NEW DESIGN APPROACH FOR WOOD BRITTLE FAILURE

MECHANISMS IN TIMBER CONNECTIONS

ABSTRACT: Timber construction has experienced considerable progress in recent years. In such progress,

apart from the implementation of new engineered timber products, the advancement of timber joints has played

a significant role. The design procedures for timber connections in most design codes are based mainly on the

yielding capacity of the fasteners using the European Yield Model (EYM). While the EYM theory provides

accurate predictions for connections that fail in a ductile fashion, it does not take into account the failure of the

connections due to the brittle rupture of wood as the consequence of fasteners group effect. Such a significant

gap in the design of connections also applies to the New Zealand (NZS 3603) and Australian (AS 1720.1) timber

design standards. A new design approach is presented which allows the practitioners to predict the connection

capacity associated with different brittle wood failure mechanisms. An extensive testing regime has been

conducted on high load-transfer capacity joints using timber rivets under longitudinal and transverse loadings

on New Zealand Radiata Pine laminated veneer lumber (LVL) and glulam. The results verify the proposal and

prove its reliability. A design guide was also developed which could eventually become a design clause in the

next revision of the New Zealand timber design standard NZS 3603.

KEYWORDS: Connection, wood capacity, fastener capacity, timber rivet, failure mode, LVL, glulam

1 INTRODUCTION 1.1 GENERAL

Timber construction has experienced considerable

progress in recent years. In such progress, apart

from the implementation of new engineered timber

products, the advancement of timber joints has

played a significant role. Structures are detailed

from a combination of hinged and moment

resisting joints. Therefore, connections are often

the most critical components of any type of

structure. Evaluation of timber buildings damaged

after extreme wind and earthquake events have

shown that weak connections (Figure 1) are one of

the major causes of problems [1].

Figure 1: Connection brittle failure at Siemens Arena, Ballerup, Denmark

1 Pouyan Zarnani, Postdoctoral Fellow in Structural

Timber Engineering, Department of Civil and

Environmental Eng., University of Auckland.

Email: pouyan.zarnani@auckland.ac.nz 2 Pierre Quenneville, Professor of Timber Design,

Department of Civil and Environmental Eng., University

of Auckland. Email: p.quenneville@auckland.ac.nz

1.2 BACKGROUND

The design procedures for timber connections in

most design codes are based mainly on the yielding

capacity of the fasteners (Figure 2) using the

European Yield Model (EYM). While the EYM

theory [2] provides accurate predictions for

connections that fail in ductile fashion, it does not

take into account the failure of the connections due

to the brittle rupture of wood as the consequence of

fasteners group effect. Such a significant gap in the

design of connections also applies to the New

Zealand NZS 3603 [3] and Australian AS 1720.1

[4] timber design standards.

Figure 2: Fastener failure modes in single shear for timber connections (involving fastener bending and/or member bearing)

Based on the state-of-the-art of wood strength

prediction models for parallel-to-grain failure in

connections using dowel-type fasteners, different

methods consider the minimum, maximum or the

summation of the tensile and shear capacities of the

failed wood block planes. This results in

disagreements between the experimental values and

the predictions. It is postulated that these methods

are not appropriate since the stiffness in the wood

blocks adjacent to the tensile and shear planes

differs and this difference leads to uneven load

distribution amongst the resisting planes [5,6]. For

instance, in a block tear-out failure (Figure 3), the

Pouyan Zarnani1, Pierre Quenneville2

contribution of the bottom or lateral shear planes to

the wood resistance cannot simply be considered as

a function of their respective area as the connection

load is not shared uniformly among the resisting

planes due to the unequal stiffness of the adjacent

wood volumes loading the fasteners.

Figure 3: Wood block tear-out failure in two joints with identical shear and tensile resisting planes and different bottom and edge distances

In the case of perpendicular-to-grain loading, the

available models for the prediction of the splitting

failure of dowel-type connections are determined

generally based on a crack growth of the entire

member cross-section. These models can be

appropriate for stocky or rigid fasteners installed

through the full thickness of the wood member.

However, for slender dowel-type fasteners such as

timber rivets, particularly when the penetration

depth of the fastener does not cover the whole

member thickness, partial splitting combined with

rolling shear can happen rather than splitting of the

entire width (Figure 4).

Figure 4: Cross-section view of wood splitting perpendicular-to-grain: (a) full width failure mode, (b) partial width failure mode

2 NEW DESIGN APPROACH 2.1 WOOD BLOCK TEAR-OUT

RESISTANCE PARALLEL-TO-GRAIN

The wood block tear-out resistance under parallel-

to-grain loading is predicted using the stiffness-

based model proposed by Zarnani and Quenneville

[7]. The proposed analysis for wood strength is best

explained using the analogy of a linear elastic

spring system in which the applied load transfers

from the wood member to the failure planes in

conformity with the relative stiffness ratio of each

resisting volume adjacent to the individual failure

plane (Figure 5).

Figure 5: Proposed spring model assuming linear elasticity

The difference in the loads channelled to the tensile

and shear planes is a function of the modulus of

elasticity (E) and modulus of rigidity (G), the

volume of wood surrounding each of the failure

planes (bottom, end and edge distances - dz, da and

de) and also of the connection geometry (Figure 6).

For details regarding the determination of the

stiffness of the resisting planes, refer to Zarnani

and Quenneville [7].

Figure 6: Simplified analytical model

By predicting the stiffness of the wood surrounding

each of the failure planes (Kh, Kb and Kl), one can

predict the proportion of the total connection load

applied to each plane, KKR ii . By further

establishing the resistance of each of the failure

planes as a function of a strength criterion, one can

verify which of the failure planes governs the

resistance of the entire connection due to being

overloaded.

It should be asserted that the strength of the shear

planes cannot be higher than the tensile capacity of

the adjacent wood volume where the load is

channelled to these resisting planes. If the attracted

load by the resisting shear planes is larger than the

tensile capacity of the associated wood volume,

then the wood block torn out from the member

would be as wide as or as deep as the member.

Therefore, as shown in Figure 7, three different

failure modes are possible and the wood load

carrying capacity of the connection (Equation 1)

equals the minimum of the resistances

corresponding to these failure modes, Pw,a, Pw,b and

Pw,c.

Figure 7: Different possible failure modes of wood block tear-out

Pw = np . min (Pw,a, Pw,b, Pw,c) (1)

where

Pw,a = min (pwh, pwb, pwl) , Mode (a)

Pw,b = min (pwh, pwb) , Mode (b)

Pw,c = min (pwh, pwl) , Mode (c)

in which

pwh = wood resistance for failure

of head tensile plane

(2)

pwb = wood resistance for failure

of bottom shear plane

(3)

pwl = wood resistance for failure

of lateral shear planes

(4)

Due to the absence of the lateral and bottom shear

planes in the failure mode (b) and (c) respectively,

the following considerations are necessary for the

resistance calculation of these failure modes:

I. For failure mode (b), the Kl/Kh and Kl/Kb ratios

should be considered as zero; and the tensile

and shear areas as wide as the member.

II. For failure mode (c), the Kb/Kh and Kb/Kl ratios

should be considered as zero; and the tensile

and shear areas as deep as the member (in the

case of double-sided joint, half of the depth

should be considered).

In Equations (2) to (4), ft,m and fv,m are the wood

mean strength in tension and in shear along the

grain (MPa). Ath, Asb and Asl are the areas of the

head, bottom and lateral resisting planes with

respect to the wood effective thickness, tef,

subjected to tension and shear stresses. Also, Cab

and Cal are the ratios of the average to maximum

stresses on the bottom and lateral shear planes

respectively [7]. np is the number of the plates

equal to 1 or 2 for one-sided or double-sided joints,

respectively.

It is also important to note that the connection

resistance corresponding to each failure mode is a

summation of the critical plane failure load, plus

the load carried by the other planes. Thus, when

one plane fails, then the entire connection load

transfers to the remaining planes in accordance

with their relative stiffness ratios. It could be

possible that the occurrence of the first failure of

one plane does not correspond with the maximum

load of the connection. Therefore, it is

recommended that after the failure of one plane,

one recalculates the wood resistance for the

remaining planes to verify whether the residual

planes can resist a higher load by defining: Kh/Kb

and Kh/Kl = 0 after the head tensile plane failure;

Kb/Kh and Kb/Kl = 0 after the bottom shear plane

failure; and Kl/Kh and Kl/Kb = 0 after the lateral

shear planes failure.

In the case of fasteners which are inserted into

predrilled holes, the area corresponding to the

cutting diameter is to be subtracted from the

resisting plane surfaces. This affects the strength of

the tensile and shear resisting planes and not their

stiffnesses.

2.2 WOOD SPLITTING RESISTANCE

PERPENDICULAR-TO-GRAIN

The wood splitting strength in perpendicular-to-

grain loading is predicted using the model proposed

by Zarnani and Quenneville [8]. The proposed

approach is based on two different possible crack

formations on the member cross-section: with

partial splitting on each side of the member

corresponding to the effective embedment depth, tef

(Figure 4b) or with full width splitting (Figure 4a).

In fact, for connections with a large penetration

depth in slender members, the governing failure

mode will be the full width splitting, and as the

ratio of member thickness to penetration depth

increases, the transition of wood failure mode from

full to partial width splitting will occur (Figure 8).

Figure 8: Occurrence zone of possible failure modes of wood splitting

Therefore, the ultimate splitting resistance of the

connection is determined as the minimum strength

corresponding to these two failure modes and is

given by Equation (5):

1 b l

t ,m th

h h

K Kf A ( )

K K

1 h l

v ,m ab sb

b b

K Kf C A ( )

K K

1 h b

v,m al sl

l l

K Kf C A ( )

K K

Pw = min (Ps,tef, Ps,b) (5)

The wood capacity for partial width splitting, Ps,tef,

is predicted using a stress-based analysis (Equation

6) and involves the perpendicular-to-grain tensile

capacity of the splitting surface of the wood

corresponding to the effective embedment depth, tef

and the crack length that propagates along the

member. The crack length along the member is

considered as the summation of the joint net section

width, wnet, and the symmetrical crack growth on

the left and right sides of the joint as a factor of the

effective depth, he (Figure 9).

Figure 9: Effective crack length on either side of the joint

In Equation (6), ftp is the tensile strength

perpendicular-to-grain; da,L and da,R are the

unloaded end distance on the left and right side of

the joint, respectively; Ct is a coefficient function of

the unloaded edge distance and the connection

length.

Ps,tef =npCtftptef [wnet+min(βhe,da,L)+min(βhe,da,R)]

(6)

The predictive equation presented for wood

splitting in the entire member cross-section, Ps,b

(Eq. (7)) is adopted from the fracture mechanics

based model developed by Van der Put and Leijten

[9]. The significant difference is the application of

the η factor which accounts for the effect of

unloaded end distance and the connection width.

(7)

in which,

and Cfp is the fracture parameter. For more details

regarding the wood splitting model, refer to

Zarnani and Quenneville [8].

3 EXPERIMENTAL PROGRAM As demonstrated over numerous studies, small-

diameter fasteners have shown a significant

advantage over large-diameter ones such as bolts

which cause large localized stresses and force

brittle ruptures in the timber. Of this family of

fasteners, the timber rivet is a well-established

example in timber connection technology [10].

Timber rivets are hardened nails with a rectangular

cross section of 3.2 by 6.4 mm. They are available

in three standard lengths; 40, 65 or 90 mm (Figure

10). Timber rivets are always driven with the long

axis parallel to the wood grain. Rivets are tight-fit

fasteners since the head is wedged into the steel

plate hole. Timber rivets provide high load-transfer

capacity, high stiffness and easy to install steel-

timber connections [11]. Rivets are part of the

Canadian CSA-O86 [12] and American NDS [13]

wood standards. However, there is no closed-form

solution for the strength prediction of this type of

connection. Therefore, timber rivets were used in

the current joint failure tests to verify the model.

Figure 10: Timber Rivet

3.1 TEST SETUP

Laboratory tests were set up to prompt wood

failures and maximize the amount of observations

on the brittle mechanism by considering higher

capacity for the fasteners compared to the wood.

Specimens were manufactured from New Zealand

Radiata Pine LVL grade 11 and GL8 grade glulam.

An extensive testing regime was conducted

including 32 test groups under parallel-to-grain

loading and 24 test groups under perpendicular-to-

grain loading. 3 replicates were tested for each

group of specimens for LVL and 4 replicates for

glulam. The specimens had riveted plates on both

faces of timber, resulting in a symmetric

connection. The steel side plates were 8.4 mm thick

of 300 grade (Fy = 300 MPa) with predrilled 6.8

mm holes to ensure adequate fixity of the rivet

head. The effect of geometry parameters such as

connection width and length, fastener penetration

depth, loaded and unloaded edge distances, end

distance, and member thickness were evaluated.

For more details regarding the connection

configurations, refer to Zarnani and Quenneville

[8,14].

The testing protocol outlined in ISO 6891 [15] was

followed. The tension load was applied to the

specimens using a displacement controlled loading

system. The deformation of the connection was

, 2

1

es b fp

e

hP bC

h

h

, ,min( , ) min( , )

2

e a L e a R net

e

h d h d w

h

L=40 mm

L=65 mm

L=90 mm

measured continously with a pair of symmetrically

placed linear variable differential transformers

(LVDT). The loading rate was adjusted to 1

mm/min and kept constant until the occurence of

failure in both or either side of the riveted

connections. A typical specimen in the testing

frame is shown in Figure 11.

Figure 11: Typical specimens in testing apparatus: (a) longitudinal loading, (b) transverse loading - mid-span, (c) transverse loading - end of member

3.2 MATERIAL PROPERTIES

All specimens were conditioned to 20°C and 65%

relative humidity to attain a target 12% equilibrium

moisture condition (EMC). The wood had an

average density of 605 and 465 kg/m3 for LVL and

glulam members respectively. For the connection

capacity parallel-to-grain, the average tensile and

shear strengths evaluated were 34.3 MPa

(coefficient of variation (COV)=12%) and 6.8 MPa

(COV=10%) for RP-LVL and 24.1 MPa

(COV=24%) and 4.2 MPa (COV=15%) for RP-

glulam (samples from inner laminations)

respectively [14]. For the stiffness properties, based

on data available in the literature, an average ratio

of modulus of rigidity to modulus of elasticity

(G/E) is considered equal to 0.045 and 0.069 for

LVL and glulam respectively in order to make the

planes’ stiffness equations independent of G and E

values. The fracture parameter value, Cfp, reported

in Jensen et al. [16] and tensile strength

perpendicular-to-grain values, ftp, evaluated by

Song [17] were used in the proposed splitting

model. The average Cfp and ftp were 22.7 N/mm1.5

and 2.06 MPa (for the tangential direction with a

COV=18%) for RP LVL and 18.4 N/mm1.5

and

1.99 MPa (at 45° to the radial direction with a

COV=24%) for RP glulam.

3.3 TEST OBSERVATIONS

Test series with a tightly spaced rivet pattern

exhibited a brittle/mixed failure mode. For parallel-

to-grain loading, a sudden failure happened where a

block of wood bounded by the rivet group

perimeter was pulled away from either one side or

both sides of the specimens (Figure 12a). As shown

in Figure 12b, in brittle failure mode, the failed

block thickness, tblock corresponds to the elastic

deformation of the rivets (tef,e) since no plastic

deflection was observed. However, in the mixed

failure mode (Figure 12c), the tblock is significantly

lower with observable small deformation of the

rivets. In the mixed failure mode cases, the load-

carrying capacity of the wood is based on the

stiffness and strength of the tensile and shear planes

corresponding to the bearing length of the fastener

(tef,y) depending on its yielding mode.

Figure 12: Wood failure parallel-to-grain: (a) block tear-out bounded by the rivet cluster perimeter, (b) brittle failure - rivets within the elastic range, (c) mixed failure - rivets with small deformation

For the tests subjected to a perpendicular-to-grain

loading, splitting of the wood occurred along the

row of rivets next to the unloaded edge and

propagated towards the timber member ends until

reaching the unstable zone (Figure 13a). The

splitting crack formed either through the entire

member width (for thin members with high

penetration-to-thickness ratios), as shown in Figure

13b or with a depth similar to the rivet effective

embedment depth (for thick members with low

penetration-to- thickness ratios), as shown in

Figure 13c.

Figure 13: Wood failure perpendicular-to-grain: (a) crack propagation along the top row of rivets, (b) full width splitting, (c) partial width splitting

4 VERIFICATION OF PROPOSED

DESIGN APPROACH 4.1 RIVETED JOINT SUBJECTED TO

PARALLEL-TO-GRAIN LOADING

The connection ultimate capacities were calculated

using the proposed stiffness-based model. For the

brittle failure mode, the elastic deformation of the

rivet was considered to determine the tef,e [7]. The

estimated tef,e was 27.1, 45.5 and 58.9 mm for a

rivet penetration (Lp) equal to 28.5, 53.5 and 78.5

mm correspondingly. In the case of mixed failure,

the tef,y was predicted based on the bearing length

corresponding to the governing yielding mode of

the rivets equal to 28.4, 40.4 and 24.7 mm for the

different penetration lengths. The tef,y value for the

longer rivet is lower compared to the other ones.

This is due to the formation of the two plastic

hinges (see Figure 2f) for this length of rivet during

the yielding failure, whereas for the smaller rivet

sizes, there is one plastic hinge (see Figure 2d).

Figure 14 shows the strength predictions of the

experimental groups compared to the test results.

One can note that the proposed analysis results in

precise predictions with a coefficient of

determination (r2) of 0.89, a mean absolute error

(MAE) of 7.8% and a standard deviation (STDEV)

of 10.3%.

Figure 14: Predictions vs. observations for the joint load-carrying capacity under parallel-to-grain loading

4.2 RIVETED JOINT SUBJECTED TO

PERPENDICULAR-TO-GRAIN

LOADING

There was a considerable difference between the

predicted wood splitting strength and the rivet

yielding resistance for the tested joints, therefore,

the splitting failure was predicted to occur in a

brittle fashion corresponding to the rivet elastic

deformation [8]. The estimated tef,e was 24.2, 40.1

and 51.0 mm for the Lp equal to 28.5, 53.5 and 78.5

mm respectively. The proposed design approach

leads to a relatively good agreement between the

predictions and the test results with a coefficient of

determination (r2) of 0.78, a mean absolute error

(MAE) of 17.2% and a standard deviation

(STDEV) of 15.1% (Figure 15).

One can note that the proposed design approaches

result in precise predictions with high coefficients

of determination (r2) and low mean absolute errors

(MAE) and standard deviations (STDEV). This

new design approach can be extended to other

small dowel-type fasteners such as nails and screws

for connection design improvement and failure

mode prediction [18].

Figure 15: Predictions vs. observations for the joint load-carrying capacity under perpendicular-to-grain loading

5 FROM THEORY TO PRACTICE 5.1 DEVELOPED DESIGN GUIDE

A design guide has been developed [19] which

outlines the new method to establish capacities for

timber rivet connections (Figure 16). The guide

provides an aid for engineers for designing timber

rivet connections in structural wood products made

of Australasian Radiata Pine, including sawn

timber, glulam and laminated veneer lumber

(LVL). The design guide includes the design

checks for both wood and fastener load-carrying

capacities under different loading directions;

parallel- and perpendicular-to-grain. Furthermore,

it allows practitioners to predict the potential brittle

(wood block tear-out/splitting), ductile (rivet

yielding) and mixed failure modes of the

connection. To illustrate the design process, a

series of examples are also included.

Figure 15: Developed design guide

The guide is accessible through EXPAN (formerly

STIC) administered by the Engineered Wood

Products Association of Australasia (EWPAA).

5.2 CASE STUDIES

5.2.1 Carterton Event Centre

Timber rivets have been used for the first time in

New Zealand in 2011, in the truss connections of

the Carterton Event Centre (Figure 16). The trusses

were constructed using LVL members supplied by

Juken New Zealand Ltd. The auditorium trusses

were up to 24.6m long and 4.8m high, so they were

delivered to site in two pieces.

Figure 16: Riveted connections of Carterton Event Centre (McIntosh Timber Laminates Ltd.)

The riveted connection uses easily man handle-able

components that allowed the required mid-span

splice connection to be completed without

specialist lift equipment.

Use of the rivets allowed the fabricator, McIntosh

Timber Laminates of East Tamaki, to save over

$30k on this project when compared to the detailed

bolted connection option. It also allowed for

adjustments on site not possible with bolted

fastenings. The timber rivets were found very user-

friendly both in the work shop and onsite with

significantly less visual impact and material cost

compared to conventional fasteners [20].

5.2.2 Trimble Building in Christchurch

In another project by TimberLab Solutions Ltd.

(formerly McIntosh Timber Laminates Ltd.), using

rivets in the connections of the structure and energy

dissipating system of the building (Figure 17)

demonstrates the advantages of this timber fastener.

The Trimble Building in south-west Christchurch

was damaged by the September 2010 and February

2011 earthquakes. The new building was a design-

build project, which was undertaken by Mainzeal

and Opus, commencing in February 2012.

Figure 17: Riveted connections of Trimble Building in Christchurch (c/o A. Buchanan)

The building holds over 6,000 m2 of office space

over two levels and utilises LVL Pres-Lam frames

in one direction and Pres-Lam walls in the other to

resist seismic loads. The LVL of this building was

supplied by Carter Holt Harvey Ltd. The principal

structural engineer from Opus, and the structural

design team leader for the Trimble project found

that the compact timber rivets provide high strength

and stiff connections to take significant seismic

loads [21].

5.2.3 Portal frame for URM Building seismic

retrofit

In this recent project, EQStruc Ltd implemented

efficiently LVL timber portal frames to resist

seismic loads for an unreinforced masonry structure

about 45m long by 12m wide (Figure 18). To

provide strong and stiff joints between the timber

elements, rivet connections were designed with

controlled deflections. The client, a first-time user

of timber portal frames in an unreinforced masonry

(URM) building seismic retrofit, found that the

rivet connections offered a cost-effective solution.

Figure 18: Riveted connections of portal frames for URM building seismic retrofit (EQStruc Ltd.)

5.2.4 Kaikoura District Council Building

Recently, Innovations Nelson Ltd. used timber

rivets in the hold-down connections of the CLT

shear walls of the Kaikoura event centre (Figure

19). Elastic deformation of the rivet connections

were limited to 0.5 mm to satisfy the targeted

system ductility only through the extension of the

energy dissipaters under earthquake events.

Innovations Nelson Ltd. is working on another

similar project related to a 2-story commercial

building in which rivets will be applied to provide

high capacity and high stiffness hold-down

connections.

Figure 19: Riveted connections of CLT shear walls (Innovations Nelson Ltd.)

6 CONCLUSIONS Through the emergence of advanced engineered

timber products as reliable structural alternatives

along with scientific developments in timber

engineering, many international building standards

are revising their codes to accept high-rise timber

buildings. In such heavy structures where the joints

need to transfer large loads either parallel- or

perpendicular-to-grain, the brittle group failure of

the joint is more susceptible. The design procedures

for timber connections in most design codes are

based mainly on the European Yield Model which

is only applicable for the ductile failure of the

connections but not the brittle failure modes. Such

a significant gap in the design of connections also

applies to the New Zealand NZS 3603 [3] and

Australian AS 1720.1 [4] timber design standards.

A new design approach has been presented to

determine the wood resistance in connection brittle

failure mechanisms. The proposed method has been

verified using extensive testing regimes conducted

on riveted joints under longitudinal and transverse

loadings on New Zealand Radiata Pine LVL and

glulam. Based on the proposed design model, an

efficient connection design can be made by

decreasing the difference between the capacity of

the wood and the rivets. This new design approach

can be extended to other small dowel-type fasteners

such as nails and screws for connection design

improvement and failure mode prediction.

ACKNOWLEDGEMENT The authors wish to thank the New Zealand

Structural Timber Innovation Company (STIC) for

funding this research work. The authors wish also

to express their gratitude to University of Auckland

undergraduate students Samuel Wong and Shuai

Ma, who did the experimental work related to

glulam splitting as part of their final year project.

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