-------
Nozzle Pressure (psi)
Figure 3-6. Variation of mean droplet diameter with fuel
.pressure for heptane and No. 2 fuel oil using
the 1.5 gal/hr nozzle.
3-13

-------
The conclusions are that, for a given nozzle and fuel, (1) atomization
is principally a function of atomizing pressure and fuel flow, and (2) a
limited number of measurements are sufficient to approximately characterize
the variation of mean diameter with flow and fuel pressure.
3.3.2 Relation of PRE to Atomization Performance
In the previous work in our laboratory [6], both the on-design and off-
design atomizer conditions were used in the TFR to determine the influence of
atomization quality on DRE. Fuel flow rate of n-heptane was the independent
variable and the air flow rate (17.3 1iters/second) was maintained constant.
For the "on-desi gn" condi tion, Delavan WDA 60° series nozzles of various
capaci ties (5.7, 3.8, 2.8, and 1.9 1i ters/hour) were used to maintain
constant atomization quali ty as fuel flow was charged at constant nozzle
pressure (125 psig). (The data of Figure 3-3a show the assumption of
constant atomization quality was, in fact, reasonable.) The "off-design"
condition was obtai ned by varying the fuel flow of one nozzle (5.7 1 iters/
hour) at constant air flow rate.
The on-design data are presented in Figure 3-7. In Figure 3-8a, the
off-design results are compared directly to the on-design results. As
shown, the compound destruction efficiency and combustion performance was
substantially degraded over the on-design performance. The atomization
characteri zation of the nozzles used for these tests show clearly the
degradation in atomization performance associated with the off-design
operation {Figure 3-8b). The Sauter Mean Diameter increases as percent
theoretical air departs from the on-design condition of 130 percent TA.
More important, the percent mass associated with large droplets
(dp>160 measurement) increases dramatically from near zero percent if the
total mass at 130 percent TA to over 40 percent at 320 percent TA.
Under the present program, additional testing was conducted to obtain
further evidence of the association of atomization with destruction
efficiency. The TFR was operated with No. 2 fuel oil doped to 3.0 weight
percent with an equimolar mixture of test compounds used in the previous
3-14

-------
Figure 3-7. Test compound emissions from the TFR
as a function of theoretical air
n-heptane [6], (c) Copyrighted by The
Combustion Institute. (Reproduced by
permission.)
3-15

-------
a) Exhaust CO and Fraction Remaining
80(	
m °
§ 5 60
2 I
ON-DESIGN
100 150 200 250 300
PERCENT THEORETICAL AIR
350
Figure 3-8. Impact of atomizer performance on fraction of
test compound and CO in the exhaust, and
nozzle SMD n-heptane.
3-16

-------
b) Nozzle Performance
60 -

Q
ja
 |i>>>
o
o
4->	i©
C «~H
-------
study. The No. 2 fuel oi1 was supplied through a 3.8 1 iters/hour
(1.0 gallons/hour) WDA Delavan 60° nozzle at the rated pressure of 200 psig
for the "on-design" condition. To achieve "off-design" conditions, an
oversized WDA Delavan 60° nozzle (5.7 liters/hour) was used to supply the
same fuel flow rate. Theoretical air was the independent variable.
Figure 3-9 shows the fraction of each of the compounds that escaped
destruction. The on-design nozzle results show results consistent with the
previous TFR data, namely:
t A range of high DRE values are indicated at stoichiometrics between
100-200 percent theoretical air.
•	At low theoretical air the increased waste emissions indicate a
failure mode due to fuel-rich pockets breaking through the flame.
•	At high theoretical air the increased waste emissions indicates a
quenching failure mode in which the high air flow is quenching
portions of the flame prior to complete reaction.
Comparison of the on-design and off-design plots shows that the
emissions at the rich and lean failure modes are not significantly different.
However, the DRE in the region between 100 and 200 percent theoretical air
has degraded markedly from the previous high efficiency. An examination of
the nozzle characterization data should correlate to these results.
Figure 3-10 illustrates the droplet size distribution obtained for the
on-design and off-design nozzles used in the present study. Use of an
oversized pressure atomized nozzle for the off-design condition results in
low fluid pressure, and low atornization energy. Thus, the off-design results
show the dropsi ze is shi fted toward 1arger values. The key to interpreting
the effect of the shift in droplet size is found in the evaporation time plot
of Figure 3-10. This plot shows evaporation time as a function of droplet
diameter for No. 2 fuel oil, based on the "d^ law" [26]. The on-design data
shows approximately 10 percent of the mass is above 160 microns. According
to the evaporation rate plot, this 10 percent will require more than 50 msec
3-18
-------
Figure 3-9. Test compound emissions from TFR as a function of
theoretical air (No. 2 fuel oil).
3-19
-------
1000
600 400	200 100 80 60 40	20
Diameter, microns
10
¦
\ 1


-r-p-
1 1
1

Design Operation
¦a
*
•







Off-Design
—r
800 400 200 100 80 60 40	20
Diameter (microns)
Figure 3-10, Droplet size distributions and estimated evaporation time
as a function of diameter.
3-20
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to evaporate. The off-design data indicate that fully 46 percent of the mass
is greater than 160 microns. A1so note that the 1argest size class (250-560
microns) has increased from 2 to 16 percent of the total mass. Since the
evaporation times for this category range from 100-700 msec, it is evident
that the effect of moving from on- to off-design operation is a substantia?
increase in the evaporation time of a significant fraction of the fuel.
Thus, the change in atornization quality that accompanied the use of the
oversized nozzle induced an atomization failure mode; the DRE, which was much
greater than 99.99 percent was reduced to the order of 99.9 percent.
3.3.3 Mechanisms of Failure
Two general mechanisms can be identified by which poor atomization can
influence DRE. In the first, droplets which are too large to evaporate in
the available time penetrate to the reactor wall. The liquid evaporates and
exits the reactor along the cold boundary layer at the wall. In the second
mode, the droplets penetrate through the flame-zone without fully evaporating
until well into the postflame region. Here, mixing or temperature may not be
sufficient to ensure complete destruction.
Estimation of the maximum droplet diameter for which the droplets avoid
striking the wall involves 1) determination of fraction of the hydrodynamic
energy released by the nozzle that is converted into droplet velocity, and 2}
determination of the aerodynamic drag on the droplets as they simultaneously
evaporate and burn. While such calculations cannot be performed to a great
degree of accuracy, the estimation indicates that the threshold diameter for
stri ki ng the wall is approximately 200-300 microns. This is consistent with
the shift in DRE behavior associated with the spray degradation and it
indicates the following methodology for evaluating the atomization adequacy
of full-scale nozzles:
• Evaluate atomization quality in cold flow on either the actual
waste stream or on a surrogate stream of identical properties. If
possible, both dropsize and droplet velocity information should be
obtained.
3-21
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• Use the spray information to eval uate the adequacy of the match
between the combustion chamber and the spray pattern.
The manner in which the spray data would be used to evaluate the adequacy of
the match is not well defined at this juncture, but a general direction is
clear. The spray data, in conjunction with the fuel properties, will allow a
characteristic evaporation time to be calculated. Second, the nozzle design
and incinerator dimensions will then provide the necessary input to establish
whether (1) droplets directed toward the walls will contact the walls, and
{2) droplets directed away from the walls will evaporate. Required to
establish this information is a characteristic time model for representative
incinerator designs.
Areas of research evolve from this scenario in order to delineate a
realistic and practical evaluation methodology:
1.	The need for droplet velocity information in addition to droplet
SMD measurements needs to be ascertained. The methodology will
benefit if simple Sauter Mean Diameter (SMD) data, coupled perhaps
with spray angle, are sufficient,
2.	The relationship between droplet size and fuel properties on one
hand, and characteristic evaporation times on the other, must be
established for fuels representative of incineration feed stock.
3.	A major effort is required to apply and test characteristic time
modeling to the configuration and conditions representative of
practical incinerators.
4.	Finally, the methodology established must be tested first at bench
scale, and then in scaled units.
3-22
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4.0	SECONDARY ATOMIZATION
This section describes the experiments which evaluated the effect of
secondary atomization on ORE in subscal e liquid spray combustion.
4.1	Background and Objectives
One mechanism by which a liquid injection incinerator could fail to
achieve acceptable ORE is through poor liquid atomization quality. As
discussed in the previous section, poor atomization can cause a significant
fraction of the liquid to appear as large droplets. These can penetrate to
the wall or pass through the flame without completely evaporating.
Droplets that reach the wall would eventually evaporate at lower
temperature than the bulk of the incinerator. Thus, the probability of
destruction is reduced. Also, droplets that penetrate the flame-zone would
release waste that would experience a lower temperature path than those
released in the flame.
Large-scale atomizers generally fail to provide acceptable atomization
quality for two reasons:
1.	The liquid is unusually viscous or it contains sol ids (i.e.,
si urry) . This can be a particular problem for waste streams
because of the wide variability in liquid properties and solid
loadings that occur.
2.	Portions of the nozzle have degraded during use such that design
operation cannot be obtained. Again, this problem can be
aggravated for waste streams thro ugh their corrosive nature or
because of their solid content.
In either case secondary atomization has been proposed as a means of reducing
droplet size after the liquid has left the nozzle.
4-1
-------
Secondary atomization is the term used to describe the fragmentation of
droplets in hot gas due to internal generation of vapor. The presently
accepted mechanism is as follows [27]:
•	The process requires at least two miscible components of
substantially varying vol atility to be present.
•	During heating the volatile component is depleted in a thin
boundary layer at the droplet surface.
•	Since the low volatility compound predominates at the droplet
surface, the surface temperature rises to a value approximating the
boiling point of the low-volatility compound.
•	Heat transfer from the surface to the interior is much more rapid
than mass transfer of the light component from the interior to the
surface. Thus, the temperature of the droplet interior may reach
the point where fractionation occurs at the droplet center.
•	As the fractionation process generates gas the droplet expands into
a bubble which in due course ruptures. The rupture shatters the
droplet into many small fragments.
Thus, the objective of inducing secondary atomization is to cause large
droplets to break into small ones in the flame, and thereby to increase the
net evaporation rate. If poor atomization or droplet penetration to the wall
is the process limiting DRE in a particular application, then secondary
atomization may be a tool to improve DRE. The volatile component could be
obtained by selective blending of waste streams or by the addition of a pure
volatile compound. Note that we do not consider here the use of emulsions,
which is an approach that has been investigated elsewhere [28].
The objective of the work reported here was to examine the potential
that high concentrations of volatile waste compounds in an auxiliary fuel can
promote secondary atomization. A second objective was to demonstrate whether
4-2
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and compare the ORE In the small-scale reactor for conditions where secondary
atomization was present against conditions for which it did not occur. These
tests were performed under a previously characterized atomization failure
mode.
4.2	Approach
The first objective was addressed by investigating the effect of
compound concentration in No. 2 fuel oil auxiliary fuel on secondary
atomization. A series of compounds were selected for screening to represent
a range of boiling points from very volatile to numbers representative of
No. 2 fuel oil (210-26Q°C). The screening tests were performed in the slip
flow reactor (Section 2.1). The following compounds were selected for
screening (shown with their normal boiling points):
•
Dichloromethane:
39°C
•
Acryloni trile:
77°C
•
Benzene:
30°C
•
Isopropanol:
82 °C
t
Benzal Chloride:
205°C
The procedure was to first demonstrate that the compounds were miscible with
No. 2 fuel oil up to 40 percent by weight. Second, the degree of secondary
atomization was estimated visually and assigned a value in an approach
similar to that used at Princeton [29], Third, the degree of secondary
atomization was evaluated as a function of compound concentration. Fourth, a
series of tests were performed in the turbulent flame reactor in which DRE
for the various compounds was compared under conditions where the degree of
secondary atomization was known from the slip-flow screening studies.
4-3
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4.3
Results and Discussion
Each of the compounds was screened in the slip reactor at 0.5, 2, 5, 10,
20, and 40 weight percent in the No. 2 fuel oil. The results are presented
graphically in Figure 4-1 as a plot of secondary atomization intensity vs.
concentration for each of the compounds. The results indicate:
•	Secondary atomization is active only for compound concentrations
above 2 percent, except for isopropanol, which showed some activity
at 2 percent but none at 0.5 percent.
§ For any secondary atomization to occur there must be some
difference between the boiling points of the constituents. For
example, benzal chloride, which has a boiling point comparable with
that of No. 2 fuel oil, showed no activity at any concentration.
•	The resul ts i ndicate that in ten si ty is not enti rely a function of
boiling point differential. For example, i sopropanol has a boiling
point of 82°C, but it induced a substantially more active reaction
than dichloromethane (39°C). Thus, other factors than boiling
point differential (e.g. compound polarity) are related to
intensity.
To ensure that the disruption observed in the reactor was indeed secondary
atomization, high-magnification shadow photographs were employed to visualize
the process. An example of this process is shown in Figure 4-2. Sequence
1-4 shows the droplet, here a double drop, expanding into a bubble. Between
sequences 4 and 5 (approximately 0.2 msec) the bubble has ruptured and the
remaining liquid has started to disperse as small droplets.
Based on these results, two compounds were selected for testing in the
turbulent flame reactor: isopropanol and benzal chloride.
The objective of the turbulent flame reactor work was to evaluate
whether secondary atomization intensity, as determined in the siip flow
4-4
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Waste Concentration in Fuel
Figure 4-1. Effect of waste concentration on
secondary atomization intensity.
4-5
-------
Figure 4-2. Sequence of Hycam pictures showing expansion
and rupture of fuel droplet.
4-6
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screening experiments, could directly affect ORE. The experimental matrix is
shown in Table 4-1. The experiments were designed to determine the effect of
compound concentration on DRE for 1) a compound for which no secondary
atomization occurs across the entire concentration range, and 2) a compound
for which no secondary atomization occurs at low concentrations, but a
strong response is obtained at high concentrations. Thus, the first compound
yields the concentration dependence in the absence of secondary atomization.
Any strong additional concentration dependence for the second compound can be
a ttr i b uted to an i nc rease in secondary atomization i ntensi ty with
concentration.
TABLE 4-1. SECONDARY ATOMIZATION TESTS IN THE
TURBULENT FLAME REACTOR1
Compound
Concentration

0.5
2.0
10.0
Benzal Chloride
Isopropanol
None
None
None
Sporadic
None
Viol ent
^Labels in table are intensity of secondary atomi-
zation from slip flow screening.
The test condition corresponded to the off-design atomization condition
illustrated in Figure 3-10. In all other respects, the TFR was set for high
efficiency operation (120 percent theoretical air, 0.8 swirl number). Thus,
the only variables were test compound type and concentration.
The results for DRE of the test compounds are shown in Figure 4-3.
Waste penetration {fraction of original waste escaping the reactor) is
plotted against the percent waste in the fuel for the two test compounds.
Benzal chloride shows an approximately one order of magnitude decrease in
penetration between 0.5 and 10 percent waste concentration. Since no
-------
+¦>
u
m
s_
10
10"
¦o
OJ
&
i.
+J

V)
10
-4
10
s
Benzal Chloride
Isopropanol
s
Below Detection
, . t
0.5	. 1
Waste in Fuel (percent)
10
Figure 4-3. Comparison of compound penetration for benzal chloride and isopropanol as a function
of compound concentration in the auxiliary fuel. Results are for the turbulent flame
reactor operating under an atomization failure condition.
-------
secondary atomization takes place for this compound, the concentration effect
on penetration must be due to other factors (see discussion in Section S).
For isopropanol, however, the effeet of concentration is much more
pronounced. Between 0.5 and 10 percent concentration DRE improves from less
than 99.9 percent to greater than 99.9999 percent. Significantly, this
increase in DRE occurs concurrently with an increase in secondary atomization
intensi ty from none to violent. Thus, at least a substantial portion of the
difference in behavior between benzal chloride and isopropanol can be
attributed to the secondary atomization behavior of isopropanol.
In addition to DRE, the overall combustion efficiency was also
influenced by the dopants. Figure 4-4 shows CO and hydrocarbon emissions (as
measured by a flame ionization detector) as a function of dopant concen-
trations. Note that dopant type and concentration were the only variables;
in all other respects each of the experiments were identical. The data show
three key points:
1.	Increased isopropanol concentration increased combustion
efficiency. This occurred concurrently with increased secondary
atomization intensity.
2.	Increased benzal chloride concentration decreased combusti on
efficiency. This occurred in the absence of secondary atomization.
3.	At 0.5 percent, where secondary atomization was absent for both
compounds, the CO and hydrocarbon emissions for the benzal chloride
were lower than those for isopropanol.
This final point shows that other dopant dependent mechanisms than secondary
atomization are active.
Thi s work suggests that the DRE of liquid injection incinerators
operating under atomizer limited conditions can be improved by the blending
of small amounts of high volatility liquids into the waste stream. The
blending agent may be a second waste stream of markedly different vol atility
4-9
-------
o
o
200
150
100
50
i	r-"i1 'i
i > i i | f ¦¦
—r i i i i ri j 	

Isopropanol
Dopant
%o.


A

//'
...... i	

, L±J	1	

0 		_ 	
0,1	1.0	10.0
Waste Feed Concentration {Percent)
200
150
a.
ex
o
X
o
o
100
50

	y	i 	i—
Benzal Chloride Dopant

-o



1 1, 1 1 1 1 1 1,1 .1 II
—«—»	i -i—t- I-	
0.1	1.0	10.0
Waste Feed Concentration (Percent)
gure 4-4. Emissions of CO and hydrocarbons from the TFR
as a function of dopant concentration.
4-10
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rather than a pure organic liquid. These blending agents may be particularly
appropriate for slurry atomization, whose primary atomization quality is
usually limited.
4.4	Implications and Concl usions
Secondary atomization has been demonstrated to be a potential means of
improving incineration efficiency in situations where ORE is dominated by
atomization quality. For secondary atomization to occur, the following
requirements must be met:
§ A component of high volatility, relative to the fuel, must be
present. The component may be either miscible with the fuel or it
may be present as an emulsion.
•	The volatile component must be present at a concentration of the
order of at least a percent.
•	The volatile component may be either a waste compound or a
nonhazardous blending agent. In other studies water emulsions have
been used in place of a miscible volatile component.
During field testing, researchers at Midwest Research Institute [29]
noted that DRE appeared to be correlated with waste concentration in the
feed. Figure 4-5 shows a compilation of the field data for a series of units
and tests. The plot shows considerable scatter, as might be expected from
the large number of test conditions and devices included; however, the trend
of increasing DRE with increasing waste concentration in the feed is both
readily apparent and statistically significant. One explanation offered for
this behavior is that as waste concentration increases, secondary atomization
becomes more active and DRE is improved. One problem with this scenario is
that below approximately 1.0 percent concentrations volatile additives do not
induce secondary atomization. If secondary atomization is not active over
the left two-thirds of the plot, then this scenario cannot explain all of the
correlation {see further discussion in Section 5).
4-11
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10'
10"
10'
o _4
e io
o
10'
10
10"
"I I I n |
"i—i i mi f	r-r i mtr;	1—i'1'i 11'm
A AT
A


A
/^\A
A*	A
a
% &
^	A A
A^ ^A
-I	I 1 11 I I I	1 » i ' i <» <	2—1	1.1 t.U,	I	' i i n " I
10	100	1000	10,000
Waste Feed Concentration (Mass ppm)
100,000 nrorooo
Figure 4-5. Waste penetration as a function of waste concentration
in the feed stream. Data compiled for many units and
operating conditions [29].
4-12
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The results presented here have shown the following;
•	Secondary atomization is a potential means of improving DRE in
situations where DRE is limited by liquid atomization.
¦„
•	To first order secondary atomization potential can be predicted
from boiling point differential between the fuel and the waste.
•	A minimum of approximately 1.0 percent of the volatile component is
necessary to induce secondary atomization.
•	Use of emulsified water as a volatile agent may be feasible for
those cases in which streams contain waste concentrations too low
to induce secondary atomization.
•	Secondary atomization is probably not the phenomena that controls
the DRE vs. waste concentration correlation noted from the MR I
field data.
4-13
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5.0	EFFECT OF COMPOUND CONCENTRATION ON DRE
5.1	Background and Objectives
Under EPA contract Midwest Research Institute (MRI) and others have
performed extensive field tests on a wide variety of practical incineration
devices. The objectives of these tests were to characterize the waste
destruction performance of present incineration technology and to determine
if any common factors correlate waste destruction among full-scale units. To
address the second objective MRI has performed an extensive statistical
treatment of their data [29]. The most significant statistical correlation
found was the relationship between waste penetration {= 1 - DRE/100) and
waste concentration in the original feed stream. This relationship is shown
in Figure 5-1. The data show considerable scatter; this is not surprising
since the data represent a wide variety of operating conditions, wastes, and
incinerator designs. Nonetheless, the correlation (represented by the solid
line on the figure) is statistically significant.
One item of significance is that all points above the horizontal dashed
line represent noncompliance under the 99.99 percent DRE rule. These results
indicate that current technology has difficulty meeting the licensing
regulations when the waste represents less than 1000 ppm of the feed stream.
This finding has significance with respect to waste streams contaminated by
low coricentrati ons of extremely hazardous materials (e.g. dioxin or
chlorophenol contaminated pesticides).
A second significant point deals with the mechanisms that give rise to
the correl ati on. In any thermal destruction device a certain amount of the
waste feed compounds escape destruction. The understanding of this mechanism
or mechanisms is critical to:
• The design and modification of incinerators for improved
efficiency.
5-1
-------
Figure 5-1. Waste penetration as a function of waste concentration
in the feed stream. Data compiled for many units and
operating conditions [29].
,5-2
-------
•	The understanding of what makes one compound more difficult to
destroy than another in identical circumstances,
•	The understanding of the portions of the incinerator environment
that dominate PIC formation. This is necessary to design
appropriate laboratory- or sub-scale experiments to develop waste-
Pi C chemistry.
The field correlation shown in Figure 5-1 provides insight into the mechanism
controlling waste release; whatever model is developed, it must be constant
with the behavior of the correlation.
The objective of the work performed in the present study was to
determine whether the correlation is representative of flame-zone or post-
flame behavior. The specific issues addressed are:
•	Does a laboratory scale spray flame reproduce the DRE vs. waste
feed concentration behavior noted in the field tests?
•	Is the correlation characteristic of high efficiency operation or
failure mode operation?
•	Do compound rankings change as waste feed concentration changes.
5.2 Approach
The approach was to use the turbulent flame reactor to simulate the
processes occurring in incinerator flame zones. As described in Section 2,
this unit is a 100,000 Btu/hr liquid spray combustor. The flame zone is
surrounded by water-cooled walls to minimize post-flame reactions. In the
present application the reactor was fired on No. 2 fuel oil doped with a
soup of five test compounds. The soup was an equilimolar mixture of
acrylonitrile, benzene, chlorobenzene, chloroform, and 1,1,1-trichloroethane.
The soup was added to the No. 2 fuel oil in various concentrations for
testing: 30 ppm, 300 ppm, 3000 ppm, 3 percent, and 30 percent by weight.
5-3
-------
Two reactor conditions were selected for testing. Both utilized on-
design atomization. The two conditions were selected to represent the lean
and rich limits of the high efficiency window. These limits were defined by
the points where CO started to rise. The stoichiometrics selected were 120
and 150 percent theoretical air.
The sampling and analysis system used in these tests is described in
detail in Appendix A. The estimated measurement system detection limit
corresponds to 99.999 DE (C/C0 = 10"^) with 3 percent soup in the No. 2 fuel
oil. Thus, at a 30 ppm soup doping level, the detection limit is C/Co = 10~2
while C/C0 = 10""6 is detectable with 30 percent soup doping.
5.3	Results and Discussion
The DRE results for four of the compounds at 150 percent theoretical air
are shown in Figure 5-2. This stoichiometry represents a point on the fuel -
lean side of the high efficiency "window." The results indicate that DRE
values generally increase at higher waste concentrations for each of the
compounds. No values for chloroform are shown because for several cases
chloroform recoveries exceeded the feed stream value; i.e., the reactor was a
net chloroform producer. Note that the effect of waste concentration on DRE
is less apparent if only the points at 3000 ppm waste concentration and below
are considered. Conversely, if only the points at and above 3000 ppm are
considered the dependence is amplified. One conceivable explanation is that
secondary atomization is the phenomena responsible for the improvement in DRE
in the turbulent flame reactor. As discussed in Section 4, secondary
atomi zati on would be expected to be active only above ca. 1 percent waste
concentration. Thus, if secondary atomization is responsible, the waste
concentration parameter would influence DRE much move strongly above 10,000
ppm waste.
Similar resul ts for 120 percent theoretical air are shown in Figure 5-3.
These data are for a point on the fuel-rich side of the high efficiency
window. The results again indicate an increase in DRE with waste concen-
tration in the feed. The effect is more pronounced at and above 3000 ppm
5-4
-------
Figure 5-2. Waste penetration as a function of waste concentration
,in the feed stream for 150 percent theoretical air.
5-5
-------
WASTE FEED CONCENTRATION {VOLUME PPM)
Figure 5-3. Waste penetration as a function of waste concentration
in the feed stream for 120 percent theoretical air.
5-6
-------
waste than at and below 3000 ppm. The ORE values are generally similar
between Figures 5-2 and 5-3, within the scatter associated with this type of
measurement. However, below 300 ppm waste feed concentration the 150 percent
theoretical air case showed approximately an order of magnitude higher
emi ssions.
The excess chloroform measurements that were noted in both data sets can
be attributed to formation of chloroform as a PIC during reaction of the
other compounds. Indeed, chloroform has been identified as a common PIC in
field testing [26] and in pilot scale waste studies [27]. Its appearance is
generally associated with wastes containing chlorine-substituted light hydro-
carbons,such as 1,1,1-trichloroethane in the present experiments. To prove
that chloroform was a PIC in the current experiment the 150 percent theore-
tical air case was repeated without the chloroform in the feed. The DRE
results for this repeat are shown in Figure 5-4. With a few exceptions the
data repeat the general trends noted in Figure 5-2:
•	Moderate influence of waste concentrations in the feed on DRE below
3000 ppm.
•	Strong influence above 3000 ppm.
Furthermore, chloroform was tentatively identified in the reaction products
(based on retention time) at concentrations comparable with those observed in
the original (Figure 5-2) test.
The effect of changing the auxiliary fuel is shown in Figure 5-5 where
heptane replaced No. 2 fuel oil. The results indicate a less strong effect
of waste concentration on DRE. If secondary atomization is accepted as the
cause of the DRE variation in the previous data, then it follows that a more
volatile auxiliary fuel (heptane) will reduce droplet disruption intensity.
Thus, secondary atomization would be minimized and DRE would show less
variation with compound concentration.
5-7
-------
Waste Feed Concentration (ppm)
Figure 5-4. Waste penetration as a function of waste concentration
in the feed stream for 150 percent theoretical air.
Repeat data without chloroform feed.
5-8
-------
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i r n i rt-|	1—r—rrrm j
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o
—j—i iTinrp
—1—1 I 11 11 |	1—1 T 1 III -
10"1




o :
10"2

 \
O Acrylonitrile
V A
O

o 10"3
o
o
—
~ 1,1,1-TrichloroethaneX^
0 \
A Toluene v \
O Chlorobenzene
A
Estimated Sampling -
and Analysis System I
Detection Limit
Kf4
-



a :
10"5
— '




10"6
	1	
» i mill _i i i Mini
	r_iJ_uuuL
I 1 Mlllll
1 1 1 llll 1 \ 1 1 llll
10	100	1000 10,000 100,000 1,000,000
Waste Feed Concentration (ppm)
Figure 5-5. Waste penetration as a function of waste concentration
in the feed stream for heptane auxiliary fuel at 150
percent theoretical air. Chloroform not in feed stream.
5-9
-------
5.4
Implications and Conclusions
These experiments were performed to determine if the flame zone of a
liquid injection incinerator (specifically a sub-scale flame zone) would be
capable of reproducing the correlation developed from the MR I field data
(Figure 5-1). The objective was to provide insight into a phenomena that is
directly related to the processes that limit the DRE of field units. The
resul ts of thi s study show that DRE is positively correl ated with waste
concentration in the feed. The results al so suggest that the infl uence of
feed concentration is stronger at waste concentrations above 3000 ppm than
below, a trend not noted in the field data.
An examination of Figures 5-2 through 5-5 shows that no single compound
order predicts the relative destruction of the test compounds. Table 5-1
shows the number of times each compound appeared in the indicated ranking by
emission concentration. Acrylonitri 1 e is generally the most refractory
compound and chlorobenzene the easiest compound to destroy. In the previous
lab-scale incineration study [7] chloroform was the most refractory compound;
here chloroform was shown to be a PIC formed from the other chlorinated
compounds.
TABLE 5-1. CONCENTRATION RANKINGS' RANK BY
HIGHEST EMISSIONS1

1
2
3
4
Acryl oni trile
10
5
3
0
1,1,1-Trichloroethane
5
3
5
5
To!uene
1
8
6
3
Chiorobenzene
2
2
4
10
^Number of times each compound appeared in the
indicated ranking.
5-10
-------
As briefly mentioned above, one explanation for the observed behavior is
that secondary atomization becomes active at high waste concentrations. In
Section 4 we described how No. 2 fuel oil, when doped with volatile waste
compounds, can undergo in-flame droplet fragmentation due to the vaporization
of the more volatile components within the droplet. This fragmentation was
shown to result in improved DRE. In the present experiment the improvement
in DRE was associated with waste feed concentrations where secondary
atomization would become active (>3000 ppm). Also, because of its higher
volatility, heptane auxiliary fuel would not be expected to undergo secondary
atomization when doped with the present waste compounds. The corresponding
DRE data, Figure 5-5, do not indicate a consistent improvement in DRE with
waste concentration at and above 3000 ppm. Thus, the heptane results are not
inconsistent with a secondary atomization hypothesis.
The field data correlation, Figure 5-1, cannot be supported in its
entirety by a secondary atomization hypothesis. For waste feed concen-
trations below ca. 0.5 percent no secondary atomization would be expected.
Thus, the substantial variation of DRE below this inlet concentration must be
explained in another way.
A first step toward investigating the cause of the correlation is to
evaluate attributes of the correlation. In Figure 5-1, the diagonal dashed
line corresponds to a constant waste emission concentration of about 7 ppt.
If this dashed line can be assumed to represent the true correlation, and the
scatter to represent nonrelated second-order effects, then the conclusion is
that the waste emission concentration of field units is, to first order,
constant. Thus, waste emission concentrations are relatively independent of
feed concentrations over many orders of magnitude.
One approach is to determine what form of reaction kinetics are con-
sistent with this correlation. If a general reaction rate expression is
assumed:
dC/dt = -KCn	(5-1)
5-11
-------
where C = waste concentration
K = rate constant
n = order of reaction with respect to C.
Equation 5.1 is integrated to yield
Cf/C0 = exp(-kt) for n = 1
(5-2)
1-n 1-n
Cf - C0 = (n-l)kt for n = 1
where Cf, C0 = final and initial waste concentrations
t = total time
Under no circumstances are Cf and Co independent for fixed k and t. Thus,
simple kinetics do not reproduce the correlation behavior.
A situation in which Cf would be independent of C0 is under thermo-
chemical equilibrium. As long as the overall elemental input (e.g. C, H, CI,
0, etc.) is relatively invarient the final concentrations will be independent
of the initial feed concentrations. However, for an overall fuel-lean
incinerator, equilibrium predicts essentially zero emissions of waste type
compounds. However, a real incinerator is not at a single uniform
stoichiometry but rather a range of stoichiometrics that varies from fuel-
lean to fuel-rich. A hypothetical distribution is plotted in Figure 5-6A,
which shows that even though the overall stoichiometry is fuel-lean, a
certain local fraction of the gas is fuel-rich due to mixing limitations in
large scale equipment. Figure 5-6b shows equilibrium benzene concentration
as a function of stoichiometry, (Here benzene is used as a simple example of
how most waste compounds would be expected to behave). Thus, extremely fuel -
rich pockets would have high benzene concentrations independent of residence
time or initial feed composition. Figure 5-6c is a conceptual product of
Figures 5-6a and 5-6b which shows the mass of benzene emitted as a function
of equivalence ratio. It represents a "window" in which benzene
concentrations are high and, simultaneously, a significant amount of gas
exi sts with its stoichiometry in the window. The total emissions of benzene
would be obtained by integrating Figure 5~6c over stoic hi ometry. The
5-12
i
-------
1000
E
CL
Cl
-------
concentration obtained by this procedure has the required property that it is
invarient with feed concentration.
The results of the work reported in thi s section lead to the fol lowing
conclusions:
•	Waste DRE in subscal e flame zones is increased as waste
concentration in the feed fuel is increased.
t Waste DRE shows the most dramatic increase at high waste
concentrations. This suggests that secondary atomization may be
the cause of the correlation in the turbulent flame reactor.
•	Secondary atomization is probably not the phenomena responsible for
all of the DRE variation with waste concentration in the field
tests. This is because the tests show that the variation continues
at waste concentrations where secondary atomization should not
occur.
•	Application of thermochemical equilibrium to a mixing limited
incinerator yields a model that reproduces the first order aspects
of the field data; i.e., emission concentrations independent of
feed concentrations.
5-14
-------
6.0	PIC FORMATION
This task was directed towards determination of the identify and source
of hazardous PIC1s observed in the exhaust of the turbulent flame reactor.
6.1	Background and Objectives
One present concern for application of incineration technology is that
the hazard associated with a waste stream may not be removed even though the
original waste compounds are destroyed. Transformation of the waste into
hazardous products of incomplete combustion (PICs) can potentially aggravate
the hazard associated with the waste stream. For example, a hazardous but
nontoxic waste can be partially transformed into chlorinated dibenzo-p-
dioxins or dibenzofurans upon incineration.
Field testing on practical incinerators burning multicomponent wastes
have shown that a broad distribution of Appendix VIII compounds appear in the
exhaust [30]. The key question is defining the source of these emissions.
Potential ^sources are [30]:
•	Chemical transformation of waste to PIC.
•	Low DRE of Appendix VIII compounds present in trace amounts in the
auxiliary fuel.
•	Transformation of nonhazardous auxiliary fuel constituents into
Appendix VIII compounds.
•	Trace Appendix VIII compounds in the combustion air that experience
a low DRE in the flame.
•	Trace Appendix VIII compounds that are stripped from the scrubber
water.
6-1
-------
Midwest Research [30] examined each of these but was unable to conclude which
mechanism was dominant. Limited characterization of feed and waste streams
indicated that certain PICs could be accounted for by each of the mechanisms.
The turbulent flame reactor {TFR) testing described in the preceding
sections offers an opportunity to examine certain of the mechanisms. The
advantages of the TFR for these tests include:
•	The waste stream is well characterized because it is made up of
mixtures of reagent grade chemicals. Field testing rarely involves
pure or well characterized waste streams.
•	The auxiliary fuel stream is al so a well characterized light
hydrocarbon liquid {No. 2 fuel oil).
•	The TFR does not use a scrubber before the sample point. Thus,
PICs observed in the sample stream cannot arise from the scrubber
water.
The TFR affords the opportunity to examine trace organic product
distributions from a turbulent spray flame whose feed streams are better
characterized than field incinerators.
The objective of this task was to complement the data generated in the
previous sections by examining, for a limited number of conditions, the full
organic product distribution in addition to the DRE data. The key questions
were;
•	Are waste-PIC relationships readily identifiable.
•	Can PICs be attributed to the fuel oil.
6.2	Approach
For these tests the TFR was operated on No. 2 fuel oil doped with
2-chlorophenol. Only a single doping compound was used to avoid ambiguity in
6-2
-------
determining waste-PIC relationships. The TFR was operated over a range of
stoichiometries for what was otherwise a high-efficiency condition. Samples
were obtained in the usual manner using the VOST train. Those samples were
analyzed by GC-MS by S-CUBED. Each Tenax trap was connected and desorbed at
175°C for 10 min. onto a liquid nitrogen cooled cryo trap. 1.2 g of gaseous
external standard perfluorotholuene (PFT) was then injected onto the trap.
The cryo trap temperature was raised to 220°C and the data acquisition
started. A GC program of 40° (10 rain) -285°C at 8°/min was used with a 30 m,
DB-5, narrow bore, then film fused si1ica capillary column. The mass
spectrometer was repetitively scanned from 35-400 AMU in 1 sec.
This provides an identification and a semiquantitative analysis of all
of the volatile organic compounds present on the traps. Low volatil ity
compounds (generally polycyclic hydrocarbons) cannot be recovered by this
procedure.
6.3	Results and Discussion
The conditions under which samples were obtained for GC-MS analysis are
detailed in Figure 6-1. The CO and total hydrocarbon traces indicate the
presence of the high efficiency window. Sample A was obtained under an
apparent fuel-rich failure condition. Sample B and C were obtained at
the fuel-rich and fuel-lean extremes of the high temperature window,
respectively.
A total of 57 organic species were identified In the three pairs of
traps. Thus, a large number of product species were identified in the
combustion products of a simple hydrocarbon fuel and a single Appendix VIII
compound. However, examination of the concentrations of the various products
showed that only a small number of compounds were present under all
conditions at relatively high concentrations. These are listed in Table 6-1.
(For reference, note that zero ORE of the chlorophenol would result in 163
micrograms/gm and 99.99 percent ORE would result in 0.0163 micrograms/gm).
Thus, each of these compound concentrations is substantially above what would
be the legal limit for chlorophenol release if it had been a P0HC. The
6-3
-------
9000
50 60 70 80 90 100 110 120 130 140 150 160 170
Percent Theoretical Air
Figure 6-1. Variation of CO and total hydrocarbons with stoichiometry.
Points A, B, and C indicate the conditions under which
GC-MS analysis was obtained.
6-4
-------
appearance of thiophene Is key to understanding the source of these
hydrocarbon compounds. Since the doped waste compound is sulfur free, the
thiophene must have appeared from the sulfur in the fuel oil. Since No. 2
fuel oil can, by code, contain up to 1 percent sulfur, the thiophene would be
the resul t of unreacted organic sulfur escapi ng destruction or reformed
sulfur that avoids oxidation to SO2. By inference, the other compounds shown
in Table 6-1 can be assigned as unreacted No. 2 fuel oil constituents or
products of partial decomposition of the fuel oil.
TABLE 6-1. MAJOR COMPONENTS IDENTIFIED

Sample
ABC
Methyl Ethyl Ketone
To!uene
Benzene
Thiophene
Methylene Chloride
Dimethyl Benzene
.75 1.7 2.9
0.10 0.39 0.93
0.13 * *
0.020 0.027 0.019
0.053 * *
0.0042 0.024 0.24
^Interference with MEK, cannot be quantified.
Values are micrograms/gram flue gas.
Table 6-2 shows the remaining compounds identified on the cartridges.
In general these compounds are present in very small concentrations {see
Appendix C). The series of fluorenated hydrocarbons cannot be attributed to
the doped waste compound and thus must originate from the fuel or the
combustion air. Thus, at least a large fraction of these trace organic
compounds arise from the combustion of the light fuel oil rather than the
partial reaction of any Appendix VIII compounds.
6-5
-------
TABLE 6-2. TRACE ORGANIC COMPOUNDS FOUND
ON THE CARTRIDGES
methane,
benzene,
dichlorodifluoro
ethylmethyl***
ms thane.
2-methylundecane
trichlorofluoro
benzene,
ethane.
ethyl dimethyl***
1,1,2-trichloro
2-methy1naphthalene
1,2,2-trifluoro
naphthalene
1-butanol
trimethyl benzene***
ethene, trjchloro
cyclohexane,
2-pentene-3,4-dimethyl
ethylmethyl***
cyclohexane, methyl
2-propyl-l-heptanol
eye 1opentane,
4-propylheptane
1,2,3-trimethyl
3-methyl -1 -hexene
1-hexene,
nonane
3,5,5-triraethyl
pentalene, octahydro-
2-pentanane, 4-methyl
2-methyl
acetic acid,
isooetanol
propylester
3-methylhexane
cyclopentane
cyclohexane, butyl
1,T,3,4-tetramethyl
1-hexene-
3-methyl heptane
3,5,5-trimethyl
cyclohexane,
cyclohexane,
dimethyl***
diethyl***
benzene,
decene
chloropentafluoro
decane
acetic acid,
decane, 4-methyl
butyTester
benzene,
octane
1,2,3-trichloro
ethene, tetrachloro
cyclopentane,
cyclohexane,
1-methyl-
trimethyl***
2-(2-propenyl)
ehlorobenzene
decahydro naphthalene
ethylbenzene
undecane
2-methylnonane
3-ethylheptane

cyclohexane, propyl

diethyl phthalate
6-6
-------
These results Indicate that organic products of incomplete combustion
can arise from any organic compound present in the feed stream. This is
independent of whether it is associated with the hazardous waste compound,
the non-Appendix VIII compounds in the waste, the auxiliary fuel, or the
burner air. Thus, conditions which maximize the combustion efficiency {of all
organic compounds into CO2) will be expected to minimize the formation of
hazardous compounds, independent of source. The field tests [30] generally
indicated that waste and PIC emissions were of similar magnitude for each
facility. The PIC concentrations measured in this study were also generally
comparable with the waste compound emissions found in the previously
described tests. Thus, it appears that combustion efficiency, DRE, and PIC
destruction are coupled problems, and that design and operating criteria that
address one aspect will address each.
6-7
-------
7.0	CONCLUSIONS
This study was directed toward a series of related fundamental issues in
the hazardous waste area. The principal conclusions are summarized as
fol 1 ows:
1.	Waste Atomization: For degraded atomizers, the principal cause of
poor waste destruction efficiency is the increase in the fraction of very
large droplets. The extreme delay in evaporation associated with these large
droplets can lead to unreacted material reaching the wall or penetrating
through the flame zone. Design to avoid this behavior is more difficult for
hazardous waste incineration than for conventional combustors because:
t A large amount of empirical experience has been obtained on liquid
fuel combustion.
• The atomization properties of waste streams (viscosity, surface
tension, presence of solids) can vary considerably.
The results suggest a design methodol ogy in which atomization quality is
directly measured in cold flow. The size and trajectory of the largest
droplets are compared to the combustion chamber geometry to determine the
initial suitability of the design.
2.	Secondary Atomization: Some materials may have sufficiently
poor atomization properties to prevent acceptable spray fineness at any
conditions. The use of a volatile waste dopant was shown to induce in-flame
droplet fragmentation and to improve DRE. This suggests that use of volatile
dopants, or the blending of different waste streams can be used to avoid poor
DRE due to penetration of large droplets through the flame.
3.	Compound Concentrati on : Field test data show a remarkable
correlation between compound concentration in the feed and DRE. Testing in
the turbulent flame reactor also showed this correlation. However, the
pattern for the subscale flame indicated that secondary atomization was a
7-1
-------
potential cause of the behavior at higher concentrations. This does not
explain the subscale variation of DRE with waste concentration at low waste
concentrations , nor does it fully ex piai n the field data . A mechanism
involving mixing limited equilibrium chemistry was proposed for the field
data.
4. PIC Formation: The yield of trace organic compounds was measured
from the turbulent flame reactor. The results indicated that:
• PIC concentrations were comparable with waste emissions.
t Incomplete combustion of the auxiliary fuel rather than true PICs
from the doped waste dominated the apparent PIC emissions.
Thus, Appendix VIII PICs can arise from any of the hazardous or nonhazardous
constituents of the waste stream or the auxiliary fuel. The implication is
that conditions that promote high combustion efficiency will favor reduced
PIC emission.
7-2
-------
8.1
1
2
3.
4
5
6
7
8
9
10
11
12
REFERENCES
Resources Conservation and Recovery Act, Public Law 94-580, 1976.
Environmental Protection Agency: "Hazardous Waste and Consolidated
Permit Regulations." Federal Register 45:138, July 16, 1980.
Environmental Protection Agency: "Incineration Standards for Owners
and Operators of Hazardous Waste Management Facilities." Federal
Register 46:138, July 16, 1980.
Environmental Protection Agency: "Incineration Standards for Owners
and Operators of Hazardous Waste Management Facilities." Federal
Register 46:264, January 2, 1981.
Trenholm, A. P., P. Gorman, B. Smith, and D. Oberacker: "Emissions
Test Resul ts for a Hazardous Waste Incinerator RIA." Ninth Annual
Research Symposium on Hazardous Waste, EPA-600/9-84-O15, NT IS
PB84-234525, U.S. ~tPA, 19B4, p. THE	
Kraml ich, J. C., M. P. Heap, W. R. Seeker, and G. S. Samuel sen:
"Flame-Mode Destruction of Hazardous Waste Compounds," 20th Int. Symp.
Combust., The Combustion Institute, Pittsburgh, PA, 1985, p. 593.
Kraml ich, J. C., M, P. Heap, J. H. Pohl , E. M. Poncelet, fi. S.
Samuel sen, and W. R. Seeker: Laboratory Scale Flame-Mode Hazardous
Waste Thermal Destruction Research, fc.PA-6U0/2-84-086, Nl" IS PB84-184902,
U.S. tPA, 1984.	
Duval 1 , D. S., and W. A. Rubey: Laboratory Evaluation of the High
Temperature Destruction of Kepone and Related Pesticides"
hPA-600/2-79-299, NI'IS PB 264892;—U.S. EPk, 1979,		
Duval 1 , D. S., and W. A. Rubey: Laboratory Evaluation of High
Temperature Destruction of Poly-chlorinated 8iphenyls and Related
Compounds, EPA-600/2-77-228, OTIS PB 279139, U.S. EPA, 19771	
Dellinger, B., D. S. Duval! , D. L. Hall, and W. A. Rubey: "Laboratory
Determinations of High Temperature Decomposition Behavior of Industrial
Organic Materials," Presented at the 75th Annual Meeting of the APCA,
New Orleans, LA, 1982.
Dellinger, B., J. L. Torres, W. A. Rubey, D. L. Hall, J. L. Graham, and
R. A. Carnes: "Determination of Thermal Stability of Selected
Hazardous Organic Compounds," Hazardous Waste 1:137, 1984.
Lee, K. C., J. L. Hansen, and D. C. Macauley: "Predictive Model of the
Time-Temperature Requirements for Thermal Destruction of Dilute Organic
Vapors," Presented at the 72nd Annual Meeting of the APCA, Cincinnati,
8-1
-------
13.	Lee, K. C., N. Morgan, J, L. Hansen, and G. M. Whipple. "Revised Model
for the Prediction of the Time-Temperature Requirements for Thermal
Destruction of Dilute Organic Vapors and its Usage for Predicting
Compound Destructibil ity," Presented at the 75th Annual Meeting of the
APCA, New Orleans, LA, 1982.
14.	Tsang, W. and W. Shaub: "Chemical Processes in the Incineration of
Hazardous Waste," Presented at the ACS Symposium on Detoxification of
Hazardous Waste, New York, NY, 1981.
15.	Cudahy, J. J. and W. Troxler: "Incineration Character!*sties of RCRA
Listed Hazardous Wastes," Journal of Hazardous Materials 8:59, 1983.
16.	Ni hart, R. K., J. C. Kramlich, G. S. Samuel sen, and W. R. Seeker:
"Continuous Performance Monitoring Techniques for Hazardous Waste
Incinerators," EPA-600/2-89-021, U.S. EPA, May 1989.
17.	Kramlich, J. C., W. D. CI ark, W. R. Seeker, G. S. Samuel sen, and C. C.
Lee: "Engineering Analysis of Hazardous Waste Incineration - Continuous
Monitoring of Incinerators," ASME/AIChE National Heat Transfer
Conference, Denver, CO, 1985.
18.	Clark, W. D., M. P. Heap, W. Richter, and W. R. Seeker: "The
Prediction of Liquid Injection Hazardous Waste Incinerator
Performance22nd National Heat Transfer Conference, Niagara Falls,
NY, 1984.
19.	Bridle, T. R., B. K. Afghan, H. W. Campbell, R. J. Wil sinson, J,
Carron, and A. Sachdev: "The Formation and Fate of PCDD's and PCDF's
During Chiorophenol Combustion," Presented at the 77th Annual Meeting
of the APCA, San Francisco, CA, 1984.
20.	Berglund, R. N. and B. Y. H. Liu. "Generation of Monodisperse Aerosol
Standards." Environ. Sci. Technol. 147, 1973.
21.	Baker, R. J., P. Hutchinson, E. E. Khal il , and J. H. White! aw:
"Measurements of Three Velocity Components in a Model Furnace With and
Without Combustion," 15th Symposium (International) on Combustion,
The Combustion Institute, Pittsburgh, pa, iy/b, p. bw.
22.	Jungclaus, G. A., P, G. Gorman, G. Vaughn, G. W. Scheil , F. J. Bergman,
L. D. Johnson, and D. Friedman: "Development of a Volatile Organic
Sampling Train (VOST)," 9th Research Symposi um on Hazardous Waste,
EPA-600/9-84-015, NTIS PB84-"Z3¥E2b, U.S;~OT7T§84. p. 1.		
23.	Edwards, J. B. Combustion: Formation and Emission of Trace Species,
Ann Arbor Science, Ann Arbor, MI, 1974.	~~
8-2
-------
24 Dietrich, V. E.: "Dropsize Distribution for Various Types of Nozzles."
Proceedings of the 1st International Conference on Liquid Atomization
and bpray Systems, The Fuel Society of Japan, Tokyo, Japan, 1982, p.69.
25.	Jones, A. R.: "Factors Affecting the Performance of Large Swirl
Pressure Jet Atomizers." Report R/M/M1G54, Central Electricity
Generating Board, Marchwood, England, 1978.
26.	Glassman, I.: Combustion. Academic Press, New York, 1977.
27.	Wang, C. H., X. Q. Liu, and C. K. Law: "Combustion and Microexpl osi on
of Freely Falling Mul ticomponent Droplets," Combust. Flame 56:175,
1984.		
28.	Lasheras, J. C., A. C. Fernandez-Pello, and F. L. Dryer. "On the
Disruptive Burning of Free Droplets of Alcohol/n-Paraffin Solutions and
Emulsions," 18th Symposium (International) on Combustion, The
Combustion Institute, Pittsburgh, PA, 1981. p. 293.	"
29.	Trenholm, A., P. Gorman, and G. Jungclaus: Performance Evaluation of
Full Scale Hazardous Waste Incinerators, Volume 2, "LPA-6U0/Z-84-181 b7
1st I IS PBSb-129518, 198C			
30.	Trenholm, A., R. Hathaway, and D. Qberacker: "Products of Incomplete
Combustion from Hazardous Waste Incinerators," 10th Symp. Inciner.
Treatment Hazard. Waste, EPA-500/9-84-022, NT IS PB85-116291, U.S. EPA,
imr.	
8-3
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APPENDIX A
VOLATILE ORGANIC ANALYSIS
The following discussion covers the volatile organic analytical
equipments, the operating techniques, and the quality assurance procedures.
A. 1 Analytical Equipment
Sampling and Adsorption: The equipment is illustrated in Figure A-l.
Exhaust gas samples were collected from the stack by an uncooled 6.35 mm OD
{0.25 in.) stainless steel probe. After leaving the stack, the sample passed
through a glass valve and into the Volatile Organic Sampling Train {VQST).
The sample first passed through a glass coil condenser being cooled by
circulating ice-water before entering the first Tenax cartridge. The
cartridge consists of an 8.5 mm long by 16 mm OD Pyrex trap with 6.35 mm OD
(0.25 in.) ends packed with 2.0 grams of Tenax-GC (40-80 mesh). The
adsorbent was held in place with small plugs of si 1 ianized glass wool. The
sample line was connected to the cartridge by 111 tra-Torr fittings. Next, the
sample passed through a water trap, a straight water-cooled condenser, the
Tenax/charcoal cartridge, and a second water trap. The Tenax/charcoal
cartridge was packed with Tenax-GC {40-80 mesh) and activated charcoal (60-80
mesh) at a 1:1 weight ratio. The sample subsequently passed through a gas
dryer, a rotameter, and a dry test meter. All connections upstream of the
cartridge were either 316 stainless steel or 6.35 mm OD (0.25 in.) Teflon.
Desorption and Analysis: This system is based on a Perkin-Elmer Sigma-2
gas chromatograph with a Sigma-10 Integra tor/data station. The column is a
0.5-long, 3.18 m OD (1/8 in.) Teflon tube packed with Porapak-Q. The 30 cc/
min. carrier gas flow is maintained through the Tenax cartridges, the column,
and the analytical trap, as shown in Figure A-2, 320 Thermal Desorption
System.
• A-1
-------
Figure A-l. Sampling and VOST adsorption system.
-------
N-
CD
r*\
Thermal
Desorptian
Chamber
/
Heated
Line
Flow During
Flaw to
GC/MS
M
o-i
Desorptian
Flow- During
[ Adsorption
1 X 2 X 3 J( 4
»W
•	» i • ~
He or N2
2E-Q
• • <$ • i • »# •
Analytical Trap
with Heating Coil
(0.3cm diameter
by 25 cm long)
Vent
o 3% SP-2100 (1 cm)
© Tenax (7.7cm)
© Silica Gel (7.7cm)
4^	Charcaa i (7.7 cm)
Figure A-2. Schematic diagram of trap desorntion/analysis system.
A-3
-------
A. 2 Operating Procedures
Prior to use, the Tenax arid Tenax/charcoal cartridges are conditioned
under a 20 cc/min. helium flow at 200 C for 45 min. Both before and after
sampling the tube ends are covered with eelophane and refrigerated.
After the reactor condition has been set, the cartridges are placed in
the adsorption train, the cooling water started, and the sample flow opened.
A sample flow rate of 0. 23 1 i ters/mi n. has been found sati sfactory as
discussed below. At the conclusion of sampling (2.3 liters or 10 rain.) the
cartridges are removed and refrigerated.
For analysis, the cartridges are connected into the GC carrier gas line
and placed in the desorption block. The block is raised to 120°C and held at
this temperature for 5 min. This time and temperature have been found
sufficient to quantitatively desorb all of the test compounds. During
desorption the GC oven is maintained at room temperature. As the compounds
desorb, they are collected at the inlet of the GC column. The compounds do
not start to separate at room temperature. If this was not the case, peaks
would be broadened by the deposition of newly desorbed compound at the column
inlet after separation had started. At the conclusion of the 5-min.
desorption time the column was heated to 120°C, held at this temperature for
25 min., and programmed at 5 C/min. to 150°C when samples containing heptane
or No. 2 fuel oil were analyzed; or to 180°C for all other cases. Under the
present analytical conditions, benzene and 1,2-dichloroethane have nearly
identical retention times; because of this, these two compounds were never
used in the same experiment. The FID gas flows were 71 cc/min. hydrogen and
442 cc/min. air.
The integrator output record consists of a chromatogram detector signal
trace and a table of integrated peak areas. The trace is compared with the
tabulated output to insure that the compound peaks are free of interferences,
that the baseline boundary conditions were correctly constructed by the
integrator, and that the presence of unanticipated peaks could be noted.
Integrated peak area values were converted into ppm by use of the calibration
A-4
-------
curves and destruction efficiencies were calculated using the ppm anticipated
from zero destruction operation.
A.3 Calibration Techni ques
Test compound calibration was performed by three independent procedures.
This was necessary to determine the inherent scatter and reproducibility of
the measurement technique, and to locate any of the calibration tests. The
three techniques were 1) direct syringe injection of test compound onto the
GC column; 2) syringe injection of liquid test compound into a Tenax
cartridge, followed by routine desorption and analysis; and 3) preparation of
known standards in a dilution tank, followed by routine Tenax sampling and
analysis of the tank contents.
Direct Column Injection: Liquid samples were withdrawn with a
calibrated microliter syringe and directly injected onto the column through
the injection septa. Chromatograph conditions were identical to the nominal
operating procedure except:
1.	The Tenax cartridge desorber was not a part of the system.
2.	Rather than follow the prescribed oven temperature programming, an
isothermal temperature of 120°C was used (except for chlorobenzene
where 180OC was used {except for chlorobenzene where 180°C was
used}. The isothermal temperatures increased the speed at which
calibrations were performed. The programming is necessary during
normal sampling to separate light hydrocarbons from the earliest
test compound peaks. Because of the linearity of the FID
amplifier, the change in temperature programming does not affect
the calibration.
Table A-l lists the FID mass sensitivity for the various compounds tested,
relative to methane.
A-5
-------
TABLE A-l. MASS SENSITIVITY RELATIVE TO METHANE FOR THE FID
Compound
Relative
Sensi-
tivity
Compound
Relative
Sensi-
tivity
Acrolein
0.54
Ethyl Acrylate
0.59
Phenol
1.22
Hexachlorobenzene
0.0
Benzene
1.23
Toluene
1.09
Carbon Disulfide
0.0
Vinyl Chloride
0.69
Acrylonitrile
0.59
Methyl Ethyl Ketone
0.68
1,1*1-Trichloroethane
1.64
Chlorobenzene
0.77
Chloroform
0.09


-------
Syringe Injection Onto Tenax: Liquid test compound samples were
directly injected onto packed Tenax cartridges. These were subsequently
desorbed and analyzed. The results in the present study were used as an
initial, qualitative test for breakthrough volume. After injection, helium
was drawn through the cartridge. For some tests a second cartridge was
placed behind the first. Breakthrough was detected by the appearance of test
compound on the second cartridge and, simultaneously, by the loss of response
from the analysis of the first cartridge.
Dil ution Tank: Dilute samples were prepared by evacuating an 11-liter
glass tank and injecting a known amount of sample into the tank as the tank
was rapidly repressurized. After repressurization the tank was al 1 owed to
equilibrate and the tank pressure and temperature were noted for calculation
of the correct dilution factor. A portion of the tank contents are pumped
through the Tenax sampling system. The remainder of the calibration test is
identical to the normal sampling and analytical procedure.
A. 4	Calibrations
Figures A-3 through A-6 show sample calibration plots for benzene,
chl orobenzene , chloroform, and ac ryl oni tril e. These are plotted as
microliters liquid vs. integrator peak area. The two plots shown on each
graph correspond to the direct injection calibration and the Tenax
calibration. The close agreement between the two curves for all compounds
indicates that breakthrough volume was not exceeded for these compounds.
Sample Storage: Each SC analysis required about 1.5 hours. Thus, a
backlog of unanalyzed samples occasionally accumulated. In no case was a
sample held longer than 24 hours before analysis. However, a series of tests
were performed to determine if 24 hours was a safe period for storage. A
standard gas was prepared and adsorbed onto two cartridges in parallel. The
first was analyzed immediately and the second was stored under nominal
storage conditions for 24 hours and analyzed. The results, shown in
Table A-2, indicate that no significant loss of test compound occurred. The
scatter between the two analyses is typical of the measurement technique.
-------
Figure A-3, Calibration curves for benzene.
ft-8
-------
i—r
,A Direct Injection
, O Tenax
i—i—ntr
Q)
S-
£-
O
R3
S-
O)
ai
2000
1000
0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0
Volume (microliters)
Figure A-4. Calibration curves for chlorobenzene.
A-9
-------
Figure A-5. Calibration curves for chloroform.
Volume (microliters)
Figure A-6. Calibration curves for acrylonitrile.
A-10
-------
TABLE A-2. SAMPLE STORAGE STARILITY
Compound
Analysis
Inrniediate
(ppm)
After 24
Hours (ppm)
Acrylonitrile
16.7
18.2
Chloroform
16.2
14.5
1,2-Dichloroethane
16.2
16.9
Chlorobenzene .
8.77
9.5
-------
A.5	Uncertainty, Accuracy, and Preci si on
The problem of sample repeatability and precision really involves two
questions:
•	What is the repeatability of the analytical technique given a time-
steady, known concentration to measure?
•	What is the time-steadiness of the experiment, assuming a perfect
measurement technique?
The first of these questions were addressed by the repeated analysis of known
calibration standards. An example of such a series is shown in Table A-3 for
benzene. The resulting standard deviation of the relative error is 2.1
percent. Thus, for the measurements to be accepted at the 90 percent
confidence interval, the relative error is approximately + 4.2 percent. The
error for all calibration data as a group indicated an approximate + percent
at the 90 percent confidence interval for the Tenax procedure.
The data indicate that the inherent time unsteadiness of the experi-
mental DE measurements was small under nonoptimum conditions and substantial
under optimum conditions. For the optimum conditions the observed time
unsteadiness exceeded the +_ 5.0 percent uncertainty associated with the
analytical system and, thus, led to the conclusion that optimum DE
measurements and rankings were random, time unsteady, and were probably
related to the statistical nature of a turbulent flame. However, none of
these optimum data were used to establish the ranking presented in this
study.
Quality Assurance/Quality Control requirements are applicable to this
project. The data contained in this report are NOT supported by QA/QC
documentation as required by the United States Environmental Protection
Agency's Quality Control Policy.
A-12
-------
TABLE A-3, CALIBRATION REPEATABILITY DATA FOR BENZENE
Response (peak area)
Relative
Error
Expected
Obtained
1141
1101
.0350
1141
1126
.0131
1470
1399
.0483
1470
1451
.0129
2129
2090
.0183
2129
2138
-.0042
3118
3146
-.00898
3118
3153
-.0112
3118
3136
-.0058
A-13
-------
APPENDIX B
RAW DATA
This section presents the raw test compound data and describes the
procedure for converting the raw data into destruction efficiencies.
B. 1	Sample Chromatograms
A selection of chromatograms typical of various turbulent flow reactor
conditions are presented below. The numbers printed next to each peak are
the retention times in minutes. A chromatogram typical of high-efficiency
turbulent flow reactor operation is shown in Figure B-l. Few peaks are
present and only one peak, associated with a fuel fragment at 45.18 min. is
measurable. Figure B-2 shows a chromatogram for moderate DE turbulent flow
reactor data. A 1 ow-DE turbulent flow chromatogram is shown in Figure B-3.
At very low efficiency operating conditions significant quantities of fuel
and fuel fragments are released by the flame in addition to the test
compounds. These conditions can result in an uninterpretable chromatogram,
as shown in Figure B-4.
B.2	Calculation Procedures
Chromatogram peak areas were related to moles for each compound by use
of the calibration figures in Appendix A. This number is converted into mole
fraction based on dry gas by:
v (moles compound) (24.63 liters/mole) /r, ,,
Mole Fraction (Dry) = .	(sample volume - liters)	 (B>1)
The mole fraction water in the wet combustion gas is estimated from a
complete combustion model and the dry mole fractions are corrected to burner
(wet) mole fractions by:
Burner Mole Fraction = (Dry Mole Fraction) - (1 - Water Mole Fraction) (B.2)
B-1
-------

Figure B-l. Chromatogram for a high-DE turbulent flow reactor condition.
-------
B-3
-------
T y * I
t « zi-*
B-4
-------

p-

&
*

CO
04
M
v£i
V
S^jaas.-arKSa'SK"~—
-------
The mole fraction of compound that would be present at the burner exit if
efficiency were zero is calculated by:
Zero Efficiency _ (Fuel Flow) (Moles Compound/Mass Fuel)	/R ^
Mole Fraction ~	Total Reactor Molar Flow	1 '
The (Moles Compound/Mass Fuel) is calculated from the fuel mixture
composition and the (Total Reactor Molar Flow) is calculated from the fuel
flow, the air flow, and the complete combustion model. The fraction
unreacted compound is the ratio of Equation B-2 and B-3.
B*3	Raw Data Tables
Tables B-l through B-4 summarize the data from the vol atil e organic
analysis.
B-6
-------
TABLE B-l. PIC FORMATION—TEST CONDITIONS
Sample No.
Compound
SR
CO (ppm)
C02 (%)
02 (%)
HC (ppm Propane)
92802
No. 2 Fuel Oil
0.55
19,000
7.1
7.1
3,750
92705

0.6
4,060
13.7
2.8
60
92)04

0.7
120
131
3.9
50
92701

1.0
no
9.0
8.5
20
92702

1.2
1,970
7.6
10.3
200
92801

1.6
5,350
5.8
12.2
2,000
100102
No. 2 Fuel Oil
+0.1% Chlorophenol
0.55
12,420
5.8
12.8
7,400
100101*

0.65
13,400
11.4
4.4
3,500
92804*

0.7
254
11.6
' 4.9
82.5
92803

1.0
224
8.5
8.9
100
92805*

1.2
302
7.4
10.8
55
92806

1.6
2,400
7.9
10.6
350 J
*GC-MS Data resolved for those conditions.
-------
TABLE B-2. SECONDARY ATOMIZATION
Sample
No.
Compound
Compound
Concentration
CO
(ppra)
C02
tt)
°2
(%)
HC (pprn
Propane)
Destruction
Removal Effi-
ciency {%)
92102
Isopropanol
0.5%
160
11.0
5.8
125
99.821
92103

2.0%
124
12.0
4.7
32
99.998
92104

10.0%
77
12.1
4.8
62.5
100
92502
92503
92602
Benzal Chloride
0.5%
2.0%
10.0%
124
140
165
10.9
11.2
10.5
6.2
5.8
6.2
50
85
110
97.636
98.766
99.781
-------
TABLE B-3. COMPOUND CONCENTRATION DATA {CHLOROFORM FREE)
Sample
No.
Compound
Concentraion
Fuel
CO
(ppm)
CO o
(%)
0?
(%)
HC (ppm
Propane
Destructi
Acrylon-
i tri1e
on Remova
Triehio-
roethane
1 Effici
Toluene
ency (%)
Chioro-
benzene
100401
30 ppm
No. 2
Fuel Oil
240
9.1
9.2
100
94.62
99.999
100.00
i
99.999
100402
300 ppm

-
-
-
-
75.77
89.694
88.90
86.89
100403
3000 ppm

1333
7.2
11.4
500
94.52
99.995
97.75
99.39
J00501
3%

338
9.1
9.1
200
99.912
99.999
99.999
99.999
100502
30 ppm
Heptane
157
7.2
10.7
48
97.07
68.04
71.79
99.999
100503
300 ppm

106
9.0
7.8
70
58.56
99.999
99.29
99.999
100504
3000 ppm

90
9.0
8.0
40
99.999
99.46
99.93
99.999
100506
31

-
-
-
-
95.39
99.999
99.995
99.999
-------
TABLE B-4. COMPOUND CONCENTRATION DATA (WITH CHLOROFORM)







Destruction Removal Efficiency
{%)
Sample
Compound
CD
CO
CO?
°?
HC (ppm





No.
Concentration
Ol\
(ppm)
(S)
(%)
Propane
Acrylon-
Chlor-
Trichlo-

Chloro-






itrile
oform
ethane
To!uene
benzene
91902
30 ppm
1.2
120
11.2
5.0
90
99.236
0.00
99.629
92.710
99.999
91903
30 ppm
1.5
150
9.5
7.6
30
78.62
0.00
81.159
89.764
99.999
91803
300 ppm
1.2
112
9.5
7.5
34
94.081
65.284
99.999
77.84
95.516
91804
300 ppm
1.5
650
8.6
8.9
85
91.595
0.00
47.13
54.309
99.999
91805
3000 ppm
1.2
73
10.4
6.2
28
99.677
88.578
99.993
99.897
99.999
91806
3000 ppm
1.5
1465
7.9
9.5
150
94.597
0.00
91.830
89.457
97.208
91807
3%
1.2
450
11.1
5.0
52
98.954
99.146
99.565
99.007
98.668
91904
3%
1.2
108
12.0
3.9
35
98.984
99.869
99.999
99.996
99.999
91905
3%
1.5
69
9.8
7.2
25
99.996
90.295
99.750
99.990
99.999
92001
30%
1.2
191
11.4
5.8
24
99.9999
99.925
99.995
99.998
99.998
92002
30%
1.5
300
13.3
3.5
40
99.998
99.712
99.998
99.999
99.999
-------
APPENDIX C
GO-MS RAW DATA
The following section is a transcript of the analysis obtained from the
GC-MS analysis of the three cartridge traps used in the present study.
C-1
-------
TABLE 1
TENAX TRAP DATA
COMPOUND		SAMPLE AMOUNT (ual	
928Q4A 92804B 92805A 92805B 100101A 100101B 4VBLK
co2
5.7
0.75
11
1.1
6.6
0.50
methane,






dlchlorodifluoro
*
0.09


#

methane,






trichlorofluoro

1.15

0.10

0.76
ethane,






1,1,2-trlchloro






1,2,2-tr1fluoro

0.04

0.06


methane, dlchloro
*
**


#
0.62
benezene

**



1.5
methylethylketone






(MEK)

21
2.5
32

8.8
thlophene

0.32

0.23

0.24
1-butanol

1.0




ethene, trlchloro

0.15




2-pentene-3,4-d1methy1





0.05
cyclohexane, methyl


0.71


0.20
cyclopentane,






1,2,3-tr1methyl

0.02
0.41



1-hexene,






3,5,5-triraethyl

0.02




2-pentanone, 4-methyl

0.15




toluene

4.6
9.1
1.9
0.51
1.6
acetic acid,






propylester





0.07
cyclopentane






1,1,3,4-tetramethyl





0.006
3-methyl heptane


0.09



cyclohexane,






dimethyl***


0.07

0.02

benzene,






chloropentafluoro

0.04

0.003


acetic add.






butylester

0.04

0.03


octane

0.02
0.23



ethene, tetrachloro

0.03
0.08
0.01

0.003
cyclohexane.






tr1methyl***

0.16
0.25
0.16
0.03
0.04
chlorobenzene

0.02

0.009


ethylbenzene

0.09
0.44
0.07
0.05
0.005
2-methylnonane


0.13






C-2



-------
TABLE 1 - TENAX TRAP DATA (Continued)
COMPOUND		
benzene,
ethylmethyl***
2-methylundecane
benzene,
ethyldimethyl***
2-methyl naphtha1ene
naphthalene
trimethy> benzene***
cyclohexane,
ethylmethyl***
2-propyl~l-heptanol
4-propy1heptane
3-methy1-1-hexene
benzene, dimethyl***
nonane
pentalene, octahydro-
2-methyl
Isooctanol
3-methylhexane
cyclohexane, butyl
1-hexene-
3,5,5-trimethyl
cyclohexane,
diethyl***
decene
decane
decane, 4-methyl
benzene,
1,2,3-tr1chloro
cyclopentant,
1-methyl-
2-(2-propenyl)
decahydro naphthalene
undecane
3-ethylheptane
cyclohexane, propyl
diethyl phthalate
SAMPLE AMOUNT fug)
92804A 928Q4B 92805A 92805B 100101A 100101B 4VBLK
0.03
0.20
0.28
0.14
0.06
0.10
0.16
0.07
0.004
0.02
0.04
0.01
0.21
0.10
0.02
0.009
.002
.14
0.
0.
0.54
2.6
0.65
0.17
0.07 0.01
0.004
0.04
0.13
0.03
0.002
0.005
0.02
0.15
0.01
0.28
0.01
0.009
0.004
0.03
0.008
0.004
0.003
0.005
0.006
0.009
0.01
0.02
0.006
0.004
* Co-elutes with CO2, unquantitatable.
** Co-elutes with MEK, unquantitatable.
*** Isomers present, reported as a single entry.
(1595V—7 J
-------
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