A 11 IDS 7 A b 3 E 7

NBSIR 88-3767

MBS

PUBLICATIONS

Products of Wood Smolder and

Their Relation to Wood-Burning

Stoves

T. Ohiemiller

W. Shaub

U.S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Fire Research

Gaithersburg, MD 20899

May 1988

Sponsored by: —QC

U.S. Department of Energy 100

Washington, DC 20585 . U56

#88

1988

NBSIR 88-3767

PRODUCTS OF WOOD SMOLDER AND

THEIR RELATION TO WOOD-BURNING

STOVES

Research Information Center

National Bureau of Standards

Gaithersburg, Maryland 20899

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T. Ohlemiller

W. Shaub

U.S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Fire Research

Gaithersburg, MD 20899

May 1988

Sponsored by:

U.S. Department of Energy

Washington, DC 20585

U.S. DEPARTMENT OF COMMERCE, C. William Verity, Secretary

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director
I
TABLE OF CONTENTS

Page

List of Tables IV

List of Figures V

Abstract 1

1. INTRODUCTION 2

2. EXPERIMENTAL APPARATUS AND PROCEDURE 4

2 . 1 Sample Configuration 4

2.2 Wood Combustion Chamber 5 2.3 Associated Equipment 6

2.4 Test Procedure 8

2 . 5 Condensate Analyses 10

3. WOOD SMOLDER: RESULTS AND DISCUSSION 11

3 . 1 Overall Behavior of Wood Smolder 11

3 . 2 Effect of Varying Air Flow Rate on

Overall Behavior 13

3.3 Composition of the Evolved Products 18

3.4 Condensate Fingerprints 24

3.5 Rate of Heat Release . . . 26

3.6 Temperature Profiles: Implications for

Heat Transfer 28

3.7 Gas Phase Pyrolysis and Oxidation Kinetics 32

4. CONCLUSIONS 1 37

5. ACKNOWLEDGEMENTS 38

6. REFERENCES 39

7. APPENDIX A 44

8. APPENDIX B 53

iii
LIST OF TABLES

Page

Table 1. Normalized Area Percentage of Some Tar

Components 42

Table 2. Summary of Rate Constant Data 43

i

iv
LIST OF FIGURES

Page

Figure 1. Wood sample configuration and dimensions. Actual

orientation of sample in wood combustion chamber is

same as that shown here 56

Figure 2. Schematic of wood combustion chamber with sample in

place. Windows at indicated positions on front

plate protrude outward away from the U- channel in the

sample 57

Figure 3. Schematic of flow system supporting the wood

combustion chamber 58

Figure 4. Arrangement of thermocouple insertion points on the top

side of the wood combustion chamber. These same spacings

are assumed to hold at the actual depth of the

thermocouple junctions 59

Figure 5 (a) . Post- ignition smolder propagation rate as a function of

air flow along the U-channel for white pine. The

cross-hatching indicates the approximate position of

the borders for extinction and for flaming 60

Figure 5 (b) . Post- ignition smolder propagation rate as a function

of air flow along the U-channel for red oak. The line

with the arrows indicates that in the two tests at or

to the right of the line, flaming eventually developed 61

Figure 6. Maximum temperature on the interior surface of the

U-channel as a function of air flow for red oak and

white pine 62

Figure 7. Post- ignition weight loss rate for red oak and white

pine as a function of air flow in the U-channel 63

Figure 8. Apparent equivalence ratio of the wood smolder process

averaged over all test time subsequent to igniter

removal. This is shown for both red oak and white

pine as a function of air flow 64

Figure 9. Mass of wood gasified per mass of oxygen consumed

averaged over all test time subsequent to igniter

removal. This is shown for red oak and white pine

as a function of air flow 65

Figure 10. Composition of the exhaust gases from a test of red

oak at 13.3 cm/sec air velocity 66

Figure 11. Composition of the exhaust gases from a test of white

v
pine at 13.6 cm/sec air velocity 67

Figure 12. Composition of the exhaust gases from a test of

white pine at 15.3 cm/sec air velocity 68

Figure 13. Composition of the exhaust gases from a test of

white pine at 23.0 cm/sec air velocity. Flames

erupted at the time indicated by the vertical

line and the test was terminated 69

Figure 14. Composition of the exhaust gases from a test of

white pine. The air flow velocity was cut abruptly

at the time indicated by the vertical line from

22.2 to 14.4 cm/sec. The flow was cut because

flaming appeared imminent ' 70

Figure 15. Composition of the exhaust gases from a test of

white pine. The air flow decayed slowly throughout the post ignition period from a starting value

of 21.2 cm/sec to a final value of 17.4 cm/sec. The

decay was caused by narrowing of a point in the

exhaust line due to charred deposits 71

Figure 16. Composition of the exhaust gases from a test of

white pine at an air flow velocity of 8.7 cm/sec.

The smoldering was virtually extinguished at the

time the test was terminated 72

Figure 17. Composition of the exhaust gases from a test of

red oak at an air velocity of 22.2 cm/sec. Flames

erupted at the three points indicated. The first

two flaming episodes were terminated in seconds by

cutting the air flow to the chamber; it was resumed

within minutes 73

Figure 18. Composition of the exhaust gases from a test of

red oak at an air velocity of 10.4 cm/sec 74

Figure 19. Composition of the exhaust gases from a test of

red oak at an air velocity of 18.7 cm/sec. The test

was terminated when flames erupted *. 75

Figure 20. Average composition of products from the surface of

the wood for the post- ignition period. Data for

white pine as a function of air flow velocity 76

Figure 21. Average composition of products from the surface

of the wood for the post- ignition period. Data

for red oak as a function of air flow velocity 77

Figure 22. Average yield of products from unit mass of

vi
gasified wood for post- ignition period. Data for

white pine as a function of air flow velocity . .

.

Figure 23. Average yield of products from unit mass of

gasified wood for the post- ignition period. Data

for red oak as a function of air flow velocity .'.

Figure 24. Estimated, post- ignition period, average rate of

heat release as a function of air flow velocity.

Data for red oak and white pine

Figure 25. Cross-sectional profiles of red oak at two air

flow velocities with superimposed isotherms.

Point at which maximum temperature was first

reached is also indicated. Note also the gaps at

mid- depth due to shrinkage of the two pieces of

wood making up the bottom of the U- channel

Figure 26. Calculated profiles of three principal heat

transfer modes for raising temperature of wood

ahead of the point at which the maximum surface

temperature is achieved. Profiles are shown for

two air flow velocities, superimposed on a crosssection of the base of the U- channel. The

quantities in parentheses are the integrals

under the curves and they pertain to a one

centimeter width of the channel periphery

Figure 27. Predicted (1/e) lifetimes of selected species in

wood tar as a function of temperature. Predictions are for pyrolytic and oxidative degradation

78

79

80

81

82

83

vii
'

-

.
PRODUCTS OF WOOD SMOLDER

AND THEIR RELATION TO WOOD-BURNING STOVES

T. J. Ohlemiller and W. Shaub

Abstract

The smoldering combustion of solid wood is a process pertinent to both fire

safety and to the generation of air pollutants in wood burning stoves;

fundamental aspects of this process are examined here in order to provide

insights into both of these problem areas. The wood configuration employed

here was designed to permit self -sustained smolder with visual access. The

sample was in the form of a U-shaped channel 74 cm long with 6.4 cm thick

walls . This sample was contained in a steel chamber which confined incoming

air to flow along the interior of the channel. Smoldering was initiated on

the interior surface of the upstream end of this channel and it propagated in

the same direction as the airflow (forward smolder) . Tests were conducted

with both red oak and white pine and both woods behaved quite similarly. In

separate tests the air flow velocity (referred to the initial cross section of

the channel) was varied from about 9 to 22 cm/sec. At the low end of this

range, the smoldering process was prone to extinction; at the high end it was

increasingly likely to transition into flaming combustion. The smolder

velocity, peak temperature, rate of heat and product evolution all increased

over this flow range in- an essentially linear manner. Analysis of the

temperature profiles in the wood pointed to a dominance of radiative transfer

in the smolder propagation process at low air flow rates and roughly equal

roles for radiation and convection at higher flow rates. The principal

species in the tar evolved from smoldering wood resemble, in large degree,

those from previous forced wood gasification experiments and those obtained

from a wood stove operated in a smoldering mode. This reinforces the previous

conclusion that air pollutants arise directly from wood smolder and change

little in passing through the firebox of a stove. Estimates of the kinetic

rates of destruction of several of these species by pyrolysis and by oxidation

further reinforce this idea. These same estimates imply that pollutant

1
species generated in a smoldering mode cannot be destroyed without either a drastic re-design of the wood stove or the use of some form of afterburner.

1. INTRODUCTION

This report is a sequel to a previous publication entitled "Products of Wood

Gasification" [1] . Both that and the present report address the gasification

of wood in an attempt to shed light both on the fundamentals of wood

combustion and on the origins of the pollutant species that are emitted from

wood-burning stoves. In the previous study, the non- flaming gasification of

wood was examined in a highly idealized configuration -- the samples were 4

cm. cubes irradiated on one face. The effects of the large number of

parameters in that problem were examined one by one. Thus the effects of heat

flux and duration, ambient oxygen level, wood grain orientation, wood type and

moisture content, and sample thickness were all examined. For realistic

ranges of each of these variables taken individually, the changes in the

proportions of the various products (both gaseous and liquid) were rather

limited, being generally in the range of a factor of two or three.

The condensible products formed in that study were "fingerprinted" by means of

capillary gas chromatography. These fingerprints showed substantial

similarity to those from condensible products captured in the exhaust stack of

a wood stove operated in a smoldering mode (inlet air cut down to the point

that flames were absent and the only combustion was due to surface oxidation

of the wood char) . This mode is frequently used to produce extended burns

from a single load of wood and it produces copious air pollution. The

fingerprint comparison just noted suggests that the origin of most of these

air pollutants is at and below the wood surface during its gasification and

that little further chemical change occurs in the stove before these species

exit the stack (for this smoldering mode of burning). That, in turn, implies

that the pollution problem is inherent in the use of wood in the smolder mode;

only some post-smolder intervention (e.g., an after-burner) would help to

eliminate the pollutants.

2
The previous results are all based on the use of forced gasification (due to

the externally applied heat flux) and an essentially one -dimensional heat flow

geometry. This is a potentially deceptive over-simplification of the actual

behavior of wood in a fire or in a wood stove. In these real-world cases, the

smolder zone is multi -dimensional and it spreads outward from a limited

portion of the fuel where it began. Furthermore, it does so in a selfsustaining manner that does not require some external heat input. Since this

produces a different temperature distribution and a different blend of

pyrolysis and oxidation regions, one might well expect that the product

distribution could differ from that seen in the one - dimens ional , forced

gasification cases. Examining this issue is one motivation for the present

work.

A second motivation for studying the self -sustained smolder case is that it

contains within it a relation between the rate of air supply and the rate of

wood gasification which is of use in estimating the rate of pollutant or toxicant emission from this mode of combustion. This relation will be applied

here in estimating the feasibility of . destroying the pollutants from wood

stoves by gas phase reactions.

In estimating the feasibility of destroying pollutant species by gas phase

reactions, one needs estimates of the relevant rate processes. These

estimates have been obtained here by the methods of chemical kinetics

.

A third motivation for studying the self - sustained smolder of solid wood is

that, despite its relevance to both wood stoves and to fire safety, it has not

been examined before in any quantitative detail. For the purposes of fire

safety hazard assessment, one needs, at a minimum, the rate of heat release

from the smoldering wood and the rate of toxicant generation (principally CO)

.

However, one would also like to have some insights into the factors which

control the rate of smolder spread over the wood. All of these points will be

addressed in varying degree.

3
2. EXPERIMENTAL APPARATUS AND PROCEDURE

2.1 Sample Configuration

There is no standard configuration for wood being burned in a wood stove. Any

geometry that attempts to closely mimic a pile of split logs is clearly going

to introduce complexities that only serve to make interpretation more

difficult. The objective here is a simple geometry with which it is easy to

interpret the behavior of a spreading, self-sustaining smolder reaction wave.

A single flat slab of wood which is thermally thick (actual thickness greater

than the depth of a thermal wave originating at the slab surface) will not

sustain smoldering combustion unless an external heat flux supplements the

heat from wood char oxidation. Preliminary experiments established that this

external heat flux needs to be greater than about 1 W/cm2

. Although a flat

slab geometry would certainly be desirable for its simplicity, it is not

practical for the present case. The added heat flux would have to irradiate

only the active reaction zone; this would require that it move with the

spreading reaction zone at a rate whose value would be hard to pin down.

Instead of trying to impose a moving heat source, it is better to choose a geometry which produces its own; this assures truly self-sustaining smolder

spread. Two flat slabs of wood with an air gap between them have been used

in the past [2]

.

Smoldering on each interior slab face provides the

sustaining flux for the face opposite. However, unless the air gap is wide,

this configuration provides limited visual access to the smolder zone.

The configuration that was used here is shown in Figure 1. It is a flat

bottom U-shaped channel; smoldering spreads along all three interior surfaces

of the U. This approximates a closed channel which is sometimes used as an

idealization of the flow passages through a burning bed of fuel particles

[3,4]; the geometry used here then can be viewed as an idealization of that

occurring in a wood stove with a fuel load on a grate and an air flow coming

up through the grate into the bottom of the fuel bed. An observer looking

into the open side of the U has a good view of the smolder behavior on the

4
bottom of the U and an oblique view of the other two surfaces. The original

intent was to exploit this direct view of the bottom of the U by using an

imaging infrared radiometer to track the surface temperature distribution

there. This proved to be impractical due to obscuring window deposits.

However, the bottom of the U remained a natural focus for assessing the

behavior of the smolder spread process.

Note that the sides of the U channel each consist of two pieces of wood 3.2 cm

thick. Full thickness (6.4 cm) wood was not available in either of the two

wood types used (white pine and red oak) . The two pieces were glued together

(Titebond Wood Glue, Franklin Chemical Industries) 1 and further held by a small number of wood screws, especially near the end to be ignited. The three

composite side pieces were similarly glued and screwed together to form the U

channel shown in Figure 1. There was some tendency for the U channel to

attempt to warp during a test in such a way as to partly close the open side

of the U; this was prevented by placing a series of four rigid steel spacers

in this open side at intervals along the channel.

The height- to -width ratio of the U channel is such that, if the long sidewalls

were uniformly hot, the radiation from them would produce an essentially

uniform flux over the lateral expanse of the bottom of the U. Preliminary

tests indicated that, at least for the igniter size used here, the noted ratio

could not be decreased much below the value used here without precluding

smolder propagation but this subject deserves further study in the future.

2.2 Wood Combustion Chamber

The wood sample was enclosed in a combustion chamber in order to provide for

control of the flow of air and combustion products; Figure 2 shows a schematic

of this chamber. The actual orientation of the sample was the same as that

1 In order to adequately describe equipment , it is occasionally necessary

to identify commercial products by manufacturer's name. In no instance does

such identification imply endorsement by the National Bureau of Standards nor

does it imply that the particular product or equipment is necessarily the best

available for that purpose.

5
shown in Figure 1. The walls are 1.27 cm steel and they form an enclosure

only slightly larger than the wood sample itself. Any gaps (above and below

as well as toward the open end of the U) were filled with layers of ceramic

fiber insulation wrapped in aluminum foil; thus the outer surfaces of the long

sides of the U were backed by 1.3 cm of insulation while the bottom of the U

had no insulation. Flow was thereby constrained to pass only along the

interior of the U channel.

A controlled flow of air at ambient temperature entered the left end of the

chamber in Figure 2 through a 1.26 cm ID pipe fitting; no attempt was made to

condition the air stream with regard to turbulence. However, instead of

simply entering as an axial jet, the air was deflected by a plate in front of

the pipe opening into four streams tangential to the end plate of the chamber.

This caused mild turbulence in a stream that would otherwise be judged to be

laminar from the Reynolds number in the U channel

.

The chamber has a series of five Vycor windows set in water-cooled frames

which provide visual access to the interior of the U channel. As noted above,

these had a definite tendency to become coated with condensate during the

course of a test but this generally did not preclude visually following the

glowing combustion zone.

For the tests reported here, the igniter was always placed at the upstream end

of the U channel, producing smolder spread in the same direction as the air

flow. This is called forward smolder. (Other cases will be examined in the

future.) The igniter consisted of a set of three separately controlled

electrical heaters, each spaced about 6 mm from an interior surface of the U

channel. They were each operated from a separate constant temperature

controller. They heated a length of the channel of about 12 cm.

2.3 Associated Equipment

Exhaust gases exited from the right end of the chamber as seen in Figure 2.

The collector cone at that end and the subsequent transfer lines were heated

as shown schematically in Figure 3; the lines could not be kept isothermal but

6
were generally in the range from 150 to 200° C. This was not sufficient to

preclude some deposition of heavy tars which subsequently formed a thin char

coating on the interior of all the stainless steel transfer lines. This

appeared to represent a very minor loss mechanism for products on a weight

basis though it could represent species of interest for air pollution studies.

The somewhat complex arrangement of plumbing downstream of the wood chamber

reflects two factors: 1) it was not feasible to push flow through the chamber

while the igniter was in place; instead it was pulled through by the pump

indicated in Figure 3; 2) it was not desirable to send the exhaust flow

through the gas analyzers continuously, thus they were sampled intermittently.

In light of these there were two flow paths used most of the time. During

ignition all exhaust flow went through the pump (Air-Vac TD 260 air driven

venturi pump) . The inlet closure plate was removed from the left end of the

chamber and room air was drawn onto the U channel. There was no superatmospheric pressure in the chamber to force gases into the analyzers so no gas analyses could be done during the ignition interval. After ignition, the

left closure plate was installed and the indicated components fed laboratory

air to the chamber. The internal pressure in the chamber was thus slightly

elevated (1.02 to 1.06 atm.). The main flow path was then through the water

cooled heat exchanger indicated schematically in the lower right area of

Figure 3. This device was utilized merely to lessen the load of condensate

going into the laboratory exhaust system. At intervals the full flow from the

chamber was switched from the heat exchanger path to the path with a large

condenser coil (surrounded by a dry ice/ethanol slush at -72°C) followed by a filter and a dry gas test meter (Singer DTM-115 or DTM-325) for volumetric

flow measurement. This permitted sampling of the condensate for later gas

chromatographic fingerprinting and also for assessing its mass flow rate. The

dry gas test meter results, together with separate determination of the

average molecular weight of the gas, permitted calculation of the total mass

flow rate from the wood chamber. At other times, part of the flow was

diverted, again through a condensate trap and a filter, to a set of gas

analyzers. These included a Beckman polarographic oxygen analyzer, an

Infrared Industries Model 702 C0/C02 analyzer and a Beckman Model 402 total

hydrocarbon analyzer. The oxygen analyzer was calibrated with room air before

7
each measurement; the other two analyzers were calibrated before each smolder

test

.

In addition to product analyses , information was obtained on the internal

temperature profiles in the wood as the smolder reaction zone moved past the

region of measurement. Nine thermocouples ( chrome 1/alumel in 1.02 mm dia.

stainless steel sheaths) were embedded in the bottom of the U-shaped wood

channel. Their junctions were on the centerline of this piece and they

penetrated the wood from the side which is on top in Figure 1. They were

arrayed in three groups of three thermocouples as diagrammed in Figure 4.

Since the thermocouples had to penetrate about 10 cm of wood, from the chamber

wall inward, special precautions were taken to assure that the junctions ended

up spaced the same as the entry points as shown in Figure 4. A 7 . 5 cm long

drill guide was used to orient a 3 mm dia. pilot drill that stopped about 2.5

cm short of the intended depth of the thermocouple junction. The remaining

depth was then drilled just slightly larger in diameter than the thermocouple

sheath to assure its good contact with the wood. The thermocouple outputs

were recorded at ten minute intervals by a datalogger.

The design of the experiment include following the weight of the sample inside

of the chamber by means of weighing both together. The chamber is an order of

magnitude heavier than the sample but a highly sensitive balance was obtained

to overcome this problem. Unfortunately, the total weight somewhat exceeded

the balance capacity and it was necessary to displace some of the weight onto

a pivot point (through which the inlet and exhaust lines for the chamber

passed) . The arrangement proved to be unworkable since thermal expansion of

the exhaust line caused a varying force on the balance comparable to the

sample weight loss. There was no opportunity to correct this arrangement

during the course of the present experiments and so the balance provided no useful data; this will be corrected in future tests.

2.4 Test Procedure

The U channel was assembled typically a few days before a test from wood that

had spent one or mord months in conditioned storage (50 ± 5% R. H. at

8
23 ± 2°C) . The wood itself was originally kiln-dried; both the white pine and

red oak used were knot free.

After the sample was inserted in the chamber, the transfer lines were brought

up to temperature. Since the collector cone at the end of the chamber was

also heated, this caused some pre-heating of the right end of the sample. The

smolder reaction zone never got more than about halfway along the U channel

during a test so this remote preheating should have had little effect.

When the lines were adequately heated, the igniter was brought up to

temperature over a period of a few minutes. The igniter temperature was

typically 450°C for white pine and 500°C for red oak. Higher temperatures

might well have accelerated the ignition process but they run the risk of

flaming ignition as well. Ignition was, in fact, quite slow requiring about

lh hours for white pine and about three hours for red oak. The air flow

during this time was in the same range for all of the tests, 90 (+5, -10)

g/min. Because the balance system did not work, the weight loss from the

sample at the end of the ignition period could not be determined directly; *it

was estimated from the fact that the volume of wood consumed at this point was

about the same in all cases

.

After the igniter was removed, the inlet end of the chamber was sealed and the

incoming air flow was set at the desired value. The intent was always to hold

it constant from this point onward but there were occasional complications.

In two cases, gradual accumulation of charred residue in the smallest diameter

region of the exhaust gas transfer line caused a slow decay in the flow rate.

In a few other cases near the borderline for transition from smoldering to

flaming, the flow was deliberately cut back when it appeared that flaming was

imminent due to bright, localized glowing of the char. In all but one case

the flow was restored to its original level within a few minutes.

The rate of spread of the smolder reaction zone after ignition was quite slow,

even in high air flow cases. Thus it was possible to follow its progress for

up to seven additional hours and still have the reaction front move less than

2/3 of the way along the U channel. In general this post- ignition time was

9
made as long as feasible in order to lessen the relative role of the ignition

period in the overall sample weight loss. Visual observations were recorded

as to the position of the leading edge of the smolder front and, where

possible, the length of the glowing char oxidation zone.

Up to three condensate samples were taken during the post- ignition smolder

period. Typically, a substantially greater number of permanent gas analyses

were made in this same time period. The maximum quantity of condensate that

could be captured in the cold trap was sometimes severely limited by a tendency for ice to block flow through the trap; this will be fixed in future

tests

.

Smoldering was terminated by replacing the air flow with pure nitrogen. The

sample was removed from the chamber the next day and weighed to determine

total weight loss (ignition plus post- ignition)

.

Time limitations precluded an extensive series of tests; ten fully

instrumented runs were made after nine preliminary runs that explored the

necessary sample configuration, minimum air flow and ignition conditions.

2.5 Condensate Analyses

Ethanol was used to flush the condensate from the glass cold trap. This

mixture was then analyzed in two ways. The first involved capillary gas

chromatography to obtain fingerprints that could be compared with those from

the previous test configurations [1]

.

For this reason the analyses were done

in, as nearly as possible, the exact same manner. The same type of column

was used (J & W DB-1701, although it was necessary to replace it with a new

unit) together with the same carrier flow and temperature programming. The

fingerprints thus obtained consisted of the chart produced by a Hewlett

Packard 3390A Chromatographic Integrator. This chart is a record, about six

feet in length, of the pattern of peak elution from the column over a period

of two hours. The peak pattern contains features due to relative retention

times and characteristic peak shapes that make it possible to compare

fingerprints with those obtained previously without a need to repeat any of

10
the mass spectral analyses that were reported before [1]

.

It is worth

repeating an observation made in the previous work. The majority of the

condensible organic species in the cold trap are not volatile enough to come

through the chromatograph column, evidently because they have polymerized. It

was previously estimated that the about 80% by mass does not elute from the

column.

The second chromatographic analysis on the condensate from the cold trap was

intended to quantify the amount of condensible organic species (tar) in the

trap. This was done indirectly by analyzing for the two well-defined

components in the mixture, ethanol and water on a modified Porapak QS column.

In spite of some improvements in technique such as the use of a surfactant to

homogenize the tar/ethanol/water mixture, the results of these analyses had

very poor reproducibility. This was due in large part to the small fraction

of tar in the mixture which made its determination by difference less

accurate

.

3. WOOD SMOLDER: RESULTS AND DISCUSSION

3.1 Overall Behavior of Wood Smolder

The igniter and, subsequently, the smolder process the igniter induced first

pyrolyzed the wood to a char then oxidized that char so that only a fluffy,

gray-white ash remained. Typically the igniter was removed when this ash was

beginning to form on a front about 2h to 3 cm downstream from the end of the

igniter; the wood around the igniter was gasified to an estimated average

depth of about 2h cm at this point. If the igniter was removed too soon, the

smolder process extinguished within a few minutes. This happened occasionally

and the igniter was simply put back into place and left for another half hour

or so

.

After the igniter was removed and the air flow was set at the desired value,

the progress of the smolder front in the downstream direction was noted at

intervals. The same sequence noted above was occurring, now stretched

longitudinally along the length of the U channel -- heat from the char

11
oxidation zone upstream was fed downstream where it dried, pyrolyzed and

charred the wood, then induced the oxidation of the char leaving only an ash

residue. The progress of this complex reaction zone was basically steady, i.

e., the smolder velocity on the inner surface of the U channel was constant on

average. The leading edge of the char oxidation zone was not flat or even,

typically, because of the random disturbances introduced by shrinkage cracks

in the wood. However, plots of the position of the leading edge of the ash or glowing zone as a function of time over the course of a several hour burn gave

rather good straight lines for most tests. This constant average rate of

movement has been seen in other two-dimensional, forward smolder processes

[5,6] but it was less expected here because the thermal wave in the solid

phase is very long (longer than the sample) and the total distance burned over

the entire test duration is short compared to this thermal length. Evidently

the wood is, in effect, evenly pre -heated over the length burned here. It

will be seen later that the progress of the char oxidation front is not

limited by the disappearance of oxygen and, instead, appears to be limited by

the rate of heat transfer to the unburned wood.

The char oxidation zone extended upstream for a substantial distance from its

leading edge; it was of the order of 20 cm long and probably varied with the

air flow rate but good data on this are not available. It was over the length

of this char oxidation zone that the channel cross section opened up

considerably. This zone was terminated on the upstream end by quenching,

particularly at the bottom of the U with its lack of insulation behind the

wood. Typically, more than a centimeter of charred and uncharred wood were

left when the reaction ceased on the upstream end. Because of the varying

channel cross section in the reaction zone, all flow velocities given below

are referred to the original U channel dimensions (with the gas at room

temperature)

.

As noted above, the wood cracked randomly as it charred, though the cracks

were oriented predominantly either along the longitudinal axis of the channel

or transverse to it, presumably due to the wood grain orientation. The cracks

did not seem to play a major role in the propagation process except, perhaps

at very low air flow velocities. There the char reaction seemed to move

12
preferentially in the cracks, possibly because they afforded a lesser

radiative heat loss rate. Shrinkage of the wood also caused separation of the

two pieces which made up each side of the U channel. Again this did not seem

to play a significant role for propagation under most conditions. It provided

more crevices for the smolder to hide in at low air velocities. It may have

played some role also in the development of flaming at high air flow

velocities. The transition to flaming seemed to be associated with the

development, over a period of several minutes, of regions of the char (several

centimeters in longitudinal extent) which became hotter (as determined by

their brighter glow) than surrounding regions of the char. This could have

been facilitated by oxygen penetration into the gap between the two layers of

wood causing heating from both sides for the innermost layer; there are no data to confirm this speculation more directly.

3.2 EFFECT OF VARYING AIR FLOW RATE ON OVERALL BEHAVIOR

Smolder propagation rate . The air inflow rate was varied from the extinction

limit to the flaming limit, a range which proved to be rather narrow for the

present wood chamber configuration. Figure 5 shows the variation of the

surface smolder propagation rate with air flow rate between these limits. As

expected, the smolder spread rate increases with increased air flow into the

upstream end of the U channel. This is consistent with the behavior of

forward smolder in other configurations [4,5]. The increased flow of air

accelerates the rate of transfer of oxygen to the wood char surface on the

interior of the U channel. This, in turn, increases the rate of char

oxidation at and below the surface. The attendant increase in the rate of

heat evolution in the surface region raises its temperature thereby increasing

the rate of heat transfer to unburned portions of the wood downstream and this

causes the smolder front to move downstream more rapidly. Figure 6 shows that

the char surface temperature does indeed increase substantially with increased

air flow rate. These temperatures were the peak values seen by the initially

submerged thermocouples (the shallowest in Figure 4) ; the thermocouples

penetrated the char surface when it receded due to char consumption.

13
The smolder spread rates for white pine and red oak are essentially equal for

a given air flow rate. The least squares lines fitted to each set of data

points differ, as seen in Figure 5, but, in view of the data scatter, these

lines cannot be said to be significantly different. The data in Figure 7 derive from a different measurement, that of weight loss averaged over the

post- ignition period2

. The average weight loss rate for red oak is 1.6 to 1.8

times greater than for white pine. This is consistent with essentially equal

smolder propagation rates if one factors in the fact that red oak is about 1.6

times more dense than white pine. Given this density difference, however,

equality of smolder spread rates is unexpected. In one -dimensional forward

smolder through a permeable bed of wood fibers, Ohlemiller and Lucca [7] found

the propagation rate of the char oxidation zone to be limited by oxygen

consumption and char oxidation stoichiometry. If the same mechanism applied

here, it would imply that the spread rate over oak would be about 40% slower

than that of white pine. On the other hand, Ohlemiller [6] found that the

rate of heat transfer to the unburned fuel can play a direct role in the

propagation rate for two-dimensional smolder of permeable fuel layers. With

this in mind, one notes that the thermal conductivity of wood is proportional

to density [8] and then expects that oak might propagate smolder faster than

white pine. It will be seen below that heat transfer plays a central role

here but that conduction through the wood does not; thus the equal propagation

rates can be rationalized.

Flaming limit . Transition to flaming combustion was encountered at about the

same air flow velocity (20 to 25 cm/sec) for both types of wood. The

transition process was unpredictable as to its time of occurrence. It appears

that it becomes an increasingly likely event as one approaches the upper end

of the air flow range in Figure 5. Thus it will occur earlier in a test when

the flow rate is 25 cm/sec than when it is 20 cm/sec.

2 The sample weight was only measured before and after a test. The weight

loss during ignition was estimated from the observation that the volume of

wood gasified at the end of the ignition period was approximately the same

from test to test. The weight of this wood volume (approx. 1000 cm3

) was

subtracted from the total weight loss to estimate post- ignition weight loss.

14
As noted previously, the transition process (i.e., the initiation of flaming

in the gas phase) was typically associated with localized regions of greater

than average char surface temperature. Transition to flaming has three

requirements: (1) a local gaseous mixture of fuel and oxygen that is above the

flammability limit at the local temperature; (2) a heat source within this gas

mixture sufficiently hot to ignite the gases; (3) if the flame is to be

sustained beyond a momentary flash, a continuing source of fuel gases, usually

assured by heat feedback from the flame to the solid fuel surface. If the

local temperature and the gas mixture volume are sufficient, self heating can

replace the extra heat source called for in (2)

.

In the present complex

configuration, (3) could possibly be met by a combination of gases generated

by smoldering and by flaming.

It seems quite probable that the cracks in the char provided local spots which

favored the build-up of greater than average concentrations of flammable gases

since they would be subject to lesser mass transfer rates than regions of the

char surface forming the interior walls of the U channel. In this regard, it

is pertinent to address what is known about the concentration of flammable

gases in the U channel. Gas analyses were done only on the average

composition leaving the channel; these are discussed in greater detail below.

For now it is relevant to note that the average CO concentration in the

exhaust gas was in the range from 2 to 3%; the total hydrocarbon concentration

was in the range from h. to 1%. At room temperature the lower flammability

limit of CO is 12.5% [8]. The total hydrocarbons were not analyzed in the

current work but in the previous tests with the same woods they were found to

consist largely of methane, ethane and ethylene [1]. The lower flammability

limits for these species at room temperature range from just under 3% to about

6% [8]

.

The lower flammability limit decreases very slowly with increasing

temperature; e.g., that of ethane drops from about 3% at room temperature to

just under 2% at 550°C [8]. As will be seen below, there is some oxygen

depletion along the length of the U channel; again this has only a small

effect on the lower flammability limit [8]. On the basis of CO and total

hydrocarbons, then, the average gas composition is not flammable for any of

the test conditions.

15
There is another potentially significant source of fuel vapors, the

condensible material denoted as tar previously. This constituted about 5% of

the exhaust gas mass. If all of this material was in the gas phase in the

same locale as the other two fuels, it would almost certainly make a flammable

gas mixture. The most likely place for this to occur, however, is in the

pyrolysis zone downstream of the char oxidation zone; here the temperatures

are too low to ignite the mixture

.

In any event, it is clear that the hydrocarbons alone do not need a great deal

of concentration enhancement above the exhaust level average to be within

their lower flammability limit. They are undoubtedly significantly more

concentrated in the boundary layer near the char surface and in cracks in the

char. Thus, in the absence of direct local measurements, one can still make

the reasonable inference that flammable mixtures exist at many points along

the hot char surface. If so, the factor limiting the appearance of flaming is

the temperature; the char must get hot enough in a region adjacent to one of

these flammable mixtures to ignite it. Momentary flash ignitions were not

observed so evidently requirement (3) above is readily met, though it is not

clear just how. Direct sampling of the gases within the U channel is needed

to test the ideas expressed here.

If temperature is the controlling variable in transition to flaming with wood

in this configuration, one might expect some apparatus dependence for the air

velocity at which the transition occurs. The temperature achieved for a given

rate of heat generation (corresponding to a given rate of air inflow) is

somewhat dependent on the rate of heat loss which, in turn, depends on

apparatus design. It is interesting to note that Alexopoulos and Drysdale

obtained comparable air flow velocities for transition to flaming with forward

smolder upward in an insulated chimney lined with wood-based fiber board [9]

.

Evidently the critical air velocity value does not vary rapidly with heat loss

conditions

.

Extinction limit . The lower limit to smolder propagation, extinction, was

only actually seen with white pine. The smoldering in the lowest air flow

case for red oak in Figure 5 was clearly quite weak, however. Interestingly,

16
the extinction with white pine occurred after the sample had smoldered for

about six hours beyond the end of the ignition interval. Form the propagation

rate in Figure 5(a), it is apparent that, even in this time, the smolder front

moved only a few centimeters downstream. The preponderant smolder process

then was inward penetration into the wood and it was probably stopped by

quenching when the remaining thickness of wood could no longer provide

sufficient insulation. This suggests that this lower limit could probably be

extended somewhat by using thicker wood walls and providing more external

insulation for them.

Effective fuel to air ratio . The ratio between the rate of wood gasification

and the rate of air inflow is a parameter one needs in order to estimate the

rate of evolution of pollutant species from wood smolder. This parameter is

configuration dependent since it affected both by the degree of contact

between the wood and the air and also by their relative directions of

movement. This last refers to the issue of whether the air flows in the same

direction as the smolder front moves (as it does here) , the opposite

direction, or some combination of the two plus transverse flow as well; these

concepts are essentially the same as those embodied in the idea of updraft,

downdraft or crossdraft burning modes in a wood stove [10] . For the two

simpler cases (flow strictly with or against the direction of smolder

propagation) as they occur in a bed of porous fuel , it has been shown that

this fuel to air ratio differs considerably [7]. Configuration dependence has

not yet been investigated in the present context but the result just noted

implies that caution is necessary in using the present result for other

configurations. Where data are available for wood stoves, this ratio has been

found to vary by at least a factor of four [11].

Figure 8 shows the results obtained here for the ratio of post- ignition weight

loss rate to average air inflow rate. Given the scatter in the oak data, it

cannot be said that this ratio changes significantly with air flow velocity.

The ratio once again shows more red oak being gasified with the ratio for the

two woods being approximately in the ratio of their densities. The values of

the fuel to air ratio obtained here fall at the high end of the range seen for

wood stoves [11] . In wood stoves operated with a very limited air supply rate

17
(where one might expect a predominance of smoldering) the ratio tended to be

lower than the range seen here; there it varied from 0.04 to 0.10 [11]. The

stove tests used 5 by 10 cm and 10 by 10 cm oak brands stacked in the fire

box; the air flow was apparently a crossdraft.

As will be seen below, oxygen consumption was not complete in the

configuration used here. When the fuel consumption data are recomputed on the

basis of the oxygen actually consumed, one obtains the results in Figure 9.

Again more red oak was gasified relative to white pine in approximately the

ratio of the density of the two woods. These comparisons between the two

woods in all cases suggest that the smolder zone moved down the U channel

gasifying the two woods at an equal volumetric rate (for a given air inflow

rate) independent of wood type.

The data in Figure 9 reflect the incompleteness of the wood oxidation in the

present smoldering combustion mode. Stoichiometric oxidation of the two woods

requires that the ratio in Figure 9 be in the range 0.5 to 0.7; these

stoichiometric values are based on pitch pine and white oak, the closest woods

for which elemental compositions were available [12]. It is this high degree

of incompleteness which makes this combustion mode so polluting. This is

reflected also in the composition of the evolved gases which is examined

below.

3.3 Composition of the Evolved Products

Time -dependent composition of the gases. Figures 10 through 19 show the

variation of the major products with time from post- ignition to the end of

each test.

The first point to note is that, in all of the tests, there was a considerable

amount of oxygen passing along the entire length of the sample without being

consumed. The reason for this is basically the same as it is in many wood

stoves; the contact between the incoming air stream and the hot char surface

is not sufficient. For the present configuration one can readily calculate

that an oxygen molecule entering the reaction zone on the centerline of the U

18
channel will convect more than the full length of this zone before it can

diffuse laterally to the char wall. The original intent was to run the air

flow sufficiently slowly that there would be time for this diffusion but this

called for air flow rates below the extinction level in this apparatus

.

A more subtle point concerning the remaining oxygen has to do with the failure

of the reaction zone length to grow sufficiently to consume all of that

oxygen. If there is plenty of oxygen in the gas at the leading front of the

reaction zone, one might expect the front to move downstream until it consumes

all available oxygen and then move onward only at the rate at which completion

of fuel consumption on the upstream end of the reaction zone permits; this is

what happens in forward smolder through a permeable fuel bed [7]

.

Channel configurations with flow and reaction zone movement in the same

direction have been studied in the context of coal and oil shale gasification

[13,14] and in the context of flame propagation in fuel-lined ducts [15]. In

the first two contexts, complete consumption of oxygen appears to be typical,

at least for the conditions described in these references. For fuel -lined

ducts with flaming combustion, cases were seen in which oxygen consumption was

small and in which the consumption was total; however, in the case of small

oxygen consumption, the fire either died out eventually or accelerated to a case of total oxygen consumption. Roberts and Clough [15] presented a simplified analysis of the energy conservation in the duct which implied that

two propagation rates are possible but the lower rate, which leaves residual

oxygen, is unstable; the solution achieved depends on the vigor of the

ignition process. The assumptions in this model do not fit the present

problem well. De Ris [16] suggested that the cases of incomplete oxygen

consumption were the result of flow stratification in which a layer of cooler,

oxygen- containing gas flowed along the bottom of the duct and escaped

reaction. In the absence of flow and oxygen profiles within the U channel, it

is not certain if such behavior is pertinent here, but it seems doubtful. No

stratification of the smoke flow was noted during the tests. Furthermore, the

reaction front on the bottom of the channel did not lag behind other portions

.

It seems more likely that the rate of heat transfer to the unburned fuel

downstream of the char oxidation zone in the present configuration was not

19
enough to permit a sufficient extension of the reaction zone so that all

oxygen could be consumed; this will be examined further below.

Another point to be noted about the time -dependent gas composition data in

Figures 10 to 19 is that, in most cases, the composition is very nearly

steady. This goes along with the previous observation that the rate of

propagation was essentially constant. There are two notable exceptions, the

tests in Figure 17 and Figure 19; in both cases the air flow rate was in the

region of the flaming limit. The decaying oxygen penetration of the reaction

zone suggests that the length of the zone was increasing toward the zero

penetration case discussed above; unfortunately, confirming data on the length

of the reaction zone as a function of time for these cases are not available.

The rate of propagation of the reaction front was not accelerating in either

case; it was as steady, on average, as other cases of slower air flow rate.

For the reaction zone length to grow longer, however, it is only necessary for

the rate of longitudinal movement of the smolder front to increase relative to

the rate of lateral penetration of the front into the wood. These two rates,

which are the orthogonal components of the smolder velocity vector normal to

the isotherms in the solid, may not have reached a steady balance in these two

cases. This is a situation which could be illuminated by a two-dimensional

smolder propagation model which is not available at present.

Also note that in Figures 17 and 19 there is a an increase in the ratio of C02 to CO with increasing time. This is small but evidently real; for example, in

Figure 19 the ratio goes from the range 2.4 - 2.6 up to 3.0. This could

reflect increased flow time available for gas phase CO oxidation to C02 as the

smolder zone lengthened. The residence time (flow time) is about one second

in these cases and the peak gas phase temperature is about 650° C.

Extrapolation of Dryer's [16] kinetics for the homogeneous oxidation of 'wet'

CO indicates that some oxidation is possible in these conditions. The

characteristic blue flame indicating rapid CO oxidation was not seen here.

The heterogeneous reduction of C02 to form CO evidently requires somewhat

higher temperatures to be appreciable for the times available here [17].

Definitive characterization of the behavior of these gases for this particular

20
system would require sampling of the composition of the gas phase along the

length of the U channel.

Time-averaged compositions . Comparisons with the results of our previous,

small scale, forced gasification study of these same two woods is most readily

done on the basis of the average composition of the evolved products. The

current results are given in Figures 20 and 21 for white pine and red oak,

respectively. They are expressed there as fractions of the total product mass

in the post- ignition period. The scatter in both of these Figures is such

that no real trends with air flow rate can be said to exist. This is in spite

of a 200° C range in surface temperatures for the air flow range indicated

(Fig. 6). The closest analog to the present series of experiments in the

previous small scale forced gasification experiments is that in which the

incident heat flux to the surface was varied. Increasing the heat flux there

from 2.5 to 6.9 W/cm2 caused the peak surface temperature to go from about

550° C to about 750° C [18] . For the most part, the variation in the fractional

distribution of evolved products over this flux range was not strong but there

were significant increases in the fractions of both CO and C02

. These

increases were mainly in the range of 50% but for red oak the increase was

markedly stronger for CO. These comparisons between the two types of

experiments are rather strained because the forced gasification tests were

highly transient in nature and much shorter in time scale (10-15 min.); this

biases the products toward those formed early at lower temperatures than the

above peak surface values would indicate. This is particularly evident in the

case of water which comprised 40 to 60% of the evolved products in the forced

gasification experiments and only 20 to 30% of the products here. The wood

samples in both test series were equilibrated to the same ambient conditions

so that their water contents were quite comparable.

The actual origins of the various major products shown in Figures 20 and 21 is

not completely clear. If the char were a pure pyrolytic graphite, its

oxidation in the present temperature range would yield only C02 [19]

.

The

char formed from cellulose contains amorphous carbon as well as significant

amounts of hydrogen and oxygen [20]

this is very likely also true of the char

contributed here by the hemi-cellulose and lignin components of the woods.

21
Oxidation of cellulosic char yields equal amounts of both CO and C02 in the

absence of inorganic ash; differing ash components alter this ratio

appreciably [21] . The previous forced gasification experiments were performed

in nitrogen as well as air; those results imply that in the current tests the

majority of the CO. and C02 arises from char oxidation. The remainder comes

both from pyrolysis of the wood components (particularly holocellulose) and

from secondary degradation of the tar components

.

Degradation of wood tar was studied by Boroson [22] . Tar generated by

pyrolysis of the components of wood can undergo two classes of reaction as it

passes out through the bulk of the sample (through the char layer typically on

the sample exterior) and into the surrounding gas phase. The first class

comprises homogeneous pyrolytic reactions, probably unimolecular in nature.;

the second class comprises heterogeneous reactions between the tar components

and the char surface over which they pass as they exit the sample. The

homogeneous reactions typically convert a fraction of the tar to gases, the

major one being CO. The fraction converted is dependent on temperature and,

at a given temperature, is essentially fixed after about one second

(particularly at temperatures above 600° C) . The homogeneous reactions are

very fast and affect only a fixed fraction of the tar independent of times

longer than a few milliseconds and independent of temperature, at least in the

range from 400 to 600° C. The fraction of tar undergoing this heterogeneous

reaction was estimated to be about 1/3. The products were mainly CO, C02 and

char.

In the context of the present experiments, Boroson' s results imply that

anywhere from zero to 50% (over the temperature range in Figure 6) of the

initial tar generated within the pyrolysis zone of the wood was degraded by

homogeneous pyrolysis. One can estimate that it takes 0(1-10 sec) for the

gases generated within the sample to pass through the char layer; the

subsequent time in the gas phase passing along the length of the reaction zone

in the U channel is 0 (1 sec)

.

Thus essentially all of the homogeneous

degradation of the initial tar occurs within the sample. The same holds true,

of course, for the heterogeneous degradation. The combined degree of

22
degradation of initial tar thus varies from 1/3 to about 5/6, based on

Boroson's results applied to the present context.

There is a discrepancy between the above inferences based on Boroson's results

and the results in Figures 20 and 21. Those Figures show a quantity of

surviving tar that is both substantial in amount and constant with air flow

(and, therefore, temperature). The above arguments imply that the amount of

tar should decrease by about a factor of two as from the lowest to the highest

air flow rate. Tar determinations were not very accurate (standard deviation

about ±50%) so this could mask any such trend. On the other hand, since CO,

in particular, is a major product of the tar degradation, one should see a substantial increase in its weight fraction with increasing air flow rate; the

CO analyses are believed to be quite accurate. The reason for this

discrepancy is not known.

Figures 22 and 23 show the post- ignition average product information in a different manner, i.e., on the basis of mass of product generated per unit

mass of wood gasified. This is the form that is most useful for determining

the rate of evolution of toxic or pollutant species from smoldering wood in

other contexts where the behavior 5 of wood might be expected to be similar.

Note that the sum of the products for any given flow condition need not add up

to one even though one gram of wood went into generating them; oxygen from the

inflowing air is also incorporated in the products. The average summation of

the product masses was , in fact, 1.17 grams per gram of wood gasified; the

average deviation is rather large, 0.12. This variability reflects a varying

mass balance on the system. The principal errors are believed to reside in

the determination of the total mass flow out of the chamber as a function of

time and in the estimate of the amount of wood gasified during the ignition

process. In addition, as noted previously, the tar fraction values used here

have a ±50% uncertainty. The water values are the average of those determined

from the water transducer in the exhaust line and the water content of the

condensate trap; these typically were about 20% apart. Given these

uncertainties and the evident scatter in the data, it is again not possible to

say that the products vary in any significant way with air inflow rate.

23
Similarly, one cannot assert that the products from the two different types of

wood are significantly different.

The weight ratio of C02 to CO in the previous, forced gasification experiments

was in the range from 1.5 - 2.5 to 1. Here it is about 3 to 1 . In stove

burns done in the first year of this program, the ratio was substantially

higher, from 6-9 to 1 [23]; this approaches the value of 10 to 1 seen for

flaming combustion in a free-standing fireplace [24] so one cannot rule out

the role of flaming combustion in those early stove burns. It seems probable

that CO oxidation to C02 in the gas phase is the reason that the numbers here

and in stove burns exceed those seen in the forced gasification experiments.

This comparison also implies that such oxidation is considerably less in the

present configuration than it is in a wood stove operated in a slow burning

mode [23]

.

3.4 Condensate Fingerprints

This subject was pursued much more extensively in our previous work [1] where

an effort was made to identify many of the species eluted through a capillary

gas chromatographic column. Time constraints severely limited the follow-up

work with the present condensates (tar) seen in the cold trap. The goal here

was only to compare the patterns of species quantity and retention time

(i.e., the "fingerprint") that one gets as an output from the chromatograph.

Even this presents some technical difficulties, especially in the first 10% or

so of the fingerprint where solvent pooling effects 3 lead to non-reproducible

retention times.

For white pine, visual matching of the fingerprints indicates a strong degree

of similarity between those obtained here and those obtained in the forced

gasification experiments. It appears that nearly all of the same components

are present in the two types of condensates but there are a small number of

3 The solvent carrying the tar onto the chromatographic column condenses

to a liquid at the front end of the column. This liquid spreads unevenly and

irreproducibly under the influence of the column carrier gas causing a variable initial distribution of the tars on the front end of the column.

24
exceptions among the larger peaks, most of which are unidentified. One which

has been identified is levoglucosan, a primary product of cellulose

degradation. In the forced gasification tests, it was nearly always one of

the prominent species in the tar; here its presence could not be confirmed.

Changes in the relative proportions of the eluted species with method of

generation are harder to judge precisely. Table 1 shows some data along these

lines for eight of the major species in the eluted condensate. Since it is

not possible to assure that one is injecting exactly the same total mass of

tar onto the chromatograph column with each different sample , the area

percentage numbers presented there have normalized by the area of one of the

most reproducible peaks, that for creosol. The area percentages shown in the

Table are calculated after both methanol, the solvent in the previous tests,

and ethanol, the solvent in the present tests, have been subtracted out. Area

percentages are only an approximate indication of mass concentration in the

tar because detector response factors are not included.

Table 1 shows normalized area percentages for the condensate from the base

case of forced gasification and for low and high air flow cases of selfsustained smolder (present work) . For white pine one sees there variability

in the tars from the present tests of as much as a factor of two to three.

This may not be significant, since that degree of variability was seen in the

previous tests repeated under identical conditions. Acetic acid was

problematical probably because it occurs in the first ten percent of the

fingerprint, that most affected by solvent pooling disturbances; it was highly

variable and thus was omitted. For the other species, between the base case

for forced gasification and the present results, one sees variations of up to a factor of five (furaldehyde and phenol) . Comparable variations appear to be

present in other unidentified peaks. This implies that the chemical history

of the tars in the two types of experiments differs significantly.

Unfortunately there is not enough information available to indicate the source

of these differences. Experiments similar to those of Boroson [22], but with

tar analyses, are needed to clarify this point.

25
The same qualitative statement about visual matching of fingerprints between

the previous and present results holds for red oak, although here there may be

more species where the proportions are different between the two types of

test. This is not so apparent from Table 1 where the basic story about

relative proportions appears to be about the same as for white pine.

Interestingly, the visual match between fingerprints obtained here and those

obtained from a wood burning stove (for mixed oak species) in the previous

study [1] is as good as or better than that noted previously between the wood

stove and the forced gasification results. This reinforces the tentative

conclusion arrived at previously that there is very little gas phase chemistry

occurring in a wood stove when it is operating in a smoldering mode (referred

to previously as an "overnight" mode of operation) . Again this implies that

the pollutant species escaping from the stack of a wood stove operated in this

manner arise directly from the wood smolder process in the condensed phase and

then pass through the firebox and up the stack without much alteration. The

gas phase kinetics discussion below tends to reinforce this idea further.

3.5 Rate of Heat Release

The rate at which smoldering combustion releases heat is of interest both in

the context of wood burning stoves and in the context of fire safety; in the

latter, one needs the heat release rate in order to predict the transport of

combustion products in the buoyant plume which rises above an unconfined

combustion source [25]

.

The present experiment did not include any means for

directly measuring the heat evolved from the wood smolder process , but this is

not always necessary. For flaming fires, it has been shown that one can

accurately estimate the rate of heat of release from the rate of oxygen

consumption; the two are related by a factor (13.1 kJ of heat per gram of

oxygen consumed) which is nearly constant for most organic materials of

interest [26]. This technique has not been previously applied to smoldering

combustion and, indeed, it is not completely clear how accurately it can be

applied. Smoldering is characterized by incomplete combustion; this is

apparent in the Figures just discussed. Huggett [26] has shown that the above

factor converting oxygen consumption to heat release can apply with reasonable

accuracy not only when C02 and H2 0 are the combustion products but also when

26
partial oxidation products (alcohols, aldehydes acids) are formed. The

accuracy is less when alcohols and aldehydes are products, however; the

absolute error depends on the quantities of each but it only approaches 20%

when these are the only products. Here there is a. very complex and illdefined mixture of tars being generated, some of which are undoubtedly

partially oxidized; one can only assume, on the basis of Hugget's

calculations, that they are not a major source of error.

A further factor to consider is the formation of CO; its heat evolution per

gram of oxygen consumed is only 6.9 kJ/g. In the present experiment, not all

of the CO comes from oxidation; some is generated directly by pyrolysis of the

wood or tar. A direct measure of this pyrolytic fraction is not available; in

the previous forced gasification experiments it was about one -fourth of the

total CO.

The results shown in Figure 24 are average heat release rates over the postignition period. It has been assumed there that all CO comes from oxidation

and that all oxygen consumed in partial oxidation of tar components yielded

the full heat release characteristic of C02 formation. As expected, the rate

of heat release increases with increased oxygen inflow. What is not expected

is that the two types of wood are indistinguishable. Recall that Figure 7

showed the rate of mass loss for red oak to be higher than that for white pine

by a factor about equal to their density ratios (1.6). Figures 22 and 23 show

the product distributions for the two woods to be about the same. Then red

oak should be generating oxidized products at about 1.6 times the rate for

white pine and therefore should be generating more heat in about the same

ratio. If this is the case, it is hidden by the considerable data scatter in

Figure 24.

To estimate the rate of heat release from forward smolder a different mass of

solid wood, one could presumably scale directly with area of wood likely to be

involved in the manner here. In practice this could be difficult to estimate

but one could at least get an order of magnitude estimate. The same scaling

would apply, as a zeroth approximation, for the rate of species generation.

27
3.6 Temperature Profiles; Implications for Heat Transfer

The set of nine thermocouples imbedded in the wood that formed the base of the

U channel provided sufficient information that one can infer the approximate

thermal profiles in the wood. Figure 25 shows such profiles for red oak at

low and high air flow velocities. The isotherm patterns are shown

superimposed on top of tracings of the wood as seen in cross-section; the

plane shown bisects the base of the U channel. The positions of the isotherms

have been determined by linear spatial interpolation between the nine

instantaneous temperatures at nine known locations. The resulting pattern was

shifted longitudinally so that the 300° C isotherm sits on top of the

approximate demarcation "line” between charred and uncharred wood. (This is

actually not sharp enough to be termed a true line.) Note the crack between

the two pieces of wood which form the base of the U; the crack does not appear

to have much effect on the isotherms. One can also see the irregular

regression of the surface caused by both by shrinkage of the wood as it

charred and by conversion of some of that char to ash (not shown) . It is

apparent from the top of Figure 25 that the reaction zone, in progressing from

left to right past the fixed thermocouple positions, will eventually expose

them to the surface of the oxidizing char and then to the free space within

the enlarging U channel. Thus the peak value of the char surface temperature

can be measured by the imbedded thermocouples. The isotherms have also been

extrapolated to the surface to provide an approximate measure of the surface

temperature distribution ahead of the maximum value.

A striking aspect of the isotherms is their very acute angle with respect to

the interior surface of the U channel. This implies that the heat transfer in

the wood is nearly normal to the channel surface; that in turn implies that

heat conduction through the wood plays a minor role in driving the reaction

zone forward along the U channel.

It is possible to use the above information to do a more quantitative

assessment of the relative contributions of the various modes on heat transfer

in this smolder propagation process. This could facilitate extending the

present results to other situations.

28
Three modes of heat transfer are considered: conduction through the wood in

the longitudinal direction, convection by the gas flow along the interior of

the U channel, and radiation along the interior cavity of the channel. Kanury

and Blackshear [27], on the basis of measurements on cellulose cylinders,

suggested that an additional mode of preheating is condensation of water and

pyrolysis vapors moving down the temperature gradient in the porous structure

of the condensed phase. Such movement would be essentially longitudinal in

the present grain orientation. Our previous data (Figures 22 and 23) show

that the quantities of tar and water are comparable; however, water, with its

high heat of vaporization, is expected to be the more significant component in

this regard. In fact the temperature data obtained in the present tests show

only a water plateau. This tends to be most pronounced near the outer surface

of the base of the U channel, probably because of the lack of thermal

insulation there. It does not appear that neglect of this has an appreciable

impact on the calculations that follow. Water condensation could also help to

preheat the inner surface of the channel, supplementing the convection and

radiation discussed below. However, the water condensation front is well

downstream of the region to be considered here; tar condensation is neglected

because it does not appear to be significant.

The longitudinal heat conduction flux through the wood is readily estimated

from the temperature profiles shown in Figure 25. This will be done in the

plane where the peak surface temperature is just reached. One needs to

compute, as a function of depth below the inner surface of the channel, the

value of the following expression.

-k(T) dT/dx

Here k(T) is the temperature -dependent thermal conductivity of the wood; the

second term is the temperature gradient in the longitudinal direction. The

gradient in the plane of interest is estimated from the temperature profiles.

Thermal conductivity data are not available above 100° C; the linear

temperature dependence below this temperature is extrapolated for the present

purpose [27]; thus, for red oak, along the grain, one has:

29
k(T) = 0.0039 (T/300) , W/cm°C

The results, which are shown in Figure 26, are discussed below.

The convective heat flux from the gas flowing along the channel to the wall

(downstream of the peak temperature point) is more difficult to estimate

because a direct measure of the gas temperature is not available. It will be

assumed that the gas acts as a plug flow which is initially all at the peak

temperature of the charred wood surface. Since it is quite unlikely that the

gas over the whole channel cross section gets to this temperature, this will

yield an over estimate of the convective heat flux. Given this assumption,

one can calculate the temperature of the gas as it proceeds downstream from

the initial peak temperature plane by solving the following differential

energy balance equation:

dTG/dx = -hAw [ TG" Tw( x)]

This simply expresses the assertion that the gas temperature decays as it

loses heat to the wall by convection. Here m^ is the mass flow rate of the

gas, CG is the heat capacity of the gas, TG is the gas temperature, h is the

convective heat transfer coefficient, is the channel sidewall area per unit

length of channel (the open side of the channel, which has about 14% of its

side wall area, is assumed adiabatic for lack of better information; there are

recirculation zones between the channel and the windows along the open side

which make this a reasonable approximation) , Tw (x) is the temperature

distribution along the channel wall obtained in Figure 25 by extrapolating the

in-depth profiles to the surface as shown; these surface temperatures results

are well- fitted by a straight line or a parabola. The above equation can then

be solved analytically for TG

. The resulting expression is used to calculate

the convective heat flux to the wall:

h [Tg (x) -Tw (x)

]

This profile along the longitudinal direction is also shown in Figure 26.

30
Estimation of the radiative flux from the sidewalls of the U channel to the

base of the U requires certain assumptions also. The principal assumptions

are that the temperature profile of the sidewalls is the same as that

determined from Figure 25 for the base of the U and also that, once the peak

surface temperature is achieved on the surface, the surface remains at that

temperature for all points upstream (up to the other end of the char oxidation

zone) . Visual observations of the glowing of the channel walls suggest that

these assumptions are reasonable but they have not been quantitatively

checked. It is also assumed that the walls remain flat throughout the

isothermal region whereas they actually curve irregularly outward as the char

is gasified by oxidation; this assumption is less of an approximation at the

higher air flows

.

With the preceding assumptions one can show that the longitudinal distribution

of radiant flux along the centerline of the base of the U channel is given by:

q_ n

(x) = 2aT (F-F')+ RAD w , max

Za\

'A

max

[T

w (x')n 2 2

L i+(x-x') 2 2 2

Z +a +(x-x')

dx'

Here a is the Stefan -Boltzmann constant, Tw MAX is the peak wall temperature,

F is the radiative view factor from the isothermal portion of the sidewall to

a point on the axis of the U channel base, F' is a correction to the preceding

view factor when the point on the axis of the U channel base is beyond the < end

of the isothermal portion of the sidewall, Z is the half -width of the U

channel, Tw (x') is the linear or parabolic fit to the surface temperature in

the non- isothermal region, x' MAX is the maximum downstream position to which

the integration of the flux from the non- isothermal sidewall region is taken, x is the longitudinal coordinate, x' is a dummy longitudinal coordinate of

integration. The view factors F and F' , which have the same functional form, are obtained from a standard reference [28]:

F = (1/2*) [ arctan(b/i) - (i/(a2 + i

2

) %) arctan(b/(a2 +i2 )* ) ]

Here b is the length of the isothermal sidewall region and a is the height of

the sidewalls.

31
The necessary integration in the above equation is best done numerically. The

resulting longitudinal profiles of radiative flux along the central axis of

the base of the U channel are shown in Figure 26

.

Figure 26 shows calculated results for a high and a low air inflow velocity.

The flux distributions are shown superimposed on a cross-sectional plane

through the central axis of the base of the U channel; the starting point at

the left of the Figure corresponds to the point where the surface temperature

first reaches its maximum value and the distributions stretch downstream from

that point. Also shown in the Figure are the integrals of the areas under

these profiles which measure the total heat transferred to the wood by each

mode per unit width of the wall. It is immediately clear that conduction heat

transfer downstream through the wood is a minor contributor to the overall

wood heat-up process. This helps explain how the two different kinds of wood,

with their differing thermal conductivities, can have essentially the same

smolder propagation velocity. One has to be more cautious in assessing the

relative importance of the other two heat transfer modes; there are

significant approximations in these numbers which are difficult to quantify.

It is probable that both convection and conduction have been somewhat

overestimated, given the assumptions noted above. It is probably reasonable

to conclude that radiation is dominant at low flow velocities and that the two

modes are comparable at high flow velocities. This is not necessarily what

one might expect, however, and this issue is worth revisiting when more solid

data are available on surface and gas temperatures in the U channel. This

issue is worth pursuing because it makes inferences about other configurations

and sample dimensions possible. For example, if radiation is the dominant

heat transfer mode, to a first approximation, smolder will propagate just as fast against the air flow as with it. These issues will be examined in future

tests

.

3.7 Gas Phase Pyrolysis and Oxidation Kinetics

A further goal of the present study is to estimate the conditions of

temperature, time and oxygen level required to destroy the various species

evolved from wood burning. Such results could give the wood stove designer

32
target conditions which his design must shoot for in order to assure minimal

air pollution.

Attempt at direct measurement . The original intent in this study was to

perform experimental measurements of the rates of gas phase thermal pyrolysis

and oxidation for about half a dozen of the species most prominent in the cold

trap condensate. A laminar flow reactor was constructed for this purpose;

such devices are relatively simple and inexpensive, which probably accounts

for their continuing popularity for making measurements such as those planned

here. Such reactors are typically constructed of small diameter tubes (in the

present case, 1 mm and 2 mm ID) to assure good temperature uniformity over the

tube cross section (this is necessary with highly exothermic oxidation

reactions) . The tube surface is heated uniformly to some desired temperature

and the gas mixture of interest is passed through the tube at a measured flow

rate so that the residence time is well-defined. The composition of the

products leaving the reactor is then analyzed; from various series of such

experiments one can infer the kinetic rate parameters for the degradation of

the compound of interest.

Since the reactor tube is small in diameter, the gas molecules have a significant probability of striking the wall one or more times during their

passage through the tube. In order to minimize the probability that such wall

collisions would result in catalysis, the wall was made from pure gold, a material noted for its inertness in many contexts. When it became apparent

that the wall was in fact causing heterogeneous reactions despite the use of

gold, the reactor was designed to accept two different diameters of tube so that, in principle, the homogeneous and heterogeneous contributions to a measured net change in the concentration of the species of interest could be

separated; only the homogeneous contribution is pertinent to the present

study. Further tests with this double tube reactor revealed unexpected

complexity, however: pyrolytic reactions appeared to be accelerated by the

wall while oxidative reactions were retarded. This raised the question of the

separability of the two modes of reaction (wall vs. gas phase); it was quite

possible that they were interacting in a non-linear manner. In view of these

complications , the limited time and the availability of an alternative means

33
to estimate the kinetic rate parameters, this experimental effort was

terminated.

Prediction of the rate parameters. The rate of disappearance of a chemical

species in an isothermal, well-mixed system containing oxygen is the result of

two parallel modes of chemical change, thermal pyrolysis and oxidative attack.

The first process is unimolecular under the conditions here and thus depends

only on the concentration of the species of interest. The second process is

potentially the result of attack by a variety of radical species on the target

molecule and there are, therefore, several further parallel paths that depend

both on the concentration of the target molecule and on the concentration of

the attacking radical (bimolecular reactions). Thus one can write for the net

rate of disappearance of the target molecule:

d[M]/dt = -ku [M] - [MJSk^ [RJ

i

where [M] is the concentration of the target molecule (mol/liter), t is time,

ku is the rate constant for unimolecular decomposition, i is the rate

constant for the i th reaction that involves attack by radical Ri

. Rearranging

this slightly, one sees that the rate of disappearance of the target molecule

is characterized by a time constant r expressed as follows:

r = l/{ku + SkbJRJ}

i

In an isothermal flow, this is the time for the concentration of the target

molecule to decay to 1/e of its initial value. In the absence of oxygen, only

the first term in the denominator applies since only unimolecular

decomposition occurs. In the presence of oxygen both terms apply.

For the present purpose of estimating the rate of disappearance of pollutant

species in the gas phase of a wood burning stove, one can assume that the

principal oxidizing radical in all cases is the hydroxyl radical [29, 30].

One then needs an expression for the concentration of this radical as a

34
function of temperature. The following applies within about a factor of three

for a wide range of fuel -air mixture ratios [30]

[OH] = 10' 1 7 exp( -17600/T) (mol/liter)

where T is the absolute temperature in Kelvins. This is an equilibrium value

of OH concentration established by various rapid elementary reactions in the

radical pool.

In order to proceed further and estimate the net rate of target molecule

disappearance, one needs estimates also for the rate constants. Data for ku and kbi were obtained from the literature or were developed by estimation

techniques described in Appendix A of this report. These data are summarized

for the selected compounds reported in Table 2. The species there include

several identified in our previous study together with a selection of relevant

aromatic species. For each chemical compound, the selected values for the

rate constants, ku and kbi , together with computed values for (0H) EQ were

incorporated into the above equations to develop values for r. These computed

values for each selected compound are shown in Figure 27, in which the

predicted time- temperature degradation behavior of these compounds is

illustrated for 63% decomposition, the so-called "l/e M degree of degradation.

Corresponding degradation behavior for any other degree of decomposition can

be obtained by recomputing r for each temperature for pyrolysis or oxidation

according to the expression,

r = k“ 1 ln(l/(l-a)

)

In the equation above, a is the degree of decomposition which can be specified

by dividing the percent of decomposition by 100.

Figure 27 shows separately the temperature -dependent time to destroy a target

molecule by either pyrolysis or oxidation. Note that the requisite oxidation

conditions for all of the species fall within a band that is about one order

of magnitude wide; the pyrolysis conditions vary much more widely with the

35
chemical identity of the species. Except at high temperatures, oxidation is

predicted to be the primary path of destruction.

Also indicated in Figure 27 is the approximate domain of time and temperature

in which one can expect typical woodburning stoves to operate if the air

supply is somewhat restricted. The upper temperatures certainly include

flaming combustion; the lowest temperatures pertain to the pure smoldering

mode examined in this study. The residence times are estimated from stove

volume and typical air flow rates. It is worth noting that for this range of

residence times the predicted temperatures necessary to oxidize the various

pollutant species are in approximate agreement with the rule of thumb

sometimes used by stove designers (1100°F - 870 K; ref. 31).

It should be quite clear from Figure 27 that the temperature- time domain of

operation of a wood stove, especially in the smoldering mode, is far from that

necessary for effective destruction of the pollutant species generated by the

wood burning process. On this Figure, the only realistic way to move in order

to achieve pollutant destruction is upward in temperature, since the

horizontal scale is logarithmic. To do this one needs to get all of the gas

in the stove essentially up to the same temperature as the actively oxidizing

char surface. With the usual random arrangement of wood logs in the firebox,

there is simply not enough contact between the gas and the solid phases to

approach this thermal equilibrium. Such contact requires re-sizing of the

fuel down to small chunks, forming a packed bed through which all of the

incoming air flow must pass. However, there are problems with using a packed

bed fuel configuration. If the air moves through the bed in the same

direction as the propagating smolder oxidation zone (forward smolder), the

gases must pass through cooler regions of the fuel bed where they will emerge

contaminated with pyrolysis vapors. If the air moves in the opposite

direction (reverse smolder) , there is a tendency for the smolder reaction zone

to be incomplete in its consumption of the solid [7]

.

Pre-heating the air is

of some help in boosting the emerging gas temperature [31], but extensive preheating requires significant re-design of the woodstove.

36
The conclusion which the preceding discussion rather insistently points to is

that effective elimination of pollutants generated in a woodstove operated in

a smoldering mode requires either some drastic re-design of the stove (not

envisioned above) or the use of some form of afterburner on the exhaust.

Given the large gap in Figure 27 between the typical operating conditions for

smoldering and the estimated requisite conditions for pollutant destruction,

it is unlikely that refinements in the estimates of the latter would alter

this conclusion.

3.8 Conclusions

In the apparatus used here, the forward spread of smolder over solid wood was

self-sustaining only over a rather narrow range of air inflow rates. The

lower limit is probably extendable by further thermal insulation around the

sample and/or by the use of thicker wood; the upper limit, due to the

transition to flaming, is probably more inherent in the behavior of wood and

thus is probably independent of the apparatus design.

Both red oak and white pine behave in a similar manner in this forward smolder

mode and it appears that the smolder process tends to sweep out a constant

volume of wood as a function of time, independent of the wood properties. The

minor role for heat conduction through the wood in causing smolder to

propagate is one reason for this; the equal temperatures” achieved on the

oxidizing char surface is the other. The latter fact evidently reflects equal

amounts of heat evolved per gram of oxygen reaching the char surface of either

wood. The essentially linear increase in smolder velocity, surface

temperature , rate of heat and product release with increases in air inflow

rate also reflect a pivotal role for the rate of oxygen supply to the

oxidizing char surface. However, the failure of the char oxidation zone to

grow to the point where it was long enough to consume all incoming oxygen

implies that oxygen supply rate has a less than total control over the rate of

smolder propagation. This is in contrast to the complete dominance oxygen

supply rate has in one -dimensional forward smolder through a packed fuel bed.

In the present configuration, the failure to consume all oxygen points to a key role for downstream heat transfer rate in the propagation process. This

37
heat transfer was seen to be dominated by radiation, at least at low air flow

rates; at higher flow rates convection appears to be equally important.

The major products of the wood smolder process appear to be essentially the

same here, in the previous forced gasification experiments and in a wood stove

operated in a smoldering mode. The ratio of CO to C02 does vary among these

various gasification circumstances and this appears to reflect varying degrees

of CO oxidation. This oxidation appears to be the only significant change

that occurs to the products evolved from smoldering wood as they pass through

a wood stove. Estimates of the kinetic rate parameters for the oxidation and

pyrolysis of several of the species in the tar confirm that they are not

subject to destruction at the combination of temperatures and times

encountered in conventional wood stoves. Gas phase destruction is possible

only at higher temperatures and these are probably achievable only through

thee use of an auxiliary heat source such as an afterburner, if a stove is to

be operated in a smoldering mode.

Acknowledgements

Mr. Scott Dolan provided extensive assistance in setting up and performing the

experiments as well as in the chromatographic analyses. This work was

supported in part by the Department of Energy, Office of Buildings and

Community Systems, under Contract No. DE-AI01- 76PR06010

.

38
REFERENCES

[1] Ohlemiller, T. , Kashiwagi , T. and Werner, K. , "Products of Wood

Gasification", National Bureau of Standards NBSIR 85-3127, April 1985 .

[2] Ohlemiller, T. and Rogers, F. , "Smoldering Combustion Studies of Rigid

Cellular Plastics", Princeton University Department of Mechanical and

Aerospace Engineering Report No. 1432, May, 1979 (Final report to the

Products Research Committee)

.

[3] Thring, M. W. , The Science of Flames and Furnaces . Chapman and Hall , London (1952) p. 134 ff.

[4] Spalding, D., Some Fundamentals of Combustion . Butterworths , London

(1955) p. 98 ff.

[5] Palmer, K. , Combust. Flame 1, (1957) 129.

[6] Ohlemiller, T. , "Forced Smolder Propagation and the Transition to

Flaming in Cellulosic Insulation", National Bureau of Standards NBSIR

85-3212, October 1985.

[7] Ohlemiller, T. and Lucca, D. , Combust. Flame 54, (1983) 131.

[8] Zabetakis , M. , "Flammability Characteristics of Combustible Gases and

Vapors", Bureau of Mines Bulletin 627 (1965).

[9] Alexopoulos , S. and Drysdale, D., University of Edinburgh, private

communication, 1987.

[10]

Cooke, M. , Allen, J., and Hall, R. , "Characterization of Emissions from

Residential Wood Combustion Sources", in Residential Solid Fuels.

Environmental Impacts and Solutions . Cooper, J. and Malek, D. (editors),

Oregon Graduate Center, Beaverton, Oregon (1982), 139.

39
[11] Anon., "Residential Wood Heater Test Report", Phase II, Volume I,

Tennessee Valley Authority, Energy Use Test Facility, CHattanooga,

Tenn. , August 1983.

[12] Tillman D. , Rossi, A., and Kitto, W. , Wood Combustion: Principles

Processes and Economics . Academic Press, New York (1981), p.43.

[13] Magnani, C. and Farouq-Ali, S., Soc. Petroleum Eng. J. . October, 1975,

p.425.

[14] Krukovskii, V., Pitin, R. and Farberov, I., "Underground Processing of

Fuels", translated from Russian and published by the Israel Program for

Scientific Translations, Jerusalem .Israel (1963), p. 91 ; available

from the Federal Scientific and Technical Information Clearinghouse,

Washington, D.C.

[15] Roberts, A. and Clough, G. , Combust. Flame 11 . (1967), p.365.

[16] Dryer, F. , "High temperature oxidation of carbon monoxide and methane in

a turbulent flow reactor" , Princeton University Aerospace and Mechanical

Sciences Dept. Report No. T-1034 (1972).

[17] Laurendeau, N. , Prog. Energy Combust. Sci. 4, (1978), p. 221.

[18] Kashiwagi. T. , Ohlemiller, T. and Werner, K. , Combust. Flame 69, (1987),

p. 331.

[19] Ahmed, S. and Beck, M. , Carbon 23 . (1985), p.513.

[20] Degroot, W. and Shafizadeh, F. , Carbon 21, (1983), p.61.

[21] Shafizadeh, F. , Bradbury, A., Degroot, W. and Aanerud, T. , Ind. Eng.

Chem. . Prod. Res. Dev. 21, (1982) p.97.

[22] Boroson, M. , "Secondary reactions of tars from pyrolysis of sweet gum

40
hardwood", Ph. D. thesis, Dept, of Chemical Engineering Massachusetts

Institute of Technology, May, 1987.

[23] Gann, R. , "Comparison of principal combustion products from a wood stove

as a function of height in the chimney", Fiscal Year 1982 Final Report

from the National Bureau of Standards on DOE Project No. 133a,DE-A01

- 76PR06010 , October 1982.

[24] Muhlbaier, J., "A characterization of emissions from wood-burning

fireplaces", in Residential Solid Fuels , ibid, p.164.

[25] McCaffrey, B. , "Purely Buoyant Diffusion Flames: Some Experimental

Results", National Bureau of Standards NBSIR 79-1910, October 1979.

[26] Huggett, C., Fire and Materials 4. (1980), p.61.

[27] Kanury, A. and Blackshear, P. , Combust. Sci. Technol, 1, (1970), p.339.

[28] Eckert, E. and Drake, R. , Analysis of Heat and Mass Transfer . McGraw-Hill, New York, (1972), p.622.

[29] Benson, S. Twenty-First Symposium (International) on Combustion (1986),

proceedings in press.

[30] Tsang, W. and Shaub , W. , "Chemical processes in the incineration of

hazardous materials", in Detoxication of Hazardous Waste . (J. Exner,

ed.), Ann Arbor Science, Ann Arbor Michigan (1982), p.41.

[31] Dobson, L. , "High-tech non-catalytic woodstove design considerations",

in Proceedings of the International Conference on Residential Wood

Energy, Reno, Nevada, March, 1986.

41
Normalized

Area

Percentage

of

Some

Tar

Components*

bOOPdEo

co p

r^ CM o NO VO CO 42 f—

m p • o CM CO CO O • • CM O • • • • co o p rH o 1“

p d

£2 t—

42

CO PE d O o P

4-1 G 00 ON d f—

ON P o I—

m G • VO o 00 CM r^ m d

00 o • CM • • • • • P £ rH o f—

00 O o o 3 O CO rP iP P d

CP

co

4«S CM d

00 00 o VO ON CM d * 00 • o CO VO m 00 . a * • • o • • • • • e> VX) o r—l o CM »P ?p

ip fH i—

rP H

coWH P

42 bO

Ch P d d P SOO d COd

4-1 £ O P d *P

a>P P d d G CP d G 4-J G •P Ip 42 d •H P a o rH d p P G

(2 42 T“^ < G P o d d a bO d d O 1—1 d >n42 d p d •p • C G r—

O o o bO P O G d co •'“> CO rP o iP o d CO •H 3 3 O d 3 G O d c •p o p d 4-J P 3 d G •iH o • E P d d d d O P < r—l 42 P

4-i IP p o 25 42 3 p o CO 3 o H E * d u CP CP O o < ip IP H o * 0d o • (D C T3 p d a; d 42 p p 3

co a) i—

a) £ a) bOd d « P C P G d d P O P d POP d C EOP

4-1coppdo

•p

4-1 •H

COd P 4-1 bO d d p c6 G o d

co p p C O p ON 00 00 o vO r^ CO p P

in CM • o in <± d O

CO • • o • i • • 42 P 4-1 £2 P o rH l o O rH p d

CO 42 4-1 O P O P

at CO d d p C c > d o, 0 H VO o •p > •H

CP f—r»* ON o vX> •sl- r—l p d p

NO p • o O ip CM 00 d £ •p d CO • • o • • • • • P o G p £2 p o rH rH o o o d 42 C

*tH CO p o

•V O 3 G d d p P E P o CO o 3 C d p « m o m o •o CO CO a p

IP * CO • o VO CO E o • • o • i • • • O d

Cp o o rH i o o rH O P COdo

a>

cod

42

a> 42P

cod

•fH
TABLE 2: Summary of Rate Constant Data

Compound

m #

benzene 1

toluene 2

phenol 3

o- xylene 4

m- xylene 5

p -xylene 6

naphthalene 7

phenanthrene 8

anthracene 9

biphenyl 10

benzo (a) pyrene 11

2 -me thy1-

naphthalene

12

2 , 3 - dimethyl

naphthalene

13

o-cresol 14

m-cresol 15

p-cresol 16

benzaldehyde 17

furfural 18

furfuryl

alcohol

19

creosol 20

vanillin 21

U

( sec

5 . 75X10 1

6

exp ( - 58400/T)

3 . 16X10 1

5

exp ( -44400/T)

3 . 16X10 1

5

exp ( -43500/T)

2ku ( 2

)

ku (4)

ku (4)

1 • 33ku (1)

1 • 67ku (1)

1.67ku (l)

1.67ku (l)

1 • 83ku (1)

ku (2)

2ku (2)

ku

( 2 ) +ku( 3

)

ku d4)

ku (14)

10 1

6

exp( -44000/T)

3 . 16X101

5

exp( - 36400/T)

3 . 16xl015 exp(-38700/T)

1 . 17x10 1

6

exp ( - 33100/T)

ku (20)

-kbi (T>300°C[573°K]

)

(L-mol" 1

• sec

' 1 1 . 9 6xl03 T

2 exp ( - 344/T)

4 . 57xl03 T

2 exp (+11/T)

1 . 81xl03 T

2 exp ( - 100/T)

1 . 05xl04 T

2 exp (

- 35/T)

1 . 03xl04 T

2 exp( -127/T)

1 . 05xl04 T

2 exp ( - 9 9 /T

)

6 . 75xl03 T

2 exp ( - 969/T)

4 . 97xl03 T

2 exp (

- 1000/T)

6 . 63xl03 T

2 exp ( - 1000/T)

3 . 92X103 T

2 exp ( - 1000/T)

5 . 66xl03 T

2 exp (

- 1000/T)

3 . 31xl04 T

2 exp (- 100/T)

4 . 88xl04 T

2 exp (- 100/T)

2 . 53xl04 T

2 exp (

- 100/T)

3 . 61xl04 T

2 exp (- 100/T)

2 . 77xl04 T

2 exp (- 100/T)

1.02xl04 exp (- 100/T)

kt, i (17)

kbiW *

3 . 98xl04 T

2 exp ( -100/T)

6 . 63xl03 T

2 exp ( - 100/T)

^Numbers in parentheses refer to the compound with that number; thus ku (1)

means the value of ku for compound 1.

43
APPENDIX A

ESTIMATION OF KINETIC RATE CONSTANTS

Rate constants, ku and Kbi , were either determined from literature reports or developed by estimation methods for the twenty one organic compounds listed in

Table 2 of the main report. The details of the development of these data are

described below. It is unlikely that errors in estimating ku and kbi would

lead to conclusions contrary to those developed in this report. An order of

magnitude error in estimating the rate constants would not result in a corresponding shift in the time- temperature curves, such that they would

overlap the smoldering wood-burning stove domain illustrated in Figure 27 of

the main report. Care should of course be exercised in utilizing the

estimated rate constants for purposes other than that intended in this study.

Rate Constant Development:

1 . Benzene

Rate constants for both ku and k^ i have been reported in the literature for

the temperature domain of interest (1,2). These reported values are used

without modification.

2 . Toluene

Rate constants for both ku and kbi have been reported in the literature for

the temperature domain of interest (2,3). These values are used without

modification.

3 . Phenol

A rate constant for ku has been reported in the literature (4)

.

This value is

used without modification. Atkinson (2) has reported a value for kbi at 298 K

(2. 83x10" 11 cm3 -molec" 1

• sec" 1

) . The low temperature mechanism is dominated by

an addition reaction to form a complex which can either decompose back to the

original reactants or stabilize. At higher temperatures additional reactions

44
are possible, e.g., metathesis (hydrogen abstraction) or elimination

(displacement of hydrogen). Available evidence, based on other reaction

systems appears to favor the metathesis reaction (see for example the work of

Madronich and Felder (5) . Examination of rate constants for hydroxyl radical

reaction with other substituted benzene compounds (2) at T=298 K and at T=1000

K (e.g., data for toluene, o-xylene, m-xylene, p-xylene (2)) suggests that

(1000)/!^ i (298) has a value of the order of 1 and that Ea /R has a value of the

order of 100 K. Using these suggested relationships based upon examination of

literature data, and Atkinson's formulation for an expression describing the

temperature dependence of kbi , (AT2 exp ( -b/T) ) , it can be shown that

kh i -(3 . 0x10” 1

7

exp( -100/T)xT2

) ) cm3 -molec" 1

• sec" 1 is an estimate for reaction

of hydroxyl radical with phenol at higher temperatures than about 450 K. The

lower temperature cutoff is approximate and reflects the complex nature of the

reaction mechanism at intermediate temperatures (2)

.

4. o-

,

m-

,

and p-xylene

An estimate for ku for these compounds is that ku is about twice that reported

for toluene, since, in the former cases, each compound has two methyl group

substituents and, in the latter case, toluene has one methyl substituent.

Rate constants for kbi for each compound have been reported in the literature

(2)

.

These values are used without modification.

5 . Naphthalene

An estimate for ku for naphthalene is that ku is -1 . 3xku (benzene) . The rough

adjustment is to allow for the differing number of hydrogen atoms. More

refined estimates are possible but are unnecessary for the purposes of this

study. A rate constant for kbi has been reported in the literature (2) and is

used without modification.

6 . Phenanthrene

An estimate for phenanthrene is that ku =l . 6 7ku (benzene) . The adjustment

45
follows the reasoning given for estimation of ku (naphthalene) . A rough

estimate for khi can be made as follows:

Using reported ionization potentials for benzene (IP=9.24 eV) and napthalene

(1=8 .15 eV) (6) plus kbi values for these compounds for at T=1000 K (2,7), a plot of log10 kbi (1000 K) versus IP can be constructed. It has been observed

(2) that such linear plots result for homogeneous series of compounds. Then

using the reported value of IP=7.86 eV (7) for phenanthrene , an estimated

value for kbi (1000 K) for phenanthrene can be derived from such a plot. This

is found to be k^ i =4 . 99x10" 12 cm3 -molec" 1• sec

-

. For unsubstituted aromatic

compounds it appears that as an estimate, E/R - 1000 K, and using a fitting

function (2) of AT2 e -b/T

, a value for A can be found, A=8 . 25x10" 1

8

cm3 -molec'

1

• sec

-

• deg

. The rate constant expression that can be constructed from this

analysis (see Table 2 of main text) can be used to check the literature value

reported for kbi (phenanthrene) at T=748 K (2.2xl0-2cm3 -molec" 1

• sec

-

(2)

.

The calculated value found from the estimated expression is found to agree

well with the literature value, the former being found to be ± (phenanthrene

,

estimated)=2 . 4x10" 1

cm3 -molec• sec" 1

•

. The estimated expression for k^ i was

assumed useful without modification for T > 600 K. Below this approximate

temperature range, competition with other reaction channels causes increasing

uncertainties in the use of this expression.

7 . Anthracene

An estimate for anthracene is the ku is -1 . 6 7ku (benzene) . The adjustment

follows the reasoning given for estimation of ku (naphthalene) . A value for

kbi has been estimated according to the method outlined for phenanthrene. In

making this estimate a value of IP=7.41 eV was used for anthracene (6).

8

.

Biphenyl

An estimate for biphenyl is that ku is -1 . 6 7ku (benzene) . The adjustment

follows the reasoning given for estimation of ku (naphthalene) . A value for

kbi has been estimated according to the method outlined for phenanthrene. In

46
making this estimate a value of IP=8.27 eV was used for the biphenyl molecule

( 6 ).

9. Benzo (a) pyrene

An estimate for benzo (a) pyrene is that ku is «1 . 8 3ku (benzene) . The adjustment

follows the reasoning given for estimation of ku (naphthalene) . A value for

kbi has been estimated according to the method outlined for phenanthrene . In

making this estimate a value of 7.63 eV was used for the IP of benzo (a) pyrene

(6). The value required for kbi at T=298 K which is required in the

calculation method outlined above was obtained from a plot of log^lc^ at

T=298K versus IP for benzene and naphthalene, resulting in a value at T=298 K

of kjj

i (benzo (a) pyrene) = 3. 16x10' 11 cm3 -molec" 1

• sec' 1

. The IP used for

benfco (a) pyrene in the calculations is an average of three reported values:

7.56, 7.60 and 7.73 eV (6)

.

10. 2 -methyl naphthalene

An estimated value for ku is ku — ku (toluene), since both compounds have one

methyl group from which a hydrogen atom can be eliminated. A value for i was derived using the reasoning applied in developing a value for kbi (phenol).

In making the estimate, a value of 'khi (2-methyl naphthalene)= 5.23X10' 11 cm3 - molec' 1

• sec" 1 at T=295 K was utilized (2).

11 . 2,3- dimethyl naphthalene

An estimated value for ku is ku - 2ku (toluene) . The reasoning for this

estimate follows that for o -xylene. A value for kbi was derived using the

reasoning applied in developing a value for kbi (phenol). In making the

estimate a value of t (2 , 3-dimethyl naphthalene)= 7 . 68x10' 1

1

• cm3 -molec" 1 sec'

1 at T=295 K was utilized (2)

.

12. o-

,

m-

,

p-cresol

47
Since the cresols each have one methyl and one hydroxyl substituent groups

having approximately comparable elimination rate constants, following the

reasoning applied to the xylenes, an estimated value of ku for the cresols is

ku — ku (toluene) + ku (phenol) . Estimated values for k^ i were derived using

the reasoning applied in developing a value for kbi (phenol). In making the

estimates for kbi at T=298 K, the values used were 4.0xl0-11

, 5.7xl0-11 and

4.4xl0_11 respectively for the ortho-, meta- and para-cresol isomers (2).

13 . Benz aldehyde

The activation energy for ku can be estimated from a calculation of heat of

reaction estimated from thermodynamic parameters abstracted from the

literature (3,8) in which it is assumed that a hydrogen atom is eliminated

from the formyl substituent rather than from the aromatic ring. This pathway

is favored by the formatioh of the resulting resonance stabilized radical (9)

.

A pre- exponential factor of 10+16 sec" 1 is estimated (13). An estimated value

for kbi was derived using the reasoning applied in developing a value for

kbi (phenol). In making the estimate a value for kbi of 1.6xl0_11

cm3 -molec" 1• sec" 1 at T=298 K was used (10).

14 . Furfural

An approximate lower limit for the activation energy for ku was estimated by

first examining differences in the heats of reaction calculated (3,8) for

elimination of OH from (CH3 )0H and the compound (C 6 H5 )CH2 0H, respectively.

The difference in heats of reaction was taken as an approximate aromatic ring

stabilization energy correction factor. This correction factor (also

derivable from comparison of methane and toluene) was applied to an estimated

heat of reaction for hydrogen atom elimination from formaldehyde to obtain the

activation energy for hydrogen atom elimination from the formyl group

substituent in furfural. This activation energy may be an underestimate, a higher value being derived assuming Ea for ku (furfural) is equal to the

activation energy associated with the unimolecular dissociation of

benzaldehyde . Again, however, this difference in estimates has no effect upon

the possibility of overlap in the time- temperature curves shown in Figure 27

and discussed in this report. The pre -exponential factor was assumed to be

48
slightly less for furfural than for benzaldehyde (5)

.

An approximate estimate

for kbi (furfural) is that it is approximately equal to kj^

i (benzaldehyde) , the

two molecules having the same substituent group attached.

15. Furfuryl alcohol

The activation energy for ku was assumed comparable to the activation energy

for hydroxyl radical elimination from (C6 H5 )CH2 OH. For the latter reaction,

the bond dissociation energy is about 77 kcal per mole (8)

.

The preexponential factor was set similarly to that for furfural. The bimolecular

rate constant was assumed equivalent to that for phenol by analogy to the

comparison between kbi for benzaldehyde and furfural.

16 . Creosol

The rate constant for ku is assumed equivalent to that for the molecule

(C6 H5 )0-CH3 for the indicated carbon-oxygen bond rupture. The latter rate

constant has been derived in a shock tube study (11) . An estimated value for

kbi was derived using the reasoning applied in developing a value for

khi (phenol). In making the kbi estimate, a value for kbi (creosol) at T=298 K

had to be derived. This was calculated according to the procedural method

developed and described by Zetzsch (10). As discussed in Atkinson's report,

the rate constant can be expressed by Zetzsch' s method as,

l°gi 0 kbi(298) = -11.64-1.31 ScrJ

In the equation above, substituent constants, cr+ , are selected which produce

the largest negative value for the substituent constant summation term, Za+

. The values of a

+ used for estimating kbi (creosol) were taken from Zetzsch'

s

paper except for a

+ of the m-hydroxyl group, a+ - 0.01, which was found by

comparing values for ct + for m- , p-amino, p-hydroxy then interpolating between

substituent values for the amino- and me thoxy- group values using the known

value of a

+ for p-hydroxy (12)

.

17. Vanillin

49
The rate constant, ku , for vanillin was assumed equivalent to that for

creosol, the same bond undergoing rupture. An estimated value for ; was

derived using the reasoning applied in developing a value for kbi (phenol). As

with creosol, in making the k^ estimate, a value for k^ L (vanillin) at T=298 K

had to be derived. As with creosol, Zetzsch's method was applied (2,10), with

a

+ substituent constants obtained from reports of Brown and Okamoto (12) and

Zetzsch (10), except for a

+ for the formyl substituent, - 0.46, which was

estimated by making comparisons, as was done for creosol.

50
REFERENCES. APPENDIX A

[1] Hsu, D.S.Y., Lin, C.Y. and Lin, M.-C., 20th Symposium (International)

Combustion . The Combustion Institute, Pittsburgh, PA, pp623-630 (1984)

[2] Atkinson, R. , Chem . Rev . 85 . 69-201 (1985); and references cited

therein.

[3] Benson, S.W. and O'Neal, H.E., Kinetic Data on Gas Phase Unimolecular

Reactions . NSRDS -NBS - 21 , National Bureau of Standards, Washington, D.C

(1970)

.

[4] Shaub , W.M. and Tsang, W. , Environ. Sci. Tech. 17 . 721-730 (1983).

[5] Madronich, S. and Felder, W. , J . Phvs . Chem . 89 . 3556-3561 (1985).

[6] Rosenstock, H.M. , Draxl, K. , Steiner, B.W. and Herron, J.T., J . Phvs

.

Chem. Ref. Data 6 (Suppl. No. 1), (1977).

[7] Biermann, H.W._, MacLeod, H. Atkinson, R. , Winer, A.M. and Pitts,

J.N.Jr., Environ. Sci. Tech. 19 . 244-248 (1985).

[8] Benson, S.W., Thermochemical Kinetics . John Wiley (1976).

[9] Stein, S.E., Center for Chemical Physics, NBS, private communication

(1986)

.

[10] Zetzsch, C., Abstract A-ll (p29) in: 15th Informal Conf. on

Photochemistry . Stanford, CA, June 27-July 1 (1982).

[11] Hsu, D.S.Y., private communication

[12] Brown, H.C. and Okamoto, Y. , J. Amer. Chem. Soc. 80, 4979-4987 (1958).

on

51
[13] Tsang, W. and Shaub , W. in: Exner, J.H. (editor), Detoxication of

Hazardous Waste . Ann Arbor Science, Ann Arbor (1982).

52
APPENDIX B

APPLYING SIMPLIFIED WOODSTOVE MODELS

Viewed as a chemical reactor, a wood stove is an extremely complex device. It

turns two reactants, one fed batchwise (wood) and one fed continuously (air),

into a two -phase mixture of hundreds of products (smoke) . The interior of the

reactor is characterized by poor mixing and sharp gradients in temperature and

chemical reaction rates. Modeling this system so as to predict the rate of

evolution of a given product is a task beyond the current state of the art, in

the absence of drastic simplifying assumptions. Here the problem is briefly

addressed in terms of the two simplest possible models of a chemical reactor,

the isothermal plug flow reactor and the perfectly stirred reactor [1]

.

The isothermal plug flow reactor can be visualized as a simple tube along

which the reactants flow as they are converted to products . The reactants are completely mixed and any heat they evolve as they react does not change the

local temperature significantly. (If the reactor is adiabatic instead of

isothermal, one has, in effect, a pre-mixed laminar flame.) Such a model is a very crude approximation of the behavior of the gas phase in a wood stove,

especially since a plug flow path through the firebox is hard to identify.

If, nevertheless, one wished to apply such a model to make some very rough

estimates of product evolution rates, the results in this report and in the

previous related report are of some use in this regard for the smoldering

operation mode. Suppose that one has estimated, within an order of magnitude,

the effective plug flow velocity and path length through the stove of

interest. One then needs the concentration of wood gases which follow this

path in accompaniment with the assigned air inflow rate. Figure 7 in the main

body of the report gives the relation obtained here between the rate of air

inflow and the rate of wood gasification. This applies strictly only to the

configuration examined here but it should apply rather closely to a set of

parallel channels subject to comparable air flow per channel; this provides a rough basis for applying this result to a less organized load of wood. One

then has an estimate of the mixture of fuel and air entering the gas phase

plug flow reactor.

53
Suppose that the component of interest is in the tar. Figures 20 and 21 give

an estimate of the fraction of gasified wood which is tar; if the component is

one of those identified in the previous work (ref. 1 in main body of this

report) , the data accompanying Figure 18 of that report (see also the

discussion of that Figure) provide an estimate of the fraction of that

component in the tar. Thus one has an estimate of the concentration of the

component of interest entering the plug flow reactor.

To calculate the amount of this component which survives passage through this

reactor requires that one pick a temperature for the reactor and then utilize

the estimated kinetic rate parameters in Table 2 of the main body of this

report. For the estimated time that the gas will spend at this temperature

and the net pyrolytic and oxidative kinetic rate constant obtained from Table

2, one then gets the fraction of the species which is destroyed by plugging

these numbers into the last equation in Section 3.7 and solving for the value

of a

.

A perfectly stirred reactor is another idealization of reality. Here it is

assumed that the reactants mix instantly throughout the reactor volume;

composition and temperature are thus uniform throughout. The reactants, on

average, spend a time in the reactor equal to the reactor volume divided by

the volumetric flow rate of the reactants. From a simple species balance one

can show the change in the concentration of the reactant between entering and

exiting the reactor is [1]

AY = (Vr /F) -RR(T)

Here AY is the change in mass fraction of the species of interest, VR is the

estimated volume of the gas phase in the well-stirred portion of the firebox,

F is the mass flow rate of the reactants into the gas phase of the firebox and

RR(T) is the volumetric rate of consumption of reactant mass, obtainable at

any chosen temperature T from the rate data in Table 2. Thus the application

of this simplified model, using the data developed in this study, is quite

similar to that for the plug flow model.

54
References , Appendix B

[1] Smith, J., Chemical Engineering Kinetics . Second edition, McGraw-Hill

Book Company, New York (1970), Chapter 3

55
e*02

56

Figure

1.

Wood

sample

configuration

and

dimensions.

Actual

orientation

of

sample

in

wood

combustion

chamber

is

same

as

that

shown

here.
Thermocouple

entry

points

57

Figure

2.

Schematic

of

wood

combustion

chamber

with

sample

in

place.

Windows

at

indicated

positions

on

front

plate

protrude

outward

away

from

the

Uchannel

in

the
Air

Figure 3. Schematic of flow system supporting the wood combustion chamber.

58

heated
0 0 cc

_Q (-O

T3 I O D

co©EEc

0)co

'©c©E

"O<

59

Figure

4.

Arangement

of

thermocouple

insertion

points

on

the

top

side

of

the

wood

combustion

chamber.

These

same

spacings

are

assumed

to

hold

at

the

actual

depth

of

the

thermocouple

junctions.
0

RIR VELOCITY (CM/SEC)

5 10 15 20

RIR INFLOH RRTE (G/MIN)

Figure 5(a). Post- ignition smolder propagation rate as a function of air flow along

the U-channel for white pine. The cross-hatching indicates the approximate position

of the borders for extinction and for flaming.

60
RIR VELOCITY (CM/SEC)

0 5 10 15 20

Figure 5(b). Post- ignition smolder propagation rate as a function of air flow along

the U- channel for red oak. The line with the arrows indicates that in the two tests

at or to the right of the line, flaming eventually developed.

61
0

AIR VELOCITY (CM/SEC)

RIR INFLOW RATE (G/MIN)

Figure 6. Maximum temperature on the interior surface of the U-channel as a function

of air flow for red oak and white pine.

62
AVERAGE

AIR

VELOCITY

IN

UNBURNED

CHANNEL

(cm/s)

CM CM

LH

CD

O

i_n

^r

O

rvj

(uiuj/ 6) 3iyy ssoi 1HDI3M 3Ddy3AA

63

AVERAGE

AIR

INELOW

RATE

(g/min)

Figure

7.

Postignition

weight

loss

rate

for

unction

of

air

flow

in

the

U-channel.
o o o

LD

OJO

OJ

LDO

LDO

V) \E

CJ

>-

h-I—

uo

_ju>

0-

_l

Lu

1—i CE

M013NI did 6 / Q3 I J I SdD QOOM 6

*3

S-i c

<U CO >O

cOo

<U

b£) T3

0) a>

J-l S-i 0)> pC

CO p»o

co XI

co

<d S-i o o o 4-1

S-i a

5-1 O

a)

A

-o co

i—!0 co s •rH 0)

CO TJ •rH 0 x: o H

<u . ,c r—

•P cd>

4-i o o s

a) o 5-1 •rH St p 5-1 0

CO 0) fH

S-i •P 4-1

•rH a> C 5-i o b£> •rH c •rH cO a)

r"H o 4-1 cd 4-> 0 >

•H •P C 3 c 0

cr a) •H

<u 3 •P cr o P a) c C co 3

a)

A 4-1 P 3

(0 co cO aa a) CO < e cO

•t—i u (!) C

aD AJ •rH CO a

<0 as u 4-> a> 3 -U &0 fH *rH

•rH r—< r;

(JH Cd 5

64
Q3WnSN0D

2 0 6 / QBIJISbD dOOM 6

65
OJ OJ

o

o

r\-

O

CD

O

o

in

o

o

^r

o

o

m

o

o

ru

o

o

o

Ld

21

hH

H

r/. 3wnnoA) NomsoduoD sud

66

Figure

10.

Composition

of

the

exhaust

gases

from

a

test

of

red

oak

at

13.3

cm/sec

air

velocity.
o

o

in

o

o

^r

o

o

m

o

o

pj

o

o

o

pj pj

Ld

Z

i—

ho: 3wnnoA) Noiiisodwco sud

VO

<u o

}-i a; 3 w

bO \ •H S

fa O

67

11.

Composition

of

the

exhaust

gases

from

a

test

of

white

pine

at

13.

air

velocity.
OJ OJ

o

o

m

o

o

'^r

O

O

m

o

o

OJ

o

o

o

U

21

1—

I

co

(X 3WniOA) NO I 1 I SOdWOO SUD

68

Figure

12.

Composition

of

the

exhaust

gases

from

a

test

of

white

pine

at

15.

cm/sec

air

velocity.
OJ OJ

o

o

LD

O

^r

o

o

m

o

o

oj

o

o

o

Ld

21

i—

hr/. 3wnnoA) NomsodwoD sud

P CO bO \ T3

•h E C

fc o m

69

re

13.

Composition

of

the

exhaust

gases

from

a

test

of

white

pine

at

23.0

ec

air

velocity.

Flames

erupted

at

the

time

indicated

by

the

vertical

line

the

test

was

terminated.
OJ OJ

O

in

o

o

rr

O

m

o

o

ru

o

o

o

Ld

r

f—

f-

(X 3WniOA) NO 1 1 1 SOdMOD SUD

. 70

Figure

14,

Composition

of

the

exhaust

gases

from

a

test

of

white

pine.

The

air

flow

velocity

was

cut

t

abruptly

at

the

time

indicated

by

the

vertical

line

from

22.2

to

14.4

cm/sec.

The

flow

was

cut

because

flaming

appeared

imminent.
o

o

in

o

o

"dO

o

m

O

o

PJ

o

o

o

pj pj

Ld

h-

(X 3WmOA) NOIlISOdWOD SUD

71

Figure

15.

Composition

of

the

exhaust

gases

from

a

test

of

white

pine.

The

air

flow

decayed

slowly

through-out

the

post

ignition

period

from

a

starting

value

of

21.2

cm/sec

to

a

final

value

of

17.4

cm/sec.

The

decay

was

caused

by

narrowing

of

a

point

in

the

exhaust

line

due

to

charred

deposits.
o

o

LD

O

^r

o

o

m

o

o

oj

O

o

LJ

hin

C\J

o

C\J

i_n in

(X dWHlOA ) NOIlISOdWOD SdD

72

Figure

16.

Composition

of

the

exhaust

gases

from

a

test

of

white

pine

at

an

air

flow

velocity

of

8.7

cm/sec.

The

smoldering

was

virtually

extinguished

at

the

time

the

test

was

terminated.
C\J OJ

o

o

LH

O

^r

O

m

o

o

pj

o

Ld

H

(X 3WniOA) NOIlISOdNOD SLID

4-1 CO

5-1 •rH 4-4 CD

5-i X •i—l a) 4J

cd XH O

3 4J

cd

• £

4-> O

cd a) rH

4J 4-1 X cd

cd o 5-1 0 •1—

•rH T) cd T3 3

CD •r4 CD

5-1 X CO 4J

4-1 4J O c bO

•rH 3

4-> o •H

CO a, 4-> <D 4J

4-> CD 3

<u o

cd 5-i x: S u X3 0

5-1 a) CO 4-1 x: T3

4J 3

CO O

<D 4-> O

co Cd <D

cd CO bO X5

<u 3

4J 4-> •rH to (O, 3 3 -a

cd S-i <D X <D 4-> X cd

a) CO 3

CD •rH

a) e E X cd 5-4 4J t—

CD

Cti 4J

4-1O CD

5-1 3 o <D O <D £

•r4 CO

4-1 \ CO •H E <D

CO O T3 0 O

0- CM CO e • •rH o CM P« u CM CD

4-1 bO

• O 3

•H

i—t E

4-> cd

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73

resumed

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o

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pj

o

o

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h-

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74

ure

18.

Composition

of

the

exhaust

gases

from

a

test

of

red

oak

at

an

air

ocity

of

10.4

cm/sec.
o

PJ OJ

U

1—

h-

(y. 3UniOA) NOIlISOdNOD SdD

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,ure

19.

Composition

of

the

exhaust

gases

from

a

test

of

red

oak

at

an

air

ocity

of

18.7

cm/sec.

The

test

was

terminated

when

flames

erupted.
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LD

OJ

o

OJ

in

o

in

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76

Figure

20.

Average

composition

of

products

from

the

surface

of

the

wood

for

the

postignition

period.

Data

for

white

pine

as

a

function

of

air

flow

velocity.
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• o••••••

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77

Figure

21.

Average

composition

of

products

from

the

surface

of

the

wood

for

the

postignition

period.

Data

for

red

oak

as

a

function

of

air

flow

velocity.
in

pj

o

pj

in

in

V) \

E

CJ >

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78
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79

Figure

23.

Average

yield

of

products

from

unit

mass

of

gasified

wood

for

the

postignition

period.

Data

for

red

oak

as

a

function

of

air

flow

velocity.
W

cd

<uW

cd<D r—

0)

5-i P

cd • (D 0) X P

<4-1 a

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a)

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p

o

b£> p

•P 3

(Jh 4-1

( SllbJM ) 3SLJJ“l3d 1U3H JO JldJ JDdJJAU

80
CROSS-SECTIONAL PROFILES OF WOOD WITH

SUPERIMPOSED ISOTHERMS

10.4 cm/s

Air flow

22.2 cm/s

Air flow

it

CD

1

Figure 25. Cross-sectional profiles of red oak at two air flow velocities with

superimposed isotherms. Point at which maximum temperature was first reached is also

indicated. Note also the gaps at mi'd-depth due to shrinkage of the two pieces of

wood making up the bottom of the U- channel.

81

-1/2

cm
HEAT TRANSFER TO WOOD IN FORWARD SMOLDER

Air velocity=22.2 cm/s

Total heat=48 watts

Radiation(27 watts)

Convection

(20 watts)

HEAT 0.5

FLUX Wood \

Conduction(1.2 watts) 50cm

Air velocity=10.4 cm/s

Total heat=13.6 watts

Radiation (11 watts)

HEAT 0.5

FLUX

Convection

(2 watts)

50cm

Conduction (0.6 watts)

Figure 26. Calculated profiles of three principal heat transfer modes for raising

temperature of wood ahead of the point at which the maximum surface temperature is

achieved. Profiles are shown for two air flow velocities, superimposed on a crosssection of the base of the U-channel. The quantities in parentheses are the integrals

under the curves and they pertain to a one centimeter width of the channel periphery.

82
PREDICTED

LIFETIMES

OF

SELECTED

SPECIES

(>l) 3dniVH3dW31

c

o

•rH

4-> cd X3

cd

cd }-i bO

Pm 4->

CM

<UU

3

<4-4 Oc

o

•rH

4-) o

txO C

•H 3

Cti 4-4

83

LOGARITHM

OF

TIME

(sec)

TO

DECAY

TO

(1/e)

OF

INITIAL

CONCENTRATION
NBS-114A (REV. 2 »ac) U.S. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

SHEET (See instructions)

1. PUBLICATION OR

REPORT NO.

NBSIR-88/3767

2. Performing Organ. Report No. 3. Publication Date

May 1988

4. TITLE AND SUBTITLE

Products of Wood Smolder and Their Relation to Wood~Burning Stoves

5. AUTHOR(S)

Thomas J. Ohlemiller and Walter Shaub

6. PERFORMING ORGANIZATION (If joint or other than N BS, see in struction s) 7. Contract/Grant No. NATIONAL BUREAU OF STANDARDS -

U.S. DEPARTMENT OF COMMERCE

GAITHERSBURG, MD 20899

9.

SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS ( Street. City. State, ZIP)

. Type of Report& Period Covered

U.S. Department of Energy

Washington, D.C. 20585

10.

SUPPLEMENTARY NOTES

| | Document describes a computer program; 5F-IB5, FIPS Software Summary* is attached.

11. ABSTRACT (A 200-word or 'tess factual summary of most significant information. If document includes a significant

bi bliography or literature survey, mention it hei&)

The smoldering combustion of solid wood is a process pertinent to both fire

safety and to the generation of air pollutants in wood burning stoves. The

wood sample was in the form of a U-shaped channel 74 cm long with 6.4 cm thick

walls. Smoldering propagated in the same direction as the airflow (forward

smolder) . Tests were conducted with both red oak and white pine and both

woods behaved quite similarly. In separate tests the air flow velocity was

varied from about 9 to 22 cm/sec. At the low end of this range, the

smoldering process was prone to extinction; at the high end it was

increasingly likely to transition into flaming combustion. The smolder

velocity, peak temperature, rate of heat and product evolution all increased

over this flow range in an essentially linear manner. Analysis of the

temperature profiles in the wood pointed to a dominance of radiative transfer

in the smolder propagation process at low air flow rates and roughly equal

roles for radiation and convection at higher flow rates. The principal

species in the tar evolved from smoldering wood resemble, in large degree,

those from previous forced wood gasification experiments and those obtained

from a wood stove operated in a smoldering mode. Estimates of the kinetic

rates of destruction of several of these species by pyrolysis and by oxidation

point to a need for some additional heat source if these species are to be

destroyed by gao phas e proc e s ses — 12. KEY WORDS (Six to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semicolon s) air pollution; combustion product sr ‘smoke; smoldering combustion; wood; wood stoves

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