25, No. 4, 1986
602
0.12
7
4
0.00
-
0 0.2
ME
E '
0.04
0.1
0
0
2
4
6
i
A
0
io
8
t-
0
20
40
T I M E (rnin)
60
TIME
00
100
(rnin)
i
MF
1
~~~~~~~~
2
0
4
6
IO
8
0
T I M E (rnin)
io 08 06 04
MA I
02
40
60
80
100
TIME (min)
Figure 14. Comparisons of model correlations with experimental data for TMPTH evaporation. Test conditions: sample size, 40 pL; gas circulation, 20 mL/min; metal, low-carbon steel; temperature, 245 "C.
-
20
1
i
t
I O
00
Figure 16. Comparisons of model predictions with experimental data for TMPTH evaporation. Test conditions: sample size, 60 FL; gas circulation, 20 mL/min; metal, low-carbon steel; temperature, 200 "C.
soluble metal concentration and its effect on reaction rates should improve the accuracy of these predictions. This proposed model simplifies a complex set of oxidation-polymerization reactions into a simple reaction sequence coupled with evaporation. The preliminary correlations and predictions show that the model can accurately predict the oxidative behavior of lubricants under various environments. This model, with modifications to account for the effects of metal surfaces and additives that influence oxidation, can be used effectively in further investigations designed to provide a better understanding of the complex reactions that occur in boundary lubrication.
Figure 15. Comparisons of model predictions with experimental data for TMPTH oxidation. Test conditions: sample size, 60 pL; air circulation 20 mL/min; metal, low-carbon steel; temperature, 200 "C.
Nomenclature TMPTH, trimethylolpropane triheptanoate LCS, low-carbon steel GPC, gel permeation chromatography LMW, low molecular weight HMW, high molecular weight Mi, mass of component i Xi, mass-transfer coefficient for component i k , k * , reaction and evaporation rate constants AR, area of evaporation surface Pi, partial pressure of component i in the liquid Pio, partial pressure of component i in the vapor a, Arrhenius constant, frequency factor EA, activation energy R , gas constant T, absolute temperature Registry No. T M P T H , 78-16-0. Literature Cited
in soluble metal concentration can decrease the oxidation-polymerization rates. In the comparisons of predictions with experimental data shown in Figures 15 and 16, the model predictions at 60 p L were based on data collected at 40 pL. The model does not consider the change in soluble metal concentration when the sample size is increased from 40 to 60 pL. This may be the reason the model overpredicts both oxidation and polymerization rates at 60 pL. Modifications in the model to account for
Ali. A.: Lockwood, F.; Klaus, E. E.: Duda, J. L.; Tewksburv, E. J. ASLE Trans. 1979, 22, 267. Bailey, J. L. M.S. Thesis, The Pennsylvania State University, University Park, PA. 1983. Cho, L. F.: Klaus, E. E. ASLE Trans. 1981, 2 4 , 276. Cho, L. F.: Klaus, E. E. SA€ Tech. Pap. Ser. 1983, No. 831 679. Calaianne, J. M.S. Thesis, The Pennsylvania State University, University Park, PA, 1978. Cvitkovic. E.; Klaus, E. E. ASLE Trans. 1979, 22, 395. Czarneckl. J. R. M.S. Thesis. The Pennsvlvania State Universitv. Universitv. Park, PA, 1971. Emanuel, N. M., Ed. The Oxidation of Hydrocarbons in the Liquid Phase: Pergamon: Oxford, England, 1965.
06
-
04
I
I
I
c
MP
0 2
01
I 0
IO
1 20
I 30
I 40
I
I
50
60
70
TIME (min)
I .
603
. 25, 603-609 Ind. Eng. Chem. Prod. Res. D ~ v1980,
Naidu, S. K.; Klaus, E. E.; Duda, J. L. Ind. Eng. Chem. Prod. Res. D e v . 1984, 2 3 , 613. Reich, L.; Stivala, S. S. AutoxMstion of Hydrocarbons and Po&o/eflns; Marcel Dekker: New York, 1969. Reid, R. C.; Prausnitz, J. M.; Sherwood, T. K. The Properfies of Gases and Liquids, 3rd ed.;McGraw-Hili: New York, 1977. Ugwuzor, D. I . M.S. Thesis, The Pennsylvania State University, University Park, PA, 1982. Zuidema, H. H. The Performance of Lubricating Oils; Reinhold: New York, 1952.
Hamilton, E. J.; Korcek, S.; Mahoney, L. R.; Zinbo, M. Int. J . Chem. Kinet. 1980, 12, 577. Jensen, R. K.; Korcek, S.; Mahoney. L. R.; Zinbo, M. J . Am. Chem. SOC. 1979, 101, 7574. Jensen, R. C.; Korcek, S.; Mahoney, L. R.; Zinbo, M. J . Am. Chem. Sac. 4981, 103, 1742. Klaus, E. E.; Ugwuzor, D. 1.; Naidu, S. K.; Duda, J. L. Proceedlngs of the JSLE Internatlonal Tribologv Conference; Tokyo, Japan, 1985; p 859. Lahijani, J.; Lockwood, F. E.; Klaus. E. E. ASLE Trans. 1982, 2 5 , 25. Lockwood, F. E.; Klaus, E. E. A S E Trans. 1981, 2 4 , 276. Naidu, S. K. M.S. Thesis, The Pennsylvania State University, University Park, PA, 1981. N a b , S.K. Ph.D. Thesis, The Pennsylvania State Unhrersity, University Park, PA, 1985.
Received for reuiew December 2 , 1985 Accepted April 8, 1986
Synthesis of Styrene/Maleimide Copolymers and Physical Properties Thereof Eugene R. Moore' and Dale M. Plckelman Styrene MOMing Polymers, Dow Chemical U.S.A., Midland, Mlchlgan 48667
A series of carefully prepared copolymers of styrene and maleic anhydride @/MA) have been reacted with ammonia and converted to their respective styrene/maieimide (S/MI) derivatives. This paper discusses the preparation of these copolymers and presents extensive data on T , heat distortion temperatures, thermal stability, and melt viscosity. Some physical property data are also presenled. Some new S/MA data concerning solubility parameter, density, coefficient of expansion, refractive index, and physical properHes are also included and used for comparison.
of well-characterized S/MA copolymers further influenced our decision to proceed with this study by reacting those available copolymers with NH3. The reaction of solid S/MA copolymers with ammonia has in the past been used to modify the nature of foamed S/MA (Moore and Nakamura, 1970). In this case the NH, was allowed to diffuse into the cooled foam.
Introduction Polystyrene is well-known for its usefulness as a moldable and extrudable thermoplastic. Often applications for a clear thermoplastic, however, require higher heat resistance than that provided by polystyrene. Copolymers of styrene and maleic anhydride are known to have in-. creases in heat resistance proportional to the MA content for example, was found to increase (Moore, 1986). The Tg, linearly at 2.7 O F per added percent MA. Copolymers with styrene and maleimide have since been found to be even more heat resistant. While the copolymers for this study could have been produced from styrene and maleimide monomers, they were instead produced by first forming parent S/MA copolymers, then reacting the solid copolymer with gas-phase NH3, followed by heating under vacuum to complete conversion to the imide, as shown in Figure 1. This route has some potential advantage over production using the monomer formed from reaction of MA and NH,, since NH, can also react across the double bond of MA and thus produce a mixture of maleimide monomer and byproduds. This reaction mixture would likely require purification before polymerization. An additional problem develops with MI because, once purified, the MI monomer must be handled with care to prevent homopolymerization. MA monomer cannot homopolymerize and thus does not offer this problem. This is fortunate, since the melting point of maleic anhydride is 52 "C (Weast and Selby, 1966))well above the acceptable storage temperature for most monomers capable of freeradical polymerization. It has been very convenient to handle MA in the liquid state when polymerization is done on a larger scale. Maleimide monomer has a melting point of 93 "C (Weast and Selby, 1966), which may force it to be handled only in solution. The availability of a series 0196-4321l86l1225-0603$01.50l0
Experimental Section Copolymer Production. The base S/MA copolymers (Moore, 1986) were produced by free-radical polymerization in a well-mixed reactor with methyl ethyl ketone (MEK) as a solvent and with small amounts of either azobis(isobutyronitri1e) or benzoyl peroxide usually added as a free-radical initiator. A few copolymers were made by thermal initiation alone. Polymerization temperatures ranged from 76 to 172 "C. Higher molecular weights and higher MA contents both tended to require the lower temperatures. Devolatilization was carried out continuously to give strands of molten polymer, which were then cooled and chopped into granules. Three series of S/MA copolymers were produced, having 4-,8-, and 12-CPsolution viscosity (10% in MEK at 25 "C). Each viscosity series contained samples with aim compositions of 0,5,18, 25,33, and 48% MA. The granules were ground to about 60 mesh prior to reaction with gas-phase NH,. Conversion to Imide. About 200 lb of each copolymer (except the 0%) was converted to the maleimide form by reaction with gaseous NH3 at elevated temperature. Reaction was carried out in a large pressure-vacuum oven that held about 50 lb of the polymer in shallow trays with heated shelves where it could be heated while it was alternately pressurized with ammonia and then evacuated to remove the water of reaction. The reaction was generally carried out over a period of 2 or 3 days while the 0
1986 American Chemical Soclety
