EVALUATION OF TENSILE FRACTURE IN 
ASPEN USING FRACTOGRAPHIC AND 
THEORETICAL METHODS 
John A. Akande' and George H. Kyanka 
Wood Products Engineering Department 
SUNY, ESF, Syracuse, NY 13210 
(Received June 1989) 
ABSTRACT 
Based on the linear elastic fracture mechanics (LEFM) concept, apparent mode I fracture toughness 
K(cLR was calculated for unnotched aspen (Populus tremuloides Michx.) specimen (5 in. x 0.188 in. 
x 0.094 in. dimension) tested in tension parallel to grain. The purpose was to relate wood fracture 
toughness to average ray spacing. Observations along the longitudinal-radial plane of fracture showed 
the ray/fiber interface as a notable zone of crack propagation. Step-like progression of fracture lines 
occurred from one ray to another thus leading to an hypothesis that the average ray spacing could 
reasonably estimate the critical flaw dimension. Using this model, the calculated k(cLR was close to 
available published data. In additional experiments, it was revealed that k(cLR is sensitive to decay 
weight losses. Within 2% to 16% weight loss examined, loss of k., LR was found to be more pronounced 
with a brown rot fungus when compared to white rot fungi. 
Keywords: Ray, fracture toughness, brown rot, white rot, decay weight loss. 
INTRODUCTION 
Failure of materials, including wood, are often related to their structural ar- 
rangements and associated imperfections. These materials have flaws or discon- 
tinuities which grow, sometimes to catastrophic levels, upon the application of 
stress. Bariska and Kucera (1985) were quick to note that wood anatomy is one 
most important factor affecting its failure morphology. The importance of rays 
in wood deformation processes has particularly generated interest among wood 
researchers. Keith and Cote (1968) reported that boundaries between rays and 
the axially oriented cells are weak points where structural changes primarily oc- 
curred. Keith (1971, 1974) defined the weak points as ray margins while Cote 
and Hanna (1983) described the same area as the ray/fiber interface. Cote and 
Hanna also recognized a stepwise fracture progression from one ray to another 
in red oak subjected to radial shear. To evaluate the role of the rays during wood 
failure, this study sets out to correlate the geometrical layout of wood rays in 
aspen (Populus tremuloides), using the linear elastic fracture mechanics concepts, 
to its fracture toughness. Unnotched tension parallel to grain specimens were 
prepared and tested to failure. Test specimens were allowed to undergo a self- 
initiating fracturing process with a plan that fractography will provide necessary 
clues to apply known theoretical assumptions. It is conceivable that failure across 
the necked down region of a tension parallel specimen could propagate along the 
L T or LR fracture modes. Fractography of an L T fracture plane is expected to 
reveal systematic fracturing from one growth ring to another, which was not 
I Current address: Appalachian Hardwood Center, Division of Forestry, West Virginia University, 
Morgantown, WV 26506-6125. 
Wood and Fiber Science. 22(3), 1990, pp. 283-297 
© 1990 by the Society of Wood Science and Technology 

284 

WOOD AND FIBER SCIENCE, JULY 1990, V. 22(3) 

identified in this study. Fracture progression from one ray to another was, however, 
prominent along the longitudinal-radial plane and, for this reason, k., LR calculation 
was done as if the fracture pathway was coplanar. The assumption was that the 
theoretical stress along the shear plane is negligible. The same assumption had 
been made in previous studies on wood fracture toughness (Porter 1964; Debaise 
et al. 1966). Schniewind and Centeno (1973), Jeronimidis (1976), and Patton- 
Mallory and Cramer (1987) noted that very little systematic work has been pub- 
lished concerning the fracture toughness of wood across grain. This is felt to be 
due to the difficulty in explaining a situation where the introduced notch suddenly 
turns to fracture along the longitudinal plane. A simplified solution to this problem 
may be obtained by using wood fractography as demonstrated in this paper. 
Fracture toughness is a material property that should be sensitive to changes in 
wood strength. Since fungal decay is known to produce significant deterioration 
in wood properties, additional experiments were conducted, using the same model, 
to evaluate the influence of certain decay fungi on wood fracture toughness. 
EXPERIMENTAL 
Materials 
Tension parallel to grain specimens (Fig. 1) were prepared from kiln-dried 
sapwood of Populus tremuloides. Selected specimens were all straight-grained and 
free of defects. These specimens were evaluated as sound wood controls and also 
to study the effect of Trametes versicolor (L.: Fr.) Pilat (SUNY ESF, 40), Bjer- 
kandera adusta (Willd.: Fr.) Karst. (SUNY ESF, 58) and Gloeophyllum trabeum 
(Fr.) Murrill (SUNY ESF, 32) decay at 2%, 5%, 10% and 16% weight loss levels. 
Sample size for each trial was 15, and control specimens were prepared exactly 
the same way as their decayed counterparts, which were subsequently innoculated 
with test fungi and decayed for varying periods of time to achieve the qesired 
weight loss levels. Specimens to be decayed were all moistened and sterilized prior 
to decay experiments. 
Mechanical test 
Test specimens were subjected to axial tension using a floor model TTB Instron 
testing machine. The loading rate was 0.05 inch per minute. To maximize gripping 
friction, the machine tensile test grip surfaces were faced with sandpaper. All tests 
were performed at the laboratory testing temperature of 72 F and 4.2% EMC. 
Load-deflection curves (e.g., Fig. 2) were automatically plotted on the testing 
machine until the first indication of failure was recorded. Usually this happened 
to be the maximum load (Pmax). The failure stress (J" was finally calculated by 
dividing Pmax with the cross-sectional area of the necked-down region 

(J" = Pmax/O.O 18 psi 

(1) 

Fractography 
Fracture surfaces were examined with an ETEC Autoscan scanning electron 
microscope. The fracture surfaces were in a dry state and were merely sputter- 
coated with gold, using a Technics instrument, to enhance contrast and eliminate 
charging effects from high electron potential. Microscopic details of the fracture 

Akande and Kyanka- TENSILE FRACTURE IN ASPEN 

(2) 

285 

	
	
I 

o 
<l: 

o 
--' 
t 

p~ 

---.~ 	DEFLECTION 

FIG. I) Tension II grain specimen design. 2) Typicalload-defiection curve in tension parallel. 

surfaces were examined at 20 K v accelerating voltage. Whenever the object on 
the cathode ray tube will be better demonstrated by stereo photography, adjacent 
pairs of photographs were taken at ± 5° tilt angle. 

Determination of ray spacing 

An Eberbach micro-projector was utilized to enlarge microtomed (20 ,urn thick) 
cross sections from the necked down region of the tension specimens. One hundred 
microtomed sections were prepared and stained with sufranin-aniline. These were 
mounted on glass slides for closer examination under the projector, which was 
magnified 85 x . The average distance between two adjacent rays was obtained by 
dividing the dry state width of the tension specimen necked down region by the 
average number of rays on this narrow cross section. An average of 26 rays were 
counted on each of the cross-sections examined. 

Wood decay 

The purified malt agar-block method (ASTM D-20 17 1988a) was utilized, and 
the decay chambers were made from 16-ounce French square bottles. Two of the 
tension specimens could lie flatwise in the chamber. Fungal incubation was done 
at 28 C for varying time periods (9 days to 10 weeks) to obtain similar weight 
losses for the different fungi. Weight losses were calculated as percentages of the 
original oven-dry weights of the test specimens. 

286 

WOOD AND FIBER SCIENCE, JULY 1990, v. 22(3) 

Specific gravity determination 
After every test specimen was completely failed, five specific gravity samples 
were cut from each treatment category. Their specific gravities were then deter- 
mined using the mercury displacement method outlined in ASTM D-239 5, 1988b. 
The average specific gravity for each treatment group was later calculated. 
THEORY 
Fracture criteria 
Basic linear elastic fracture mechanics concepts relate material properties, flaw 
size, and applied loads to the stress conditions prevailing around crack tips. 

	a2 = 2E"Yshra 	(2) 	
where 			
E 	elastic modulus, 		
"Ys 	surface energy, 		
a 	crack length, 		
a 	fracture stress. 		
This equation satisfies the Griffith (1925) criterion for brittle materials and it is 
one of the earliest propositions concerning the LEFM concept. Irwin (1957) equat- 
ed 2"Ys to G which is the strain energy release rate. 

a2 = GEhra 

(3) 

The quantity (GE)'h is the tendency of the crack to propagate and this is recognized 
as the stress intensity factor K. K also represents the fracture toughness of the 
material which is a property useful in predicting failure loads of wood components 
that contain stress concentrations. 

K = ay.;;:a 

(4) 

Fracture toughness K is recognized to be a function of the specimen geometry, 
applied stress, and the size, location and orientation of the critical flaw (Tada et 
al. 1973). Hence; 

after Pellicane et al. (1982). 
Where 

K = a(7ra)'IJ. f(a, b, c) 

(5) 

K = Fracture toughness, 
f = dimensionless function of specimen geometry, 
	a 	critical (effective) flaw size, 
	b 	specimen width, 
c = constant related to crack location in the specimen. 
Since failure normally occurred at the necked down region of tension parallel to 
grain specimens (see Fig. 1), the specimen width can be taken as nominally con- 
stant. Also because no notch was preintroduced into specimens utilized in this 
study, the elements band c in Eq. 5 may be eliminated to make the most important 
consideration the effective flaw size. If the stress value a at which the critical flaw 

Akande and Kyanka= TENSILE FRACTURE IN ASPEN 

287 

FIG. 3. Stereo scanning electron fractograph of a fractured tip of a sound specimen after failure. 
Notice that the fracture plane formed a step-like progression from one plane of the rays (R) to another. 
The ray/fiber interface appears to be a weak zone where shear failure occurred. 
size is created is ascertained, then it is possible to calculate K. During the fracture 
process, it was observed that wood in tension generally behaved brittle. It was 
also clear that wood cells produced abrupt failure when decayed and this en- 
couraged the belief that decayed wood will comply with LEFM assumptions. As 
a matter of fact, Sinclair et al. (1978) have demonstrated a positive linear rela- 
tionship between fracture toughness and traditional toughness (using a Forest 
Products Laboratory toughness testing machine) in beetle-killed pine wood at 
different levels of decay. 
Fractographic model 
Failure in tension parallel to grain frequently occurred across grain (i.e., along 
the transverse plane). A close look at stereo fractographs of the fracture plane 
(Fig. 3) indicated that failures having shear components occurred along the ray/ 
fiber interfaces. When fracture lines eventually cut across the longitudinally ori- 
ented fiber cells, step-like progression of the fracture plane resulted. The fracture 
line obviously passed through sloping planes, across the fiber cells (see Figs. 4 
and 5a, b), until it becomes full (effective) crack length between rays. By the time 
the fracture line moved from one ray to another in sound wood, an average angle 
of approximately 45° (estimated visually) was projected to the transverse plane. 

288 

WOOD AND FIBER SCIENCE, JULY 1990, V. 22(3) 

FIG. 4. Scanning electron fractograph showing a step-like fracture pattern from plane (A) of ray R 
to another plane (B) in sound wood. The "step angle" was steep. tw indicates transwall failure of the 
longitudinally oriented fibers and vessels. Note that the cell lumens are generally opened during 
transwall failures. 
This step angle appeared to be influenced by the failure angle of individual cells 
illustrated in Fig. 6. After examining several fractographs, it was apparent that 
microflaws exist in wood. Upon the application of stress, small flaws within 
individual fibers would coalesce to form larger cracks whose length becomes 
critical to failure at the measured distance between rays. The critical crack is 
equivalent to the ray spacings and is analogous to the voids between butt joints 
in laminated wood beams. The stress condition at the periphery of the crack tip 
is essential to an understanding of the failure process in wood which contained a 
defined locus of stress concentration like the ray/fiber interface. Any crack tip 
bordering these zones would propagate in an unstable fashion. 
The fracture behavior of wood cells can be analyzed by taking a technique from 
Williams and Birch (1976), who showed that shear failure is less significant when 
analyzing mixed mode fracture in anisotropic media loaded in tension and that 
the fracture behavior can be evaluated as pure mode I. A situation where cracks 
turned to run parallel to the grain direction could be analyzed as pure shear mode 
along the ray/fiber interface. The theoretical stress on this plane of fracture equals 
zero, and the true fracture stress is considered to be resisted by the longitudinally 
oriented cells. Consequently, KII and KIll becomes negligible while KI calculation 
employs the numerical flaw size prior to unstable crack propagation along ray/ 
fiber interface. Since shear failure is not considered significant under tension 

Akande and Kyanka= TENSILE FRACTURE IN ASPEN 
289 
	
-2b 
~ 
a«b 
(Sa) 
t-2a -t 
¢ 
( So ). 

											- 	
									~ 			
												
	r- •........ 						•.. 				
r- 								.., 				
				•.. 				~ 				
				41 								
				~ 								
											rt: 	
								ray 			
					rJ.y 						
		rJ.y 									
1-2a ~ 

y 
	Sound wood ce 115 	Decayed wood ce 115 
FIG. 5. Step-like fracture seen from tangential-longitudinal plane. 
parallel to grain, K, calculation can be done as if the critical crack had propagated 
in a coplanar fashion. 

With regard to wood decay, microscopic evidence in this research (Fig. 7) 
indicated that fracture jumps also occurred from ray to ray, but many of the 
longitudinally oriented cells offered little or no resistance and, therefore, broke 
abruptly at low angles (about 10° or lower) to the transverse plane (see Fig. 8). 

The assumption that the average spacing between rays corresponds to the critical 
flaw dimension is similar to the assumption made by lung and Murphy (1983) 
while examining the effect of butt joints on fracture behavior of parallel laminated 
veneer members. The ply thickness was assumed by these workers to correspond 
to the critical flaw size in their calculation of fracture toughness. This same as- 
sumption was made by Smith and Penny (1980) while analyzing the fracture 
behavior of butt-jointed laminated wood beams. Figure 9a and b show the sim- 
ilarity between lung and Murphy's model and the present model. To calculate 
the apparent fracture toughness in butt-jointed laminated wood, lung and Murphy 
(1983) used the following equation 
	K, = uy:;rt· F(d/t)g(dm/t) 	(6) 

290 

WOOD AND FIBER SCIENCE, JULY 1990, V. 22(3) 
A 

A 

9.' 

Non-decayed cell 

Decayed cell 

N.B. BC is the fracture plane and L ABC is the reference angle. 
FIG. 6. Idealized diagram showing the 'step angles' to failure on individual wood cell. 

FIG. 7. Scanning electron fractograph showing a longitudinal-radial fracture plane in wood decayed 
by Gloeophyllum trabeum to 5% weight loss. Notice a step-like progression of the fracture plane. The 
longitudinally oriented fibers (f) and vessels (V) failed in an abrupt transwall (tw) fashion at angles 
almost parallel to the transverse plane. hy shows hypha in vessel lumen. R indicates ray cell. 

Akande and Kyanka- TENSILE FRACTURE IN ASPEN 
291 

FIG. 8. Scanning electron fractograph showing a longitudinal-radial fracture plane in wood decayed 
by Bjerkandera adusta to 10% weight loss. Observe how the ray cell (R) was degraded (arrows) and 
how abrupt transwall fracture (tw) of vessels (V) and fibers (F) occurred at angles almost perpendicular 
to the grain direction. 
(8) . 
Butt 
joint 
Jung & Murph 
model (1983). 
+ 
+ + 
+ 
+ 
+ 
(b) . 
			I--- 	
		I- 		
	2a 			
	=---.:. 			
~ 				
L 
LT 
		_ t _ 	
		1::=:= 	
~ 	k 	
y 			
Present 
model. 
2a = ray spacing i.e. flaw 
in wood critical to 
failure. 
FIG. 9. Models for fracture toughness determination. 
t = veneer thickness 

292 
where 

WOOD AND FIBER SCIENCE, JULY 1990, V. 22(3) 

	a 	tensile stress, 
	t 	veneer thickness (synonymous with ray spacing in this investigation), 
	d 	specimen width, 
	m 	butt joint spacing, 
	F 	correction factor relating K to t/d ratio for an isotopic material (Tada et 
al. 1973), 
	g 	function to correct for interaction of edge butt joint and butt joint in the 
next adjacent lamina. 
The function F(d/t)g(dm/t) is similar to f(a, b, c) in Eq. 5 and it relates to crack 
geometry, location, and size. The value of these functions for a body loaded in 
tension and containing a central crack is always close to unity (see Pellicane et 
aI. 1982). In a situation where fractography has already provided essential infor- 
mation about the geometry, location, and size of critical flaws, Eq. 6 can be 
condensed to K) = aV7rt, which is the same as Eq. 4, except that t now becomes 
the critical crack length or ray spacing on the fractographs. 
CALCULATION OF KIc LR 
As a first order approximation, the load needed to calculate KIc corresponds to 
the first indication of instability on the load-deflection diagram. As already in- 
dicated, this was Pmax in many of the instances and its average value dropped 
for each of the treatment groups as decay progressed. The average spacing between 
rays is also felt to be equivalent to a critical crack embedded in wood. It was 
therefore equated to 2a. Jung and Murphy (1983) already demonstrated that Klc 
may be insensitive to the exact crack location as long as the crack remains in the 
material interior and perpendicular to the line of applied uniaxial stress. 

2a = 0.0072 in. 

.'. a = 0.0036 in. 

• (7) 

A close look at the fracture line across the longitudinally oriented cells (fig. 4) 
showed that the fracture line is not exactly perpendicular to the fiber axis. In other 
words, the crack length is somewhat different from a, and the effective crack length 
a' = a/cos B (see Fig. 6) where B is the "step angle" across the longitudinally 
oriented cells. Thus, for sound wood where B :::::: 45°, cos B = 0.707 and a' = 
0.0051 in. This same dimension can be taken for 2% weight loss since fractographic 
evidence (Fig. 10) showed that the wood cells still remained sound. Beyond 2% 
weight loss, the worst flaw theory will be assumed because cells that broke abruptly 
from decay were likely to influence fracture pathway. For decayed cells, an angle 
of failure B :::::: 10° will be assumed (see Fig. 6). Cos B = 0.985, therefore a' = 
0.0037 in. 
The fracture stress a psi was calculated on force per unit nominal cross sectional 
area. The stress values and corresponding fracture toughness values obtained, by 
using Eq. 4 are presented in Table 1. 
Effect of changing specific gravity 
Wood specific gravity changes with decay. It will not be correct because of this 
change to calculate the estimated true fracture stress based on a nominal cross- 

Akande and Kyanka= TENSILE FRACTURE IN ASPEN 

293 

FIG. 10. Scanning electron fractograph of a longitudinal-radial fracture plane in wood decayed by 
Trametes versicolor to 2% weight loss. The step angles of failure across the longitudinally oriented 
cells were steep just like that of sound wood. 
sectional area obtained for sound wood specimens. The true area of the specimen 
cross section is bound to change as the wood specific gravity drops with decay. 
The estimated true fracture stress 0"* was calculated as force per unit area of the 
actual cell-wall substance present. 

where 

O"*psi = P maxi Aw 

(8) 

Aw = area of cell wall substance 
The actual area of cell wall substance was calculated based on this eq. from Siau 
(1971). 

Va = I - S.G(O.667 + O.OIM) 
TABLE I. Nominal fracture stress a psi and (corresponding K'cLR psi vrn.). 

% Wt. Joss 	Sound wood 	T. versicolor 	B. adusta 	G. trabeum 	
0 	12,362 (1,564) 				
2 		12,177 (1,541) 	11,806 (1,494) 	8,097 (I,Q25) 	
5 		11,942 (1,287) 	10,965 (1,182) 	6,552 (706) 	
10 		10,730 (I, I 57) 	9,828 (1,059) 	5,192 (560) 	
16 		8,0 II (863) 	8,072 (870) 	4,35 I (469) 	

(9) 

	294 	WOOD AND FIBER SCIENCE, JULY \990, V. 22(3) 
TABLE 2. Wood specific gravity SG and (corresponding areas of cell-wall substance Aw in."). 

%Wt.loss 
o 
2 
5 
10 
16 

Sound wood 	T. versicolor 	B. adusta 	G. trabeum 	
0.447 				
(0.00537) 				
	0.440 	0.429 	0.440 	
	(0.00530) 	(0.005 I 5) 	(0.00530) 	
	0.430 	0.400 	0.410 	
	(0.00520) 	(0.00480) 	(0.00490) 	
	0.392 	0.370 	0.350 	
	(0.00470) 	(0.00440) 	(0.00420) 	
	0.299 	0.320 	0.330 	
	(0.00360) 	(0.00380) 	(0.00400) 	

where 
	Va 	fractional void volume, 
	S.G 	specific gravity at moisture content M 
It was assumed in deriving Eq. 9 that the specific gravity of water = 1 and that 
the specific gravity of cell-wall substance = 1.5. Since the wood specimens tested 
in this work were all at oven-dry conditions, M ~ O. Average S.G of the test 
samples was measured and the results are presented in Table 2. 
At oven-dry conditions, 

Va = 1 - 0.667S.G 

(10) 

where 0.667S.G is the fractional wood volume V w s same as the fractional volume 
of wood substance which can take a wide range of values from 0 to 1. Once V w 
is known by using S.G at oven-dry conditions, the area of cell wall substance A, 
can be calculated. 

(11) 

where 0.018 in- as previously mentioned is the nominal cross sectional area of 
the test specimen. 
TABLE 3. Mean load at failure P max Ib, standard deviation, and (true stress <T*psi calculated from the 
average load values). I 

% Weight loss 	Sound wood 	T. versicolor 	B. adusta 	G. trabeum 	
0 	222.5 ± 38.6 				
	(41,434) 				
2 		219.2 ± 33.9 	212.5 ± 38.5 	145.8 ± 34.7 	
		(41,359) 	(41,262) 	(27,509) 	
5 		215.0 ± 23.6 	197.4 ± 33.2 	117.9±42.1 	
		(41,346) 	(41,125) 	(24,061) 	
10 		193.0 ± 38.4 	176.9 ± 30.7 	93.5 ± 24.5 	
		(41,064) 	(40,205) 	(22,262) 	
16 		144.2 ± 38.9 	145.3 ± 36.2 	78.3 ± 22.6 	
		(40,056) 	(38,237) 	(19,575) 	
1 Average value obtained from a sample size of 15 for each test condition. 			

Akande and Kyanka- TENSILE FRACTURE IN ASPEN 

TABLE 4. K*,/R psi vm. and its (log transformation). 

295 

% WI. Joss 	Sound wood 	T. versicolor 	
0 	5,243 		
	(3.720) 		
2 		5,234 	
		(3.719) 	
5 		4,457 	
		(3.649) 	
10 		4,426 	
		(3.646) 	
16 		4,318 	
		(3.635) 	

B. adusta 

5,222 
(3.718) 
4,433 
(3.647) 
4,334 
(3.637) 

4,121 
(3.615) 

Gc trabeum 

3,481 
(3.542) 

2,593 
(3.414) 

2,400 
(3.380) 

2,110 
(3.324) 

Aw = 0.018 X 0.667S.G = 0.012S.G 

(12) 

The area of cell-wall substance Aw calculated for each of these S.G values is also 
given in Table 2. The estimated true fracture stress 

a*psi = Pmax/O.O 12S.G 

(13) 

Results obtained for the true fracture stress a*psi at different decay weight loss 
levels are presented in Table 3 and the apparent fracture toughness values K*lc LR 
obtained by using the estimated true stress data are given in Table 4. Figure 11 
shows the relationship obtained between Log KIcLR and percent weight loss with 
decay. As weight loss increased due to fungal decay, the observed resistance to 
fracture dropped as expected. 

	

Upper bound (Sound wood) 
	3.8 	
	3.7 	
	3.6 	
Log K *JcLR 	3.5 	
	3.4 	
	3.3 	
	0 	

5 

10 

15 

20 

25 

% Weight loss 
FIG. II. K *Ic LR as a function of decay weight losses. 

	296 	WOOD AND FIBER SCIENCE, JULY 1990, V. 22(3) 
TABLE 5. Reported K,/R data/or woods. 

Source 
Schniewind & Centeno (1973) 
Murphy (1978) 
Present study 

Wood specimen 
Douglas fir 
(edge notched beam) 
(static bending) 
Douglas fir 
(slit notched beam) 
(static bending) 
Aspen 
(tension parallel) 

Mean ~LR 
psi \I'in: 
2,450 
1,466 
1,564 

Moisture content % 
12 
II 
Oven-dried 

Letting x represent the percentage weight loss, the least square regression fits 
for the different fungal treatment are given as follows: 

Trametes versicolor: 
Bjerkandera adusta: 
Gloeophyllum trabeum: 

Log K*IcLR = 3.7026 
Log K*JcLR = 3.7061 
Log K*IcLR = 3.5289 

0.0049x 
0.0063x 
0.0138x. 

(14) 
(15) 
(16) 

DISCUSSION AND CONCLUSIONS 
Relatively few data exist on LR and LT fracture modes in wood because of the 
difficulty in monitoring crack tip displacements which often get arrested or turn 
abruptly to cause splitting along the grain (Schniewind and Centeno 1973). In this 
study, a reasonable estimate of the amount of energy absorbed in an LR mode 
of fracture across grain was obtained by relating fractographic evidences to ex- 
perimental stress data. Results obtained by using nominal fracture stress in this 
work are close to available data for other woods (see Table 5). It is felt that any 
error involved in calculating the K,c LR value is marginal and may be related to 
the visual selection of step angles from one ray to another. 
Comparison ofthe fracture surfaces with the fracture toughness values indicates 
that the geometry of the surface profile is important to an analysis of the energy 
absorbed during wood fracture. When the fracture plane is relatively flat due to 
abrupt breakage of cells in decayed wood, KicLR value falls. For a jagged fracture 
plane resulting from angular failures across sound wood cells, a tougher fracture 
energy was encountered. The declining trend in fracture toughness with decay 
weight loss is similar to that reported with increasing loss of specific gravity by 
Sinclair et al. (1978). Attempts were not made to calculate fracture toughness 
along other modes of crack propagation, in this work, because the fracture stresses 
leading to failure in these modes were not at our disposal. Generally, crack pro- 
duced in tension often follows a steep angle across sound wood fibers until its tip 
reaches the weak interface between ray and fiber cells where rapid crack propa- 
gation occurred. The region of the rays was also identified to be weak in decayed 
wood but the step angles across the longitudinally oriented cells were relatively 
flat compared to that in sound wood. Based on fractographic evidence which 
projected the ray/fiber interface as a weak region, it is postulated that the spacing 
between rays can be important in obtaining fracture toughness values for wood. 
The step-like progression of fracture from one ray to another also supports this 

Akande and Kyanka= TENSILE FRACTURE IN ASPEN 

297 

view. Before this postulate is, however, accepted for engineering application, 
further research is recommended to calculate KIc LR values for different wood 
species (using the distance between rays as the critical flaw dimension) so that the 
results can be compared to published data. In doing this, it is suggested that wood 
samples be selected from specified locations in the log. This is felt to be important 
because the spacing between rays may widen from pith to bark in some species 
and this can influence the fracture toughness values. 
REFERENCES 
AMERICAN SOCIETY FOR TESTING AND MATERIALS (ASTM). 1988a. Standard method of accelerated 
laboratory test of natural decay resistance of woods. D 2017-81/86. Annual book of ASTM 
standards. Construction. Section 4, vol. 04.09 Wood, Philadelphia, PA, pp. 323-326. 
---. 1988b. Standard test methods for specific gravity of wood and wood-base materials. D 2395- 
69/77. Annual book of ASTM standards. Construction. Section 4, vol. 04.09 Wood, Philadelphia, 
PA, pp. 359-366. 
BARISKA, M., AND L. J. KUCERA. 1985. On the fracture morphology in wood. Part 2: Macroscopic 
deformations upon ultimate axial compression in wood. Wood Sci. Techno!. 19: 19-34. 
C<hE, W. A., AND R. B. HANNA. 1983. Ultrastructural characteristics of wood fracture surfaces. 
Wood Fiber Sci. 15(2):135-163. 
DEBAISE, G. R., A. W. PORTER, AND R. E. PENTONEY. 1966. Morphology and mechanics of wood 
fracture. Materials Res. Stand. 6:493-499. 
GRIFFITH, A. A. 1925. Proceeding of the 1st International Congress for Applied Mechanics. Delft 
(Waltman). 55 pp. 
IRWIN, G. P. 1957. Analysis of stresses and strains near the end ofa crack traversing a plate. J. Appl. 
Mech. 57(22):361-364. 
JERONIMIDIS, G. 1976. The fracture of wood in relation to its structure. Leiden Botanical Service 3: 
253-265. 
JUNG, J., AND J. F. MURPHY. 1983. An investigation of the fracture of butt joints in parallel-laminated 
veneer I. Wood Fiber Sci. 15(2): 116-134. 
KEITH, C. T. 1971. The anatomy of compression failure in relation to creep-inducing stresses. Wood 
Science 4:71-82. 
--. 1974. Longitudinal compressive creep and failure development in white spruce compression 
wood. Wood Science 7: 1-12. 
--, AND W. A. COTE. 1968. Microscopic characterization of slip lines and compression failures 
in wood cell walls. Forest Prod. J. 18(3):67-74. 
MURPHY, J. F. 1978. Using fracture mechanics to predict failure in notched wood beams. Proceedings, 
1st International Conference on Wood Fracture. Banff, Alberta, pp. 159-173. 
PATTON-MALLORY, M., AND S. M. CRAMER. 1987. Fracture mechanics: A tool for predicting wood 
components strength. Forest Prod. J. 37(7/8):39-47. 
PELLICANE, P. J., J. BODIG, AND J. R. GOODMAN. 1982. Simulation of the tensile strength-fracture 
toughness relationship. Wood Science 14(4): 168-177. 
PORTER, A. W. 1964. On the mechanics of fracture in wood. Forest Prod. J. 13:325-331. 
SCHNIEWIND, A. P., AND J. C. CENTENO. 1973. Fracture toughness and duration of load factor I. Six 
principal systems of crack propagation and duration factor for crack propagating parallel to grain. 
Wood Fiber 5(2):152-159. 
SlAU, J. F. 1971. Flow in wood. Syracuse University Press, Syracuse, NY, 131 pp. 
SINCLAIR, S. A., G. IFJU, AND J. A. JOHNSON. 1978. Changes in toughness of wood from beetle-killed 
shortleaf pine. Forest Prod. J. 28(7):44-47. 
SMITH, F. W., AND D. T. PENNY. 1980. Fracture mechanics analysis of butt joints in laminated wood 
beams. Wood Science 12(4):227-235. 
TADA, H., P. C. PARIS, AND G. R. IRWIN (ED.). 1973. Stress analysis of cracks handbook. Del Research 
Corp. Hellertown, PA. 
WILLIAMS, J. G., AND M. W. BIRCH. 1976. Mixed mode fracture in anisotropic media. ASTM, 1916 
Race Str. Phil., PA 19103. Technical Bulletin 601:125-137.