DETAILED SEISMIC ASSESSMENT AND IMPROVEMENT PROCEDURE FOR VINTAGE FLEXIBLE TIMBER DIAPHRAGMS

Currently there is little guidance available on an experimentally-validated detailed seismic assessment procedure for vintage flexible timber diaphragms such as are routinely encountered in New Zealand unreinforced masonry buildings. The results from recent testing of full-scale diaphragms are presented and interpreted with particular attention given to the definition of shear stiffness and shear strength values, whilst acknowledging that the recommendations derive from a small data set. References are provided to information previously published elsewhere to justify the theoretical framework adopted, and the procedure is linked to ASCE 41-13 for guidance regarding diaphragm scenarios that have not been studied by the authors. A procedure is provided to account for the effects on diaphragm response of supplementary stiffness due to masonry end walls. The performance of several diaphragms that were improved with either overlays or underlays is reported as potential proof-tested standard solutions. The assessment procedure is demonstrated by providing a mock worked example of a detailed diaphragm assessment.


INTRODUCTION
New Zealand unreinforced masonry (URM) buildings are typically between 80 and 120 years old, with their timber diaphragms therefore of the same age and usually constructed with native timber species.It is known that the nails used to construct these diaphragms are soft when compared to currently manufactured nails, but equally it is noted that the timber has often hardened from a century of drying and that the diaphragm appears to have 'locked up' from a century of use.Whilst it is recognised that vintage timber diaphragms are flexible, anecdotal evidence and observations from the 2010/2011 Canterbury earthquakes suggest that for many URM buildings large diaphragm displacements were not observed.
Recorded seismic response from instrumented buildings having flexible diaphragms has provided support for past analytical studies, such as by Tena Colunga and Abrams (1996), that have shown that design criteria based on rigid diaphragm assumption are not necessarily conservative when applied to buildings with flexible diaphragms, especially when determining the magnitude of seismic input, as demonstrated by Raggett and Rojahn (1991).Hence there is currently a significant lack of clarity regarding a suitable detailed seismic assessment procedure for vintage flexible timber diaphragms.
Information is presented below to explain the source of the recommendations proposed for a revision to the NZSEE 'Red Book' (NZSEE 2006) that pertain to the detailed seismic assessment of vintage flexible timber diaphragms.It is acknowledged that the stiffness values for straight sheathed diaphragms may require future modification if additional in-situ testing is undertaken that indicates that alternative stiffness values are more appropriate.Brignola et al. (2012) have provided detailed information on relevant past testing related to the seismic performance of flexible timber diaphragms, plus the effect that the stiffness of the wall-diaphragm connections has on the global stiffness of the diaphragm.They have also presented details related to existing diaphragm assessment procedures in USA and Italy, and a prior procedure used in New Zealand.

DISPLACEMENT ACCEPTANCE CRITERIA
In order to develop a detailed seismic assessment procedure for vintage flexible timber diaphragms as found in New Zealand URM buildings it is necessary to first establish a displacement acceptance criteria.Three distinct criterions are promulgated, associated with: (1) limiting the extent of diaphragm distortion; (2) controlling diaphragm displacements to an extent that they minimise detrimental influence on outof-plane wall deformations; and (3) limiting global building drift ratios.

Limiting diaphragm distortion
Flexible timber diaphragms are known to deform as shear beams (Wilson et al. 2013a(Wilson et al. , 2013b) ) and to exhibit orthotropic behaviour when loaded in their plane, due to the orientation of the joists (see Figure 1(a)).The diaphragm mid-span displacement can be related to the maximum permissible rotation of nail joint connections (with maximum rotations occurring at the diaphragm boundaries where shear deformations are largest), with nail joint rotation being dictated by the spacing s (in mm units) between the nail couple and the slip in each nail e n (in mm units) resulting from the applied shear force (see Figure 1(b)).For single straight sheathed diaphragms loaded perpendicular to the orientation of the joists, as per NZSEE (2006) Eq. 11A(2), the relationship is given by: (1) where L is the diaphragm span.For straight sheathed diaphragms loaded parallel to the orientation of the joists the corresponding relationship is: (2) As noted in NZSEE (2006) section 11.3.1, the flooring of existing timber diaphragms is usually 25-50 mm thick, with board widths of 100-200 mm.The specific spacing between nails within a nail joint should be measured during a detailed inspection of the diaphragm, but is typically about 75% of the board width and in the order of s = 100 mm.
Limited measured data exists regarding nail slip characteristics for New Zealand vintage flexible timber diaphragms, with Wilson et al. (2014b) having reported data for nail connections extracted from several relevant diaphragm samples.A further complication is to establish the tolerable condition of the most extremely loaded nail joints that merits definition of the diaphragm acceptance criteria.Hence it is suggested that 'first yield' (or departure from elastic response) is too restrictive as a criterion for the most severely loaded joints, and instead Wilson et al. (2014b) report the nail slip at maximum connection strength as having a mean value of 8.1 mm.Given that there is some variance amongst measured results, a reduction factor of 0.75 is proposed and a nail slip of e n = 6.0 mm is recommended.Strictly speaking Equations 1 and 2 do not apply once inelastic response occurs within the most severely loaded nail connections of the diaphragm, but recalling that the majority of nail connections within the diaphragm will remain in their elastic state, the relationships are deemed to still be appropriate.Hence it follows that for an assumed s =100 mm Equation 1 would provide Δ max = L/33.33 for loading oriented perpendicular to the joists and Δ max = L/53.33 for loading oriented parallel to the joists.For a URM building having plan dimensions of 8 m  12 m, a wall thickness at diaphragm height of 230 mm, and with joist spanning the short dimension, the maximum acceptable midspan diaphragm displacements associated with limiting diaphragm distortion would be 226 mm for loading perpendicular to joist and 216 mm for loading parallel to joist.

Limiting wall instability
Based on experimental results reported by Vaculik (2012) and by Griffith el al. (2007), the diaphragm maximum in-plane displacement measured with respect to the diaphragm URM side walls shall not exceed 50% of the thickness of the adjacent URM end walls.
For multi-storey buildings consideration must be given to the situation that potentially arises when diaphragms at different storey heights deform outof-phase.The worst case scenario would develop when the diaphragms on two adjacent levels are 100% out-of-phase (see Figure 2(a)).However, this deformation is considered to be a higher mode with smaller associated displacements and hence is deemed to be too severe as an acceptance criteria, and instead the critical condition is associated with the assumption that the diaphragm at any given storey height is deformed to its maximum mid-span displacement while the diaphragms on adjacent levels are undeformed.
For solid URM walls the wall thickness is likely to be approximately 230 mm for a 2 leaf wall, and so this criterion would result in a maximum diaphragm mid-span displacement of 115 mm (see Figure 2(b)).For solid walls having a greater thickness the acceptable diaphragm mid-span displacement would be proportionally increased.For cavity construction with adequate cavity ties installed, the inner masonry wythe is usually the load bearing wythe and this criterion will require the maximum acceptable diaphragm displacement to be limited to 50% of the thickness of the inner wythe.For a cavity wall having a single brick thickness of 110 mm this criterion would result in a maximum permissible mid-span diaphragm displacement of only 55 mm.This criterion may also govern for unusual building configurations.

Limiting building drift
NZS 1170.5:2004clause 7.5.1 limits inter-storey deflection to 2.5%, including contributions from global building deformation due to foundation rotation (see Figure 2(c)).However, the treatment of building deformations attributable to flexible foundation response is outside the scope of the reported study.
The total mid-span diaphragm displacement relative to the ground (excluding foundation rotation) is composed of deformations due to the in-plane loaded URM side walls and to deformation of the diaphragm with respect to the side walls, with the relative contributions of the two deformation modes being dependent on the characteristics of the diaphragm and wall-diaphragm connections, but also on the characteristics of the URM side walls, such as wall cross-section details and the geometry of wall penetrations.
The information in Figure 3 is reproduced to show that in general the in-plane loaded side walls will deform with small drifts.Note that from Figure 3(b) it can be established that the strength plateau was reached at a roof-height displacement of approximately 10 mm, corresponding to a drift of 0.16% or an inter-storey displacement of 6.2 mm for a wall height of 4,000 mm.
The suggested approach is to assume that in general the URM side walls may be treated as approximately rigid for the purposes of diaphragm assessment, but that for cases where the side walls are particularly slender, heavily penetrated by openings, or there is significant foundation flexibility, more detailed study is required to establish the limiting diaphragm displacement to ensure that a total diaphragm mid-span drift of 2.5% is not exceeded (see Figure 2(c)).Consequently, for a URM building with a storey height of 4,000 mm this criterion will result in an acceptable diaphragm mid-span displacement of 100 mm.

Criterion assumed to govern for the interpretation of physical test data
As the walls of URM buildings typically have inter-storey heights ranging from 3,500 mm to 4,500 mm, a representative height of 4,000 mm has been adopted here and the acceptable mid-span diaphragm displacement is therefore adopted as 100 mm for the treatment presented below.It is noted that this 100 mm mid-span diaphragm displacement will typically correspond to less than 50% of the cross-section thickness of adjacent solid URM walls.However, for an inner load bearing wythe of a cavity wall that has a single brick thickness (usually approximately 110 mm) the limiting mid-span diaphragm displacement will be reduced to 55 mm, emphasising the importance of clearly determining the wall cross-section characteristics.For cavity construction where the inner load bearing wythe is two bricks thick (typically 230 mm), the 2.5% drift criteria for the mid-span diaphragm displacement limit will govern for walls having a height of less than or equal to 4,600 mm.

DETERMINING EQUIVALENT SHEAR STIFFNESS FOR STRAIGHT SHEATHED DIAPHRAGMS
The force-displacement response of straight sheathed diaphragms is known to be nonlinear, and has been shown to be adequately modelled by second-order curves such as suggested by ABK (1982).Shown in Figure 4(a) is the result of a cyclic test performed by Giongo et al. (2013) on a vintage straight sheathed timber floor as typically encountered in New Zealand unreinforced masonry buildings.It can be seen that the as-built vintage straight sheathed flexible timber diaphragm deformed to lateral displacements exceeding the 100 mm acceptance criteria without exhibiting a strength plateau, and hence the diaphragm response can be approximated as an equivalent elastic system over the displacement range of interest.As shown in Figure 4(b), care needs to be exercised when defining the secant stiffness for the diaphragm so that a realistic determination can be made regarding whether the diaphragm has displaced to less than or greater than the acceptable displacement when subjected to ultimate limit state loads.mechanics of flexible timber diaphragms most correctly match that of a shear beam (as has previously been assumed in most assessment procedures).A feature of a shear beam is that both the shear stiffness and the shear strength are proportional to the section depth B. During diaphragm in-plane deformation due to earthquake shaking the deformed profile of the diaphragm is assumed to be approximately parabolic and hence the acceleration profile along the diaphragm is also assumed to be parabolic.For in-plane horizontal diaphragm accelerations due to tributary weight from the two masonry boundary end walls, the diaphragm self-weight, and any imposed live load on the diaphragm, the total lateral force on the diaphragm is defined by C(T d )W trib where C(T d ) is the horizontal design action coefficient for the diaphragm.C(T d ) is dependent on diaphragm period, which in turn is dependent on diaphragm shear stiffness.Consequently, reliable assessment of vintage flexible timber diaphragms requires accurate estimation of the diaphragm shear stiffness.
For a diaphragm of length L and depth B and assumed to have a rectangular cross-section, the diaphragm mid-span displacement when loaded with a parabolic load having a total value of C(T d )W trib is defined by Wilson et al. (2013b) as: (3) where G d is the diaphragm equivalent shear stiffness.Note that for diaphragms that are bounded on one or both sides by comparatively flexible timber framed partition walls (when compared to the in-plane stiffness of URM solid or cavity walls) there is no experimental test data currently available regarding likely response, and such a scenario will require a specific study.Hence when considering experimental data for sections of diaphragm displaced to 100 mm, the above equation can be re-expressed to establish the experimentally measured secant shear stiffness at 100 mm as: (4) where F is the total in-plane force applied to the diaphragm when displaced to 100 mm.As noted in Figure 4(b), the interpretation of physical test data for a mid-span displacement of 100 mm (the governing displacement limit when bounded by solid URM walls) will lead to a slightly conservative estimate of diaphragm equivalent shear stiffness for situations where the diaphragm is bounded by single brick wythe cavity URM walls and the limiting diaphragm displacement is 55 mm.

INTERPRETING TEST DATA TO OBTAIN EQUIVALENT SHEAR STIFFNESS VALUES
It is recognised that it is not possible to make strong recommendations on likely equivalent shear stiffness values when only a small set of field diaphragm test data exists, but Giongo et al. (2013) have published the results from in-situ testing of two sections (9.6 m × 5.6 m and 9.6 m × 4.7 m) of a vintage straight sheathed timber diaphragm that were both loaded perpendicular to the orientation of the joists, with the joists oriented parallel to the 9.6 m dimension.The building within which the diaphragm sections were housed was located in Whanganui (lower North Island of New Zealand) and was constructed in 1913, making the diaphragms 99 years old at the time of testing.Both diaphragm test sections were in poor condition, as shown in Figure 5.In order to adequately load the diaphragms it was necessary to remediate the walldiaphragm anchorages loaded in shear using 16 mm epoxygrouted anchorages at spacings no greater than 2.0 m.This test data is significant because the diaphragm was constructed of native timber matai floorboards (130 mm × 22 mm) overlying native rimu timber joists (50 mm × 300 mm) with an average spacing of 450 mm.Two wire nails (≈ 2.93 mm × 40 mm) were present at each floorboard-to-joist intersection, with an average spacing of 100 mm.In the region where the joists were overlapped (see Figure 6) there were two wire nails that were considered to provide negligible connection between the joists.
As indicated in Figure 6, the test set-up was self-reacting by enabling loads to be applied back to the building corners and the loading mechanism was designed to approximately replicate the parabolic load distribution anticipated during earthquake excitation.The loading system utilised two opposing cable systems to allowed reversed cyclic loading to be undertaken, with the cables of the opposing loading system being slack for each half cycle of loading.Low friction bearing pads were positioned between the joists and the underlying mid-span supporting timber beam such that potential sources of friction were alleviated.
For the two lateral load tests on the diaphragm that is depicted in Figure 5 and Figure 6, the secant stiffness values at 100 mm of mid-span displacement are reported in Table 1.From Table 1 the effect of joist lap conditions on diaphragm shear stiffness  is evident.Both diaphragms reported in Table 1 had discontinuous joists, but during testing of Diaphragm B the effect of discontinuous joists was more pronounced due to the greater number of floor boards that were not continuous at the location of joist lap.Wilson (2012) used Finite Element modelling to numerically investigate the effect of joist continuity on diaphragm shear stiffness and found that diaphragms in poor condition and having continuous or lapped joists are 27.5% stiffer than diaphragms with discontinuous joists.This result is confirmed to some extent by the 13% increase in shear stiffness for Diaphragm A where the effect of discontinuous joists was less pronounced due to the greater number of floor boards that were continuous at the location of the joist lap.2, where it is evident that the shear stiffness of newly constructed timber diaphragms was notably less than the values reported in Table 1 for in-situ testing of vintage flexible diaphragms.This outcome is principally explained by recognising the increased timber density of aged and therefore seasoned timber, which may result in timber densities being elevated by 50%, and the associated influence of timber density on the performance of nailed connections.
From Table 2 it can be established that the ratio of stiffnesses for testing parallel and perpendicular to the joist orientation is approximately 1.45/1.00.This finding is consistent but a little higher than the ratio of 1.30/1.00obtained by Wilson (2012) using Finite Element modelling.From consideration of these two results, it is assumed that a ratio of 1.33/1.00suitably    1) by using 2.85 mm × 75 mm nails, and measured improved stiffness as shown in Figure 7.This finding is consistent with anecdotal evidence that the renailing of vintage flexible timber diaphragms elevates their stiffness.On this basis it is recommended that the stiffness values presented in Table 3 be elevated by 20% when the diaphragm is re-nailed at every floorboard-to-joist intersection using modern nails applied by nail gun.
Timber roofs of URM buildings were frequently constructed with both a roof lining and a ceiling lining, and roof linings in New Zealand vintage URM buildings have been observed to sometimes be constructed of either diagonal sheathing or double layers of sheathing.Hence, it is likely that such roof diaphragms are significantly stiffer than are mid-height floor diaphragms.For diaphragms constructed using sheathing types other than straight sheathed (such as may be encountered in roofs) the diaphragm shear stiffness values listed in Table 3 should be multiplied by the values given in Table 4, which are derived from ASCE 41-13.When roof linings and ceiling linings are both assumed to be effective in providing stiffness their contributions should be added.

EFFECT OF DIAPHRAGM PENETRATION
Wilson (2012) undertook Finite Element modelling to investigate the influence of diaphragm penetrations on diaphragm shear stiffness and shear strength and established that for typically sized diaphragm penetrations (usually a rectangular shaped area which accommodates a staircase, corresponding to less than 10% of the diaphragm gross area) linear scaling of the ratio of net to gross diaphragm area was

Midspan Displacement -Δ [mm]
As-built Renailed   A special study should be conducted to establish the influence of non-typical sizes of diaphragm penetrations.Therefore a reduced diaphragm shear stiffness to account for the presence of diaphragm penetrations can be calculated as given below. (5)

ADDED STIFFNESS OF BOUNDARY WALLS
The effective diaphragm shear stiffness must be further modified to account for stiffness of the URM end walls deforming out-of-plane in collaboration with the flexible timber diaphragm.Hence: (6) where may be determined using any rational procedure to account for the added stiffness and incompatibility of deformation modes that arise due to collaborative deformation of the URM end walls displacing out-of-plane in flexure and the diaphragm deforming as a shear beam (see Figure 8).Depending on the boundary conditions, the URM end walls may resist out-of-plane deformations by developing two-way bending actions as shown in Figure 8.However, because of the lack of an established procedure to determine the out-of-plane stiffness of two-way spanning walls, simple analytical models were used to provide an estimated value of by instead assuming one-way vertically spanning wall behaviour.Figure 9(a) shows a simplified model of a flexible diaphragm excluding out-of-plane wall stiffness and Figure 9(b) shows a similar model that includes the out-of-plane wall stiffness.In these models refers to the mass of the diaphragm and refers to the mass of the out-of-plane responding URM walls, with the coefficient 2 describing the fact that there are two end walls and the co-efficient 1⁄2 describing the tributary mass of these end walls with respect to the diaphragm.
Assuming that the out-of-plane responding walls are cracked at floor levels (see Figure 10) and that they respond in oneway vertically spanning deformation modes as indicated in Figure 2(a), the stiffness contribution from these walls can be calculated by assuming cracked wall rocking action as shown in Figure 10.The contribution to diaphragm shear stiffness from out-of-plane responding walls located in the middle stories of multi-storey buildings can then be calculated as the sum of the stiffness of the cantilever walls in the floors located directly above and below the diaphragm.It is noteworthy that such an approach is conservative as the effect of the applied overburden, such as for the ground storey wall in Figure 10, is ignored.
The timber diaphragm has been modelled as a shear beam spanning between the in-plane loaded side walls, and the distributed mass of the combined diaphragm and two out-ofplane loaded URM boundary end walls has been represented by an equivalent lumped mass corresponding to a generalised single-degree-of-freedom system.
Using this modelling approach, the generalised lumped mass for a shear beam is 8/15 of the total mass, as indicated in Figure 9.The degree of freedom is located at the diaphragm mid-span, and k d represents the diaphragm shear stiffness associated with midspan displacement and a parabolic load distribution.The natural period for the two systems can be calculated as: (7) and Figure 9: Simplified analytical models of diaphragm.(8) where ( 9) and ( 10) The stiffness contribution of the URM end walls is: (11) where is the moment of inertia of the cracked wall section (corresponding to the secant stiffness at 100 mm wall displacement) and can be approximated as or 0.01 times the moment of inertia of the uncracked section (Derakhshan et al. 2013).The subscripts l and u refer to the lower and upper walls respectively, recognising that both wall height and wall thickness may vary between stories, such that appropriate values of and H should be adopted to account for such variations.For a top-storey wall, is taken as zero.
For mid-height diaphragms the ratio of the periods for the two models shown in Figure 9 is obtained as: (12) and substituting the cracked moment of inertia and performing further simplification one can obtain: (13) where and are, respectively, the effective wall thickness in the lower and upper storeys and is the ratio of the masonry Modulus of Elasticity and diaphragm shear stiffness at 100 mm maximum diaphragm displacement.i.e. (14) and has 1/m units.After further simplification can be calculated as: (15) and the stiffness improvement factor due to the stiffness of the orthogonal end walls loaded out-of-plane is therefore calculated as: (16) To demonstrate use of the above expression, consider equal to 3 GPa, and equal to 240 kN/m (i.e. ).For a ground storey height of 4.0 m and thickness of 0.350 m and a second storey wall height of 3.6 m and thickness of 0.230 m and a diaphragm length of 10.4 m and depth of 6.5 m the value of is: (17) where W trib is the total tributary weight acting on the diaphragm, being the sum of the weight of the tributary walls (ie the product of the tributary height, thickness and density of the two URM walls tributary to the diaphragm accounting for wall penetrations) and diaphragm self-weight plus live load (ψ E × Q i as per NZS 1170.5 section 4.2).

IMPROVED DIAPHRAGMS
Many techniques may be used to improve the performance of flexible timber diaphragms, associated with providing added stiffness and/or strength.The details below specifically pertain to the interpretation of experimental data for improvement techniques that have been physically tested.Improved diaphragms typically exhibit distinct ductility capacity and the results have been interpreted as equivalent elastic-perfectly plastic response, with the reported stiffness being based upon the linear branch of the response curve.structural grade plywood laid over the existing floorboards with 75 mm × 24 gauge sheet metal straps fastened to the plywood edges with ECKO SF-9215 staples at 100 mm centres, as shown in Figure 11.The staple wire had a rectangular cross-section of 1.24 mm × 1.00 mm and a leg length of 15 mm.Field nailing (approximately 300 mm centres) was applied to the plywood sheets at the locations of the joists to mitigate buckling of the panels during large diaphragm displacements, while nailing was provided at 100 mm centres around all diaphragm edges to effectively transfer shear forces.All nails were 3.15 mm (diameter) × 75 mm roundhead power driven nails.Particularly pertinent to later discussion is the orientation of the plywood panels (placed on top of the straight sheathing) and the use of staples at 100 mm centres to fix the sheets to the underlying joists.
The resultant envelopes of the force-displacement response for testing in the directions both parallel and perpendicular to joist orientation are shown in Figure 12, where an averaged bilinearised response has been overlain for each of the two separate records.
The calculation of diaphragm shear stiffness for each of the 4 retrofitted diaphragms is presented in Table 5, noting that Diaphragm 2b-PARA was tested with a 3.2×1.8m 2 diaphragm penetration.The symbol Δ y refers to the bi-linearised yield displacement as plotted in Figure 12.
Substituting the experimental results reported in Table 5 into Equation 4(accounting for the actual Δ y given in Table 5 rather than Δ= 0.1 m as written in Equation 4), it may be established that suitable stiffness values for design purposes are: for testing perpendicular to joist G d = 1,250 kN/m; for testing parallel to joist G d = 2,400 kN/m.Also, it follows that the ratio of stiffness parallel/perpendicular = 2400/1250 = 1.92.
The general form of the strength equation for improved diaphragms deforming as shear beams is given by: (19) and hence for the interpretation of experimental results for tested diaphragms (adopting ϕ=1) the unit shear resistance is: Field nailing @ approx 300 mm (fixed to the joists) Nails @ 100 mm all around diaphragm edges (fixed to the joists/blocking elements) 75 mm x 24 ga sheet metal straps fastened across all plywood edges with staples @ 100 mm crs 40 mm x 6 mm mild steel flat fastened to the blocking with 75 mm x 10 ga screws @ 100 mm crs  (20) The calculation of strength values for the diaphragms reported in Figure 12 are presented in Table 6, based upon application of Equation 20.From this data it is concluded that R d = 9 kN/m is appropriate for loading perpendicular to joists and that R d = 15 kN/m is appropriate for loading parallel to joists.Also, it follows that the ratio of stiffness parallel/perpendicular = 15/9 = 1.67.The symbol Δ u is used to describe the maximum displacement achieved before the onset of strength loss due to plastic deformations.
A final feature of the data reported in Figure 12 is that the retrofitted diaphragms had substantial ductility capacity.Whilst larger ductility values were measured for loading oriented perpendicular to joists, it is suggested that the lowest measured value be applied for both loading orientations.As shown in Table 6 it is concluded that the 15 mm plywood overlay had a reliable displacement ductility of µ = 5.A conservative value of µ = 4 has been adopted.
The improvement in diaphragm response associated with the 15 mm plywood overlay is shown in Figure 13, where it can be clearly seen that the retrofit solution resulted in a substantial increase in both shear stiffness and shear strength.

Ceiling sheathing and tin ceiling lining
One scenario that is routinely encountered in vintage flexible timber diaphragms in New Zealand is the presence of a patterned tin ceiling lining covering a sheathed ceiling diaphragm (see Figure 14).This diaphragm detailing was investigated by testing Diaphragm A (refer to Table 1) both with and without the combination of the ceiling lining and tin ceiling overlay.
When comparing the data reported in Figure 15(b) it can be determined that the combination of ceiling sheathing and tin ceiling lining significantly influenced both the strength and stiffness characteristics of the diaphragm and therefore this

mm plywood overlay
Giongo et al. (2013) instituted a 9 mm plywood overlay retrofit when testing the diaphragm in the building located in Whanganui, with specific detailing as shown in Figure 16.Attention is drawn to the symmetrical orientation of the plywood panels, resulting in isotropic rather than orthotropic diaphragm strength characteristics, and the use of gage 6 screws (30 mm long) to fix panels to the underlying flooring only, rather than attempting to fix the overlay through the flooring and into the joists.This fixing strategy eliminated the need to cut the diaphragm panels to match the joist spacing, other than on the diaphragm perimeter.Diaphragm to wall shear transfer was provided by the gage 8 screws (60 mm long) installed on the diaphragm perimeter.
When comparing the data in Table 8 with the comparable data in Table 5 for testing oriented perpendicular to joist, it would initially seem unusual that the 9 mm ply overlay was stiffer than the 15 mm ply overlay.The explanation comes from recognising that the orientation of the plywood panels was more advantageous for the configuration shown in Figure 16 and that the use of screws rather than nails resulted in stiffer connections.Conversely, as the configuration shown in Figure 16 is symmetrical the same G d value is adopted for both loading directions.Hence when comparing the stiffness ratios for the direction parallel to joist of 2,400 kN/m and    1,460 kN/m for the 15 mm and 9 mm plywood overlays respectively, it is found that 2400/1460 = 1.64 whereas 15/9 = 1.67, suggesting that for loading oriented parallel to joists the plywood thickness might be directly responsible for dictating the overall diaphragm shear stiffness.
The shear strength of the 9 mm plywood overlay is defined by R d = 96/2/2.8= 17.1 kN/m.Because the diaphragm configuration is symmetrical the strength is also assumed to be symmetrical in both directions and a conservative strength of 16.0 kN/m is assumed.Because the testing was terminated before any signs of distress were evident, it is conservatively assumed that the design has a ductility capacity of µ = 3.

Summary of shear stiffness and strength values for diaphragm improvements
The data presented above pertaining to experimentally tested diaphragm improvement techniques is summarised in Table 9.

CONCLUSIONS
A systematic approach to the detailed seismic assessment of vintage New Zealand flexible timber diaphragms is presented, incorporating all available field test data and integrating with the existing ASCE 41-13 technique when field test data is unavailable.Three tested improvement techniques were presented with a view to these results forming the basis for standard detailing.The intent is that the procedure may be further refined and developed as additional test data becomes available.
The main differences introduced in the proposed approach, when compared with procedures presented elsewhere, are the selection of the critical diaphragm displacement being dependent on out-of-plane URM end wall deformations, and the use of this criterion to establish the diaphragm secant shear stiffness.A procedure to account for possible degradation of the diaphragms is also presented.
The presented procedure is based upon developing a suitably accurate elastic procedure whilst recognising that diaphragm response can be highly nonlinear.For diaphragm configurations outside those considered here, the use of nonlinear modelling may be necessary to establish probable diaphragm deformation response (Wilson 2012).Similarly, designers should consider undertaking a sensitivity analysis to establish the extent of variation in response that is derived from assumptions regarding diaphragm condition.
An application of the proposed approach to a case study is reported in Appendix A.
joint rotation due to nail slip and nail spacing.

Figure 1 :
Figure 1: Schematics showing aspects of diaphragm response.
100% out-of-phaseAssumed deformation modes for diaphragm acceptance criteria(a) Wall-diaphragm interaction for a multi-storey URM building.Mid-span diaphragm displacement limit for URM building on a rigid foundation.(c) Mid-span diaphragm displacement limit for URMbuilding on a flexible foundation.
The most widely respected recommendations for diaphragm stiffness are the data published in ASCE 41-13, which specifies a shear stiffness for straight sheathed diaphragms of 350 kN/m.Such data derive from the studies of ABK (1982) on full-scale (6.1 m × 18.3 m) timber diaphragms.The specimens consisted of 51 mm × 305 mm (2" × 12") Douglas fir joists positioned at 610 mm spacing and having a wood framing system constructed with 102 mm × 305 mm (4" × 12") elements.The flooring was comprised of 25 mm × 152 mm (1" × 6") Douglas fir lumber sheathing, fixed to the joists using two 8d wire nails at each floorboard-to-joist intersection.Three nails were position at the end of each board.51 mm × 102 mm (2" × 4") Douglas fir blocking elements were also present.However, three concerns with direct adoption of this data are that: (1) it is not clear that the data has specific relevance to vintage diaphragms having the characteristics commonly encountered in New Zealand URM buildings, including wall-diaphragm boundary effects; (2) it is not clear what secant stiffness was used to develop these stiffness values but it is assumed that the data relates to bilinearisation of the force-displacement envelope with adoption of the initial stiffness, therefore potentially having a substantially greater magnitude than the secant stiffness to a displacement of 100 mm, as indicated in Figure4(b); (3) no guidance is provided on how to account for the orthotropic nature of the diaphragm and the influence of any possible diaphragm degradation due to insect infestation or moisture damage as is likely encountered in vintage timber diaphragms.DIAPHRAGMS AS SHEAR BEAMSWilson et al. (2013a) have confirmed that the deformation (a) Geometry of a two wythe (250 mm thick) solid URM wall.(b) Force-displacement response.
(a) Evidence of mass-loss due to insect infestation.(b) Evidence of decay due to moisture.

Figure 5 :
Figure 5: Deteriorated condition of vintage Whanganui timber diaphragm tested in-situ.

Wilson
et al. (2014a) have published the results from testing of newly constructed flexible timber diaphragms, tested both perpendicular and parallel to the orientation of joists.The equivalent shear stiffness data is presented in Table

Figure 6 :
Figure 6: Test set-up for Whanganui diaphragm test; a) wall connection bracket; b) snap shackle for instantaneous loadrelease; c) new shear anchors; d) hydraulic actuator; e) central loading plate.

Figure 8 :
Figure 8: Displacement incompatibility between diaphragm and URM end walls.

Figure 10 :
Figure 10: Assumed deformation modes for one-way vertically spanning walls.

15 mm plywood overlay
Wilson et al. (2014a) have reported the performance of newly constructed Pinus radiata flexible diaphragms retrofitted with 15 mm plywood overlays, using the construction details reproduced in Figure 11.Each diaphragm measured 10.4 m × 5.5 m and was comprised of 135 mm × 18 mm straight-edge timber floorboards fastened perpendicular to 45 mm × 290 mm joists spaced at 400 mm centres and oriented along the 5.5 m dimension.The floorboards were fixed to the joists by 3.15 mm × 75 mm common bright roundhead nails driven at a spacing of approximately 95 mm.All retrofits consisted of 2,400 mm × 1,200 mm × 15 mm AS/NZS 2269:2004

Figure 13 :
Figure 13: Comparison of force-displacement response for diaphragms tested as-built and after being retrofitted with a 15 mm plywood overlay [Wilson et al. 2014a].

Figure 14 :
Figure 14: Patterned 0.3 mm tin ceiling lining underlying a straight sheathed flexible timber ceiling diaphragm.

( a )
With ceiling sheathing and tin ceiling lining.(b) Comparison with and without diaphragm ceilingimprovements.

Figure 15 :
Figure 15: Force-displacement response of diaphragm A with and without added ceiling sheathing and patterned tin ceiling lining.

Table 2 : Secant stiffness values for laboratory testing of new diaphragms loaded perpendicular and parallel to joists
Kent et al. (2005)hragm condition on diaphragm stiffness was studied byWilson (2012)using Finite Element modelling to undertake parametric analysis.Kent et al. (2005)and Sawata et al.(2008)have undertaken physical testing to show that the effects of moisture saturation and insect infestation can substantially influence the stiffness of timber connections, with the extent of the influence extending to reductions of 50% for timber in poor condition.Wilson's numerical model suggested that stiffness values for diaphragms in

Table 3 : Shear stiffness values † for straight sheathed vintage flexible timber floor diaphragms
Values may be amplified by 20% when the diaphragm has been renailed using modern nails and nail gun † † Values should be interpolated when there is mixed continuity of joists or to account for continuous sheathing at joist lap

Table 7 : Secant stiffness values for in-situ testing of vintage diaphragm loaded perpendicular to joists
Diaphragm A with sheathed ceiling lining and patterned tin ceiling overlay (see Figure15(a)).Refer to Table1for the equivalent response without the sheathed ceiling lining and tin ceiling overlay