00547 Maik Gehloff - Timber Design Society

PERFORMANCE OF MOMENT RESISTING SELF-TAPPING
SCREW ASSEMBLY UNDER REVERSE CYCLIC LOAD
Maximilian Closen 1 , Frank Lam 2
ABSTRACT: This paper evaluates the performance of a moment resisting self-tapping screw assembly with tension
and compression plates as connecting members in a typical beam to column connection subjected to reverse cyclic
loads. Throughout the test series two main connection failure modes were observed. Failure mode one shows significant
wood crushing in combination with transverse shear failure in the column member with a generally slow load decrease
and continuous rotation of the beam member. Failure mode two shows sudden failure caused by tension fracture of the
self-tapping wood screws in the beam member. The results show that this connection exceeds the bending moment
design capacity of the glulam material by a factor of 2 in positive direction and a factor of 1.9 in the negative direction.
In general strong performance with very little capacity and stiffness degradation was observed.
KEYWORDS: Stress/failure analysis, timber structures, full thread self-tapping screws
1 INTRODUCTION 123
Timber connections with stocky dowels or bolts are
critical structural members that are generally prone to
brittle failure. Brittle failure occurs due to the poor
tensile perpendicular to grain and longitudinal shear
strengths of wood. When not intended in the design these
connections can fail prematurely when subjected to
stresses other than shear. During seismic events
connections are likely to experience additional stresses
caused by bending moments that were not considered
during the design. Furthermore [1], reports the issue of
restrained shrinkage in connections with slotted-in steel
plates that use dowels or bolts as connecting members as
critical with a potential for wood splitting and significant
connection capacity reductions.
Past research [2], [3], [4] proposes reinforcements that
effectively reduce the tendency of wood splitting
perpendicular to the grain by using structural full thread
self-tapping wood screws (STS). [5] reports outstanding
performances of bolted timber moment connections with
STS reinforcement’s perpendicular to the wood grain.
The reinforcing effect of STS in wood is based on
principals that are similar to rebar’s in concrete. The
special thread shape with assymetrical-symetrical thread
1 Maximilian Closen, Department of Wood Science, University
of British Columbia, 2842 - 2424 Main Mall, Vancouver, BC,
V6T 1Z4, Canada. Email: mclosen@interchange.ubc.ca
2 Frank Lam, Department of Wood Science, University of
British Columbia, 2424 Main Mall, Vancouver, BC, V6T 1Z4,
Canada. Email: frank.lam@ubc.ca
pitches firmly bounds to the wood fibre over multiple
layers (Figure 1). The bite of the thread forms a stiff
connection to the wood. As stiff structural elements
typically attract loads; therefore, it can be assumed that
the majority of occurring stresses perpendicular to the
grain is transferred as a tensile stress along the screw
axis. Since STS are hardened after rolling on the thread a
high tensile strength is achieved. The combination of
high tensile strength and threads that firmly bite into the
wood fibre are an effective method of splitting
reinforcement.
Figure 1: Screw thread embedded in the wood
Current research has intensely addressed the reinforcing
potential and withdrawal capacities of STS in timber
connections; however, the potential of these fasteners as
primary connector has been neglected. Utilizing the
previously described properties of these screws as
primary fastener strong, reliable and cost efficient timber
connections can be created. [6] reports the application of
STS as primary fastener in rigid frame corners with

significant higher load bearing potential when compared
to glued finger joints and dowels. The configuration of
the STS for this joint type however requires modelling of
internal stress flows using the principles of strut-and-tie
systems. Because many engineers have limited access to
appropriate high cost modelling software this promising
approach may not be commonly applied in practice until
design equations are derived.
In this study an experimental investigation was
conducted to evaluate the performance of a moment
resisting STS assembly (see Figure 5). The capacity of
the assembly primarily depends on the shear and
withdrawal resistance of the STS which can according to
[7], in most cases, conservatively be estimated in major
Canadian species using the design equations presented in
[8]. A total of ten specimens were subjected to reverse
cyclic loading conditions. In the following the
performance of the moment resisting STS assembly in
terms of ultimate capacities, stiffness and ductility is
presented. In addition, a detailed description of all
observations made during testing is provided.
compression plates (ZD-plates). To allow for
construction tolerances, the housing was oversized by 2
mm in depth and 4 mm in width. These tolerances
proofed to be sufficient for the specimen assembly in the
lab. Each ZD-plate received four 10 mm x 240 mm
ASSY VG countersunk head screws driven-in at a
±30°angle to the wood grain of the beam member (see
Figure 4). Detailed information about strength
parameters and characteristics of the ASSY VG screws
such as minimum requirements on the raw wire material,
yield moment and tensile strength are summarized in [9]
with links to proposed design equations in [8] and [10].
In [11] a proposed design approach which estimates the
capacity of connections with ZD-plates can be found.
2 MATERIALS AND METHOD
All specimens for this test series were cut from the
centre lamellas of 130 mm x 912 mm Canadian Douglas
fir glulam of stress grade 20f-E. As a typical example for
the investigated connection a beam to column or column
base connection can be assumed. Both connection
application examples are typically needed in timber
moment frames or simply as a column base-fixture.
A custom made steel shoe assembly (see Figure 2) using
9.5 mm thick steel plates was welded to fit the 130 mm x
304 mm vertical beam member. On each side of the
shoe, steel brackets were welded to the vertical steel
plates for additional stiffening. To connect the vertical
steel plates a 200 mm x 304 mm bracing-steel plate was
welded in between. To be able to fit the beam member
into the steel shoe a 12 mm wide and 220 mm deep slot
needed to be cut into the lower beam end. Bearing of the
beam member on the bottom steel plate and the bracing
steel plate during load application was avoided by
providing a 20 mm gap by simply slotting the beam to a
depth of 220 mm. A second 9.5 mm thick bottom steel
plate was fastened to the horizontal column member
using 10 mm x 240 mm ASSY VG screws. A total of
twelve 12.5 mm Grade 5 bolts were used to connect the
steel shoe to the bottom steel plate through 13 pre-drilled
holes with 12.5 mm receiver nuts welded on the bottom.
Each bolt and nut was equipped with a 35 mm washer
and tightened to a pre-set torque of 200 Nm. All timber
members were fabricated using the Hundegger K2 5-axis
fully automated joinery machine that was available at the
Centre of Advanced Wood processing at the University
of British Columbia (UBC). The previously mentioned
12 mm x 220 mm slot in the beam member was cut using
the Hundegger K2 circular saw. The required 40 mm x
25 mm round housing for the receiver nut and washer at
the column member was milled using a 25 mm finger
mill. The 25 mm mill was also used to mill the column
member housing for the 27 mm x 86 mm tension and
Figure 2: Custom steel shoe and bottom steel plate
(some parts were removed from drawing for clarity)
(all measures in mm)
Moisture content measurements with a resistance
moisture meter at all specimen ranged between10.3 %
and 11.8 % prior to assembly.
To reduce the impact of wood crushing on data
recordings at the location of the horizontal tie downs five
10 mm x 240 mm ASSY VG screws were driven-in
perpendicular to the grain (see Figure 3). The tie down
steel pipe was now bearing on the screw head preventing

the wood from crushing under large loads. Because large
loads and heavy deformations were expected on the steel
shoe they were replaced after each test.
cycle was selected with a deformation of 0.05∆ followed
by a primary cycle of 0.075∆. This sequence progresses
up to 2.0∆.
Figure 3: Typical bearing reinforcement
Figure 5: Test set-up
The horizontal force component applied at the top of the
beam member was recalculated using the displacement
recording of a linear position transducer mounted
vertically under the actuator (see Figure 6). Furthermore
two cable extension position transducers were attached
to the beam member in two different heights. These
recordings were used to calculate the horizontal
movement of the member from which moment rotation
plots can be derived. A multi-purpose test system
controlled the single ended MTS actuator with 250 mm
stroke in positive and negative direction during load
application.
Figure 4: ZD-plate specifications from SWG-Production
(all measures in mm)
The entire research project was conducted at the UBC
Timber Engineering and Applied Mechanics (TEAM)
test facility. An illustration of the general test setup is
shown in Figure 5. Ten specimens were subjected to
displacement controlled reverse cyclic loading according
the test procedure developed by [12]. This protocol
consists of initiation cycles, trailing cycles and primary
cycles with equal positive and negative amplitudes. The
initiation cycles at the beginning serve to check the test
equipment and the load deformation response. Primary
cycles mark the beginning of a new displacement phase
and are followed by trailing cycles with 75% of the
amplitude of the respective primary cycle of each phase.
The loading sequence was calibrated to the average
reference deformation ∆ ultimate = 3° which was obtained
from two monotonic tests with a constant loading rate of
16 mm per minute. To account for variations between
monotonic and cyclic testing and damage accumulation
the suggested deformation ∆ =0.6∆ ultimate was applied to
determine the final loading procedure. The first initiation
3 RESULTS AND DISCUSSION
The recorded loads, displacements and geometric data
were used to derive moment rotation plots. A straight
line segment connecting the point of 0.4M max with the
origin of the moment rotation curve was used to define
the initial connection stiffness. Where applicable, a yield
moment derived from the point of intersection of the
initial stiffness line and a tangent to the moment rotation
curve having 1/6 of the slope of the initial stiffness line
was calculated. In specimen with a steep load drop after
the maximum moment was reached no yield moment
was derived.
The moment resistance of the assembly was calculated at
the centre of rotation of the steel shoe.

Figure 6: General test layout and testing equipment (all measures in mm)
Therefore, the horizontal force component was
multiplied with the respective corrected distance (C in
Figure 6) to the centre of rotation. Respective rotations
of the beam member were derived from the geometric
measurements of points B, D and E in Figure 6. The STS
assembly with tension and compression plates shows
high moment capacities with small rotations throughout
the entire test series. Because of the layout of the steel
shoe with a large number of bolts and bracings a tension
fracture failure of the screws was promoted. Mostly this
tension fracture resulted in brittle connection failures at
small rotations of the beam member. Due to the high
horizontal forces that the connection resisted during
testing, compression failures in the beam member at the
point of load application were observed. As a result up
and down shifting of the load application fixture started
to occur in progressed cycles (Figure 7). Compression
failures of the wood were also observed on the column
member at the location of the tie down steel pipes. At
this location 16 mm steel plates and 10 mm x 240 mm
full thread screws were applied to resist the expected
high compression stresses. Three specimens did however
fail regardless of the reinforcement and enlarged bearing
area in a failure combination of compression
perpendicular to grain and transverse shear (Figure 7).
Furthermore, the intense crushing and splitting of the
column member caused gaps to develop between the
concrete strong floor and the bottom of the column. At
maximum load these gaps measured approximately 15
mm to 20 mm in positive and negative direction. Finally
this resulted in higher rotations of respective specimens
(see Figure 8). The intended 20 mm gap between the
bottom of the beam member and the steel shoe proved to
be sufficient throughout the test series. No bearing of the
beam member on the steel shoe in either positive or
negative direction with potential for increasing the
moment resistance was observed.
Figure 7: Typical wood crushing at point of load
application (beam member) and compression -
transverse shear failure (column member)
Figure 8: Developed gap due to intense wood crushing
and transverse shear failure at bottom of column member

The M16 (10.9) bolts used to connect the steel shoe to
the ZD-plates of the beam member showed signs of
heavy local stresses on the thread shortly under the hexhead.
Up and down movement of the bolts was observed
throughout all tests in progressed cycles. Heavy bending
of the short, stocky bolts was found in three tests after
the specimen were disassembled (see Figure 9). In these
tests withdrawal resistance and push out failure of the
full thread screws on one side of the beam member
occurred and a gap between the ZD-plate bottom and the
wood was formed. This gap caused an increased bending
stress on the screws forcing them to yield therefore
reducing the connection stiffness and capacity. Typically
the combination of high tensile stresses and screw
bending causes brittle fractures in hardened fasteners.
In [11] failure of the thread inside the ZD-plates and
failure at the bolt thread is reported. In addition [11]
reports bending of the ZD-plates under heavy loads. This
observation cannot be confirmed. One may assume that
the shorter ZD-plates (86 mm length) used for tests in
this project compared to the 100 mm used in [11]
provide higher stiffness’s and therefore larger bending
strength.
As it can be seen in Figure 10 the M16 bolts were
pushed out (away from the steel plate surface) in a way
that no bearing of the bolt head on the steel plate was
prevented. The cab formed ranged between
approximately 2 mm and 4 mm. In general the largest
gap formed at the top bolt with decreasing gap size
toward the bottom of the beam member. A possible
reason for the loosening of the bolts may be the large
number of cycles that the specimen was subjected to and
the small vibrations that were caused by the friction
between the steel plate and the wood.
Table 1 and Table 2 summarize the results for the ten
tests which were derived from the moment-rotation
relationships and the geometric data recordings as
outlined in previous chapters. In Figure 11 a typical
moment-rotation plot from specimens with significant
wood crushing and transverse shear failure in the column
member (observed in three out of ten tests) is shown. As
a result of wood crushing and transverse shear failure no
sudden load drops as seen in Figure 12 in either cycle
can be seen. A load drop of approximately 20% of the
recorded maximum load occurred in the positive cycle.
After specimen disassembly it was found that this load
drop was caused by the fracture of one tension screw in
the top ZD-plate. The remaining screws on this side of
the specimen were still functioning to carry loads and
showed only little signs of high stresses with regard to
screw bending and screw withdrawal. In all tests the
failure of the connection initiated in the top ZD-plate and
then progressed downward toward the column member.
Figure 11: Typical Moment rotation plot derived from
specimen Mc-10 with significant wood crushing and
transverse shear failure at the column member in the
positive and negative cycle.
Figure 9: Bolt bending
Figure 12: Typical Moment rotation plot derived from
specimen Mc-4 with wood crushing at the column
member in the positive cycle and tension fracture failure
of screws with a related steep load drop in negative
cycle.
Figure 10: Bolt push-out

Beside a fracture failure of the screws in tension a
withdrawal-push-out failure of the compression screws
was observed mostly in the top ZD-plates (see Figure
13). A significant visible failure on the ZD-plate itself
was not found however slight indentations at the lid
caused by the head of the compression screw were found
throughout the entire test series. In addition, significant
bending of the compression screws indicating high shear
stresses at the interface between the ZD-plate and the
wood member occurred. In general little signs of high
stresses were observed at the bottom ZD-plate
Figure 13: Observed withdrawal-push out failure (top left) and disassembled ZD-plate with compression screw bending
and fractured tension screws after testing (top right). Screw withdrawal in column member (bottom left).Tension fracture
of 12.5 mm bolt and steel plate yielding (bottom right).
Beside the failure modes of the tension and compression
screws in the beam member withdrawal failure of the
screws driven into the column member also occurred.
Along with the withdrawal of the screws tension splitting
perpendicular to the grain at the location of the screws
tip was observed. It is assumed that large tensile stresses
were transferred through the screw shaft resulting in high
stress concentrations at the screw tip. These stress
concentrations finally cause wood splitting perpendicular
to the grain. A further failure observed at the steel shoe
was tension fracture of the 12.5 mm bolts. Accordingly
the bottom steel plate separated itself from the top steel
plate through significant yielding and screw withdrawal
(see Figure 13).

Table 1: Summary of recorded test results
Mc-1 Mc-2 Mc-3 Mc-4 Mc-5
pos. neg. pos. neg. pos. neg. pos. neg. pos. neg.
cycle cycle cycle cycle cycle cycle cycle cycle cycle cycle
max M. [KNm] 69.34 -101.8 88.81 -86.07 79.20 -82.90 89.14 -85.53 108.4 -80.87
rotation @ max M. [°] 2.84 -8.46 3.25 -2.67 3.61 -4.21 2.59 -2.51 2.99 -3.24
failure M. [KNm] 55.47 -81.46 71.05 -68.86 63.36 -66.32 71.32 -68.43 86.73 -64.70
rotation @ failure [°] 3.30 -4.34 3.90 -2.82 3.82 -5.18 3.46 -3.28 3.00 -3.67
40% max M. [KNm] 27.74 -40.73 35.52 -34.43 31.68 -33.16 35.66 -34.21 43.36 -32.35
rotation @ max M. [°] 1.27 -2.16 1.41 -1.47 1.28 -1.44 0.92 -1.08 0.91 -0.98
yield M. [KNm] - - - - - -79.62 83.90 - 101.5 -75.65
rotation @ yield M. [°] - - - - - -3.46 2.25 - 2.13 -2.29
stiffness. [KNm/°] 21.90 18.87 30.52 33.85 24.75 22.97 38.87 35.05 47.75 33.01
ductility [-] - - - - - 1.17 1.15 - 1.40 1.41
Mc-6 Mc-7 Mc-8 Mc-9 Mc-10
pos. neg. pos. neg. pos. neg. pos. neg. pos. neg.
cycle cycle cycle cycle cycle cycle cycle cycle cycle cycle
max M. [KNm] 107.8 -88.31 99.55 -88.41 95.22 -100.9 93.26 -93.98 102.5 -89.18
rotation @ max M. [°] 3.60 2.81 3.55 2.19 3.12 2.87 3.46 4.26 3.14 2.76
failure M. [KNm] 86.28 -70.65 79.64 -70.73 76.17 -80.79 74.61 -75.19 82.02 -71.34
rotation @ failure [°] 3.55 -3.27 4.24 -2.68 3.33 -3.13 3.65 -4.92 3.36 -3.32
40% max M. [KNm] 43.14 -35.32 39.82 -35.36 38.09 -40.39 37.30 -37.59 41.01 -35.67
rotation @ max M. [°] 1.44 -1.14 2.01 -0.90 1.27 -1.05 1.40 -2.08 1.13 -1.05
yield M. [KNm] 106.2 -81.31 - - 94.83 -99.52 - - 95.43 -86.82
rotation @ yield M. [°] 3.05 -2.35 - - 3.12 -2.58 - - 2.71 -2.55
stiffness. [KNm/°] 34.93 37.50 27.27 47.15 29.89 38.47 31.18 23.79 36.36 33.97
ductility [-] 1.18 1.19 - - 1.00 1.11 - - 1.16 1.08
Table 2: Summary statistics of tests results
pos. cycle
neg. cycle
Average Stdev. COV Average Stdev. COV
max M. [KNm] 93.33 12.38 0.13 -89.81 7.08 -0.08
rotation @ max M. [°] 2.92 1.02 0.35 -0.62 4.17 -6.72
failure M. [KNm] 74.66 9.90 0.13 -71.84 5.66 -0.08
rotation @ failure M. [°] 3.56 0.36 0.10 -3.66 0.86 -0.24
40% max M. [KNm] 37.33 4.95 0.13 -35.92 2.83 -0.08
rotation @ max M. [°] 1.30 0.31 0.24 -1.33 0.45 -0.34
yield M. [KNm] 96.62 9.68 0.10 -84.58 9.26 -0.11
rotation @ yield M. [°] 2.64 0.52 0.20 -2.65 0.47 -0.18
stiffness. [KNm/°] 32.34 7.50 0.23 32.46 8.42 0.26
When looking at the average ultimate capacity values
listed in Table 2 for the positive and negative cycle one
can see the high performance of this assembly. The
average ultimate moment resistances exceeded the
design moment capacity of the 130 mm x 304 20f-E
grade glulam member by a factor of 2 in the positive and
a factor of 1.9 in the negative cycle. Considering the
high stiffness of the connection with 32.34 KNm/° in the
positive and 32.46 KNm/° in the negative cycle
application of this connection in real structures seems

appropriate. As a result of eventual fracture of the STS
the ductility of this connection was relatively low. This
can be easily amended with changing the details of the
steel hardware to force the failure into a more ductile
mode by ensuring steel yielding can occur before screw
fracture and withdrawal.
4 CONCLUSIONS
The research projected presented in this paper evaluated
the performance of a moment resisting self-tapping
screw assembly under reverse cyclic load. The applied
tension and compression plates in combination with selftapping
wood screws arranged at 30° to the wood grain
of the beam member resulted in high moment resistance
and very high connection stiffness. Failure of the main
connecting members, the 10 mm x 240 mm ASSY VG
screws, was promoted through the thick bottom steel
plate and the large number of 12.5 mm bolts that were
used as connector at the bottom. The ultimate moment
resistance exceeded the bending moment design capacity
by a factor of 2 in the positive direction and a factor of
1.9 in the negative direction. Main failure modes
observed in the beam member were tension fracture and
withdrawal of screws. In the column member wood
splitting perpendicular to the grain and compression/
transverse shear failure occurred. At the steel shoe
bottom plate significant yielding and tensile failure of
bolts were observed in progressed cycles under heavy
loads.
Even though the ductility of this connection was
relatively low, the high capacity of the tension and
compression plates in combination with the STS allows
one to weakening sections of the steel shoe and therefore
promoting steel yielding instead of screw fracture.
Forcing the steel parts to yield will eliminate brittle
failure modes and increase ductility. In addition, the ease
of assembly and the potential for pre-manufacturing
instead of on site construction may broaden the field of
application for STS connections in large structures.
Creating timber connections that are capable of
exceeding the capacity of the members they connect may
significantly simplify the design of moment resisting
self-taping screw assemblies in the future. This will
form part of the on-going research and development
activities at UBC on moment resisting connections.
ACKNOWLEDGEMENT
The cooperation received from Schraubenwerk Gaisbach
Produktion (SWG) GmbH (www.swg-produktion.de) for
the supply of the tension and compression plates (ZDplates)
and the ASSY VG screws is appreciated.
[2] Blass H.J.: Schmid M.: Self-tapping screws as
reinforcement perpendicular to the grain in timber
connections. Lehrstuhl fuer Ingenieurholzbau und
Baukonstruktionen, University of Karlsruhe (TH)
Germany, date unknown.
[3] Bejtka I.: Verstaerkungen von Bauteilen aus Holz
mit Vollgewindeschrauben. Lehrstuhl fuer
Ingenieurholzbau und Baukonstruktionen,
University of Karlsruhe Germany, 2005.
[4] F. Lam., M. Gehloff., M. Closen.: Moment-resisting
bolted timber connections. Proceedings of the
Institution of Civil Engineers, Structures and
buildings 163, pages 267-274, 2010.
[5] M. Gehloff., M. Closen., F. Lam. Reduced edge
distances in bolted timber moment connections with
perpendicular to grain reinforcements. In 11 th World
Conference on Timber Engineering, 2010.
[6] M. Trautz, C. KOJ. Proceedings of the International
Association for Shell and Spatial Structures (IASS)
Symposium: Valencia, 2009. The RWTH Aachen
University.
[7] Gehloff M.: Pull-out Resistance of Self-Tapping
Wood Screws with Continuous Thread. Department
of Wood Science, The University of British
Columbia Vancouver, Canada, 2011.
[8] DIN (Deutsches Institut fuer Normung e.V.) DIN
1052:2004-08: Entwurf, Berechnung und
Bemessung von Holzbauwerken- Allgemeine
Bemessungsregeln und Bemessungsregeln fuer den
Hochbau. Beuth Verlag, Berlin, 2004.
[9] DIBt Allgemeine Bauaufsichtliche Zulassung,
Wuerth ASSY VG plus Vollgewindeschrauben als
Holzverbindungsmittel. From Deutsches Institut
fuer Bautechnik Berlin, Germany, 2006.
[10] EC 5, EN 1995-1-1: (D), Eurocode 5: Bemessung
und Konstrucktion von Holzbauten – Teil 1-1:
Allgemeines- Allgemeine Regeln und Regeln fuer
den Hochbau CEN/TC 250 Structural Eurocodes,
1994.
[11] Pruefbericht Nr. 106109: Versuche zur Ermittlung
des Trag- und Verformungsverhaltens von
biegesteifen Rahmenecken mit ZD-platten.
Karlsruher Institut fuer Technology, Versuchsanstalt
fuer Stahl, Holz und Steine Karlsruhe, Germany ,
2010.
[12] H. Krawinkler, F. Ibarra, L. A. Ayoub, R. Medina.
Development of a testing protocol for wood frame
structures. Rep. W-02, CUREE-Caltech Woodframe
Project, Stanford University Stanford, California
2001.
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