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Avoiding knife-edge countersinks in GLARE through dimpling Submitted to Fatigue and Fracture of Engineering Materials and Structures on March 4, 2004
C. Rans*, P.V. Straznicky
Department of Mechanical and Aerospace Engineering Carleton University
1125 Colonel By Drive Ottawa, Ontario, Canada, K1S 5B6
ABSTRACT
Traditional machine countersinking practices create a knife-edge condition in one or
more of the outer aluminum layers in riveted GLARE joints. Press countersinking
(dimpling) provides an alternative method of countersinking that prevents the formation
of a knife-edge; however, its application and potential benefits to fatigue performance in
GLARE are not known. This paper investigates the dimple forming process and its
application to GLARE, and the resulting benefits in fatigue crack initiation life in unfilled
rivet holes. Initial results showed that the limited formability of GLARE complicates the
dimpling process, but that dimpling shows promise as a method for increasing the crack
initiation life of riveted GLARE joints.
Keywords: Dimpling; Riveting; GLARE; FML; Knife-edge
1. Introduction
Fibre metal laminates (FMLs) are a family of hybrid laminated sheet materials
consisting of alternating metallic and fibre-reinforced plastic layers. One particular FML,
GLARE (GLAss REinforced aluminum), is considered a suitable candidate for
application in fuselage skin panels of pressurized aircraft, with its first application in the
* Corresponding author; Email: [email protected]
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Airbus A380-800 megaliner1. The major advantage of GLARE in airframe applications
is the low fatigue crack growth rate compared to monolithic aluminum alloys. The
mechanisms responsible for this superior crack growth behaviour include the plane stress
state in the thinner metallic layers; the branching effect of part-through cracks growing
through the laminate thickness; and the crack bridging properties of the fibre layers (fibre
bridging), which reduces the crack tip stress intensity factor2-5. Experiments have shown
that this behaviour can result in a 10-fold improvement in overall fatigue life over
monolithic aluminum alloys of similar thickness2. Such fatigue performance allows for
the use of smaller sheet thicknesses in fatigue critical applications (such as fuselage skin
panels), creating the opportunity for structural weight savings.
The crack initiation process in GLARE is analogous to that of monolithic
aluminum alloys; namely, monolithic aluminum sheets and aluminum layers in GLARE
with the similar stresses will have similar crack initiation lives. The lower stiffness of the
fibre layers in GLARE, however, causes the aluminum layers to carry a higher proportion
of load, resulting in a shorter crack initiation life compared to monolithic aluminum of a
similar thickness and under the same applied load. A further reduction in crack initiation
life occurs in applications where monolithic aluminum is replaced with thinner GLARE
laminates; however, improvements in overall fatigue life are still obtained due to the
superior crack growth properties of GLARE.
Mechanically fastened joints are fatigue critical locations in airframes. Large
stresses occur at fastener locations as a result of fastener bearing loads and elastic stress
concentrations associated with the fastener hole. In jet aircraft, the external aircraft
surfaces are made flush by machining countersinks for the fastener heads into the skin.
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An additional stress concentration is generated by the countersink, dependent on the
depth of the countersink relative to the skin thickness. In the extreme case where the
countersink depth exceeds the skin thickness, a knife-edge condition is formed, and the
elastic stress-concentration factor becomes 72% greater than that in a plain hole6. Due to
the small thickness of the individual layers in GLARE, the practice of machine
countersinking creates a knife-edge in the outermost aluminum layers (lamina knife-
edge). The application of thinner GLARE laminates could result in cases where the
countersink depth exceeds the laminate thickness, producing a through-thickness knife-
edge (Figure 1). The resulting stress concentrations likely accelerate crack nucleation;
thus, if the knife-edges were avoided, the crack nucleation period could be increased
thereby further enhancing the GLARE advantage.
One possible approach to avoiding a knife-edge is the use of a formed countersink
by means of dimpling. Originally developed for monolithic aluminum alloys, dimpling
was used to prevent knife-edge countersinks in sheet thicknesses below 0.064″7. The
development of reduced-countersink-depth rivets permitted machine countersinking in
sheet thicknesses down to 0.040″, reducing the need for dimpling in primary aircraft
structures. Countersink depths that avoid a lamina knife-edge in GLARE, however, are
not feasible, raising the question of the potential benefits of dimpling GLARE. In
addition to preventing the formation of a knife-edge, dimpling introduces the following
additional factors that may influence fatigue performance:
• Complex load transfer mechanism resulting from the interlocking of dimpled
sheets,
• Residual stresses resulting from dimple formation,
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• Potential for delamination damage, fibre fracture and other forming related
failures during dimple formation in GLARE.
Due to the limited use of dimpling, the above factors and their effects on fatigue
life are not well understood.
As part of a larger study into determining the fatigue performance of dimpled
GLARE lap joints8, the results from a series of dimple forming tests and crack initiation
test of unfilled dimpled holes in GLARE are presented in this paper. This investigation is
limited to biaxial Grade 3 GLARE (equal number of fibres in the 0° and 90° directions)
with a layup consisting of two aluminum layers and one fibre layer (designated GLARE
3-2/1). Table 1 shows a comparison of the basic properties of GLARE 3-2/1 and 2024-
T3 alloy. GLARE 3-2/1 was chosen for this study due to its potential application as a
fuselage skin in narrow-body aircraft, its susceptibility to through-thickness and lamina
knife-edges, and its higher formability compared to thicker GLARE laminates.
2. Dimple Forming Techniques
Dimpling is a forming process, where a countersunk is formed by plastically
deforming the sheet material under high pressure using a set of dimpling tools (a punch
and die). This process is performed after the fastener holes have been drilled and
deburred, and can be performed at room temperature (cold dimpling) or at elevated
temperatures (hot dimpling). Hot dimpling is predominantly used in thick or brittle sheet
materials, where cold dimpling would produce radial cracks. The basic geometry of a set
of dimpling tools is shown in Figure 2. The springback angles shown are often included
in the tool geometry to minimize distortion of the sheet material surrounding the dimple
due to elastic springback after dimpling.
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Two basic variations of the dimpling process exist: coin dimpling and radius
dimpling (Figure 3). The major difference between the two variations is how they ensure
proper nesting of the dimpled sheets being joined. In coin dimpling, each sheet is
dimpled independently, and the geometry of the formed dimple (dimple cone) is identical
in all sheets. To ensure proper nesting, a reduction in the thickness of the dimple cone is
necessary. Figure 4 shows that, for a standard dimple angle of 100°, a 23% reduction in
sheet thickness is required. Coin dimpling also results in a cylindrical fastener hole, as
excess material resulting from the thickness reduction in the dimple wall is formed
against the mandrel of the punch. In radius dimpling, all sheets being joined are dimpled
simultaneously. Proper nesting is assured by a variation in maximum dimple cone
diameters and angles in the mating sheets; however, some local reductions in dimple cone
thickness will also occur. As a local thickness reduction along the wall of the dimple
cone is not necessary, radius dimpling requires relatively lower loads than coin dimpling.
Radius dimpled holes also produce less cylindrical fastener holes due to varying amounts
of hole stretch in the mating sheets (Figure 3b).
3. Dimpling Trials in GLARE
A series of dimple forming tests were conducted to identify what dimpling
processes and forming loads would be suitable for GLARE 3-2/1. All tests were done to
simulate the countersinking process required for a MS20426 AD 4-4 rivet in a simple
GLARE 3-2/1 riveted lap joint (see Table 2 for rivet geometry details). Visual inspection
of the formed dimples, and inspections of cross-sections taken through the rivet hole
centre line were used to evaluate dimple nesting and to determine if any damage occurs
during dimple formation in GLARE.
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3.1 Test Set-up
Dimpling was performed on 6.0″x1.5″ GLARE 3-2/1 strip with 0.1285″ diameter
(#30 drill) rivet hole centred on the strip. In the case of radius dimpling trials, two
additional holes, spaced 1″ on either side of the centred hole, were added to accept
temporary sheet metal fasteners (clecos) to hold the two sheets together during dimpling.
Clecos were not necessary in the coin dimpling trials as this process is conducted on a
single sheet at a time. Dimpling was performed cold (room temperature tools) and “dry”
(without the use of lubricant or sealant) under load control using a servo-hydraulic test
frame with custom steel fittings designed to accept standard dimpling tools. A set of
commercially available dimpling tools (US Industrial Tools Co. part # TP4264) were
used for all dimpling trials. These tools contained no springback angle in the punch or
die, and formed a countersink conforming to a 1/8″ diameter rivet with a MS20426 style
head. Dimple forming loads of 4,000, 5,000, and 6,000lbf were selected for this
investigation based on work done in reference9.
3.2 Coin Dimpled Holes
This section presents the results from coin dimpling trials in GLARE 3-2/1. The
following observations were made and found to be similar for all forming loads:
• Coin dimpled holes did not nest neatly together, indicating that an insufficient
reduction in dimple wall thickness was obtained.
• Glass fibres were visible within the dimpled hole (Figure 5), indicating that fibre
layer was delaminating and was being squeezed out of the laminate.
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• As the dimple forming load was increased, dimple nesting improved; however,
the amount of fibre layer squeezed out into the rivet hole increased as well.
Overall, unsatisfactory nesting of the dimples was observed for all coin dimpled
GLARE 3-2/1 coupons. This is primarily a result of insufficient thinning of the dimple
cone wall required in coin dimpled holes. The 0.012″ thick aluminum layers in GLARE
3-2/1 are very thin, requiring higher forming loads to produce local thinning than thicker
monolithic aluminum sheets. The observed delamination and expulsion of the fibre layer
indicated that the major portion of the wall thinning was being achieved by expulsion of
the fibre layer. Conceivably, proper nesting of the coin dimpled holes could have been
obtained by increasing the forming load and squeezing out more of the fibre layer;
however, it was decided that this damage would be too detrimental to the fatigue life to
warrant further investigation.
3.3 Radius Dimpled Holes
The following observations were made during the radius dimpling trials:
• Radius dimpled holes nested neatly together for all dimple forming loads
investigated.
• No immediately visible evidence of delamination was present, unlike in the coin
dimpling trials.
• A slight curvature was left in the coupon after dimpling, suggesting that the
dimple was underformed (Figure 6). This curvature was slightly reduced by
increasing the forming load from 4,000lbf to 5,000lbf; however, negligible
improvement was obtained when the forming load was further increased to
6,000lbf.
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• Delamination damage, local thinning of the aluminum layers, and tight bending
radii were observed in the cross-sections taken through the dimpled holes (Figure
7).
Results from the radius dimpling trials showed more promise than coin dimpling.
Proper nesting between dimples was achieved, and although delamination damage was
detected, the fibre layer was not squeezed out of the dimple wall as in coin dimpling.
Two possible contributing factors to the delamination are the tight bending radii produced
by the dimpling tool set used in the trials, and the local thinning of the aluminum layers.
Small bending radii produce a larger deformation gradient through the thickness of the
laminate, resulting in large interlaminar shear stresses. Similarly, large interlaminar shear
stresses are produced as the fibre layers resist the elongation of the aluminum layers due
to local thinning.
The specimen curvature observed due to radius dimpling GLARE is typically an
indication of an insufficient forming load; however, no improvement in curvature was
observed for forming loads above 5,000lbf. The larger springback typically observed in
GLARE compared to monolithic aluminum10, combined with the lack of any springback
angle in the dimpling tools are likely responsible for the curvature observed with forming
loads above 5,000lbf.
4. Crack Initiation Coupon Testing
Crack initiation tests were performed on simple 10″x1″ GLARE 3-2/1 coupons
containing a single unfilled rivet hole centred on the coupon. This configuration allowed
the effects of eliminating the knife-edge on crack initiation life to be studied in isolation
from differences in load transfer and hole filling properties of dimpled and machine
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countersunk joints. Coupons containing non-countersunk, machine countersunk, and
dimple countersunk holes were tested. Based on the results of the dimple forming
experiments, only radius dimpled holes, with a 5,000lbf forming load, were considered.
Both coupons made from the sheet in contact with punch during dimpling (dimpled upper
sheet or dimpled US), and the sheet in contact with die during dimpling (dimpled lower
sheet or dimpled LS) were tested. Additionally, the machine countersunk coupons had a
reduced countersunk depth (corresponding to a NAS1097 AD 4-4 rivet) to ensure that
only a lamina knife-edge was formed in the coupon, rather than the extreme through-
thickness knife-edge case. Three coupons were tested for each configuration, with the
exception of the dimpled LS coupons where four were tested.
4.1 Test Set-up
All tests were conducted by means of a MTS 810 servo-hydraulic test frame,
using 4Hz constant amplitude sinusoidal load varying between 544 lbf and 16 lbf
(R=0.03). This load was selected to represent typical hoop loads experienced due to
pressurization of a narrow-body aircraft, based on work reported on in references11-13.
Load was applied to the specimens through a set of standard hydraulic grips. For
consistency, all specimens were tested with the aluminum layer adjacent to the 90° fibre
layer (angle is with respect to direction of load application) as the countersunk surface
(Figure 8).
Routine visual inspections with a 20X optical microscope were performed to
detect fatigue cracks. Crack initiation, for the purpose of these tests, was defined as the
presence of a 0.03″ crack, measured from the edge of the rivet hole (not the edge of the
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countersink). Vertical lines were marked at the 0.03″ crack length positions prior to
testing to facilitate the determination of the crack initiation point.
4.2 Results
Figure 9 shows the average crack initiation life observed for each type of coupon.
The dimpled US coupons exhibited approximately 90% improvement in crack initiation
life compared to the machine countersunk coupons, and a similar crack initiation life to
the non-countersunk coupons, as a result of preventing the formation of a knife-edge.
The dimpled LS coupons showed a significantly improved crack initiation life; testing of
the three out of four of the coupons was stopped after 1,000,000 fatigue cycles with no
detectable fatigue cracks. The fourth coupon failed after 710,300 cycles, without prior
detection of crack initiation. The reasons for the large discrepancy in crack initiation
performance between the dimpled US and LS coupons are not apparent; however,
differences in residual stress distribution, degree of delamination damage, and geometry
of the dimple cone between the upper and lower sheet likely play a role. Further
investigation into these differences is required.
Crack initiation occurred along the minimum section for all tested coupons.
Cracks initiated at the knife edge/rivet hole in the machine countersunk and non-
countersunk coupons. In contrast, cracks initiated at the maximum diameter of the
dimple cone in the dimpled US coupons. This shift in initiation location is likely a result
of residual tensile stresses generated during dimple formation. No data on crack
initiation location could be obtained for the dimpled lower sheet coupons.
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5. Discussion
From the dimpling trials, radius dimpling showed the greatest promise for
applications in GLARE. The level of damage produced during coin dimpling was
deemed too severe for practical application of the process. Furthermore, the logistical
issues associated with independently dimpling all sheets being joined make coin dimpling
undesirable from a manufacturing standpoint. Delamination damage between the 0° and
90° fibre layers was also observed in the radius dimpling trials; however, this damage
was much less severe than observed during coin dimpling.
The presence of any degree of delamination around a rivet hole can have a
negative impact on fatigue performance. Further testing performed in conjunction with
this study has found that the observed delamination is exacerbated during the riveting
process and results in buckling damage of the lower sheet (Figure 10). In the vicinity of
this damage load will only be carried by the unbuckled layers of the lower sheet. The
higher stresses in these layers will accelerate crack initiation and growth. Subsequent
exploratory fatigue testing of a dimple countersunk riveted GLARE lap joint detailed in
reference8 also indicated a reduction in crack growth resistance. Crack growth rates in
this specimen were equivalent to those observed in monolithic 2024-T3, indicating that
the delamination was growing ahead of any fatigue cracks and limiting the effectiveness
of the fibre bridging mechanism. Both of these factors indicate a need to prevent or
minimize the extent of delamination damage during dimpling and riveting to achieve
potential fatigue life improvements. Potentially, this could be achieved by modifying the
dimpling tools to take into consideration the limited formability of GLARE; as an
example, changes could be made to the bending radii.
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Despite the presence of the delamination, the crack initiation coupons indicated a
potential improvement to crack initiation life resulting from eliminating the knife-edge.
The large variation in crack initiation life between the dimpled US and LS coupons also
suggests that residual stresses caused by the dimpling process significantly affect the
crack initiation life. Optimization of the dimpling tools and process may allow a
beneficial residual stress field to be tailored around the dimpled rivet hole, further
improving the potential fatigue life benefit of dimpling.
Although the initial results from this study indicate a potential for improved crack
initiation life in dimple countersunk joints, several factors remain unknown. Firstly, load
transfer through a dimpled joint is not well understood. In a machine countersunk joint,
the main component of load transfer is bearing between the rivet and joint sheets. In
dimpled joints, bearing will occur between the dimple cones, with the rivet preventing
their separation. Friction between the nesting dimple cones may also represent a
significant portion of the load transfer (Figure 11). The exact contributions of these
effects on the total load transfer, and the associated stress distribution within dimpled
joints are unknown, and relatively little fatigue test data exists due to the limited
application of dimpling in monolithic aluminum. Secondly, the degree of hole filling and
associated residual stresses when riveting dimpled GLARE is not known. Müller2 has
demonstrated the importance of the residual stresses resulting from riveting on the fatigue
performance of GLARE and monolithic aluminum joints; thus any changes due to
dimpling could have an impact on fatigue life. Before the true potential of dimpling can
be assessed, these factors will have to be investigated, and the potential issues associated
with the observed delamination damage overcome.
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6. Conclusions
Based on the results presented, the following conclusions can be drawn from this
study:
1. Coin dimpling GLARE 3-2/1 using commercially available dimpling tools results
in severe delamination damage and causes the fibre layer to be expelled from the
laminate within the dimple cone. This damage was deemed too severe to warrant
further investigation.
2. Radius dimpling GLARE 3-2/1 using commercially available dimpling tools
results in delamination damage within the formed dimple cone. The severity of
this damage is lower than that produced by coin dimpling, and may be
preventable through optimization of the dimpling tool geometry.
3. Unfilled dimpled holes exhibited a longer crack initiation life than machine
countersunk dimpled holes.
4. A significant difference between the crack initiation performance between the
upper and lower dimpled sheets was observed. Differences in degree of
delamination damage and residual stress distribution may be responsible;
however, further study is required.
5. Potential for a significant increase in crack initiation life by dimple
countersinking was observed; however, further investigation into the influence of
changes in load transfer mechanism and degree of rivet hole filling are required to
assess the true potential of dimpling GLARE.
With regards to conclusions 4 and 5, further studies are currently taking place to
assess the potential of dimpling. Specifically, finite element analysis is being used to
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investigate the residual stresses generated due to dimpling and riveting of dimpled holes
and alternative dimple tool geometries are being assessed through similar dimple forming
experiments as presented here.
7. Acknowledgements
This work was in part supported by Bombardier Aerospace, the Institute for
Aerospace Research of the National Research Council of Canada, and through research
grants from the Natural Sciences and Engineering Research Council (NSERC). The
authors would also like to thank Dale Cox and Donald Gorman from the Flight Research
Lab for their technical assistance and advice, and undergraduate students Magali Duval
and Mathew Wells for their assistance with fatigue testing.
References: 1. Vlot, A. (2001). GLARE: history of the development of a new aircraft material.
Kluwer Academic Press, Dordrecht, The Netherlands.
2. Müller, R.P.G. (1995), “An experimental and analytical investigation on the fatigue behaviour of fuselage riveted lap joints: the significance of the rivet squeeze force and a comparison of 2024-T3 and Glare 3.” Ph.D. dissertation, Delft University of Technology, Delft, The Netherlands.
3. Krishnakumar, S. (1994), “Fibre metal laminates – the synthesis of metals and composites,” Mater Manuf Process, 9, (2), 295-354.
4. Slagter, W.J. (1994), “Static strength of riveted joints in fibre metal laminates.” Ph.D. dissertation, Delft University of Technology, Delft, The Netherlands.
5. Lazzeri, L. (2001), “Fatigue behaviour of riveted Glare lap joints,” Fatigue Fract Eng Mater Struct, 24, 579-589.
6. Young, J.B., Lee, K.K. (1993), “Stress concentration factors in countersunk holes,” Aeronautical Journal, 97, (968), 267-276.
7. Bruhn, E.F. (1973), Analysis and Design of Flight Vehicle Structures. Jacobs Publishing Inc, Indianapolis.
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8. Rans, C. (2003), “An experimental investigation into the fatigue behaviour of dimple countersunk GLARE riveted lap joints.” M.A.Sc. Thesis, Carleton University, Ottawa, Canada.
9. Templin, R.L., Fogwell, J.W. (1942), “Design of tools for press-countersinking or dimpling 0.040-inch-thick 24S-T sheet,” NACA TN- 854.
10. de Jong, T.W., Kroon, E., Sinke, J. (2001), “Formability”, in Fibre Metal Laminates an Introduction. (A. Vlot, J.W. Gunnink, eds.). Kluwer Academic Press, Dordrecht, The Netherlands.
11. Eastaugh, G.F. (1994), “Multiple site fatigue damage in fuselage skin splices.” M. Eng. Thesis, Carleton University, Ottawa, Canada.
12. Krizan, D.V. (1999), “Assessing the Effects of Corrosion Damage on Longitudinal Fuselage Joints.” M. Eng. Thesis, Carleton University, Ottawa, Canada.
13. Cook, J. (2001), “Influence of corrosion damage topography on fatigue properties in longitudinal fuselage lap joints.” M. Eng. Thesis, Carleton University, Ottawa, Canada.
14. DoD (1998), MIL-HDBK-5H. U.S.A. Dept. of Defence.
Page 16 of 21
Table 1: Comparison of properties3 of GLARE 3-2/1 and 2024-T3.
Material Specific Weight (lbs/in3)
Direction Tensile Ultimate
Strength (ksi)
0.2% Tensile Yield Strength
(ksi)
Tensile Elastic
Modulus (Msi)
Ultimate Tensile
Strain (%) L 104 44 8.4 4.7 GLARE 3-2/1 0.091 LT 104 41 8.4 4.7
L 66 52 10.5 19
2024-T3 0.100 LT 66 47 10.5 19 L, LT = Long (Grain) and Long Transverse directions respectively
Table 2: Rivet geometries14.
MS20426 AD4-4 NAS1097 AD4-4
Material 2117-T4 2117-T4
Do (in) 0.125 0.125
h (in) 0.042 0.027
l (in) 0.250 0.250
Figure 1: Schematic of a through-thickness knife-edge (a) and lamina knife-edge (b).
(a) (b)
100.0°
hl
Do
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Figure 2: Basic dimple tool geometry shown with springback angle (adapted from reference9). Tools used in this study did not contain any springback angle.
100°100°100°
100°112°124°
Hole Stretch(a) (b)
NOTE: Radius dimple angles shown are representative only andvary with material thickness, hole size, and dimple tool geometry.
Figure 3: Comparison of nested dimpled sheets with installed rivet: (a) coin dimpled sheets, (b) radius dimpled sheets.
Max. Die Diameter
Die Springback Angle
Punch Springback Angle
Min. Die Diameter
Die Dimple Angle
Punch Dimple Angle
Min. Punch Diameter
Max. Punch Diameter
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Figure 4: Reduction of dimple wall thickness in coin dimpled hole.
Figure 5: Coin dimpled hole showing expulsion of fibre layer.
Under-formed Dimpleforming load too low
dimple tool springback angle too small
Over-formed Dimpleforming load too high
dimple tool springback angle to large
Properly-formed Dimplecorrect forming load
correct dimple tool springback angle
Figure 6: Checking degree of dimple forming.
t
t′
t′ = t·sin(θ/2)
θ
Page 19 of 21
Figure 7: Radius dimpled hole showing delamination between fibre layers.
Al-190o
0o
AL-2P
Axis of rivet hole
Figure 8: Orientation of fibre layers relative to applied load (P) for crack initiation testing.
Page 20 of 21
158
70.8135
> 1000
0
100
200
300
400
500
600
700
800
900
1000
Cycl
esto
initi
atio
n[k
c]
StraightHole
MachineCsk
DimpledUS
DimpledLS
P = 544 lbf/inR = 0.03
Figure 9: Comparison of mean crack initiation life of coupons with unfilled rivet holes.
Figure 10: Radius dimpled and riveted GLARE 3-2/1 showing buckling damage of lower sheet8.
Page 21 of 21
Rivet-to-Sheet Bearing
PP
Dimple-cone Bearingand Friction
PP
(a)
(b)
Figure 11: Primary load transfer mechanism in (a) machine countersunk riveted joints and (b) dimpled countersunk riveted joints.
