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An experimental investigation of foundation anchorage details and base shear capacity for log buildings.


Log structures are constructed using round logs (or manufactured timbers) that are stacked vertically and have interlocking corners. Thru-rods or lag screws are used to give load path continuity. This project was conducted to evaluate the lateral force resisting pathways in log structure foundation details with a focus on seismic shear resistance provided by friction, anchor bolts, and thru-rods. The experiments used test specimens that represented two common construction details for sill log-foundation anchorage. One detail had the sill log sitting on the floor diaphragm, and the other detail had the sill log in direct contact with the sill plate. Anchor bolts and thru-rods were included in both details. The measured coefficient of friction between the sill log and the floor diaphragm was approximately 0.4. The sill log-foundation details were tested to determine force-displacement behavior and yield modes. The sill log-foundation details were loaded vertically to simulate wall and roof dead load and then forced to move horizontally with a static and then with a fully reversed, cyclic, quasi-static test protocol. The force-displacement curves showed an initial stiffness, slip, and post-slip stiffness. The hysteretic diagrams were open and boxy, which demonstrated that energy was dissipated by friction between the sill log and floor diaphragm or sill plate. Initial and ultimate yield modes were different depending on foundation details. The base shear design requirement for a representative wall was calculated following the Uniform Building Code. Both connection details had measured capacities greater than required for Uniform Building Code seismic zone 4.


In recent years, log structures have been marketed as an alternative to conventional light-frame wood structures for residential and commercial occupancies. While early log buildings were small, simple, box-like structures, newer log structures are often large, expensive structures with large wall openings for windows that frequently result in high aspect ratio wall segments that may not resist lateral forces like the earlier versions.

Log shear walls (Fig. 1), which are also bearing walls, resist lateral loading differently than light-frame shear walls. Lateral loads are transferred from the top to the bottom of light-frame shear walls through the sheathing, framing, and nailed connections, where sheathing type and thickness and nail type and spacing are important design considerations (Breyer et al. 1999). Lateral loads are transferred from the top to the bottom of log shear walls through log-log friction, thru-rods (or other hardware), and inter-wall corner connections (Haney 2000). Hence, log-log coefficient of friction, thru-rod tension, thru-rod hole size, and connection type are important. Light-frame and log shear walls also dissipate energy differently. Nail fatigue, nail withdrawal, and nail pull-through are important energy dissipation mechanisms in light-frame shear walls. Log-log slip with friction is expected to be an important energy dissipater in log shear walls.



Foundation anchorage is an important component in the seismic performance of light-frame wood and log buildings. Mahaney and Kehoe (2001) provided an extensive literature review on the subject of foundation anchorage in light-frame wood buildings. Log structure foundation anchorage logically differs from that used for light-frame wood structures in that the building mass is greater in log structures and connection geometry is different as a result of the relatively large log diameter.

The total horizontal seismic force acting at the base of a log structure is the base shear (V). The Uniform Building Code (UBC) (ICBO 1997) provides guidelines for calculating the base shear. Factors involved in this calculation include seismic zone, site characteristics, occupancy, configuration, structural system, height, and most importantly weight of the structure.

Natural hazard-load resistance of light-frame wood shear walls has been the subject of concerted research emphasis. Although many log structures are in high-wind and seismic zones, virtually no information exists on lateral load path in log structures and performance under lateral loading. Results from testing of log structure components in this project will contribute to the body of knowledge pertaining to the structural performance of log buildings.


The objectives of this study were to present experiments and analysis that address base shear performance, capacity, and design for log structures. Specific objectives were to:

* Determine mean and variability for sill log-floor diaphragm sliding friction.

* Evaluate the static capacity of the sill log-foundation anchorage system and identify the failure mechanisms.

* Assess two typical sill log-foundation connections under reversed-cyclic loading to identify the failure mechanisms and establish energy dissipation characteristics.

* Evaluate the sill log-foundation connections relative to UBC base shear requirements.

Technical background

Log structures are placed on foundations that are similar in design to those used for light-frame wood and masonry construction. The sill log is connected to the foundation using 16-mm-diameter anchor bolts. A standard anchor bolt spacing is 1830 mm, with oversized holes to facilitate placement during construction. Tension in the anchor bolts is released as the sill log shrinks as a result of drying (Scott et al. 2002), and in addition, the nuts on the sill log anchor bolts cannot be retightened due to inaccessibility. The lateral load resistance resulting from sill log friction is ignored during design just as friction is ignored between the sill plate and the floor in light-frame structures.

Two foundation anchorage details are common to the log structures industry. In detail 1, the wall sill log is positioned on the floor diaphragm (Fig. 2a) with a complete assembly having foundation anchor bolts and wall thru-rod hardware. For the purposes of this investigation, detail 1 was tested in two ways: 1) as incomplete (detail li), with the foundation anchor bolt present but the thru-rods absent; and 2) as the complete connection (detail 1c), with both foundation anchor bolts and thru-rod hardware. The thru-rods pass through the sill log from above and terminate with a washer and nut on the bottom surface of the floor sheathing.

In detail 2 (Fig. 2b), the sill log sits directly on the sill plate. Washers and nuts secure the sill plate to the foundation and coupler nuts are used to extend the anchor bolts into the sill log or connect them to wall thru-rods.

Friction is an important factor in both anchorage details. Friction is the force tangent to the interface of two bodies when one moves against the other. The friction force (F) is proportional to the normal force (N) pressing the bodies together:


F = [mu] X N [1]

where [mu] is the coefficient of friction (Jastrzebski 1976).

The Wood Handbook (FPL 1999) states that the wood-wood coefficient of friction is dependent on moisture content (MC) and surface roughness. McKenzie and Karpovich (1968) report coefficients of friction as a function of sliding speed and MC. The coefficient of friction for wet wood starts at a static value of 0.84 and declines to a value 0.36 at a sliding speed of 55 mm/sec. The coefficient of friction for 12 percent MC wood starts at a static value of 0.60 and levels off to a value of 0.46 at a sliding speed of 55 mm/sec.

Bejo et al. (2000) report coefficients of friction for wood-based structural composites as a function of grain orientation and a contact pressure ranging from 0.5 kPa to 60 kPa. Static coefficient of friction values for laminated strand lumber range from 0.84 to 0.48, while kinetic coefficient of friction values range from 0.52 to 0.22. For laminated veneer lumber, static coefficient of friction values range from 0.70 to 0.33, while kinetic coefficient of friction values range from 0.39 to 0.20. Coefficients of variation range from 0.04 to 0.28.

The coefficients of friction by Bejo et al. (2000) were obtained using small-scale specimens with an inclined plane method, or a milling machine and lathe system. The MCs, normal pressures, roughness, and contact areas are not like the conditions for a sill log in a log structure. Friction tests performed on specimens simulating sill log conditions are needed.

Test protocols for quasi-static cyclic testing were reviewed by Gatto and Uang (2001). Common test protocols include sequential phased displacement (SPD) (Porter 1987), the International Organization for Standardization (ISO) protocol (ISO 2001), and the method by the Consortium of Universities for Research in Earthquake Engineering (CUREE) (Krawinkler et al. 2000). These test protocols were devised to represent the seismic demands that are imposed on structures during an earthquake.

The CUREE protocol was developed for shear wall assemblies of light-frame wood structures (Krawinkler et al. 2000) and has been used for wall assemblies (Stahle 2001) and mechanical connections (Fonseca et al. 2002). It is based on the reference displacement (D), which is the maximum deformation the specimen is expected to sustain according to a prescribed acceptance criterion, which is developed from static testing of duplicate specimens. The protocol consists of a series of fully reversed cycles starting with a series of low amplitude initiation cycles followed by sets of cycles of increasing amplitude as described by Krawinkler et al. (2000). Comparisons of test results showed that the test protocols produced load capacities within [+ or -] 10 percent (Gatto and Uang 2001). Given this result, the CUREE protocol was chosen for the sill log-foundation connection tests to simulate performance under seismic loading.

Methods and materials

Apparatus and measurements

The loading system was a test frame, with two hydraulic actuators, one each for vertical and horizontal loading. Linear variable differential transformers (LVDTs) recorded displacements, and load cells mounted on the actuators were used to sense forces. Load and displacement data were recorded using Lab View[R] 6.1. Four specimen configurations representative of materials and connection hardware were developed for the friction, monotonic, and quasi-static tests of detail 1 and detail 2.

Friction test specimens

The detail 1 friction test (Fig. 3) was devised to measure the coefficient of friction between a sill log and the plywood floor sheathing. No foundation anchorage or wall thru-rods were present in these test specimens. The sill log was represented by a piece of Douglas-fir lumber that was cut using a bandsaw, giving a surface texture similar to the bottom surface of a sill log. The plywood floor sheathing was attached to a laminated veneer lumber (LVL) billet using 12 64-mm-long galvanized deck screws. The sill log surface was attached to the bottom flange of a steel I-beam using 16 45-mm-long galvanized deck screws. A frictionless interface at the bottom of the assembly was created by lubricating face-to-face polyethylene sheets (nominally 6 by 150 by 1220 mm). Thus, the LVL billet could be moved by the horizontal actuator, and the measured friction resistance was only that generated at the wood-plywood interface. The vertical actuator applied a constant point load of 13.3 kN, representing the dead load of the structure, to the top center of the I-beam, which was a load-distributing element as well as a fixture for the sill log surface.

Static test specimen

The detail li static test was developed to evaluate the lateral force resistance of the detail li foundation anchorage. The specimen (Fig. 4) consisted of a sill log, the floor diaphragm, and the foundation anchorage hardware. The sill log was represented by a section of Douglas-fir LVL. The floor diaphragm was intended to be typical construction and consisted of plywood sheathing, I-joists, oriented strand lumber rim board, Douglas-fir sill plates, and LVL blocking. The plywood sheathing was attached to the I-joists and rim board using construction adhesive and 2.9-mm-diameter, 60-mm-long nails. The nails were spaced 152 mm along the panel edges and 305 mm in the field of the panel along the I-joists. The rim board was attached to the I-joists and sill plate using 2.9-mm-diameter, 60-mm-long nails. One nail was driven through the rim board into each I-beam flange. Three nails were driven at each rim board corner. The rim board was toenailed to the sill plates using the same size nails at 152-mm spacing. The I-joist flanges were nailed to the sill plates using two 3.3-mm-diameter, 76-mm-long nails. The LVL blocks were attached to the plywood and the sill plate using four 2.9-mm-diameter, 60-mm-long nails. The foundation anchorage was two 16-mm threaded rods fixed to the foundation test fixture such that they would behave as though embedded in concrete. These rods passed through the sill plate and were coupled to extension rods so that the foundation anchorage extended through the plywood floor and LVL billet terminating in a counter-sunk hole in the sill log. Nuts and malleable washers were installed on the foundation anchor bolts with a 5-mm gap between the nuts and washers. The gap was intended to represent shrinkage of the sill log due to drying in service (Scott et al. 2002).



Quasi-static test specimens

One test specimen for each foundation detail was prepared for quasi-static testing. The detail li quasi-static test specimen was identical to the detail li static test specimen.

The detail 1c test configuration was the same as for the detail li specimen except that two 16-mm thru-rods were included (Fig. 5). Two 22-mm holes were drilled through the sill log and floor sheathing to accommodate the thru-rods. The two thru-rods ran through the sill log and floor sheathing and were secured wrench-snug with nuts and washers on the top of the sill log and the bottom of the floor sheathing. The locations of the thru-rods and anchor bolts are shown in Figure 5.

The detail 2 quasi-static test was designed to evaluate the lateral force resistance of the detail 2 foundation anchorage. The specimen (Fig. 6) consisted of a sill log, sill plate, and connection hardware. The sill log was represented by a Douglas-fir LVL billet, the sill plate was a Douglas-fir board. The assembled test specimen is shown in Figure 7. Two anchor bolts were fixed to the foundation test fixture and extended 75 mm through the sill plate. The sill plate was secured to the foundation test fixture using washers and nuts tightened on the anchor bolts. The anchor bolts were coupled to an additional threaded rod that extended through the sill log. The sill log was secured by placing a malleable washer and nut on each anchor bolt. A 5-mm gap was left between the nut and malleable washer to simulate the size change due to log shrinkage in service.

Test protocols

Friction tests. -- A constant 13.3-kN (10.9-N/mm of specimen length) normal force, representing the dead load of the structure, was applied to the I-beam throughout the test. For the purposes of the test, the sill log surface was stationary and the LVL billet and attached plywood were horizontally displaced according to the function D = 51 sin(0.5[pi]t), where D = displacement (mm) and t = time (sec). Data were taken at a rate of 20 Hz, and each test lasted 16 seconds. This test was replicated with 14 plywood/sill log surface pairs. The maximum displacement and velocity were [+ or -]51 mm and 76 mm/sec., respectively. The MCs of the sill log surface and plywood were measured by ovendrying coupon specimens after each test.

Static test. -- A constant 26.6-kN (10.9-N/mm of specimen length) normal force, representing the dead load of the structure, was applied to the sill log through the I-beam during the test. The LVL billet was laterally displaced in a monotonic direction at a rate of 0.125 mm/sec. The test was run until a horizontal force of 44.5 kN was reached; this was the horizontal actuator load limit. Data were taken at a rate of 10 Hz.




Quasi-static tests. -- A constant normal force of 26.6-kN (10.9-N/mm of specimen length) was applied throughout the test. The sill log was displaced according to the CUREE protocol (Scott 2003) at a frequency of 0.1 Hz. Data were taken at a rate of 100 Hz. After the monotonic and quasi-static tests, the MC of the sill log was measured.

Results and discussion

Friction test

A force-displacement plot typical of the 14 specimens is shown in Figure 8. The coefficient of friction is based on the forces in cycle 1 and cycle 4 at zero displacement in the positive stroke. Each cycle of the four-cycle test showed a drop in the coefficient of friction from the previous cycle. The average coefficient of friction for the first cycle was 0.40 with a coefficient of variation of 0.20 (95 percent CI: 0.25 to 0.56). The average coefficient of friction for the fourth cycle was 0.37 with a coefficient of variation of 0.22 (95 percent CI: 0.22 to 0.53). The data are similar to Bejo et al. (2000) and Kellog (1981). A paired t-test was used to test the hypothesis that no difference in the coefficient of friction existed between the first and fourth cycles. The hypothesis was rejected (p-value = 0.0002), meaning that the 8 percent change in the mean coefficient of friction from the first to the fourth cycle was significant.

Static test

The monotonic static test reached a force of 44.5 kN, which was the load capacity of the horizontal actuator, where the displacement was 69 mm. Even at this large displacement, the force-displacement curve was still ascending. The CUREE reference displacement (D) for light-frame shear walls is supposed to be 0.6 of the ultimate displacement, which is the displacement at 0.8 times the ultimate load on the descending force-displacement slope (Krawinkler et al. 2000). However, a reference displacement for single connections and assemblies where the load-displacement response has an extended plastic characteristic has not been suggested. This problem was encountered by Fonseca et al. (2002) with staple connections; the peak load for staples was not reached even for large displacements. Thus, an unrealistic reference displacement resulted, and as a result, they decided to use a reference displacement for staples that was similar to the reference displacement for nails. In light-frame shear walls, Langlois et al. (2004) found that the reference displacement did not drastically influence the overall wall response to cyclic loading. Therefore, an arbitrary reference displacement of 19 mm was chosen as a reasonable value for use in all quasi-static tests of the sill log anchorage details.


Quasi-static tests

Figure 9 shows detail li quasi-static test data with its backbone curve and the detail li static test curve superimposed. The monotonic static force-displacement curve and quasi-static test backbone curve show similar results. At a force of 11.1 kN, the sill log-sheathing friction was overcome and the log started to slide, showing classic stick-slip behavior of friction systems. This corresponded to a coefficient of friction of 0.4, which was consistent with the friction tests. The log continued sliding at a constant force until the anchor bolt contacted the perimeter of the hole in the sill log and then the sheathing. No permanent damage was visible at this point. Forces greater than 11.1 kN were resisted by anchor bolt bending and bearing damage in the sill log and plywood sheathing. The monotonic test forces were about 15 percent lower than the quasi-static test backbone curve, which was attributed to the variability in the assemblies used for each test.

The anchor bolts transferred force from the log to the sheathing. This resulted in sheathing embedment failure. As the force increased, the bolt was embedded into the plywood at a rate of 0.34 mm/kN. At the end of the test, the sill log did not move back to its original position when it was unloaded, which indicated permanent damage was present in the anchor bolt or the floor sheathing, or both.

Forces were transferred to the sheathing through friction and anchor bolt bending during the test. These forces were transferred to the rim board via glued and nailed connections between the sheathing and the rim board as well as sheathing and I-joists. The sheathing-rim board connection showed no evidence of distress during or after the tests.

The load path went from the rim board to the sill plate through toenail connections between the rim board and the sill plate. The toenail connections deformed as the rim board moved as much as 6 mm relative to the sill plate in the later stages of the test. When unloaded, the rim board did not move back to its original position, and post-test inspection showed evidence of nail bending and withdrawal.

The sill plate resisted the horizontal force by friction and the anchor bolts. The sill plate moved no more than 2 mm during the test. No damage was observed in the sill plate-anchor bolt connection.

The detail li monotonic static and quasi-static test showed that the overall base shear capacity for this system would be limited by plywood embedment capacity, anchor bolt bending, and rim board-sill plate toenail connection deformation. For this system, plywood embedment was the first apparent yield mode.

The detail 1c hysteresis diagram was a set of unpinched curves much like those of detail li. Again, the test was terminated when the horizontal actuator reached load capacity at 44 kN. At this force, the displacement was 23 mm. The backbone curve was similar to the detail li test up to 18 mm of displacement. At this point, detail li was losing stiffness, but detail 1c was not softening. The presence of the two thru-rods reduced the rod bearing force by half at each plywood hole, and as a result the plywood embedment capacity was not reached at the same total load. Damage to the rim board-sill plate toenail connection was observed in the detail li and detail 1c specimens, and even though system capacity was not reached, it appeared that the detail 1c assembly yield mode would have been at the rim board-sill plate toenail connection.

The detail 2 hysteresis diagram was also a set of open unpinched curves. At the actuator load capacity, the displacement was 14 mm. The sill log axial force was simultaneously transferred to the sill plate and anchor bolts and into the foundation fixture. Initial log movement occurred at a force of 13.5 N. This force was higher than that for initial log movement in the detail 1 tests. This was because the anchor bolts and friction were initially resisting the lateral force; the effects of the oversized holes were reduced by the nuts and couplers that were located inside of the sill log and fit tightly in the hole.

The anchor bolts resisted the lateral force through bending. Post-test inspection showed that two plastic hinges formed, one at the sill plate-foundation interface and the other at the sill log-sill plate interface.

The sill plate moved as much as 5 mm horizontally in each direction during the test. Post-test inspection showed crushing around both sill plate holes. A split was also observed from one hole to the end of the sill plate. This follows the results of tests reported by Mahaney and Kehoe (2001) in light-frame wood construction.

This test showed that the overall base shear capacity for the detail 2 system was limited by anchor bolt bending, and that the anchor bolts in the detail 2 test sustained more damage than in the detail 1 tests. By placing the sill log on the foundation, the sill plate also sustained damage but did not limit the capacity of the system. Placing the sill log on the foundation as in detail 2 excludes the plywood sheathing, rim board, and toenail connection between the rim board and sill plate as potential capacity and performance limitations.


Energy dissipation

The detail li, detail 1c, and detail 2 force-displacement responses were box-like shapes, not pinched hysteretic curves that are typical of dowel connections. This was due to the sill log-floor sheathing and sill log-sill plate friction, which significantly contributed to the energy dissipation of these connections. Each hysteretic loop corresponded to a quasi-static test cycle. The area of a loop plus that of each loop before it was the cumulative energy dissipation up to that point. At 40 cycles, the cumulative energy dissipation for detail li and detail 1c was the same and equal to 10.0 MNmm, while energy dissipation for detail 2 was 11.7 MN mm (Fig. 10).

Base shear requirements and capacities

For seismic design, the UBC (ICBO 1997) requires design for an earthquake load (E) given by:

E = [rho][E.sub.h] + [E.sub.v] [2]

The redundancy factor p has an upper bound of 1.5. [E.sub.h] is the load due to horizontal ground motion (base shear). [E.sub.v] is the load effect of vertical ground motion and is zero for allowable stress design.

The UBC base shear formula is:

V = [[[C.sub.v]I]/RT]W [3]

The UBC also defines an upper bound for base shear as:

V = [[2.5[C.sub.a]I]/R]W [4]

In Equations [3] and [4], [C.sub.v] and [C.sub.a] are seismic (response spectrum) coefficients. Assuming a stiff soil profile and a site in seismic zone 4 where the closest distance to a known seismic source is greater than 15 km, [C.sub.v] = 0.64 and [C.sub.a] = 0.44 (UBC, Tables 16-R and 16-Q). Most log structures are homes or standard occupancy structures where the seismic importance factor is I = 1.00 (UBC Table 16-K). In Equation [3], T = 0.111 seconds is the fundamental period of vibration calculated following UBC equation 30-8 for a log structure with an assumed height of 3 m. The response (modification) factor (R) depends on the structural system; a specific value for R has not been assigned to log structures. However, R can range from 2.8 (light steel frame) to 4.5 (masonry shear walls). The most conservative estimate for V is obtained by assuming R = 2.8. A less conservative, but more realistic value for V is obtained by using the R-value for masonry walls. The base shear is V = 2.06 W from Equation [3], and the upper bound is V = 0.393 W from Equation [4]. Hence, the upper bound for V controls for this log structure.

The total seismic dead load is W, the sum of the wall and roof weights. A Douglas-fir log that is 355 mm in diameter and 15 percent MC weighs 535 N/m. The dead load for a representative wall that is nine logs high and 2.44 m long, plus the tributary roof dead load, is 21.7 kN. Using this dead load for W in Equation [4], V = 8.55 kN. The upper bound V is divided by 1.4 to convert from strength level to allowable stress design, and with [rho] = 1.5 in Equation [2], E = 9.16 kN for a representative wall length.

Detail 1i, detail 1c, and detail 2 tests, each a representative wall length of 2.44 m, reached a lateral force of at least 44 kN. Dividing the test force by the design load gives a ratio of capacity to design of at least 4.8 for each detail. Thus, each detail appears to be consistent with the factor of safety for mechanical connections.


Friction tests with sill log surfaces against plywood floor sheathing revealed that a coefficient of friction of 0.40 was changed to 0.37 after four fully reversed cycles representing 408 mm of total travel.

Three sill log-foundation test configurations (detail li, detail 1c, and detail 2) showed factors of safety consistent with those for mechanical connections. All three tests were terminated at 44 kN; the associated displacements were 35 mm for detail li, 23 mm for detail 1c, and 14 mm for detail 2. The detail 1c test showed that the addition of thru-rods caused the yield mechanism to occur at the rim board-sill plate toenail connection. The detail 2 test showed that the sill log resting on the sill plate results in less displacement than detail 1 but can produce sill plate splitting.

Box-like hysteretic plots in all three connections showed sill log friction was a source of energy dissipation. A cumulative energy dissipation, at 40 cycles, of 10 kN mm was observed for both detail 1 tests, while energy dissipation for detail 2 was 11.7 MN mm.

The base shear design requirement for a representative log shear wall was calculated following the UBC. The connection details had capacities greater than required for the UBC seismic zone 4.

Literature cited

Bejo, L, E.M. Lang, and T. Fodor. 2000. Friction coefficients of wood-based structural composites. Forest Prod. J. 50(3):39-43.

Breyer, D.E., K.J. Fridley, and K.E. Cobeen. 1999. Design of Wood Structures. McGraw-Hill, Inc., New York.

Fonseca, F., S. Rose, and S. Campbell. 2002. Nail, wood screw, and staple fastener connections. Consortium of Universities for Research in Earthquake Engineering, Univ. of California, Richmond, CA. 161 pp.

Gatto, K. and C.-M. Uang. 2001. Loading protocol and loading rate effects literature review. Task 1.3.1. Woodframe Project Testing and Analysis Literature Reviews. The CUREE-Caltech Woodframe Project. Univ. of California, San Diego, CA. pp. 27-44.

Haney, T. 2000. How log buildings resist lateral loads. Log Building News 32:1-6.

International Conference of Building Officials (ICBO). 1997. Uniform Building Code. ICBO, Whittier, CA.

International Organization for Standardization (ISO). 2001. Timber structures--joints made with mechanical fasteners--quasi-static reversed cyclic test method. ISO/DIS 16670.3. ISO, Geneva, Switzerland.

Jastrzebski, Z.D. 1976. The Nature and Properties of Engineering Materials. John Wiley & Sons, Inc., New York. 633 pp.

Kellogg, R.M. 1981. Physical properties of wood. In: Wood: Its Structure and Properties. F.F. Wangaard, ed. Pennsylvania State Univ., Univ. Park, PA. pp. 187-224.

Krawinkler, H., F. Parisi, L. Ibarra, A. Ayoub, and R. Medina. 2000. Development of a testing protocol for wood frame structures. CUREE/CalTech Woodframe Project Rept. Stanford Univ., Stanford, CA. 85 pp.

Langlois, J.D., R. Gupta, and T.H. Miller. 2004. Effects of reference displacement and damage accumulation in wood shear walls. J. of Struct. Eng. 130(3):470-479.

Mahaney, J.A. and B.E. Kehoe. 2001. Anchorage of woodframe buildings. Task No. Woodframe Project Testing and Analysis Literature Reviews. The CUREE-Caltech Woodframe Project. Univ. of California, San Diego, CA. pp. 75-88.

McKenzie, W.M. and H. Karpovich. 1968. The frictional behavior of wood. Wood Sci. and Tech. 2:139-152.

Porter, M.L. 1987. Sequential phased displacement (SPD) procedure for TCCMAR testing. Third Meeting of the Joint Technical Coordinating Committee on Masonry Research, U.S.-Japan Coordinated Earthquake Research Program, Tomamu, Japan, pp. 1-15.

Scott, R.J. 2003. Lateral force resisting pathways in log structures. MS thesis, Oregon State Univ., Corvallis, OR. 173 pp.

________, R.J. Leichti, and T.H. Miller. 2002. Foundation anchorage in residential log structures. In: Biographies & Abstracts, Forest Products Society 56th Annual Meeting. Forest Prod. Soc., Madison, WI. p. 48.

Stahle, R. 2001. The influence of partial submersion on wood shear wall performance. Diploma thesis. Univ. of Karlsruhe, Karlsruhe, Germany. 124 pp.

USDA Forest Service, Forest Products Laboratory (FPL). 1999. Wood Handbook: Wood as an Engineering Material. Forest Prod. Soc., Madison, WI. 463 pp.

Randy J. Scott

Robert J. Leichti*

Thomas H. Miller

The authors are, respectively, Graduate Research Assistant, Oregon State Univ., Corvallis, OR 97331-5751; Associate Professor, Dept. of Wood Science and Engineering, Oregon State Univ.; and Associate Professor, Dept. of Civil, Construction, and Environmental Engineering, Oregon State Univ. This paper is based on the thesis written by the senior author. The assistance of Milo Clauson with testing is gratefully acknowledged. Funding for this research was provided by the USDA Center for Wood Utilization Research at Oregon State Univ. and the Forest Research Lab., Oregon State Univ. This paper is No. 3591 of the Forest Research Lab., Oregon State Univ. This paper was received for publication in August 2003. Article No. 9741.

*Forest Products Society Member.
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Author:Scott, Randy J.; Leichti, Robert J.; Miller, Thomas H.
Publication:Forest Products Journal
Article Type:Abstract
Geographic Code:1USA
Date:Apr 1, 2005
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