GASIFICATION OF ORGANIC CONTENT
OF SODIUM-BASE SPENT PULPING LIQUORS IN A N ATOMIZED SUSPENSION TECHNIQUE REACTOR S T E V E N P R A H A C S A N D J E A N J. 0 . G R A V E L ' Pulp ti? Paper Research Institute of Canada, Montreal, Canada
The theoretical and practical aspects of producing gases suitable for ammonia and methanol synthesis from sodium-base spent pulping liquors, with simultaneous recovery of the pulping chemicals, were explored. Theoretical equilibrium compositions were calculated and compared to experimental results obtained under pyrolytic conditions with sodium-base spent liquors, utilizing the atomized suspension technique (AST). Experimental data are presented from tests performed in a 1 X 15 foot AST reactor at 750" to 850" C. wall temperature, 0- to 30-p.s.i.g. reactor pressure, 50 to 180 pound per hour feed rate ( 1 0- to 60-second residence time), and approximately 1 to 1 organics-water ratio. Equilibria were not achieved a t above operating conditions and the degree of approach is estimated. Some kinetic considerations are given. The results indicate that good synthesis gas yields, with simultaneous pulping chemical recovery, can be achieved a t 800" to 850" C. reactor temperatures.
THE recovery of chemicals and heat values from spent sulfite pulping liquors has been the object of extensive research in the pulp and paper industry for some years. Though several processes have been demonstrated in the last few years, more economic methods of recovery are still being sought. I n a method of treatment proposed by Gauvin, the concentrated spent liquor is evaporated and thermally decomposed in the absence of air in a tubular reactor in which the liquor is fed as a spray (4). This process, the atomized suspension technique (AST) ( 5 ) , has since been the object of basic studies in heat and mass transfer, while its application to various effluent feeds and liquors has been explored at the pilot scale (6: 7). A comprehensive study of the pyrolysis of a typical neutral sodium-base liquor was carried out in the pilot plant and the yields of pulping chemicals and the recovery of heat obtainable over a range of reactor operating conditions were reported recently (7). This paper examines the composition of the pyrolysis gas obtained in these pilot plant studies, and by comparison with calculated equilibrium compositions gains a first insight into the kinetics of the system, The use of these gases in the synthesis of tonnage chemicals, such as methanol and ammonia, would appear to offer a new avenue of benefits for this process in addition to the recovery of pulping chemicals and heat considered heretofore. The potential yields of these chemicals are considered, together with the application of these results to other spent pulping liquors. Equilibrium Calculations
The calculation of the theoretical equilibrium compositions of the products from the pyrolysis of spent pulping liquors is complicated by the complex nature of the feed and the variety of products which can form in the solid and vapor phases. 'The pioneering work of Whitney, Elias, and May (76), Sillen Present address, 1440 St. Catherine St., W., Montreal 25, Canada. 180
l & E C PROCESS D E S I G N A N D D E V E L O P M E N T
and Rengemo (75), and Bauer and Dorland (2) has been recently expanded by Rostn, who has calculated equilibrium compositions for a range of spent pulping liquor compositions at high temperatures under pyrolytic and partially oxidizing conditions (8, 70-74). Although Rostn's published data cover a wide range of conditions, the composition of the feed material used in the experiments described here was outs ide this range because of a larger water-organics ratio, characteristic of neutral sulfite semichemical (KSSC) type spent liquors. Thus, new calculations had to be performed for the chemical system under study, over the range of conditions corresponding to the actual test conditions. RosCn's method and basic thermodynamic data (70, 7 7 ) were used with some additional assumptions, the essence of which was that under the range of equilibrium conditions considered, all the sodium present in the feed will convert into Na2C03, and all the sulfur will be in the form of H2S. This leaves an organic mixture of a given elementary composition, and water, which can adjust to the equilibrium conditions. The calculations were carried out on two feed compositions representative of two batches of liquor used in the tests. Because of difficulties in obtaining reliable elementary analysis of the organics in these liquors, the figures used for the calculations were obtained by back calculations based on the product analysis from the pyrolysis experiments, for which abundant data were available. O n this basis, the elementary compositions of the total solids in these two feeds, expressed on the basis of 10 carbon atoms, were: Feed A = C I O H ~ Z . ~ O ~ . ~ S ~ , O N ~ Z . ~ Feed B = CloHI7,~Ol~.~So.9Na3.o The typical solids concentrations were 59 and 58 weight %, respectively, for feeds A and B. After eliminating the N a ~ C 0 3 and H2S which are quasi-inert products under equilibrium conditions considered here, the remaining feed mixtures were
expressed, for the convenience of the further calculations by RosCn’s method (70, 77), in the following general form:
I
=
W
the equilibria of the above reactions and the applicable content conditions :
+ nw H 2 0 + noOz
PH~O=
where
I = initial mixture that participates in the reactions W = 10C (s) 6.25 H Z (g) 3.5 0 2 (g), an.?ssumed “typical” organic solid elementary composition n, = number of moles of HzO associated with 1 mole of W no = number of moles of 0 2 associated with 1 mole of W
+
+
The two feed materials, A and B, expressed in the above form showed the following compositions:
+ 16.8 Hz0 - 0.95 = W + 23.45 HzO + 0.07
1 , =W
I,
0
2
+ ‘/z
co + CO
‘/2
0 2
0 2
0 2
=
HzO
(1)
=
coz
(2)
+ 2 Hz = CHd + ‘/z
(3)
0 2
The equilibrium concentrations for the whole mixture can be calculated at suitably selected values of equilibrium oxygen partial pressure, and initial water content, n,, of interest, by solving the following simultaneous equations which represent
Table 1. n,
20 25 30
20
25
30
20 25 30
pcoz = ~C02~cO-1PcO(P02)1’2
(5)
-’”
(6)
PCHi
= KCEIiKCO-’PCOPHz”(POz)
+ pcoZ+ 10 PQ-’ P H+ ~ PH~O +2 PQ-l (6.25 + n,) pco2 + 0.5pc0 + 0 . 5 p ~ ~+0 PQ-’ (3.5 + no + 0.5 n,) + f + f =P PCO
PCH,
PCH,
PHz
PHzO
PCO
(8)
=
PCOz
(9)
(10 )
PCHi
Ki = equilibrium constant of a reaction leading to the formation of i from its elements p i = partial pressure of component i, atm. abs. P = system pressure, atm. abs. Q = total number of molecules in the gas phase, per 10 atoms of C The above set of equations was solved for a pressure of 1 atm. abs.; for temperature levels of 800”, 1000°, and 1200” K.; for nu values of 20, 25, and 30 (to cover a range wide enough for any practical NSSC feed pyrolysis); and for pol values corresponding to a range of no approximately from -2.0 to +2.0. The set of equilibrium compositions so calculated is presented in Table I. Equilibrium compositions corresponding to the exact feed compositions used here (I* and I B ) were obtained by interpolation, at 1 atm. and 800”, 1000”, and 1200” K. Rostn’s
+
PH2
(7)
=
Calculated Composition at Equilibrium of Gas Phase from Pyrolysis of NSSC Liquors of Composition W no 0 2 , at Different Pressures, Temperatures, and nw and no Values
-1ogPOz
(4)
in which, for any component i,
Initial checks for these two reactant mixtures showed that in the temperature and pressure ranges to be used in the equilibrium calculations, all the carbon should be present in the gas phase. Consequently the individual equilibria to be considered in this case can be reduced to those of Reactions 1. 2, and 3 ( 7 7 ) :
Hz
KH~OPH~(PO~)’’~
PHz0
PCO
pcoz
+ n,
HzO
PCH4
Q
no
0.093 0.054 0.038 0.090 0.071 0.053 0.049 0.0355 0.258
30.8 32.9 33.8 35 . O 36.0 37.0 42.0 43.2 44.0
-1.1 +1.1 $2.2 -2.8 -1.5 -0.3 -1.4 0 $0.9
26.4 26.2 26.1 26.4 26.3 26.2 26.2 26.1 26.0
0.300 0.272 0.252 0.320 0.309 0.289 0.299 0.280 0.261
0.367 0.420 0.442 0.394 0.422 0.447 0.462 0.490 0.510
1-Atm. Pressure, 800’ K. 0.0389 0.193 0.0346 0.216 0.032 0.226 0.333 0.163 0.0317 0.175 0.0300 0.187 0.026 0.163 0.0243 0.172 0.0227 0.179
21 . 0 20.4 20.3 20.1 20.4 20.1 20.0 19.6 20.0 19.8 19.6
0.531 0.420 0.400 0.354 0.440 0.379 0.352 0.268 0.362 0.320 0,278
0.193 0.305 0.325 0.362 0.319 0.387 0.405 0.489 0.416 0.464 0.507
1-Atm. Pressure, 1000° K. 0.1825 0.0916 0.138 0.137 0.130 0.145 0.114 0.161 0.121 0.120 0.102 0.142 0.094 0.149 0.0693 0.1734 0.0855 0.135 0.0721 0.1453 0.062 0.155
0.005 0.001 0 0 0,001 0 0 0 0.0003 0 0
35.8 36.2 36.3 36.4 41.5 41 . 0 41.2 41.2 45.3 46.5 46.1
-3.52 -0.65 0 $1 . o -1.90 -0.15 +O .40 +2.75 -1 .oo +0.6 +2.20
16.4 16.0 15.6 16.0 15.6 15.4 15.6 15.4 15.2
0.484 0,405 0.321 0.422 0.336 0.293 0.348 0.303 0.258
0.241 0.320 0.402 0.334 0.421 0.462 0.436 0.478 0.520
1-Atm. Pressure, 1200’ K. 0.206 0.071 0.179 0.098 0.148 0.128 0.156 0.0855 0.130 0.113 0.116 0.127 0.116 0,100 0.103 0.113 0.089 0.125
0 0 0 0 0 0 0 0 0
36.1 36.1 36.2 41.4 41.2 41.2 46.2 46.3 46.5
-2.90 -0.95 +1.10 -2.3 0 +1.1 -1.1 $0.2 +1.5
VOL. 6
NO. 2
APRIL
1967
181
data for n, = 15 were taken into consideration when interpolating for n, = 16.9 (Ill). For ease of comparison with experimentally determined gas-phase compositions, the equilibrium compositions obtained were recalculated to the steamfree basis. A calculation was then made to see whether any “free” carbon-Le., carbon not combined as Na2COa-should remain in the solid phase, using the formula
Ca = 10 -
Q(PCO
+ + PCO~
Table It.
K.
Concentration, Val. yo Feed A Feed B 31.9 34.6 44.7 50.3 7.0 5.5 16.3 3.5 1.7a 0 16.8 23.3 59.7 60.0 23.1 16.8 1 .o 0 0 0 14.0 19.5 57.7 57.4 28.8 23.3 0 0 0 0
Component
Free solid Ca
coz co
1000
(11)
PCEJ
Equilibrium Compositions at 1 Atm.
?W.,
Hz
where C, = free carbon remaining as a percentage of the total potential free carbon in the feed-Le., the total carbon in feed less that in the product NazCOa as calculated from the total‘ Na. The results of this calculation are shown in Table 11. All are zero except for a small percentage in the case of feed A a t 800’ K. and 1 atm. abs. Since the experimental AST runs considered here were made at approximately 2 atm., it was necessary to make a correction in the calculated equilibrium compositions. However, both the experimental data and RosCn’s calculations for 1 and 30 atm. showed that the effect of pressure in the range of interest is not substantial. Correction factors were therefore calculated from Rostn’s data for the above two pressures a t nu = 15, and applied on the assumption that changes in the equilibria are proportional to the absolute pressure ratios. The equilibrium compositions and “free” solid C values so calculated are given in Table 111.
CH4 Free solid C” 1200
Free solid C“ a
% ’ of potential free carbon in feed. Table 111.
Cmp.,
K.
Equilibrium Compositions at 2 A h .
Concentration, Vol. % Feed A Feed B 34.0 36.9 38.5 43.4 5.9 4.6 21.6 15.1 1.5“ 0
Component
800
CO2
Hz
co
CHI
Free solid C”
1200
C0.J - _
Hz
co
Experimental
Apparatus. A Aowsheet of the AST pilot plant used in the tests is given in Figure 1. The reactor vessel consists of a cylindrical pressure vessel of 316 stainless steel, 12 inches in inside diameter, and 15 feet high, equipped with an atomizing nozzle a t the top and a product draw-off pipe a t the bottom. The tube is surrounded by six banks of electrical radiant heaters, each individually controlled to maintain a set wall temperature in the corresponding zone of the reactor. The liquor-feeding system consists of two steam-heated tanks, from which the warm liquor Aows through a motorized edge-type filter, via a magnetic flowmeter to a controlled
a
CH4 Free solid Ca % of potential free carbon in feed.
-METERS
SAMPLING VALVES BY-PASS
I
!
f
FLOWMETER
FILTER
FEED TIHKS
REACTOR
Figure 1. 182
CYCLONE SEPARATOR
SCRUBBER
CONTROL VALVE
COHDENSER
AST experimental apparatus
I&EC PROCESS DESIGN AND DEVELOPMENT
0
20.1 . .~ 56.8 23.1
0 0
volume high-pressure triplex metering pump. The liquor is fed a t high pressure to a steam heater and then to an electrical heater consisting of llr-inch stainless steel tubes used as a low voltage resistance element. I t is then atomized in the reactor by flashing from a circular orifice into the relatively low reactor pressure (0 to 45 p.s.i.g.). I n the reactor, the atomized
ORIFICE
1
14.4 57.1 28.6 0
TRLiP
liquor is evaporated and thermally decomposed. T h e bulk of the solid product is collected in the cone, while virtually all the suspended fines are collected from the hot gas stream in a cyclone separator. After expansion in a pressure control valve, the water vapor is removed in a shell and tube condenser and the pyrolysis gases pass to a set of stainless steel orifice meters in parallel, a wet-test meter for orifice calibration, a set of gas sampling valves, and a scrubber for odor control. Analyses. FEEDLIQUOR.The specific gravity and solids concentration of the feed liquor were determined by TAPPI Standard Method T-629. The solids calorific value was obtained in a Parr bomb, using sodium peroxide as oxidant. In calculating the calorific value from the measurement, no ash correction was used. The ash content of the evaporated solids was determined by ashing at 600’ C. Total sodium was analyzed by flame spectrophotometry. Total sulfur was determined by wet oxidation and subsequent precipitation of BaS04. Sulfate sulfur was determined by TAPPI Standard Method T-624. GASES. The analyses were performed by absorption in standard solutions in an Orsat-type apparatus, except for the H& which was determined separately by adsorption on anhydrous copper sulfate. All paraffins were reported as methane, since spot checks showed that only very small amounts of ethane were present. Feed Material. Both feed materials were obtained from the same source, a midwestern U.S. pulp mill using the NSSC pulping method. Feed A was stored a t room or slightly higher temperatures for 2 l / ~to 3 years, feed B for 1 to 6 months prior to their use in the experiments reported here. This may have been partly responsible for the markedly lower H and 0 content of the organics in feed A.
to come directly from the wood. This was estimated by subtracting all the Na2S03 that could be present in the feed solids on the basis of feed sulfur analysis, and it was assumed that the rest of the sodium was present in the pulping liquor in the form of NazC02. After subtracting the corresponding total number of sodium, sulfur, carbon, and oxygen atoms, the remaining organic matter was considered to have come from the wood chips. Discussion
Conversions us. Equilibria in Gas Phase. T o compare the gas-phase conversions experimentally obtained in the AST runs with the calculated equilibria, the former were plotted against the latter in the following way. First the equilibrium compositions calculated for feeds A and B a t 2 atm. abs. (see Table 111) were plotted as solid lines against equilibrium temperature, Figures 2 and 3, respectively. From the product gas analyses obtained during the experimental AST runs, the components HzS, Nz, and 0 2 were eliminated, the first on the grounds of being an inert component in the system, the latter two because their presence was due to purging the instruments with Nz and/or leakage during sampling or analysis. The concentrations of the remaining components were increased proportionally.
Experimental Data. Twenty representative experimental runs were selected from the 39 published earlier (7) for investigating the gas-phase reactions. T h e main operating variables in this group of experiments were the reactor wall temperature (750’ to 850’ C.) and the feed rate (4 to 14 imperial gallons per hour). The two batches of feed material used in the experiments are treated here separately. Typical experimental data and calculated yields are shown in Table IV. I n all relevant calculations, it was assumed that the HClinsoluble material in the solid product has the following composition :
~
wt.% C
H 0 S
Ash
73 2.3 15
5.5 4.2
This was the average of four actual elementary analyses on HC1-insolubles from experiments with feed B. There was relatively little variation in the elementary composition of the HC1-insolubles from one sample to the other, except that the oxygen and hydrogen content decreased somewhat with increasing degree of gasification of the organics. No elementary analysis for the HCI-insolubles was made for runs with feed A, and the above composition was assumed for both types of feed. T h e “organic carbon in solid product, yo’’ was calculated separately for each run, on the basis of potentially available free carbon in the feed, assuming that under equilbrium conditions all of the sodium would be in the form of NazCOa. I t is likely that the actual carbon content in the case of runs with feed A was generally somewhat higher because of the higher carbon content of the organics in the feed. The gas yields given on the organics basis in Table I V were calculated in a manner to make the gasification data useful for estimating yields to be expected using spent liquors of various kinds, including kraft-, calcium-, or magnesium-base sulfite liquors. T h e “organics fed” in this case are assumed
750
800
850
900
950
1000
1050
EOUlLlERlUM TEMPERATURE,
I100
1150
1200
O K
Figure 2. Composition of pyrolysis gases from feed liquor A at 2 atm. abs.
60-
no)
0
W
a
3
50-
efi
40-
LEGEND
.-cot
0
0-H * 0-
co
A - CHq
+- ILLUMINANTS ( C p H, ETC I
800
850
900
950
1000
1050
EQUILIBRIUM TEMPERATURE,
1100
1150
1200
O K
Figure 3. Composition of pyrolysis gases from feed liquor B at 2 atm. abs. VOL. 6 NO. 2
APRIL 1967
183
Then, on the same graph as the calculated curves for gasphase component concentrations at equilibrium for the same feed, the set of points representing each corrected analysis was plotted at the point on the temperature scale which made its figure for COZconcentration lie on the curve of COz concentration at equilibrium. This method of plotting was chosen because it readily shows the relative progress of the reactions between the various experiments in terms of one gas component, COZ, and indicates the pattern of deviation of the other gas concentrations from the calculated values at equilibrium, over the full range of the experiments. The COZ concentration was chosen as the basis of comparison, because it was considered the most reliable when, as in these tests, the Orsat analyzer is used. A cursory observation of the experimental points shows that, in general, there is relatively good agreement between the experimentally found gas concentrations and those calculated for equilibrium, relative to a given COz concentration. The only consistent deviation is represented by the presence of appreciable concentrations of illuminants (unsaturated hydrocarbons). Rostn, in his initial checks for the possible gas components in the systems considered (77), calculated that any unsaturated hydrocarbons would be present in concentrations less than lo-* volume yo. Orsat analyses, spot-checked by infrared absorption on three gas samples, have shown the presence of illuminants in approximately 1 to 3 volume % concentrations, made up of varying proportions of ethylene and acetylene. I t appears that decomposition of some of the organic components present in the spent pulping liquors produces some unsaturated hydrocarbons in the initial stages of the pyrolysis, and these are not decomposed in the short residence allowed in the reactor. Martin ( 9 ) has shown that significant amounts of ethylene evolve at about 600' C. under pyrolytic conditions from cellulose. The absolute amounts, and to a lesser extent the concentrations of methane found in the experiments, tend to remain constant rather than to decrease sharply with temperature, as would be expected from the equilibrium calculations. One possible explanation of this could be a reversion of the CO and Hznear the outlet of the reactor, where the temperature was always lower than at the reactor wall. I t might also be due to C H I forming on the solid carbon particles in concentrations above the over-all system equilibriums, because of the relatively low water vapor pressure in the boundary layer. I t is likely that once methane is formed, it will not decompose readily when it gets into the water-rich gas phase for a few seconds, since the C H I molecule has considerable stability at the temperatures prevailing in the experiments performed here. The superimposition of experimental points over the equilibrium curves on the basis of COS concentration provides one measure of approach to equilibrium for the various tests. Another measure of the over-all kinetic performance of the gas phase of the system at various reactor wall temperatures and feed rates can be obtained by looking at the CO/COz ratio. The choice of this ratio is based on the relatively good accuracy of the COz determination and the relatively sharp increase in the ratio with increasing equilibrium temperature over the range of interest. The apparent equilibrium temperatures reached are presented in Figure 4 as a function of feed rate for different reactor wall temperatures and feed materials. The corresponding equilibrium temperature values, together with the difference between these and the prevailing wall temperatures, are also given in Table IV. I n the 750' to 850' C. range, the reactor temperature is the most important variable. The corresponding equilibrium tempera184
I & E C PROCESS D E S I G N A N D D E V E L O P M E N T
LEGEND FEED " A " - 750.C A F E E D "4"- 800.C rn F E E 0 " A " - B50.C o FEED'%"- 7so.c A F E E D % ' - BOO'C
Y
h
2
m
6501
,
+
9 6*5i
5OOL
I
I
L-
IO
I---
12
L_J 14
FEED RATE, IGPH
Figure 4. Equilibrium temperatures obtained vapor phase, based on CO-CO2 ratios
in
ture reached is naturally higher at higher reactor temperatures, but also the gap between the two is reduced at higher reactor temperature at least at low feed rates. The effect of feed rate is important at 800' C. and above, while the conversions are almost independent of feed rate at 750' C. A possible explanation for this is that the reduction of COZ by Hz, and the steam-carbon reaction, the primary source of CO, are very slow below 750' C. The fusion of the Na2C03-NaZSOd mixture in the solid product, which occurs at 760' to 780' C., could also have a role in the rapid acceleration in general, and C O production in particular with decreasing feed rate-Le., increasing residence time. Further studies are planned for elucidating this phenomenon. The fact that the results from feed B approach the calculated equilibria more closely than do those from feed A at similar operating conditions, suggests that aging the spent liquor organics significantly reduces their reactivity. Thus, processing of freshly collected feeds in a pulp mill may produce a somewhat higher degree of conversions than is achieved with our more reactive feed B which had been stored for several months. Conversions us. Equilibrium in Solid Phase. The degree of approach to the equilibria by the reaction in the solid phase is more difficult to estimate than that in the gas phase. According to our calculations, no solid carbon should exist under equilibrium conditions above about 540' C. in the case of feed A. I n the case of feed B this temperature limit could not be established precisely, because it fell below the minimum temperature (527' C.) for which basic equilibrium data were available. I t is only in the case of run A31, where the free carbon in the solid product is reduced to 1.7% of the potential free carbon in the feed, that the approach to equilibrium can be determined with some accuracy. At 527', l.5yo of the free carbon is expected to remain in the solid phase (see Table 111), and the 1.7% found in this solid product suggests a n equivalent temperature of about 525' C. This run was performed at 850' wall temperature, giving for the solid-phase reaction an apparent "lag" of about 325' C. The lag measured in the gas phase by the CO/COZ ratio was only 184' C., indicating that solid-phase reaction was behind by about an additional 140' C. This difference in the degree of conversion between the gas and solid phase is probably even greater in the case of the other runs, on account of higher feed rates and/or lower operating temperatures. Despite the uncertainties of
the actual difference of the degrees of conversions in the two phases, the comparisons show clearly that the controlling reactions are between the gas phase and the solid phase. T h e main heterogeneous reactions expected in the gasification of this system are:
+ H2O = CO + H2 C + 2 H2O COz + 2 H2 C
=
c+coz=2co C
+ 2 H2 = CHI
(12) (13) (14) (15 )
Also, the water gas shift reaction has an indirect role, by consuming or producing CO according to
CO
+ HzO
=
COz
+ Hz
(1 6)
I t has been shown by Gadsby, Hinshelwood, and Sykes (3) that by far the most important of these reactions, from the point of view of gasification, is Reaction 12. According to Gadsby, Reactions 13 and 14 are relatively slow. Reaction 15 should be very slow because increase in the H2 partial pressure definitely retards the over-all gasification process. I t is clear from these considerations that the best way of promoting the gasification at a given operating temperature is to reduce the feed concentration, and this has been confirmed here (7). I n a commercial process, however, it is desirable to use a highly concentrated feed, to reduce the evaporative load in the reactor. For a given water-organics ratio, the steam-carbon reaction is controlled either by strictly kinetic reasons or by diffusion rates. Furthermore, in a downflow AST reactor, the solid particles have somewhat shorter residence time in the reactor than the gas phase, because of the initial velocity imparted by the spray nozzle and the Stokes velocities of the particles. More study of the pertinent solid-vapor reactions is needed before the relative importance of the above factors can be determined. But from the above considerations, it can be expected that finer atomization would increase the rate of the heterogeneous reactions and thus accelerate the over-all conversions. Smaller drop sizes should result in lower particle velocities and thus increased residence times allowing the reactions to proceed further, regardless of whether the basic limitation is kinetic or diffusional in nature. If diffusion plays a part, the improvement will be larger, since a decrease in particle radius will increase the surface area and reduce the diffusion paths.
and lignin which in various proportions make up the organic content of the spent liquors. A simple sugar of the C8H120a type gives a maximum theoretical yield of 7 5 . 5 tons of NHa per 100 tons and polysaccharides, ( C ~ H I O Ocan ~ ) yield ~ , 83.9 tons of NH3 per 100 tons. Lignin, assuming an elementary composition of CloHlo.~O~.,, calculated from the Adler formula, can yield 129.0 tons of NH3 per 100 tons of lignin. O n the other hand, the theoretical maximum yields h o m any spent liquor will be always somewhat less than what would correspond to the ratios of sugar-polysaccharides-lignin in the organic portion of the liquor, because the base (Na, Ca, or Mg) will always tie up some potential hydrogen-producing carbon as carbonate, and most of the sulfur will be converted into HzS, thus consuming some hydrogen. The amount of NH3 which could be produced from the reactor off-gases can be estimated by assuming that it would be produced from all the HZ present as such plus all that obtainable according to Reaction 16, when it is completely shifted to the right. The amount of methanol which could, alternatively, be produced can be estimated by assuming that all CO present will be converted into CHIOH, together with that portion of the COZwhich can be converted by the available excess Ha, according to the reaction :
C02
+ 3 Hz = CHIOH + HzO
(17)
The yields of "2, or alternatively of CHaOH, in tons per 100 tons of organic matter in waste liquor, have been estimated in this way for all the experimental AST runs reported herein, and are tabulated in Table IV. The estimated yields of NH3 have also been piotted against feed rate (Figure 5). The trends are very similar to those shown in Figure 4, and show the effect of the same variables on the extent of gasification reached. Although all the above reactions have been assumed to be quantitative, which would not be the case in practice, breaking down the methane into CO, COP, and H2 by a secondary reformer, or by recycling it into the AST reactor, could on the other hand increase the synthesis gas yields appreciably over the figures used. Other Types of Spent Pulping Liquors. A limited number of tests have been made in recent years in our AST pilot plant on other types of spent pulping liquors-e.g., sodium bisulfite, sodium-base acid sulfite, sodium sulfide (kraft), calcium-base acid sulfite, ammonium-base, and ion-exchanged decationized liquors. The data obtained so far show that the
Practical Significance of Experimental Results
Production of Ammonia a n d Methanol. The gases produced from the pyrolysis of spent pulping liquors in AST reactors have the composition of typical crude synthesis gases used in the manufacture of ammonia, methanol, and other chemicals. I n the basic pulping chemicals recovery process considered heretofore, these gases are used as a fuel to heat the reactor vessel and generate the steam required to concentrate the spent liquors in a multiple-effect evaporator. The substitution of the cheapest available fuel for this duty, and the use of the reactor off-gases in the synthesis of chemicals, appear to offer the possibility of additional benefits from spent liquors over and above the recovery of pulping chemicals. The maximum amounts of I\"$, or alternatively of CHaOH, obtainable from feed A were calcuiated from the elementary composition of the waste liquor organics to be 104.1 or 97.8 tons, respectively, per 100 tons of organics. The corresponding figures for feed B were 70.2 tons of NH3 or 65.9 tons of CH30H. Maximum yields were calculated for the sugars, polysaccharides,
LEGEND FEED " A " A FEED * A " FEED ' A ' * 0 FEED'%"A FEED"8"-
d -1
40:
>
-
w
I
I 750-C
800-C 050°C
750'C 000'C
c
O
L
L
-
2 4
6
e
10
12
14
FEED R A T E , IGPH.
Figure 5. gases
Potential yields of ammonia from pyrolysis
VOL. 6
NO. 2
APRIL
1967
185
Table IV. A5
Operating conditions 1 Reactor wall temp., C . 2 Feed rate, imp. gal./hr. 3 Reactor pressure, p.s.i.g. 4 Feed solids r;te, Ib./hr. 5 Feed temp., C. 6 Feed pressure, p.s.i.g. 7 Measured outlet temp., C . 8 Est. residence time, sec. Feed analysis 9 Spec. grav., g./cc. 10 Solids concn., wt. 7 0 11 Solids calorific value, B.t.u./lb. 1 2 Ash, wt. 7' of solids 1 3 Total sodium, wt. yo 14 Total sulfur, wt. yo 15 Sulfate sulfur, wt. % Product gas analysis 16 HzS, V O ~ . % 17 COz, V O ~ .% 18 Illuminants (assmd. CzHa), vol. % 19 Hz, V O ~ . % 20 co, vol. yo 21 CHI, V O ~ . 70 22 0 2 , vol. yo 23 NI,V O ~ . % 24 Av. molecular weight
Typical Experimental Results Run No. A14 A3 7
750 4.4 6 34.5 220 670 545 30.2 1.35 58.2
750 13.0 30 96.0 210 580 640 18.3
41.5 9.80 4.44 0.02
1.33 55.5 5271 41 .O 9.68 4.17 0.81
6.5 26.9 2.7 47.3 4.9 12.2 0.0 0.0 19.0
9.5 32.0 1.9 36.0 5.1 15.3 0.0 0.0 22.4
N.A.
B7
B17
B30
750 4.5 6 34.2 170 530 540 25.8
800 4.1 10 30.2 240 810 555 27.4
800 13.5 30 107 200 480 685 16.1
1.32 55.7 4880
40.3 9.83 4.55 0.01
1.33 57.1 5118 41.8 9.05 3.88 0.51
1.34 59.5 5045 40.4 9.30 3.80 0.63
2.5 20.1 1. o 52.6 15.6 8.2 0.0 0.0 16.7
4.1 29.9 1 .o 56.1 4.7 4.1 0.0 0.0 17.9
2.8 25.2 0.8 57.4 8.8 4.0 0.0 0.0 16.8
850 4.0 10 32.4 220 590 650 25.4 1.35 59.8
N.A.
N.A.
9.22 3.70 0.40
6.0 30.6 1.2 48.8 5.2 7 .O 0.0 0.0 19.7
Yields and conversions 53.0 57.6 39.5 50.0 42.6 25 Solid yield, wt. yo of solids fed 51 .O 53.0 42.0 82.0 53.2 78.0 26 Gas yield, wt. yo of solids fed 46.8 89 96 93 98 27 Mass vield. wt. c/r of feed 99 100 32.6 1.66 5.36 53.6 38.6 28 Organic carbon in solid prod., %a 49.0 95.1 145 89.4 29 Gas yield, wt. yo of organics fedb 77.9 134 77.2 8.50 10.4 4.04 16.3 16.2 6.35 30 Hz yield, SCF/lb. organics fed* 0.88 0.84 0.57 0.67 4.84 2.53 31 CO yield, SCF/lb. organics fedb 2.19 0.74 2.55 1.15 32 CH4 yield, SCF/lb. organics fedb 1.72 0.91 34.4 29.7 14.6 66.8 59.2 22.2 33 NH, eauiv. of H2. T/100T or-zs.c 32.2 27.8 62.8 13.8 55.5 20.8 34 CHiOH equiv. of H;/T/lOOForgs.d 0.78 0.18 0.16 0.16 0.35 0.17 35 CO/CO2 ratio in gas 530 522 666 542 638 549 36 Equil. temp. corresp. to CO/C92, c 184 208 251 220 228 162 37 Approach of equilibrium, A T , C. a As % of free carbon (NatCOa excl.) in feed. 7 3 w t . yo carbon in HCI-insoluble fraction of solid product assumed in all cases. b On basis of feed Including HZ to be produced by CO HzO = COz H Z reaction. d Including C H a O H to be produced by organics, NazCOa and Na,SOa excluded. C02 3 HZ = C H a O H H z O reaction.
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various types of sodium-base liquors, under identical processing conditions, produce gas yields and gas compositions which indicate very similar kinetic behavior. Liquors with other than sodium as the dominating cation show somewhat slower reaction rates. A more detailed investigation of the effect of the cation on the gasification is planned for the near future. Equally important in the choice of liquor, however, will be the inorganic transformations in the solid phase, which will affect the recovery of the pulping base and the sulfur. The most serious limitation to higher operating temperature in the AST reactor is related to corrosion, which is also influenced by the fusion point of the ash. This corrosion limitation, in the case of sodium-base liquors, would appear from our pilotscale operating experience available to date, to be at 800' to 850' C. Further work is required to establish the limiting conditions for other bases. One interesting possibility for further studies is offered by ammonia-base liquor, because of its freedom from the ash fusion problem, and the possibility of using part of the ammonia produced from the gases for the preparation of the pulping liquors. Ac knowledgment
The authors recognize the cooperation of J. Kersch in the pilot tests and of R. Grozell in the analytical work. P. Ratki and P. Lancaster assisted in some of the numerical calculations. 186
l & E C PROCESS D E S I G N A N D D E V E L O P M E N T
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literature Cited (1) Barclay, H. G., Prahacs, S., Gravel, J. J. O., Pulp Paper M a g . Can. 65,No. 12, T-553 (1964). (2) Bauer, T. W., DorIand, R. M., Can. J . Technol. 32, 91 (1954). ( 3 ) Gadsby, J., Hinshelwood, C. N., Sykes, K. SY., PYOC.Roy. Soc. 187A, 129 (1946). (4) Gauvin, W. H., Chem. Can. 7,No. 9,48-56 (1955). (5) Gauvin, W.H., U. S. Patent 2,889,874 (1959); (to Pulp and Paper Research Institute of Canada) Can. Patent 522,789 (1958). (6) Gauvin, W. H., Gravel, J. J. O., Proceedings of Symposium on Interaction between Fluids and Particles (1962), London, Institute of Chemical Engineers, London, pp. 250-9. (7) Gauvin, W. H., Gravel, J. J. O., T a p p i 43, No. 8 (1960). (8) Koszegi, L., Rostn, E., Trans. Roy Inst. Technol., Stockholm, No.221 (1964). (9) Martin, S. B., 1964 Spring Meeting, Western States Section, Combustion Institute, Paper 64-7. (10) Rosh, E., Trans. Roy. Inst. Technol., Stockholm, No. 159 (1960). (11) Ibid., No. 160. (12) Ibid., No. 161. (13) Ibid.; No. 197 (1962). (14) Ibid., No. 223 (1964). (15) Sillen, L. G., Rengemo, T., Suensk Papperstidn. 5 5 , 662 (1952). (16) Whitney, R. P., Elias, R. M., May, M. N., T a p p i 34, 396 (1951).
RECEIVED for review March 25, 1966 ACCEPTED November 7, 1966
