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• 9750 SW Nimbus Avenue oFFI #,_^ 'OP8 G R Beaverton,OR 9 7008-71 7 2 c5�7 Z2i9_eo�o p� 503-641-3478 f I 503-644-8034 food Ski UGik'b 'J 926 February 28, 2018 RECEIVED 5970-F GEOTECHNICAL RPT FEB 202019 Tigard-Tualatin School District CITY OF TIGARD 6960 SW Sandburg Street BUILDING DIVISION Tigard, OR 97223 Attention: Debbie Pearson/DAY CPM Services, LLC SUBJECT: Geotechnical Investigation and Site-Specific Seismic-Hazard Evaluation Tigard High School Tigard, Oregon At your request, GRI completed a geotechnical investigation and site-specific seismic-hazard evaluation for the planned improvements at Tigard High School in Tigard, Oregon. The Vicinity Map, Figure 1, shows the general location of the site. The purpose of the investigation was to evaluate subsurface conditions at the site and develop geotechnical recommendations for use in the design and construction of the proposed improvements. The investigation included a review of existing geotechnical information for the site and surrounding area, subsurface explorations, laboratory testing, and engineering analyses. As part of our investigation, GRI completed a site-specific seismic-hazard evaluation to satisfy the requirements of the 2012 International Building Code (IBC), which was adopted by the 2014 Oregon Structural Specialty Code (OSSC). This report describes the work accomplished and provides conclusions and recommendations for use in the design and construction of the proposed project. PROJECT DESCRIPTION We understand the portion of Tigard High School built in 1953 will be demolished and rebuilt under the 2016 Tualatin-Tigard School District Bond Program. Based on our review of conceptual plans, we understand the 1953 portion of the school to be demolished is primarily located within the western half of the building footprint, and the Auditorium/Cafeteria Building, Main Gym, Auto Shop, and the eastern half of the existing school will remain in their current configuration. The Site Plan, Figure 2, shows the approximate locations of the new school footprint and associated improvements with respect to the existing school and associated buildings. We understand the new school will consist of a classroom wing, administration offices, expanded commons and lunchroom, kitchen, auxiliary gym, boy's and girl's locker rooms, weight room, multi-purpose physical-education space, and courtyard. The new school will include a one- to two- story structure with partially embedded lower level that occupies the southeastern half of the footprint. In general, the new school will be located within the existing building footprint; however, the new building will extend up to 150 ft farther north than the existing building. Although we have not been provided structural loads for the new building, we anticipate the column and wall loads will be on the order of 100 to 200 kips and 3 to 4 kips/ft, respectively. We anticipate the finished floor elevation for the lower level will be established at or near the lowest existing site grade, which occurs along the southern edge of the existing school. Based on our review of building sections provided by Bassetti Architects, the project architect, we estimate an excavation of about 10 ft will be required to construct the partially embedded lower level. A shoring system may be necessary to support GEOTECHNICAL ■ PAVEMENT ■GEOLOGICAL■ ENVIRONMENTAL Since 1984 , - the excavation near portions of the existing school that will remain. We anticipate the finished floor elevation for the main level will generally be consistent with the school entrance, and cuts and fills to establish grades across the remainder of the site will be minimal. We understand a new bus drop-off area will be constructed north of the new school, adjacent to SW Durham Road, and a new parking lot will be constructed immediately west of the new building in an area currently occupied by tennis courts and an independent building. We anticipate the new bus loop and parking lot will be paved with asphalt concrete (AC) pavement. SITE DESCRIPTION General The project site is developed with the existing school and associated buildings, parking areas, a bus drop-off area, and athletic fields. The new school will be bordered by the new bus drop-off area and SW Durham Road on the north, the existing school on the east, the Auto Shop and athletic fields on the south, and the Main Gym and AC-paved parking areas on the west. Review of satellite imagery and our observations at the site indicate the ground surface gently slopes downward from north to south across the site. Geology Published geologic mapping indicates the site is mantled with Missoula flood deposits, locally referred to in the project area as the Willamette Silt Formation (Madin, 1990). In general, Willamette Silt is composed of beds and lenses of silt and sand. Stratification within this formation commonly consists of 4- to 6-in.-thick beds,although in some areas, the silt and sand are massive and the bedding is indistinct or nonexistent. SUBSURFACE CONDITIONS General Subsurface materials and conditions at the site were investigated between May 30 and June 1, 2017, with four borings, designated B-1 through B-4; one cone penetrometer test (CPT) sounding, designated CPT-1; and two dilatometer (DMT) soundings, designated DMT-1 and DMT-2. The borings were advanced to depths of about 6.5 to 81.5 ft, the CPT probe to a depth of about 89 ft, and the DMT soundings to depths of about 32 to 36 ft below existing site grades. The approximate locations of the explorations completed for this investigation are shown on Figure 2. Logs of the borings, CPT probe, and DMT soundings are provided on Figures 1A through 8A. The field and laboratory programs conducted to evaluate the physical engineering properties of the materials encountered in the explorations are described in Appendix A. The tern-is and symbols used to describe the materials encountered in the explorations are defined on Tables 1A through 3A and in the attached legend. Sampling Disturbed and undisturbed soil samples were obtained from the borings at 2.5-ft intervals of depth in the upper 15 ft, 5-ft intervals to a depth of 60 ft, and 10-ft intervals below 60 ft. Disturbed soil samples were obtained using a 2-in.-outside-diameter (O.D.) standard split-spoon sampler (SPT). Penetration tests were conducted by driving the sampler into the soil a distance of 18 in. using a 140-lb hammer dropped 30 in. The number of blows required to drive the SPT sampler the last 12 in. is known as the Standard Penetration Resistance, or SPT N-value. SPT N-values provide a measure of the relative density of granular soils and the relative consistency of cohesive soils. Relatively undisturbed soil samples were collected by pushing a 3-in.- O.D. Shelby tube into the undisturbed soil a maximum of 24 in. using the hydraulic ram of the drill rig. The URN 2 soil in the Shelby tubes was extruded in our laboratory and Torvane shear strength measurements were recorded on selected samples. Soils For the purpose of discussion, the materials disclosed by our investigation have been grouped into the following categories based on their physical characteristics and engineering properties: 1. PAVEMENT 2. SURFACING 3. SILT and Silty SAND The following paragraphs provide a detailed description of the materials encountered in the explorations and a discussion of the groundwater conditions at the site. 1. PAVEMENT. Explorations B-1 through B-3, CPT-1, and DMT-1 and DMT-2 were advanced in existing paved areas and encountered approximately 2 to 6 in. of AC pavement at the ground surface. The pavement is underlain by about 2 to 18 in. of crushed-rock base (CRB) course. 2. SURFACING. Exploration B-4 was advanced in an area surfaced with about 12 in. of crushed rock. 3. SILT and Silty SAND. Interbedded layers of silt and silty sand were encountered beneath pavement and surfacing in all of the explorations and extend to the maximum depth explored of 89 ft. The thicknesses of the interbedded layers typically range from about 1 in. to 15 ft; however, 20-to 35-ft-thick layers of silt and silty sand were encountered in explorations B-3 and B-4. The soils are generally brown or gray with varying degrees of rust mottling; however, between 35 and 55 ft, the silt is tan mottled rust and the sand is red- brown. In general, the silt has a variable clay content ranging from up to trace clay to clayey and contains a variable amount of fine- to medium-grained sand ranging from trace sand to sandy; however, a 6-in.-thick layer of silty clay was encountered at a depth of 30.5 ft in boring B-3. The silty sand is fine to medium grained; however, coarse-grained sand was encountered at a depth of 80 ft in boring B-3. The natural moisture content of the silt soils generally ranges from 24 to 47%; however, a sample of silt obtained at a depth of 50 ft in boring B-4 had a moisture content of 75%. The natural moisture content of the silty sand soils generally ranges from 23 to 41%. Atterberg limits testing indicates the interbedded layer of silty clay encountered at a depth of 30.5 ft in boring B-3 has a liquid limit of 41% and a plasticity index of 22%, see Figure 9A. The relative consistency of the silt soils is medium stiff to hard, based on SPT N-values of 8 to 49 blows/ft, CPT tip-resistance values of about 20 to 240 tsf, and DMT constrained modulus values of about 130 to 950 tsf, and is typically medium stiff to stiff to a depth of 30 ft and very stiff to hard below 30 ft. The relative density of the silty sand soils is very loose to very dense, based on SPT N-values of 3 to 45 blows/ft, CPT tip resistance values of about 40 to 275 tsf, and DMT constrained modulus values of about 350 to 2,100 tsf, and is typically loose to medium dense to a depth of 35 ft and dense below. It should be noted that the relative density of sand soils with high silt contents tends to be underestimated using the SPT sampler. All of the explorations were terminated in silt and silty sand at depths ranging from 6.5 to 89 ft. A one-dimensional consolidation test was completed on a sample of the silt obtained at a depth of 25.3 ft in boring B-2. Test results indicate the silt is overconsolidated and has a relatively low compressibility in the GRQ 3 preconsolidated range of pressures and a low to moderate compressibility in the normally consolidated range of pressures, see Figure 10A. Groundwater Groundwater seepage was not encountered in boring B-1 at the time of drilling. Borings B-2 through B-4 were completed using mud-rotary drilling techniques, which do not allow measurement of groundwater levels. Our review of U.S. Geological Survey (USGS) groundwater data suggests the regional groundwater level at the site occurs at a depth of about 35 ft below the ground surface. However, our work in the area indicates perched groundwater conditions can occur in the silt and sand soils that mantle the site throughout the year. To allow measurement and periodic monitoring of perched groundwater levels at the site, a vibrating-wire piezometer was installed at a depth of 50 ft in boring B-4. On June 7, 2017, the local perched groundwater in the piezometer was measured at a depth of about 6 ft below the existing ground surface. We anticipate the local perched groundwater level typically occurs at a depth of 10 to 15 ft below the ground surface during the normally dry summer and fall months and may approach the ground surface during the wet winter and spring months or during periods of heavy or prolonged precipitation. CONCLUSIONS AND RECOMMENDATIONS General Subsurface explorations completed for this investigation indicate the site is mantled with interbedded layers of medium-stiff to stiff silt and medium-dense to dense silty sand. The interbedded layers of silt and silty sand extend to the maximum depth explored of 89 ft. We anticipate the local perched groundwater level typically occurs at depths of 10 to 15 ft below the ground surface throughout the year; however, perched groundwater may approach the ground surface during the wet winter months and following periods of intense or prolonged precipitation. In our opinion, foundation support for new structural loads can be provided by conventional spread and wall foundations established in firm, undisturbed, native soil or compacted structural fill. The primary geotechnical considerations associated with construction of the proposed building and associated improvements include the presence of fine-grained soils at the ground surface that are extremely sensitive to moisture content; the potential for shallow, perched groundwater conditions; and the close proximity of the proposed excavation to the existing school that is sensitive to settlement and lateral movement. The following sections of this report provide our conclusions and recommendations for use in the design and construction of the project. Seismic Considerations General. We understand the project will be designed in accordance with the 2012 IBC with 2014 OSSC modifications. For seismic design, the 2012 IBC references American Society of Civil Engineers (ASCE) document 7-10, titled "Minimum Design Loads for Buildings and Other Structures" (ASCE 7-10). A site- specific seismic-hazard evaluation was completed for the project in accordance with the 2014 OSSC. Details of the site-specific seismic-hazard evaluation and the development of the recommended response spectrum are provided in Appendix B. Code Background. The 2012 IBC and ASCE 7-10 seismic hazard levels are based on a Risk-Targeted Maximum Considered Earthquake (MCER) with the intent of including the probability of structural collapse. The ground motions associated with the probabilistic MCER represent a targeted risk level of 1% in 50 years G RU 4 probability of collapse in the direction of maximum horizontal response with 5% damping. In general, these risk-targeted ground motions are developed by applying adjustment factors of directivity and risk coefficients to the 2% probability of exceedance in 50 years, or 2,475-year return period, hazard level (MCE) ground motions developed from the 2014 USGS Unified Hazard Tool. The risk-targeted probabilistic values are also subject to a deterministic limit. The code-based ground-surface MCER-level spectrum is typically developed using the mapped bedrock spectral accelerations, Ss and Si,and corresponding site coefficients, Fa and Fv,to account for site soil conditions. Site Response. In accordance with Section 20.4.2 of ASCE 7-10, the site is classified as Site Class D, or a stiff-soil site, based on an estimated Vs30 of about 1,050 ft/sec in the upper 100 ft of the soil profile. However, our analysis has identified a potential risk of seismically induced settlement at the site. In accordance with ASCE 7-10, sites with soils vulnerable to failure or collapse under seismic loading should be classified as Site Class F, which requires a site-specific site-response analysis unless the structure has a fundamental period of vibration less than or equal to 0.5 sec. The design response spectrum for sites with structures having a fundamental period of less than or equal to 0.5 sec can be derived using the non- liquefied subsurface profile. For periods greater than 0.5 sec, the code requires a minimum spectral response value equal to 80°/° of Site Class E. We anticipate the new structure will have a fundamental period of less than 0.5 sec; therefore, the code- based Site Class D conditions are appropriate for design of the structure. The maximum horizontal-direction spectral response accelerations were obtained from the USGS Seismic Design Maps for the coordinates of 45.4032° N latitude and 122.7691° W longitude. The Ss and Si parameters identified for the site are 0.96 and 0.42 g, respectively, for Site Class B or bedrock conditions. To establish the ground-surface MCER spectrum, these bedrock spectral coefficients are adjusted for site class using the short- and long-period site coefficients, Fa and Fv, in accordance with Section 11.4.3 of ASCE 7-10. The design-level response spectrum is calculated as two-thirds of the ground-surface MCER spectrum. The recommended MCER- and design-level spectral response parameters for Site Class D conditions are tabulated below and discussed in further detail in Appendix B. RECOMMENDED SEISMIC DESIGN PARAMETERS(2012 IBC/2014 OSSC) Recommended Seismic Parameter Value Site Class D MCER 0.2-Sec Period 1.07 g Spectral Response Acceleration,SMs MCER 1.0-Sec Period 0.66 g Spectral Response Acceleration,Ssm Design-Level 0.2-Sec Period 0.71 g Spectral Response Acceleration,Sos Design-Level 1.0-Sec Period 0,44 g Spectral Response Acceleration,So, Liquefaction/Cyclic Softening. Liquefaction is a process by which loose, saturated granular materials, such as clean sand and, to a somewhat lesser degree, non-plastic and low-plasticity silts, temporarily lose stiffness and strength during and immediately following a seismic event. This degradation in soil properties may be substantial and abrupt, particularly in loose sands. Liquefaction occurs as seismic shear stresses propagate clii 5 through a saturated soil and distort the soil structure, causing loosely packed groups of particles to contract or collapse. If drainage is impeded and cannot occur quickly, the collapsing soil structure causes the pore- water pressure to increase between the soil grains. If the pore-water pressure becomes sufficiently large, the inter-granular stresses become small and the granular layer temporarily behaves as a viscous liquid rather than a solid. After liquefaction is triggered, there is an increased risk of settlement, loss of bearing capacity, lateral spreading, and/or slope instability, particularly along waterfront areas. Liquefaction-induced settlement occurs as the elevated pore-water pressures dissipate and the soil consolidates after the earthquake. Cyclic softening is a term that describes a relatively gradual and progressive increase in shear strain with load cycles. Excess pore pressures may increase due to cyclic loading but will generally not approach the total overburden stress. Shear strains accumulate with additional loading cycles, but an abrupt or sudden decrease in shear stiffness is not typically expected. Settlement due to post-seismic consolidation can occur, particularly in lower-plasticity silts. Large shear strains can develop, and strength loss related to soil sensitivity may be a concern. The potential for liquefaction and/or cyclic softening is typically estimated using a simplified method that compares the cyclic shear stresses induced by the earthquake (demand) to the cyclic shear strength of the soil available to resist these stresses (resistance). Estimates of seismically induced stresses are based on earthquake magnitude and peak ground-surface acceleration (PGA). The cyclic resistance of soils is dependent on several factors, including the number of loading cycles, relative density, confining stress, plasticity, natural water content, stress history, age, depositional environment (fabric), and composition. The cyclic resistance of soils is evaluated using in-situ testing in conjunction with laboratory index testing but may also include monotonic and cyclic laboratory strength tests. For sand-like soils, the cyclic resistance is typically evaluated using SPT N-values or CPT tip-resistance values normalized for overburden pressures and corrected for factors that influence cyclic resistance, such as fines content. For clay-like soils, the cyclic resistance is typically evaluated using estimates of the undrained shear strength, overconsolidation ratio (OCR),and sensitivity, or directly from cyclic laboratory tests. The potential for liquefaction and/or cyclic softening at the site was evaluated using the simplified method based on procedures recommended by Idriss and Boulanger (2008) with subsequent revisions (2014). This method utilizes the PGA to predict the cyclic shear stresses induced by the earthquake. The USGS National Seismic Hazard Mapping Project (NSHMP) was used to determine the contributing earthquake magnitudes that represent the seismic exposure of the site for the MCEc hazard level. A crustal event on the Portland Hills fault and an event on the Cascadia Subduction Zone (CSZ) were determined to represent the sources of seismic shaking. For our evaluation, we have considered a magnitude Mw 7 crustal earthquake and M.9 CSZ earthquake with code-level PGAs (PGAM) of 0.45 and 0.36 g, respectively. We have conservatively assumed a groundwater depth of about 5 ft below the ground surface, which corresponds to the anticipated highest sustained groundwater level at the site. The results of our evaluation indicate there is a potential that zones of the interbedded silt and silty sand deposit below the groundwater surface at the site could lose strength or liquefy during a code-based earthquake. Based on our analysis, potentially liquefiable soils are present about 10 ft below the ground surface and extend to a depth of about 35 ft. Our analysis indicates the potential for 1 to 2 in. of seismically induced settlement, which may occur during the earthquake and after GRU 6 earthquake shaking has ceased. Conventional geotechnical practice is to assume differential settlements may approach 50% of the calculated total seismic settlement. Discussion of seismically induced building foundation settlement is presented in the Foundation Support section later in this report. Other Seismic Hazards. Based on site topography, the risk of earthquake-induced slope instability and/or lateral spreading is low. The risk of damage by tsunami and/or seiche at the site is absent. The inferred location of the Canby-Mollala Fault borders the northeastern corner of the site (Personius et al., 2003); however, the USGS does not consider the Canby-Mollala Fault to be an active, contributing source in their Probabilistic Seismic Hazard Analysis (PSHA). The USGS considers the Portland Hills Fault, located about 12 km northeast of the site, to be the closest crustal fault source contributing to the overall seismic hazard at the site. Unless occurring on a previously unmapped or unknown fault, the risk of fault rupture at the site is low. Earthwork General. The fine-grained soils that mantle the site are sensitive to moisture, and perched groundwater may approach the ground surface during the wet winter months. Therefore, it is our opinion earthwork can be completed most economically during the dry summer months, typically extending from June to mid- October. It has been our experience that the moisture content of the upper few feet of silty soils will decrease during extended warm, dry weather. However, below this depth, the moisture content of the soil tends to remain relatively unchanged and well above the optimum moisture content for compaction. As a result, the contractor must use construction equipment and procedures that prevent disturbance and softening of the subgrade soils. To minimize disturbance of the moisture-sensitive silt soils, site grading can be completed using track-mounted hydraulic excavators. The excavation should be finished using a smooth- edge bucket to produce a firm, undisturbed surface. It may also be necessary to construct granular haul roads and work pads concurrently with excavation to minimize subgrade disturbance. If the subgrade is disturbed during construction, soft, disturbed soils should be overexcavated to firm soil and backfilled with structural fill. If construction occurs during wet ground conditions, granular work pads will be required to protect the underlying silt subgrade and provide a firm working surface for construction activities. In our opinion, a 12- to 18-in.-thick granular work pad should be sufficient to prevent disturbance of the subgrade by lighter construction equipment and limited traffic by dump trucks. Haul roads and other high-density traffic areas will require a minimum of 18 to 24 in. of fragmental rock, up to 6-in. nominal size, to reduce the risk of subgrade deterioration. The use of a geotextile fabric over the subgrade may reduce maintenance during construction. Haul roads can also be constructed by placing a thickened section of pavement base course and subsequently spreading and grading the excess CRB after earthwork is complete. As an alternative to the use of a thickened section of crushed rock to support construction activities and protect the subgrade, the subgrade soils can be treated with cement. It has been our experience in this area that treating the silt soils to a depth of 12 to 14 in. with about a 6 to 8% admixture of cement overlain by 6 to 12 in. of crushed rock will support construction equipment and provide a good all-weather working surface. Site Preparation. Demolition of the existing improvements within the limits of the proposed improvements should include removal of existing pavements, floor slabs, foundations, walls, and underground utilities (if present). The ground surface within all building areas, paved areas, walkways, and areas to receive IE El 7 structural fill should be stripped of existing vegetation, surface organics, and loose surface soils. We anticipate stripping up to a depth of about 4 to 6 in. will likely be required to construct the new bus drop-off area near the northern property boundary; however, deeper grubbing may be required to remove brush and tree roots. All demolition debris, trees, brush, and surficial organic material should be removed from within the limits of the proposed improvements. Excavations required to remove existing improvements, brush, and trees should be backfilled with structural fill. Organic strippings should be disposed of off site, or stockpiled on site for use in landscaped areas. Following stripping or excavation to subgrade level, the exposed subgrade should be evaluated by a qualified member of GRI's geotechnical engineering staff or an engineering geologist. Proof rolling with a loaded dump truck may be part of this evaluation. Any soft areas or areas of unsuitable material disclosed by the evaluation should be overexcavated to firm material and backfilled with structural fill. Due to previous development at the site, it should be anticipated some overexcavation of subgrade will be required. Structural Fill. We anticipate minor amounts of structural fill will be placed for this project. We recommend structural fill consist of granular material, such as sand, sandy gravel, or crushed rock with a maximum size of 2 in. Granular material that has less than 5% passing the No. 200 sieve (washed analysis) can usually be placed during periods of wet weather. Granular backfill should be placed in lifts and compacted with vibratory equipment to at least 95% of the maximum dry density determined in accordance with ASTM D698. Appropriate lift thicknesses will depend on the type of compaction equipment used. For example, if hand-operated vibratory-plate equipment is used, lift thicknesses should be limited to 6 to 8 in. If smooth-drum vibratory rollers are used, lift thicknesses up to 12 in. are appropriate, and if backhoe- or excavator-mounted vibratory plates are used, lift thicknesses of up to 2 ft may be acceptable. On-site, fine-grained soils and site strippings free of debris may be used as fill in landscaped areas. These materials should be placed at about 90% of the maximum dry density as determined by ASTM D698. The moisture contents of soils placed in landscaped areas is not as critical, provided construction equipment can effectively handle the materials. Utility Excavations. In our opinion, there are three major considerations associated with design and construction of new utilities. 1) Provide stable excavation side slopes or support for trench,sidewalls to minimize loss of ground. 2) Provide a safe working environment during construction. 3) Minimize post-construction settlement of the utility and ground surface. The method of excavation and design of trench support are the responsibility of the contractor and subject to applicable local, state, and federal safety regulations, including the current Occupational Safety and Health Administration (OSHA) excavation and trench safety standards. The means, methods, and sequencing of construction operations and site safety are also the responsibility of the contractor. The information provided below is for the use of our client and should not be interpreted to mean we are assuming responsibility for the contractor's actions or site safety. G Rp 8 According to current OSHA regulations, the majority of the soils encountered in the explorations may be classified as Type C. In our opinion, trenches less than 4 ft deep that do not encounter groundwater may be cut vertically and left unsupported during the normal construction sequence, assuming trenches are excavated and backfilled in the shortest possible sequence and excavations are not allowed to remain open longer than 24 hr. Excavations more than 4 ft deep should be laterally supported or alternatively provided with side slopes of 1.5H:1V (Horizontal to Vertical) or flatter. In our opinion, adequate lateral support may be provided by common methods, such as the use of a trench shield or hydraulic shoring systems. Groundwater seepage, running soil conditions, and unstable trench sidewalls or soft trench subgrades, if encountered during construction, will require dewatering of the excavation and trench sidewall support. The impact of these conditions can be reduced by completing trench excavation during the summer months, when groundwater levels are lowest, and by limiting the depths of the trenches. We anticipate groundwater seepage, if encountered, can generally be controlled by pumping from sumps. To facilitate dewatering, it will be necessary to overexcavate the trench bottom to permit installation of a granular working blanket. We estimate the required thickness of the granular working blanket will be on the order of 1 ft or as required to maintain a stable trench bottom. The actual required depth of overexcavation will depend on the conditions exposed in the trench and the effectiveness of the contractor's dewatering efforts. The thickness of the granular blanket must be evaluated on the basis of field observations during construction. We recommend the use of relatively clean, free-draining material, such as 2- to 4-in.-minus crushed rock, for this purpose. The use of a geotextile fabric over the trench bottom will assist in trench-bottom stability and dewatering. All utility trench excavations within building and pavement areas should be backfilled with relatively clean, granular material, such as sand, sandy gravel, or crushed rock of up to 1'/2-in. maximum size and having less than 5% passing the No. 200 sieve (washed analysis). The bottom of the excavation should be thoroughly cleaned to remove loose materials and the utilities should be underlain by a minimum 6-in. thickness of bedding material. The granular backfill material should be compacted to at least 95% of the maximum dry density as determined by ASTM D698 in the upper 5 ft of the trench and at least 92% of this density below a depth of 5 ft. The use of hoe-mounted vibratory-plate compactors is usually most efficient for this purpose. Flooding or jetting as a means of compacting the trench backfill should not be permitted. Excavation and Shoring General. We estimate an excavation on the order of 10 ft will be required to found the partially embedded lower level. We anticipate the majority of excavations may be made with temporary excavation slopes; however, shoring may be required where excavations will be located in close proximity to portions of the existing school, such as near the southeastern corner of the Main Gym. The method of excavation and design of excavation support are the responsibilities of the subcontractor and should conform to applicable local, state, and federal regulations. The information provided below is for the use of our client and should not be interpreted to imply we are assuming responsibility for the subcontractor's actions or site safety. Groundwater Management. Depending on the time of year, excavations may encounter perched groundwater combined with running sand. We anticipate groundwater seepage, if encountered, can be controlled by pumping from temporary sumps in the bottom of the excavation; however, depending on the amount of water entering the excavation, the use of wells or well points extending below the depth of excavation may be required. Problems associated with the control of groundwater can be reduced if the work is scheduled for the dry season,when groundwater levels are at their lowest. 11D 9 Temporary Excavations. The inclination of temporary excavation slopes will depend, in part, on the groundwater conditions encountered at the time of construction and the contractor's ability to control these conditions. In this regard, we anticipate temporary excavation slopes can be cut at 1.5H:1V if groundwater levels are maintained at least 2 ft below the bottom of the excavation. Flatter slopes will be necessary if significant seepage or running soil conditions are encountered. In addition, the excavations may need to be advanced in stages to permit groundwater levels to lower as the excavation is being made to reduce the risk of instability of the excavation sidewalls. The actual depth of each lift and the waiting period will depend on the observed behavior of the excavation sidewalls. At each stage, the groundwater surface should be allowed to draw down near the toe of the slope or bottom of the excavation before advancing to the next stage. Some minor amounts of sloughing, slumping, or running of temporary slopes should be anticipated shortly after individual stages are excavated. A blanket of relatively clean, well-graded crushed rock placed on the slopes may be required to reduce the risk of raveling soil conditions if temporary excavation slopes encounter perched groundwater. We recommend the use of relatively clean, free- draining material, such as 2-to 4-in.-minus crushed rock, for this purpose. The thickness of the granular blanket should be evaluated based on actual conditions but would likely be in the range of 12 to 24 in. In our opinion,the short-term stability of temporary slopes will be adequate if surcharge loads due to existing footings, construction traffic, vehicle parking, material laydown, etc., are maintained an equal distance to the height of the slope away from the top of the open cut. Other measures that should be implemented to reduce the risk of localized failures of temporary slopes include (1) using geotextile fabric to protect the exposed cut slopes from surface erosion; (2) providing positive drainage away from the tops and bottoms of the cut slopes; (3) constructing and backfilling walls as soon as practical after completing the excavation; (4) backfilling overexcavated areas as soon as practical after completing the excavation; and (5) periodically monitoring the area around the top of the excavation for evidence of ground cracking. It must be emphasized that following these recommendations will not guarantee sloughing or movement of the temporary cut slopes will not occur; however, the measures should serve to reduce the risk of a major slope failure. It should be realized, however, that blocks of ground and/or localized slumps may tend to move into the excavation during construction. Shoring Criteria. We recommend using shoring to support the excavation in the following areas: (1) where site constraints do not permit the excavation sidewalls to be sloped at about 1.5H:1V or flatter;and (2) where existing improvements (utilities, adjacent structures, etc.) are located within a setback zone defined by a plane that extends upward at 1.5H:1V from the toe of the excavation and an equal distance to the height of the cut from where the plane intersects the top of the slope. It is common practice in the region to use shoring systems consisting of soldier piles and lagging or interlocked sheet piles, either cantilevered or restrained with tie-back anchors or soil-nail support. Depending on the proximity of the shoring to any existing improvements requiring protection, it may be necessary to leave portions of the temporary shoring system permanently in place to limit the risk of future settlement associated with completely removing the shoring. The design of temporary shoring systems depends on the total magnitude of forces the system is designed to resist and the tolerable yielding of the system and surrounding ground. The pattern and intensity of the lateral earth pressures on the shoring will be governed by the height of the wall, soil type, the degree to which the walls are structurally supported, and whether the walls are drained. The lateral earth pressure criteria shown on Figure 3 can be used for the design of cantilevered shoring systems, assuming the shoring R 10 can be allowed to yield somewhat into the excavation during construction and settlement behind the system can also be tolerated. The lateral earth pressure criteria shown on Figure 4 can be used for shoring restrained with horizontal bracing, tieback anchors, or soil-nails to resist larger forces and/or reduce the amount of yielding of the shoring toward the excavation. Additional lateral pressures due to surcharge loads, such as existing structures, behind the shoring systems should be added to the earth pressures shown on Figures 3 and 4. These additional loads can be computed in accordance with the criteria presented on Figure 5; however, we recommend a minimum vertical surcharge pressure of 250 psf be added behind the wal Is. Additional Shoring Considerations. For a tied-back soldier pile shoring system, we recommend all tie-back anchors develop their pull-out resistance beyond a no-load zone defined by a plane as shown on Figure 4. Verification tests should be completed for at least one anchor per level for each side of the excavation. Verification anchor tests should be conducted to at least 150% of the design anchor load. The results of the tests will be used to review and revise, if necessary, the anchor design criteria. In addition, each production anchor should be proof tested to at least 133% of the design load. The shoring contractor should have a proven record of successful shoring and tie-back installations in similar materials. If shoring is required, we recommend the following monitoring and performance provisions be included in the project specifications. 1) Horizontal movement of the shoring system in the vicinity of adjacent improvements, such as structures, should be accurately measured and recorded at each stage of the excavation by the contractor. Horizontal movement should be measured at the top and at each intermediate bracing level on at least every second soldier pile, or about every 10 ft. Settlement of the ground surface near adjacent structures should be monitored at a minimum spacing of 20 ft along the building edge closest to the excavation. 2) Horizontal movement of the shoring system should not exceed 1/2 in. toward the excavation. 3) Lagging should be installed and any voids backfilled using controlled-density fill, if necessary, as the excavation proceeds. 4) The excavation should not extend more than about 1 ft below a bracing level until the tie-backs, lagging, and backfill at that level are in place. Foundation Support We anticipate the column and wall loads will be on the order of 100 to 200 kips and 3 to 4 kips/ft, respectively. In our opinion, the proposed structural loads can be supported on conventional spread and wall footings in accordance with the following design criteria. As discussed earlier, our analysis indicates 1 to 2 in. of settlement could occur following a code-based seismic event. Based on the thickness of the non- liquefiable soil that mantles the site, we estimate the risk of ground manifestation of the seismically induced settlement is generally low. For design purposes, we recommend assuming differential seismic settlement will approach 50% of the calculated total seismic settlement over the length of the building. G RO 11 The 2015 National Earthquake Hazards Reduction Program (NEHRP) document titled "Recommended Seismic Provisions for New Buildings and Other Structures" provides guidance for acceptable limits of seismic differential settlement for different types of structures and different risk categories. In our opinion and based on Table 12.13-3 of 2015 NEHRP, 0.5 to 1 in. of seismic differential settlement over the length of the building is acceptable and consistent with current standards of practice for a life safety performance level. However, the structural engineer should determine if the structure can accommodate the estimated total and differential seismic settlements. Tying the foundations together with a network of grade beams could be considered to help reduce the potential adverse effects associated with differential vertical movement. The grade beams should be designed in accordance with the guidelines presented in the 2015 NEHRP document. All footings should be established in the medium-stiff, native soil that mantles the site. The base of all new footings should be established at a minimum depth of 18 in. below the lowest adjacent finished grade. Footing widths should not be less than 24 in. for isolated column footings and 18 in. for wall footings. Excavations for all foundations should be made with a smooth-edge bucket, and all footing subgrades should be observed by a member of GRI's geotechnical engineering staff. Soft or otherwise unsuitable material encountered at foundation subgrade level should be overexcavated and backfilled with granular structural fill. Local areas of softer subgrade may require deeper overexcavation and should be evaluated by a qualified member of GRI's geotechnical engineering staff. Our experience indicates the subgrade soils are easily disturbed by excavation and construction activities. Due to these considerations, we recommend installing a minimum 3-in.-thick layer of compacted crushed rock in the bottom of all footing excavations. Relatively clean, 3/4-in.-minus crushed rock is suitable for this purpose. Footings established in accordance with these criteria can be designed on the basis of an allowable soil bearing pressure of 3,000 psf. This value applies to the total of dead load and/or frequently applied live loads and can be increased by one-third for the total of all loads: dead, live, and wind or seismic. We estimate the total static settlement of spread and wall footings designed in accordance with the recommendations presented above will be less than 1 in. for footings supporting column and wall loads of up to 200 kips and 4 kips/ft, respectively. Differential static settlements between adjacent, comparably loaded footings should be less than half the total settlement. Horizontal shear forces can be resisted partially or completely by frictional forces developed between the base of the footings and the underlying soil and by soil passive resistance. The total frictional resistance between the footing and the soil is the normal force times the coefficient of friction between the soil and the base of the footing. We recommend an ultimate value of 0.35 for the coefficient of friction for footings cast on granular material. The normal force is the sum of the vertical forces (dead load plus real live load). If additional lateral resistance is required, passive earth pressures against embedded footings can be computed on the basis of an equivalent fluid having a unit weight of 250 pcf. This design passive earth pressure would be applicable only if the footing is cast neat against undisturbed soil or if backfill for the footings is placed as granular structural fill and assumes up to 1/2 in. of lateral movement of the structure will occur in order for the soil to develop this resistance. This value also assumes the ground surface in front of the foundation is horizontal, i.e., does not slope downward away from the toe of the footing. G RD 12 • Subdrainage/Floor Support To provide a capillary break and reduce the risk of damp floors, slab-on-grade floors established at or above adjacent final site grades should be underlain by a minimum 8 in. of free-draining, clean, angular rock. This material should consist of angular rock such as 1 1/2-to 3/4-in. crushed rock with less than 2°/0 passing the No. 200 sieve (washed analysis) and should be placed in one lift and compacted to at least 95% of the maximum dry density (ASTM D698) or until well keyed. To improve workability, the drain rock should be capped with a 2-in.-thick layer of compacted, 3/4-in.-minus crushed rock. In areas where floor coverings will be provided or moisture-sensitive materials stored, it would be appropriate to also install a vapor-retarding membrane. The membrane should be installed as recommended by the manufacturer. In addition, a foundation drain should be installed around the building perimeter to collect water that could potentially infiltrate beneath the foundations and should discharge to an approved storm drain. We anticipate the finished floor elevation for the majority of the partially embedded lower level will be established below final site grades. Unless the partially embedded lower level is designed to be watertight and resist hydrostatic pressures, subdrainage should be provided for the portion of the structure established below final site grades. A subdrainage system will reduce hydrostatic pressure and the risk of groundwater entering through the embedded wall and floor slabs. Typical subdrainage details for embedded structures are shown on Figure 6. The figure shows peripheral subdrains to drain embedded walls and an interior granular drainage blanket beneath the concrete floor slab, which is drained by a system of subslab drainage pipes. All perched groundwater collected should be drained by gravity or pumped from sumps into the stormwater disposal facility. If the water is pumped, an emergency power supply should be included to prevent flooding due to power loss. In our opinion, it is appropriate to assume a coefficient of subgrade reaction, k, of 175 pci to characterize the subgrade support for point loading with 10 in. of compacted crushed rock beneath the floor slab. Retaining/ Embedded Walls Construction of the lower level will require embedded walls with a maximum height of about 10 ft. We anticipate the walls will be cast-in-place and supported on wall or spread foundations. Foundation design and subgrade preparation should conform to the recommendations provided above for foundation support Design lateral earth pressures for retaining walls depend on the type of construction, i.e., the ability of the wall to yield. Possible conditions are 1) a wall laterally supported at its base and top and therefore unable to yield to the active state; and 2) a retaining wall, such as a typical cantilever or gravity wall, that yields to the active state by tilting about its base. A conventional basement wall and cantilever retaining wall are examples of non-yielding and yielding walls, respectively. For completely drained, horizontal backfill, yielding and non-yielding walls may be designed on the basis of equivalent fluid unit weights of 35 and 50 pcf, respectively. To account for seismic loading, the earth pressures should be increased by 10 and 18 pcf for yielding and non-yielding walls, respectively. This results in a triangular distribution with the resultant acting at 1/3H up from the base of the wall, where H is the height of the wall in feet. Additional lateral loading due to surcharge loads can be evaluated using the criteria shown on Figure 5. The lateral earth pressure design criteria presented above are appropriate if the embedded walls are drained. Although the permanent groundwater level likely occurs below the base of the proposed structure, perched groundwater may occur within the shallow silty soils and existing utility trenches during periods of GD0 13 prolonged or intense precipitation. We recommend installation of permanent drainage behind all embedded walls. For walls constructed adjacent to an open cut, we recommend placing a drainage blanket of rock that contains less than 2% fines between the backfill and the face of the wall. The drainage blanket should have a minimum width of 12 in. and can be drained through a perforated pipe at the base of the drainage blanket. A typical drainage system for walls constructed without shoring is shown on Figure 6. If shoring is used, we recommend installing continuous drainage panels on the embedded wall, which is a typical practice for similar applications in the region. The drainage panels should extend to the base of the wall, where water should be collected in a perforated plastic pipe and discharged to a sump or approved storm drain. In addition, the wall design should include positive drainage measures to prevent ponding of surface water behind the top of the wall. In areas where it is not practical to completely drain the backfill and the walls will be designed as undrained and watertight structures, yielding and non-yielding walls can be designed on the basis of equivalent fluid unit weights of 80 and 90 pcf, respectively. Overcompaction of backfill behind walls should be avoided. Heavy compactors and large pieces of construction equipment should not operate within 5 ft of any embedded wall to avoid the buildup of excessive lateral earth pressures. Compaction close to the walls should be accomplished with hand-operated vibratory-plate compactors. Overcompaction of backfill could significantly increase lateral earth pressures behind walls. Pavement Design We anticipate the new bus drop-off pavement will be subjected to bus, automobile, and light truck traffic, and new parking areas will be subjected primarily to automobile and light truck traffic with occasional heavy truck traffic. Traffic estimates for the bus drop-off and parking areas are presently unknown; however, we anticipate the new pavement will consist of AC. Based on our experience with similar projects and subgrade soil conditions, we recommend the following pavement sections. RECOMMENDED PAVEMENT SECTIONS CRB AC Thickness,in. Thickness,in. Areas Subject to School-Bus Traffic(Bus 14 5 Drop-Off Area) Areas Subject to Primarily Automobile 12 4 Traffic(Vehicle Drive Lanes) Areas Subject to Automobile Parking 8 3 (Parking Stalls) The recommended pavement sections should be considered minimum thicknesses and underlain by a woven geotextile fabric. It should be assumed some maintenance will be required over the life of the pavement (15 to 20 years). The recommended pavement section is based on the assumption pavement construction will be accomplished during the dry season and after construction of the building has been completed. If wet-weather pavement construction is considered, it will likely be necessary to increase the thickness of CRB to support construction equipment and protect the subgrade from disturbance. The GRD 14 indicated sections are not intended to support extensive construction traffic, such as dump trucks and concrete trucks. Pavements subject to construction traffic may require repair. For the above-indicated sections, drainage is an essential aspect of pavement performance. We recommend all paved areas be provided positive drainage to remove surface water and water within the base course. This will be particularly important in cut sections or at low points within the paved areas, such as at catch basins. Effective methods to prevent saturation of the base course materials include providing weep holes in the sidewalls of catch basins, subdrains in conjunction with utility excavations, and separate trench-drain systems. To ensure quality materials and construction practices, we recommend the pavement work conform to Oregon Department of Transportation standards. Prior to placing base course materials, all pavement areas should be proof rolled with a fully loaded, 10-cy dump truck. Any soft areas detected by the proof rolling should be overexcavated to firm ground and backfilled with compacted structural fill. Provided the pavement section is installed in accordance with the recommendations provided above, it is our opinion the site-access areas will support infrequent traffic by an emergency vehicle having a gross vehicle weight (GVW) of up to 75,000 lbs. For the purposes of this evaluation, "infrequent" can be defined as once a month or less. DESIGN REVIEW AND CONSTRUCTION SERVICES We welcome the opportunity to review and discuss construction plans and specifications for this project as they are being developed. In addition, GRI should be retained to review all geotechnical-related portions of the plans and specifications to evaluate whether they are in conformance with the recommendations provided in our report. To observe compliance with the intent of our recommendations, our design concepts, and the plans and specifications, we are of the opinion that all construction operations dealing with earthwork and foundations should be observed by a GRI representative. Our construction-phase services will allow for timely design changes if site conditions are encountered that are different from those described in our report. If we do not have the opportunity to confirm our interpretations, assumptions, and analyses during construction, we cannot be responsible for the application of our recommendations to subsurface conditions different from those described in this report. LIMITATIONS This report has been prepared to aid the architect and engineer in the design of this project. The scope is limited to the specific project and location described herein, and our description of the project represents our understanding of the significant aspects of the project relevant to the design and construction of the new foundations and floors. In the event any changes in the design and location of the project elements as outlined in this report are planned, we should be given the opportunity to review the changes and modify or reaffirm the conclusions and recommendations of this report in writing. The conclusions and recommendations submitted in this report are based on the data obtained from the explorations made at the locations indicated on Figure 2 and other sources of information discussed in this report. In the performance of subsurface investigations, specific information is obtained at specific locations at specific times. However, it is acknowledged that variations in soil conditions may exist between exploration locations. This report does not reflect any variations that may occur between these explorations. GRI 15 The nature and extent of variation may not become evident until construction. If, during construction, subsurface conditions differ from those encountered in the explorations, we should be advised at once so that we can observe and review these conditions and reconsider our recommendations where necessary. Please contact the undersigned if you have any questions. Submitted for GRI, \� � l N 18281 'p • FSLEY SP1°.6 1/1-C2— Renews 06/2018 Wesley Spang, PhD, PE, GE Nicholas M. Hatch, PE Principal Project Engineer This document has been submitted electronically. References Idriss, I.M., and Boulanger, R.W., 2008, Soil liquefaction during earthquakes: Earthquake Engineering Research Institute, EERI MNO-12. Idriss, I.M.,and Boulanger, R.W.,2014,CPT and SPT based liquefaction triggering procedures: Department of Civil & Environmental Engineering, College of Engineering, University of California at Davis, Report No. UCD/CGM-14/01. Madin, I.P.,1990, Earthquake-hazard geology maps of the Portland metropolitan area: Oregon Department of Geology and Mineral Studies,Open-File Report 90-02. Personius,S. F., Dart, R. L., Bradley, Lee-Ann,and Haller, K.M.,2003,Map and data for Quatemary faults and folds in Oregon: U.S. Geological Survey Open-File Report 03-095. 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I. m _,1'., 1 Ual tir,\ F'1 Q J l 1( .410,0 ',. ° �� n S„ Community L up -��� -li ZD��-t_� i .� 1tlr' �1r i[jrPark�*t4-Ramp�r o �� i VETON DR a`�, E _- .i 111111, �� ,, ti `-J'" 3 3 I n C✓9ta �� 11' I / l Wyk 7 USGS TOPOGRAPHIC MAP BEAVERTON OREG.(2014) LAKE OSWEG`O,OREG.(2014) Aar- 0 1/2 1 MILE 1 I Gn fl TIGARD TUALATIN SCHOOL DISTRICT [� TIGARD HIGH SCHOOL VICINITY MAP FEB.2018 JOB NO.5970-F FIG. 1 i - 4� [y).. q ... �~ it "4 :•_ate„ ":•, _..� - " SWDURH.4M11 ROAD - � . - \446, ' - C✓ ,,/ ' NEW BUS DROP-OFF �! / �`� f! DMT-1 V B-2 _ \ - _ F. r F. .., . ,_,•,.. _ ...,:,.7-y.,i.-.4:p., - , • N s k I: �� -700 y r ,.r' AUDITORIUM i k �� �� ,- k' t CAFETERIA I - - � : MAIN GYM% °� �. >z � EXISTING SCHOOL \ t � - - - .',..,. '4,9',4*,..:;%:::::**4!:**it:::!**z..* • ‘00‘,.....1.;;00.„41,-- e1 °,° CPT-1 ':� NEW PARKING AREA ,fib y� ' / r/ �, ' , ''. ' NEW SCHOOL FOOTPRINT (4 l �� e` `'4 °3 'Al %�. �� c . t ° e 9 ; � �I _ - DILATOMETER COMPLETED BY GRI d v / e?t / (JUNE1,2017) aroma_ • �� � _ �,- • CONE PENETRATION TEST COMPLETED BY GRI (DUNE 1,2017) f� S t, _' f9 BORING COMPLETED BY GRI �: t (MAY 30-31,2017) AUTO SHOP ' 1 ate. .. , _ _ L _ ; ,J B4 SITE MAP FROM GOOGLE EARTH PRO,DATED DULY 23,2016 I;+� As DMT-2 gliti 0 120 240 ET • TIGARDTIGARD TUALATINHIGHSCHOOL SCHOOL DISTRICT �..--...„ .,. �] MID _ '- - SITE MAP ;. __ `" FEB.2018 JOB NO.5970-F FIG.2 WALL ''i SOLDIER PILE- `� H,FT 4 V ► II ► 1 h,FT ► / II. 4111 (SEE NOTE 2) >i (35 PCF)H +h (SEE NOTE 2&4) >- (PASSIVE) (ACTIVE) NOTES: 1) SURCHARGE EFFECTS FROM EXISTING STRUCTURES,CONSTRUCTION EQUIPMENT,ETC.,SHOULD BE ADDED TO THE ABOVE DESIGN PRESSURES.LATERAL LOADS ON THE SHORING DUE TO SURCHARGE EFFECTS CAN BE COMPUTED USING THE CRITERIA PROVIDED IN FIGURE 5.THE ACTUAL AMOUNT OF THIS SURCHARGE WILL DEPEND ON THE CONTRACTOR'S APPROACH TO THE WORK;HOWEVER,WE RECOMMEND A MINIMUM ADDITIONAL VERTICAL PRESSURE OF 250 PSF BE ADDED BEHIND THE WALL. 2) FOR CANTILEVERED SOLDIER PILES WITH LAGGING,BELOW THE BOTTOM OF THE EXCAVATION,PASSIVE PRESSURE ACTS OVER TWO PILE DIAMETERS(ACTUAL AREA),AND ACTIVE PRESSURE ACTS OVER ONE PILE DIAMETER(ACTUAL AREA)AND ASSUMES A MINIMUM SOLDIER PILE SPACING OF THREE DIAMETERS. 3) DESIGN PRESSURES ASSUME FULLY DRAINED CONDITIONS. 4) ACTIVE PRESSURE ACTS OVER THE ENTIRE EXPOSED SHORING AND/OR WALL AREA. 5) SOLDIER PILES SHOULD EXTEND AT LEAST 5 FT BELOW THE LOWEST ADJACENT EXCAVATION LEVEL. 10 TIGARD IN SCHOOL DISTRICT kJ! TIGARD SCHOOL HIGH SCHOOL EARTH PRESSURES FOR CANTILEVER SHORING FEB.2018 JOB NO.5970-F FIG.3 WALL A .2 H SOLDIER PILE—_,\ 30° 25H(PSF)FOR TEMPORARY SHORING H,FT (SEE NOTE 4) NO-LOAD ZONE m 0.2 H `I _ n - H!4 h,FT V (250 PCF)h(SEE NOTE 2) NOTES: 1) SURCHARGE EFFECTS FROM EXISTING STRUCTURES,CONSTRUCTION EQUIPMENT,ETC.,SHOULD BE ADDED TO THE ABOVE DESIGN PRESSURES.LATERAL LOADS ON THE SHORING DUE TO SURCHARGE EFFECTS CAN BE COMPUTED USING THE CRITERIA PROVIDED IN FIGURE 5.THE ACTUAL AMOUNT OF THIS SURCHARGE WILL DEPEND ON THE CONTRACTOR'S APPROACH TO THE WORK;HOWEVER,WE RECOMMEND A MINIMUM ADDITIONAL VERTICAL PRESSURE OF 250 PSF BE ADDED BEHIND THE WALL. 2) PASSIVE PRESSURE ACTS OVER TWO DIAMETER(ACTUAL AREA)OF THE SOLDIER PILE AND ASSUMES A MINIMUM SOLDIER PILE SPACING OF THREE DIAMETERS. 3) DESIGN PRESSURES ASSUME FULLY DRAINED CONDITIONS. 4) ACTIVE PRESSURE ACTS OVER THE ENTIRE EXPOSED SHORED AND/OR WALL AREA. 5) SOLDIER PILES SHOULD EXTEND AT LEAST 5 FT BELOW THE LOWEST ADJACENT EXCAVATION LEVEL. G R TIGARD TUALATIN SCHOOL DISTRICT TIGARD HIGH SCHOOL EARTH PRESSURES FOR BRACED SHORING FEB.2018 JOB NO.5970-F FIG.4 < X=mH > STRIP WAD,q LINE LOAD,Qi ��a T1111212 -.,,.tea / A i_ /3/2 Z=nH O For m S.0.4: a� Q r ah = — 0. H y H �- H 10.16+ n ) I, For m >0.4: ah= 2qQ Iig•SINS COS 2a) It ah ah_ QL 1.28m2n ah H (m2+n2I2 j (#in radians) LINE LOAD PARALLEL TO WALL STRIP LOAD PARALLEL TO WALL < X=mH >I POINT LOAD,Qp Z=nH OM For m S 0.4: A-111 l' VW_IIII-A' ah - H Qp 0.28n2 H ar (0. - r lir Form >OA: _ Qp 1.77m2n2 allah H2 (m2+ n2)3 / a'h=ah COS2(1.10) NOTES: ah 1. THESE GUIDELINES APPLY TO RIGID WALLS WITH POISSON'S O Mr RATIO ASSUMED TO BE 0.5 FOR BACKFILL MATERIALS. '/, A' 2. LATERAL PRESSURES FROM ANY COMBINATION OF ABOVE Oro h LOADS MAY BE DETERMINED BY THE PRINCIPLE OF SUPERPOSITION. X=mH > DISTRIBUTION OF HORIZONTAL PRESSURES VERTICAL POINT LOAD G LD TIGARD TUALATIN SCHOOL DISTRICT TIGARD HIGH SCHOOL SURCHARGE-INDUCED LATERAL PRESSURE FEB.2018 JOB NO.5970-F FIG.5 . ^� 1 /V 3/4-IN.-MINUS CRUSHED ROCK WITH SEAL WITH ON-SITE LES( SAN5%PASsiNG NO.200 SIEVE IMPERVIOUS MATERIAL (WASHED ANALYSIS) SLOPE TO DRAIN —77— a� A: • 2 IN: I IA •CONCRETE SLAB ' .• •• • , • • •_ ' VARIES VAPOR-RETARDING MEMBRANE SYSTEM ••• • •• • o (2 IN.MIN.) 8 IN.(MIN.) (SEE NOTE 1) �IIIIIIIi' If 1.5 ••• °'.•• I VARIES(2 IN.MIN.) a.' GRANULAR BACKFILL COMPACTED 1 FT •'= TO ABOUT 95%OF THE MAXIMUM (MIN.) SEE DETAIL'A'FOR TYPICAL DRY DENSITY AS DETERMINED BY (UNDERSLAB ROCK OF UP TO 2-IN.SIZE WITH 4-IN.-DIAMETER PERFORATED DRAIN PIPES ASTM D SIT RECOMMENDATIONS NOT MORE THAN 2%PASSING THE ARE TYPICALLY PLACED ON 20-FT CENTERS •698 NO.200 SIEVE(WASHED ANALYSIS) AND SLOPED TO DRAIN(SEE NOTE 2) TEMPORARY CONSTRUCTION •: DETAIL 'A' SLOPE NOT TO SCALE • fa •,• ":SS • 11/2TO3/4 GRAVEL WITH LESS THAN .- • 'a, 2%PASSING THE NO.200 SIEVE a f 1 l I 1111 y (WASHED ANALYSIS) UNDERSLAB DRAIN 4IN:DIAMETER PERFORATED PLASTIC NOTES: DRAIN PIPE,SLOPE TO DRAIN 1) A VAPOR-RETARDING MEMBRANE SYSTEM IS RECOMMENDED FOR MOISTURE-SENSITIVE AREAS AND SHOULD BE INSTALLED IN ACCORDANCE WITH MANUFACTURER'S RECOMMENDATIONS. 2) INTERNAL 4-N.-DIAMETER PERFORATED DRAIN PIPES ARE TYPICALLY PERIMETER DRAIN NOT NECESSARY IN THOSE AREAS WHERE THE FINISHED FLOOR WILL BE ABOVE EXISTING SITE GRADES. G [ jj TIGARD IN SCHOOL DISTRICT HIGH SCHOOL HIGH SCHOOL TYPICAL SUBDRAINAGE DETAILS FEB.2018 JOB NO.5970-F FIG.E. APPENDIX A Field Explorations and Laboratory Testing APPENDIX A FIELD EXPLORATIONS AND LABORATORY TESTING FIELD EXPLORATIONS Subsurface materials and conditions at the site were investigated between May 30 and June 1, 2017, with four borings, designated B-1 through B-4; one cone penetrometer test (CPT) sounding, designated CPT-1; and two dilatometer (DMT) soundings, designated DMT-1 and DMT-2. The approximate locations of the explorations completed for this investigation are shown on Figure 2. Logs of the borings, CPT probe, and DMT soundings are provided on Figures 1A through 8A. The field exploration work was coordinated and documented by an experienced member of GRI's geotechnical engineer staff, who maintained a log of the materials and conditions disclosed during the course of work. Borings Four borings, designated B-1 through B-4, were advanced to depths of about 6.5 to 81.5 ft below existing site grades. The borings were completed with hollow-stem auger or mud-rotary drilling techniques using a truck-mounted drill rig provided and operated by Western States Soil Conservation of Hubbard, Oregon. Disturbed and undisturbed soil samples were obtained from the borings at 2.5-ft intervals of depth in the upper 15 ft, 5-ft intervals to a depth of 60 ft, and 10-ft intervals below 60 ft. Disturbed soil samples were obtained using a standard split-spoon sampler (SPT).- The outside diameter of the SPT sampler is 2 in. Penetration tests were conducted by driving the sampler into the soil a distance of 18 in. using a 140-lb hammer dropped 30 in. The number of blows required to drive the SPT sampler the last 12 in. is known as the Standard Penetration Resistance, or SPT N-value. SPT N-values provide a measure of the relative density of granular soils and relative consistency of cohesive soils. Samples obtained from the borings were placed in airtight jars and returned to our laboratory for further classification and testing. In addition, relatively undisturbed samples were collected by pushing a 3-in.-outside-diameter (O.D.) Shelby tube into the undisturbed soil a maximum distance of 24 in. using the hydraulic ram of the drill rig. The soil exposed in the end of the Shelby tube was examined and classified in the field. After classification, the tube was sealed with rubber caps and returned to our laboratory for further examination and testing. Logs of the borings are provided on Figures 1A through 4A. Each log presents a descriptive summary of the various types of materials encountered in the boring and notes the depths at which the materials and/or characteristics of the materials change. To the right of the descriptive summary, the numbers and types of samples are indicated. Farther to the right, N-values are shown graphically, along with the natural moisture contents, Torvane shear strength values, Atterberg Limits, and percent passing the No. 200 sieve, where applicable. The terms and symbols used to describe the materials encountered in the borings are defined in Table 1 A and the attached legend. Electric Cone Penetration Test One electric CPT probe, designated CPT-1, was advanced to a depth of about 89 ft using a truck-mounted CPT rig provided and operated by Oregon Geotechnical Explorations, Inc., of Keizer, Oregon. During the CPT, a steel cone is forced vertically into the soil at a constant rate of penetration. The force required to cause penetration at a constant rate can be related to the bearing capacity of the soil immediately G 11110 A-1 • surrounding the point of the penetrometer cone. This force is measured and recorded every 8 in. In addition to the cone measurements, measurements are obtained of the magnitude of force required to force a friction sleeve, attached above the cone, through the soil. The force required to move the friction sleeve can be related to the undrained shear strength of fine-grained soils. The dimensionless ratio of sleeve friction to point bearing capacity provides an indicator of the type of soil penetrated. The cone-penetration resistance and sleeve friction can be used to evaluate the relative consistency of cohesion less and cohesive soils, respectively. In addition, a piezometer fitted between the cone and the sleeve measures changes in water pressures as the probe is advanced and can also be used to measure the depth of the top of the groundwater table. The probe was also operated using an accelerometer fitted to the probe, which allows measurement of the arrival time of shear waves from impulses generated at the ground surface. This allows calculation of shear-wave velocities for the surrounding soil profile. A log of the electric CPT probe is provided on Figures 5A, which presents a graphical summary of the tip resistance, local (sleeve) friction, friction ratio, pore pressure, and soil behavior type (SBT) index. The terms used to describe the soils encountered in the probe are defined in Table 2A. Shear-wave velocity measurements were recorded for the CPT-1 probe and are shown on Figure 6A. Dilatometer Test Two DMT soundings, designated DMT-1 and DMT-2, were advanced to depths of about 32 to 36 ft using a truck-mounted CPT rig provided and operated by Oregon Geotechnical Explorations, Inc., of Keizer, Oregon. DMT soundings provide additional geotechnical information to characterize the subsurface materials. The dilatometer test is performed by pushing a blade-shaped instrument into the soil. The blade is equipped with an expandable membrane on one side that is pressurized until the membrane moves horizontally into the surrounding soil. Readings of the pressures required to move the membrane to a point flush with the blade (Po- pressure) and 1.1 mm into the surrounding soil (Pi — pressure) are recorded. The test sequence was performed at 8-in. intervals to obtain a comprehensive soil profile. A material index (ID), horizontal stress index (KD), and dilatometer modulus (ED) are obtained directly from the dilatometer data. The constrained modulus (M) is then obtained from the dilatometer data. The dilatometer test results are summarized on Figures 7A and 8A. The results show the dilatometer pressure readings (Po, Pi) and three dilatometer-derived parameters: horizontal stress index (Ko), material index (ID), and constrained modulus (M). The terms used to describe the materials encountered in the sounding are defined in Table 3A. LABORATORY TESTING General The samples obtained from the borings were examined in our laboratory, where the physical characteristics of the samples were noted and the field classifications modified where necessary. At the time of classification, the natural moisture content of each sample was determined. Additional testing included dry unit weight, Atterberg limits, one-dimensional consolidation, and grain size analyses. A summary of the laboratory test results has been provided in Table 4A. The following sections describe the testing program in more detail. G R O A-2 Natural Moisture Content Natural moisture content determinations were made in conformance with ASTM D2216. The results are summarized on Figures 1 A through 4A and in Table 4A. Undisturbed Unit Weight The unit weight, or density, of undisturbed soil samples was determined in the laboratory in conformance with ASTM D2937. The results are summarized on Figures 2A through 4A and in Table 4A. Atterberg Limits Atterberg limits testing was performed for one representative sample of silt in conformance with ASTM D4318. The test results are summarized on the Plasticity Chart, Figure 9A; Figure 3A; and Table 4A. One-Dimensional Consolidation A one-dimensional consolidation test was performed in conformance with ASTM D2435 on a relatively undisturbed soil sample extruded from a Shelby tube. This test provides data on the compressibility of underlying fine-grained soils, necessary for settlement studies. The test results are summarized on Figure 10A in the form of a curve showing percent strain versus applied effective stress. The initial dry unit weight and moisture content of the sample are also shown on the figure. Grain-Size Analysis Washed-Sieve Method. To assist in classification of the soils, samples of known dry weight were washed over a No. 200 sieve. The material retained on the sieve was oven-dried and weighed. The percentage of material passing the No. 200 sieve was then calculated. The results are summarized in Figures 1 A through 4A and in Table 4A. GRD A-3 Table 1A GUIDELINES FOR CLASSIFICATION OF SOIL Description of Relative Density for Granular Soil Standard Penetration Resistance Relative Density (N-values),blows per ft very loose 0-4 loose 4- 10 medium dense 10-30 dense 30-50 very dense over 50 Description of Consistency for Fine-Grained (Cohesive)Soils Standard Penetration Torvane or Resistance(N-values), Undrained Shear Consistency blows per ft Strength,tsf very soft 0-2 less than 0.125 soft 2-4 0.125-0.25 medium stiff 4-8 0.25-0.50 stiff 8- 15 0.50- 1.0 very stiff 15- 30 1.0-2.0 hard over 30 over 2.0 Grain-Size Classification Modifier for Subclassification Boulders: Primary Constituent Primary Constituent >12 in. SAND or GRAVEL SILT or CLAY Cobbles: Adjective Percentage of Other Material(by weight) 3- 12 in. trace: 5-15 (sand, gravel) 5- 15 (sand,gravel) Gravel: some: 15-30(sand, gravel) 15-30(sand, gravel) t/a 3/4 in. (fine) sandy,gravelly: 30-50 (sand, gravel) 30-50(sand,gravel) 3/4-3 in. (coarse) Sand: trace: <5 (silt, clay) No.200- No.40 sieve(fine) Relationship of clay and No.40- No. 10 sieve(medium) some: 5 12 (silt, clay) silt determined by No. 10- No. 4 sieve(coarse) silty, clayey: 12-50(silt,clay) plasticity index test Silt/Clay: pass No. 200 sieve Table 2A: CONE PENETRATION TEST(CPT)CORRELATIONS COHESIVE SOILS Cone-Tip Resistance, tsf Consistency <5 Very Soft 5 to 15 Soft to Medium Stiff 15 to 30 Stiff 30 to 60 Very Stiff >60 Hard COHESIONLESS SOILS Cone-Tip Resistance, tsf Relative Density <20 Very Loose 20 to 40 Loose 40 to 120 Medium 120 to 200 Dense >200 Very Dense Reference Kulhawy,F.H.,and Mayne, P.W., 1990,Manual on estimating soil properties for foundation design:Electric Power Research Institute,EL-6800. Table 3A: SOIL CHARACTERIZATION BASED ON MARCHETTI FLAT-PLATE DILATOMETER TEST Description of Consistency for Fine-Grained (Cohesive) Soils Soil Type CH, CL ML, MH DMT Constrained Modulus (Mohr),tsf Consistency Ip`< 0.6 0.6 <Io`2)< 1.8 Very Soft 0 -30 0-50 Soft 30- 60 50 - 100 Medium Stiff 60- 100 100-200 Stiff 100- 175 200- 375 Very Stiff 175 + 375 + Description of Relative Density for Granular Soils Soil Type") SM, SC SP, SW DMT Constrained Modulus(Mohr),tsf Relative Density 1.8 <ID`z'< 3.3 3.3 <lo'2 Very Loose 0-75 0- 100 Loose 75 - 150 100-200 Medium Dense 150- 300 200 -425 Dense 300- 550 425 -850 Very Dense 550 + 850 + 1) Unified Soil Classification System 2) ID = Material Index Table 4A SUMMARY OF LABORATORY RESULTS Sample Information Atterberg Limits Moisture Dry Unit Liquid Plasticity Fines Location Sample Depth,ft Elevation,ft Content,% Weight,pcf Limit, % Index,% Content, % Soil Type B-1 S-2 2.5 - 24 - - - 59 Sandy SILT S-3 5.0 - 31 - - - - Sandy SILT B-2 S-1 2.5 - 27 - - - - Sandy SILT S-2 5.3 - 34 - - - - Silty SAND S-2 6.0 - 32 90 - - - Sandy SILT S-3 7.0 - 29 - - - 27 Silty SAND S-4 10.5 - 29 - - - - Silty SAND S-4 11.5 - 23 94 - - - Silty SAND S-5 12.0 - 27 - - - 25 Silty SAND S-6 15.0 - 31 - - - - Silty SAND S-7 20.0 - 31 - - - 70 Sandy SILT S-8 25.3 - 31 - - - 62 Sandy SILT S-9 27.0 - 33 - - - 67 Sandy SILT 5-10 30.0 - 35 - - - 36 Silty SAND S-11 35.0 - 33 - - - - SILT S-12 40.0 - 36 - - - - Silty SAND S-13 45.0 - 33 - - - - Silty SAND S-14 50.0 - 34 - - - 63 Sandy SILT B-3 S-1 2.5 - 36 - - - 50 Silty SAND S-2 5.5 - 38 - - - - Sandy SILT S-2 6.5 - 39 83 - - - Silty SAND S-3 7.0 - 41 - - - - Silty SAND S-4 10.0 - 32 - - - - Silty SAND S-5 12.5 - 39 - - - 26 Silty SAND S-6 15.3 - 31 91 - - - Silty SAND S-7 17.0 - 41 - - - - Silty SAND S-8 20.0 - 34 - - - 15 Silty SAND S-9 25.0 - 30 - - - - Sandy SILT S-10 30.8 - 33 - 41 22 - Silty CLAY S-11 35.0 - 38 - - - - SILT 5-12 40.0 - 47 - - - - SILT 5-13 45.0 - 37 - - - - Silty SAND S-14 50.0 - 28 - - - 28 Silty SAND S-15 55.0 - 45 - - - - SILT 5-16 60.0 - 39 - - - 58 Sandy SILT 5-17 70.0 - 42 - - - - Sandy SILT 5-18 80.0 - 32 - - - 15 Silty SAND B-4 S-1 2.5 - 36 - - - - Silty SAND S-2 5.0 - 37 - - - 33 Silty SAND S-3 7.7 - 34 - - - - Silty SAND G1 `O Page 1 of 2 Table 4A SUMMARY OF LABORATORY RESULTS Sample Information Atterberg Limits Moisture Dry Unit Liquid Plasticity Fines Location Sample Depth,ft Elevation,ft Content, % Weight,pcf Limit, % Index,% Content,% Soil Type B-4 S-3 8.2 — 40 81 — — — Silty SAND 5-4 9.5 — 38 — — — — Silty SAND S-5 12.7 - 38 — — — — Silty SAND S-5 13.2 — 32 91 — — Silty SAND S-6 14.5 — 29 — — — — Silty SAND S-7 20.0 — 30 — — — 37 Silty SAND S-8 25.0 — 31 — — — — Silty SAND S-9 30.0 — 28 — — — — Silty SAND S-10 35.0 — 44 — — — — SILT S-11 40.0 — 36 — — 82 SILT S-12 45.0 — 36 — — — Silty SAND S-13 50.0 — 75 — — — — SILT G R I Page 2 of 2 BORING AND TEST PIT LOG LEGEND SOIL SYMBOLS SAMPLER SYMBOLS Symbol Typical Description Symbol Sampler Description y 2.0-in. O.D. split-spoon sampler and Standard (,.., LANDSCAPE MATERIALS Penetration Test with recovery(ASTM D1586) FILL ii Shelby tube sampler with recovery (ASTM D1587) GRAVEL;clean to some silt,clay,and sand 3.0-in. O.D. split-spoon sampler with recovery (ASTM D3550) Sandy GRAVEL;clean to some silt and clay N Grab Sample 1g Silty GRAVEL; up to some clay and sand ll Rock core sample interval ra, Clayey GRAVEL; up to some silt and sand Sonic core sample interval SAND;clean to some silt,clay, and gravel Geoprobe sample interval 9 Gravelly SAND; clean to some silt and clay INSTALLATION SYMBOLS :'; Silty SAND; up to some clay and gravel Symbol Symbol Description 11/ Clayey SAND; up to some silt and gravel ill Flush-mount monument set in concrete SILT; up to some clay, sand, and gravel ® Concrete, well casing shown where applicable OHGravelly SILT; up to some clay and sand j'j applicable Bentonite seal,well casing shown where Sandy SILT; up to some clay and gravel = Filter pack, machine-slotted well casing shown •— • where applicable '''' Clayey SILT; up to some sand and gravel Grout,vibrating-wire transducer cable shown where applicable f /' CLAY; up to some silt, sand,and gravel • Vibrating-wire pressure transducer liff! Gravelly CLAY; up to some silt and sand I 'J 1-in.-diameter solid PVC /. Sandy CLAY; up to some silt and gravel U 1-in.-diameter hand-slotted PVC MI Silty CLAY; up to some sand and gravel Grout, inclinometer casing shown where applicable Vtold PEAT FIELD MEASUREMENTS BEDROCK SYMBOLS Symbol Typical Description Symbol Typical Description g Groundwater level during drilling and date m measured +++ BASALT t Groundwater level after drilling and date measured w =1 MUDSTONE �j Rock core recovery(%) -- �/, w - SILTSTONE / Rock quality designation (RQD, °(°) 1- a SANDSTONE Ft 0 z SURFACE MATERIAL SYMBOLS w w Symbol Typical Description 0 t9 ■ Asphalt concrete PAVEMENT P ■ Portland cement concrete PAVEMENT a z• ov l °C° Crushed rock BASE COURSE a 0 m Z 0.1 ♦ BLOWS PER FOOT J CLASSIFICATION OF MATERIAL o o a • MOISTURE CONTENT,% v ¢J w 0 ❑ FINES CONTENT,% J e HLIQUID LIMIT,% COMMENTS AND I- a a 0 PLASTIC LIMIT,% ADDITIONAL TESTS cr 0 o Surface Elevation:Not Available o z a' co m 0 50 100 Asphalt concrete PAVEMENT(6 in.)over crushed j p c` rock BASE COURSE(18 in.) st 1 3 23 —�- 2.0 10 Sandy SILT,trace day,brown mottled rust,medium r 2 e stiff to stiff,fine-to medium-grained sand 5-2 3 _AL ❑ 5• 5- —brown below 5 ft t S-3I 3 A 6.5 5 — (5/30/2017) Groundwater not encountered 10— 15— 20— r 25— N - (6 H 5 — a w — w a 30- w 0 z z a 35— F 0 — J z - 0: O - r7—40 0 0.5 1.0 • Logged By: Drilled by:Western States Soil Conservation,Inc. TORVANE SHEAR STRENGTH,TSF Date Started:5/30/17 Coordinates:Not Available ■ UNDRAINED SHEAR STRENGTH,TSF Drilling Method: Hollow-Stern Auger Hammer Type:Auto Hammer Equipment: CME 55 HT Truck-Mounted Drill Rig Weight:Drop 140 ( _ R A BORING B-1 Hole Diameter. 8 in. Drop:30 in. V Note:See Legend for Explanation of Symbols Energy Ratio:0.79 FEB.2018 JOB NO.5970-F FIG. 1A BLOWS PER FOOT �' CLASSIFICATION OF MATERIAL o a ~ o d z • MOISTURE CONTENT, c, g W o 0 FINES CONTENT,% J _1 _e lli - �r LIQUID LIMIT,% COMMENTS AND w < w 1 ¢ Q o PLASTIC LIMIT,% ADDITIONAL TESTS o 0 Surface Elevation:Not Available o o cn m 0 50 100 Mil Asphalt concrete PAVEMENT(3 in.)over crushed —42-Ti rock BASE COURSE(10 in.) r tt . Sandy SILT,brown,stiff,fine-to medium-grained sand 2 _9 s-1 r 4 A ip i 5 5---..-.. 5.0 1 Silty SAND,brown,loose,fine to medium grained • —interbedded with 6-in.-thick layer of sandy silt at p Y"r 6 ft s 2 Dry Density=90 pcf _ .. 3 8 S3 4 J>i;:: I I 10 ': .'; S-4 6 7 1 Dry Density=94 pcf S-5 2 ' 5 1 . I, I 15-iC....,:...: :: 4 10 16.0 S-61 3 I • Sandy SILT,gray,stiff,fine-to medium grained sand 7 20— —stiff to very stiff,interbedded with 1-to 2-in.-thick layers of silty sand below 20 ft S7 I 4 q 15 ♦ 0 7 T • 'A 25— •,'.. • I'1 a -• S-8 1 o — —trace to some clay below 27 ft I o.i \ s-9 5 is • u • 10 I 1.2 w I 1- 30T:k*.� 30.0 r ' o Silty SAND,brown mottled rust,medium dense,fine s--o 1 10 2 — to medium grained LL 16 w W o 2 a :.:y::.;•:, 7 g 35—•'—' 35.0 10 / P SILT,some clay,trace fine-to medium-grained 31 3 — sand,tan mottled rust,hard,contains weakly to s-11 1 is o _ moderately cemented zones up to 1/4 in.thick J o o m _ c7 i . . °-40 �Y (CONTINUED NEXT PAGE) 0 0.5 1.0 Logged By: Drilled by:Western States Soil Conservation,Inc. ♦ TORVANE SHEAR STRENGTH,TSF Date Started:5/31/17 Coordinates:Not Available • UNDRAINED SHEAR STRENGTH,TSF Drilling Method: Mud Rotary Hammer Type:Auto Hammer /-� Equipment: CME 55 HT Truck-Mounted Drill Rig Weight:140 lb + RI] BORING B-2 Hole Diameter: 5 in. Drop:30 in. V Note:See Legend for Explanation of Symbols Energy Ratio:0.79 FEB.2018 JOB NO.5970-F FIG.2A . . ' co CLASSIFICATION OF MATERIAL Z LU 1- • BLOWS PER FOOT o P. d LA- z e MOISTURE CONTENT,% L —'e_, L tc- z c- 8 a FINES CONTENT,% J til UJ 0 LIQUID LIMIT,% COMMENTS AND o PLASTIC LIMIT,% ADDITIONAL TESTS tn o Surface Elevation:Not Available "Cit gu' el ul co—' 0 50 100 19 , 1141 Silty SAND,tan mottled rust,dense,fine to medium grained,interbedded with 1-in.-thick layers of sandy I. S-12 1 19 25 _.::•::.:4 silt :.::$:,''s.;.:.• 111[ 45—:'....•, : —interbedded with 1-to 3-in.-thick layers of silt 15 45 below 45 ft S-13 20 25 I , :.:...:...:•-: ...,'..-.. 50-4'.•;°— 50.0 Sandy SILT,some clay,tan mottled rust,hard, 17 II 4119' u S-14 — ...-• contains weakly to moderately cemented zones up 32 ---- to 1/4 in.thick,interbedded with 1-to 2-in.-thick i 51'5 layers of silty sand (5/31/2017) 55— . . 60— , . 65— 'A 75— _ 1- .. 0 o - ' V! .:1 o 1-- a 0 Fc — › — LLI 4, _ o z & z 2 IE — _ o z re 0 es Fe (D 80 0 0.5 1.0 • TORVANE SHEAR STRENGTH,TSF IN UNDRAINED SHEAR STRENGTH,TSF G RA BORING B-2 FEB.2018 JOB NO.5970-F FIG.2A l z ,L • BLOWS PER FOOT ' CLASSIFICATION OF MATERIAL o a ~ 0 o z • MOISTURE CONTENT,% L Li I—u- g w w 0 0 FINES CONTENT,% = x = J J J 3 LIQUID LIMIT,% COMMENTS AND a.w < a- (17 m m 0 PLASTIC LIMIT,% ADDITIONAL TESTS o co Surface Elevation:Not Available o z vg m o 50 100 Asphalt concrete PAVEMENT(2 in.)over crushed i 0.3 —c•?-; 1 rock BASE COURSE(2 in.) J —,` Silty SAND,brown,loose,fine to medium grained 3 5 -::'; S1 1 2 • 0 n 3 5—;', —interbedded with 18-in.thick layer of sandy silt at I — • 5ft 3-2 } 2 6 Dry Density=83 pcf S-3 i 3 ` 2 I 10—:`, --medium dense below 10 ft I 2 11 ` S4 3 ■1 . 6 . 6 _14 4 55 2 V 7 r 15- ?,' 4k Dry Density=91 pcf S6 — —contains organics at 17 ft,brown mottled rust ft6 n below 17 S-7 i 6 11 i 20=` :, —loose,interbedded with 1-to 2-in.thick layers of ) 6 -9 sandy silt below 20 ft S8 6 4 • 3 I I I 1 I m 25—' -' 25.0 . Sandy SILT,some clay,gray,stiff,fine-to ss I 4 9-4 • medium-grained sand 5 0 ii _ w a 30 —clayey,medium stiff to stiff at 30 ft,sane sand r/ below 30 ft s-to 3 _s , 1 • ET ED —interbedded with 6-in.-thick layer of silty day at 5 i w - 30.5ft I u, W _ ED _ 1 0 35— —very stiff at 35 ft,tan mottled rust,contains weakly I 7 3 to moderately cemented zones up to 1/4 in.thick 5-11 to �L23 iii 1, below 35 ft 13 1 co o — _, Za — 1 K $0 _ 1 E 1 o-40 ' (CONTINUED NEXT PAGE) 0 0.5 1.0 Logged By: Drilled by:Western States Soil Conservation,Inc. ♦ TORVANE SHEAR STRENGTH,TSF Date Started:5/30/17 Coordinates:Not Available ■ UNDRAINED SHEAR STRENGTH,TSF Drilling Method: Mud Rotary Hammer Type:Auto Hammer 7� Equipment: CME 55 HT Truck-Mounted Drill Rig Weight:140 lb �/ BORING B-3 Hole Diameter: 5 in. Drop:30 in. 1 Note:See Legend for Explanation of Symbols Energy Ratio:0.79 FEB.2018 JOB NO.5970-F FIG.3A ♦ BLOWS PER FOOT Ei CLASSIFICATION OF MATERIAL 0 •o Y •? • MOISTURE CONTENT,% �` 5 w w 0 0 FINES CONTENT,% J _1 1--1-LIQUID LIMIT,% COMMENTS AND w < w 0 PLASTIC LIMIT,% ADDITIONAL TESTS 0 0Surface Elevation:Not Available o z 0 "' 00 0 50 100 SILT,some day to dayey,trace fine-to s12 s _13 medium-grained sand,tan mottled rust,stiff, 8 _ contains weakly to moderately cemented zones up to 1/4 in.Ihidc I 1 — J 45 :f Silty SAND,red-brown,dense,fine to medium 4s.0 11 3 r S13 1 14 grained t9 Ile 1 r 50—;:if: 1 14 2 S-14 1 15 7 .--::;2.-: t l <: 1 55y1: 55.0 l- SILT,some clay to clayey,some fine-to 5 _16 — medium-grained sand,gray,very stiff s15 y r I . . I — 60—— —sandy,interbedded with 1-to 2-in.-thick layers of 8 97 I I silty sand blow 60 ft s ,1s 11 j] 16 . — _ I I 65— I F O LLl w F a 70— 9 24 E - si7 110 4 i O L 14 > 1 w Ll J - 0 2 CD' I O 75— \I —. I 0 ca 2 o 0 m I C'-80 (CONTINUED NEXT PAGE) 0 0.5 1.0 • TORVANE SHEAR STRENGTH,TSF IN UNDRAINED SHEAR STRENGTH,TSF G R I BORING B-3 FEB.2018 JOB NO.5970-F FIG.3A w ♦ BLOWS PER FOOT o CLASSIFICATION OF MATERIAL o oz • MOISTURE CONTENT,% O FINES CONTENT,% S - !• •=4-LIQUID LIMIT,% COMMENTS AND a 0 PLASTIC LIMIT,% ADDITIONAL TESTS o c7 Surface Elevation:Not Available o z m "' m 50 100 Silty SAND,gray,dense,fine to coarse grained Si8 15 p111130 (5/30/2017) 81.5 85- 90- 95- 100- 105- ry 0 ui 5 - w - ✓ 0 11a— a — J W _ O Z V 2 0 115— P 5 0 J 0 Z_ 0 - 0—120 I 0 0.5 1.0 • TORVANE SHEAR STRENGTH,TSF ■ UNDRAINED SHEAR STRENGTH,TSF G RD BORING B-3 FEB.2018 JOB NO.5970-F FIG.3A . . ' o CLASSIFICATION OF MATERIAL = UJ 1--- A BLOWS PER FOOT o o ,...; CL m • MOISTURE CONTENT,% '2' a 0 FINES CONTENT,% La W « J 1--1—LIQUID LIMIT,% PLASTIC LIMIT,% COMMENTS AND o o ADDITIONAL TESTS Surface Elevation:Not Available o F.. cn cn co 0 50 100 D`j‘ Crushed rock SURFACING(12 in.) —t:$•a'.;..-'..•:. Silty SAND,brown,loose,fine to medium grained to P1 2 5 — ..•'.:: S1 1 2 ilk 110 3 --,i..:... __H 5—t',•. —very loose at 5 ft I f: •',...f 2 3 S 2 I 2 al 1 i ...,. ... ,. —;'...--: • S-3 t Dry Density=81 pcf ..,.... 5-4 34 ?A17. --II 5 , 1 -7....: . 1 t -ff. •:i.:-. i IP* Dry Density=91 pcf 15--'.: .; —gray-brown,medium dense at 15 ft 2 17 s-6 170 1 . —: :.-- ::.:-.....-!':" ::........4'.. —•::P: 20—:: 61 —gray-brown,medium dense below 25 ft 6 \i, s-8 8 IP it 0 ........% i 2 . 1 111 -'...• : ::I- • l• rS. 30---,::: —gray below 30 ft • 1 I .. .:.. S-9 1 5 E a 14 11J _..1 a _ 0 •DC-..-i..i.' 35 0 _i Jo SILT,some day to clayey,some fine-to 14 27 \., 0— medium-grained sand,tan mottled rust,very stiff, s-to 1 ii — I o contains weakly to moderately cemented zones up 16 0 to 1/4 in.thick 1 m — _ - 1 o , ' i E I °-40 (CONTINUED NEXT PAGE) 0 0.5 1.0 • TORVANE SHEAR STRENGTH,TSF Logged By: Drilled by:Western States Soil Conservation,Inc. M UNDRAINED SHEAR STRENGTH,TSF Date Started:5/31/17 Coordinates:Not Available Drilling Method: Mud Rotary Hammer Type:Auto Hammer Equipment: CME 55 HT Truck-Mounted Drill Rig Weight:140 lb BORING B-4 Hole Diameter: 5 in. Drop:30 in. Note:See Legend for Explanation of Symbols Energy Ratio:0.79 FEB.2018 JOB NO.5970-F FIG.4A w A BLOWS PER FOOT CLASSIFICATION OF MATERIAL o a O o • MOISTURE CONTENT,% w U ❑ FINES CONTENT,% a a F 2 2 o �{�LIQUID D LIMIT,%o/a COMMENTS AND w w m < ¢ ADDITIONAL TESTS o Surface Elevation:Not Available o z cn m m 0 50 100 SILT,some day to clayey,some fine-to t 6 29 I medium-grained sand,tan mottled rust,very stiff, S- 1 10 • contains weakly to moderately cemented zones up to 114 in.thick,interbedded with 1-to 2-in.-thick { t layers of silty sand l l - I 45.0 11 Silty SAND,red-brown,dense,fine to medium T 17 33 S-12 1 1s grained 17 • \ 1 " \ 50.0 8 19 SILT,some clay to clayey,some fine-to S 13 7 . medium-grained sand,tan mottled rust,very stiff L 12 (5/31/2017) 51.5 55- 60- 65— N - f U W H 5 a w a 70— it w k7 Z 0 75— c� re o J Z — 0-80 0 0.5 1.0 • TORVANE SHEAR STRENGTH,TSF ■ UNDRAINED SHEAR STRENGTH,TSF G RD BORING B-4 FEB.2018 JOB NO.5970-F FIG.4A SPT N60 SBT Tip Resistance(CH) Sleeve Friction(Fs) F.Ratio Pore Pressure(U2) (UNITLESS) (UNITLESS) (tsf) (tsf) (%) (psi) D 0 250 0 12 0 300 0 18 0 12 S0 450 I I I I I I I I I I I I 1 1IF 1 I I I 1 i I I I I I I I I 10 5 20 I 30 - Depth (ft) ' I 50 -- - - - -'IL- I 11> 70 - 80 - -yel - M 1 , r .. , TOTAL DEPTH: 89.239 ft Ill 1 sensitive fine grained M 4 silty clay to clay 7 silty sand to sandy silt E 10 gravelly sand to sand ▪2 organic material II 5 clayey silt to silty clay 8 sand to silty sand 1' 11 very stiff fine grained(*) ■3 clay M 6 sandy silt to clayey silt 1 9 sand ®12 sand to clayey sand(*) G R 0 Observed By: N.Hatch Advanced By:Oregon Geotechnical Exploration,Inc. Date Started: 06/1/17 Ground Surface Elevation: NotAvailable Coordinates: Not Available CONE PENETRATION TEST CPT-1 FEB.2018 JOB NO. 5970-F FIG. 5A SPT N60 SBT Seismic Velocity Tip Resistance(Qt) (UNITLESS) (UNITLESS) (ft/s) (lsf) 0 0 250 0 12 0 3000 0 300 I I I I I I 3 861 626 10 — 608 — 597 1728 - 20 — 612 — 665 888 30 — 628 ..Z. — 591 - - a7�ae — 2448 40 ' 2099 Depth �i3a (ft) z7z 50 .A____ _� `_ ... c" 1 — 1398 60 831 70 — _ 1377 I L 1473 ( 80 — — \ 27e- 4 —2 90 TOTAL DEPTH:89.239 ft NI 1 sensitive fine grained El 4 silty clay to clay B 7 silty sand to sandy silt a10 gravelly sand to sand 112 organic material In 5 clayey silt to silty clay •'.8 sand to silty sand '11 very stiff fine grained(*) 113 clay 6 sandy silt to clayey silt 9 sand 12 sand to clayey sand(*) GRA Observed By: N.Hatch Advanced By:Oregon Geotechnical Exploration,Inc. Date Started: 06/1/17 Ground Surface Elevation: NotAvailable CONE PENETRATION TEST CPT-1 Coordinates: Not Available (SEISMIC VELOCITY PROFILE) FEB.2018 JOB NO. 5970-F FIG. 6A • • Po, P1 ,tsf ID Kp M,tsf 0 5 10 15 20 25 0.1 1 10 0 10 20 30 40 50 0 500 1,000 1,500 2,000 2,500 ° 1 1 1 I 1 I 1 I I I 1 I 1 I 1 1 1 1 I 1 1 1 1 1 1 I 111111 I 1 IIIIII IIII 1 F 111111IIIflhI1 III I 1 I 1 1 I 1 1 1 1 1 I 1 1 I 5 10 15 ♦ Po r - • P1 - d 20 25 30 r 35 - G RI DILATOMETER SOUNDING DMT-1 FEB.2018 JOB NO. 5970-F FIG. 7A Po, P1 ,tsf ID K0 M,tsf 0 5 10 15 20 25 0.1 1 10 100 0 10 20 30 40 0 500 1,000 1,500 2,000 2,500 0 Th111111111111111111111 111111111 111111111 11111111 1111 1 I I I I 1 I I I I III III III 1 I I I I I I Iv_— 5 Po /1...... , 10 15 x — - C20 r — Al v — C _ f 25 —-• 30 4CA y — 4 s'ill i . 35 40 - - GRp DILATOMETER SOUNDING DMT-2 FEB.2018 JOB NO. 5970-F FIG. 8A 1 GROUP UNIFIED SOIL CLASSIFICATION GROUP UNIFIED SOIL CLASSIFICATION SYMBOL FINE-GRAINED SOIL GROUPS SYMBOL FINE-GRAINED SOIL GROUPS ORGANIC SILTS AND ORGANIC SILTY ORGANIC CLAYS OF MEDIUM TO HIGH OL CLAYS OF LOW PLASTICITY OH PLASTICITY,ORGANIC SILTS INORGANIC CLAYEY SILTS TO VERY FINE ML SANDS OF SLIGHT PLASTICITY MH INORGANIC SILTS AND CLAYEY SILT INORGANIC CLAYS OF LOW TO MEDIUM CL PLASTICITY CH INORGANIC CLAYS OF HIGH PLASTICITY 60 50 CH 40 X W 30 CL U • 20 MH or OH 10 CL-ML ML or OL 0 I 0 10 20 30 40 50 60 70 80 90 100 LIQUID LIMIT, % Location Sample Depth,ft Classification LL PL PI MC,% • B-3 S-10 30.8 Silty CLAY, some fine-to medium-grained sand, 41 19 22 33 F 0 Li w a Ui w 0 0 e W L7 a C W O a GR U F PLASTICITY CHART U 0 W FEB. 2018 JOB NO.5970-F FIG. 9A • 5 10 z 15 20 25 0.01 0.1 1 10 100 STRESS, TSF Indial Location Sample Depth,ft Classification '4,pcf MC, % • B-2 S-8 25.3 Sandy SILT, gray,fine-to medium-grained sand 96 31 a a a GRIt4 E2 CONSOLIDATION TEST 0 FEB.2018 JOB NO. 5970-F FIG. 10A APPENDIX B Site-Specific Seismic-Hazard Evaluation APPENDIX B SITE-SPECIFIC SEISMIC-HAZARD STUDY GENERAL GRI completed a site-specific seismic-hazard study for the proposed improvements to Tigard High School in Tigard, Oregon. The purpose of our study was to evaluate the potential seismic hazards associated with regional and local seismicity. The site-specific seismic-hazard evaluation is intended to meet the requirements of the 2014 Oregon Structural Specialty Code (OSSC), which is based on the 2012 International Building Code (IBC). Seismic design in accordance with the 2012 IBC is based on American Society of Civil Engineers (ASCE) document 7-10, titled "Minimum Design Loads for Buildings and Other Structures." Our work was based on the potential for regional and local seismic activity, as described in the existing scientific literature, and the subsurface conditions at the site, as disclosed by the subsurface explorations completed for this project. Specifically, our work included the following tasks: 1) A detailed review of available literature, including published papers, maps, open-file reports, seismic histories and catalogs, works in progress, and other sources of information regarding the tectonic setting, regional and local geology, and historical seismic activity that might have a significant effect on the site. 2) Compilation and evaluation of subsurface data collected at and in the vicinity of the site, including classification and laboratory analyses of soil samples, and review of shear-wave velocity surveys completed at the site. This information was used to prepare a generalized subsurface profile for the site. 3) Identification of the potential seismic events (earthquakes) appropriate for the site and characterization of those events in terms of a generalized design event. 4) Office studies based on the generalized subsurface profile and the generalized design earthquake resulting in conclusions and recommendations concerning the following: a) specific seismic events that might have a significant effect on the site, b) the potential for seismic energy amplification and liquefaction or soil-strength loss at the site,and c) site-specific acceleration response spectra for design of the proposed structure. This appendix describes the work accomplished and summarizes our conclusions and recommendations. Geologic Setting On a regional scale, the site is located at the northern end of the Willamette Valley, a broad, gently deformed, north-south-trending, topographic feature separating the Coast Range to the west from the Cascade Mountains to the east. The site is located approximately 74 km inland from the Cascadia Subduction Zone (CSZ), an active plate boundary along which remnants of the Farallon plate (the Gorda, Juan de Fuca, and Explorer plates) are being subducted beneath the western edge of the North American 110 B-1 plate. The subduction zone is a broad, eastward-dipping zone of contact between the upper portion of the subducting slabs of the Gorda,Juan de Fuca, and Explorer plates and the overriding North American plate, as shown on the Tectonic Setting Summary, Figure 1 B. On a local scale, the site is located in the Tualatin Basin, a large, well- defined, southeast-trending structural basin bounded by high-angle, northwest-trending, right-lateral, strike-slip faults considered to be seismogenic. The geologic units in the area are shown on the Regional Geologic Map, Figure 2B. The distribution of nearby Quaternary faults is shown on the Local Fault Map, Figure 3B. Information regarding the continuity and potential activity of these faults is lacking due largely to the scale at which geologic mapping in the area has been conducted and the presence of thick, relatively young, basin-filling sediments that obscure underlying structural features. Other faults may be present within the basin, but clear stratigraphic and/or geophysical evidence regarding their location and extent is not presently available. Additional discussion regarding crustal faults is provided in the Local Crustal Event section below. Because of the proximity of the site to the CSZ and its location within the Tualatin Basin, three distinctly different sources of seismic activity contribute to the potential for the occurrence of damaging earthquakes. Each of these sources is generally considered to be capable of producing damaging earthquakes. Two of these sources are associated with the deep-seated tectonic activity related to the subduction zone; the third is associated with movement on the local, relatively shallow structures within and adjacent to the Tualatin Basin. Subsurface and Geologic Conditions. Published geologic mapping indicates the site is mantled with Missoula flood deposits, locally referred to in the project area as the Willamette Silt Formation (Madin, 1990). In general, Willamette Silt is composed of unconsolidated beds and lenses of silt and sand. Stratification within this formation commonly consists of 4- to 6-in.-thick beds, although in some areas, the silt and sand are massive and the bedding is indistinct or nonexistent. Seismicity General. The geologic and seismologic information available for identifying the potential seismicity at the site is incomplete, and large uncertainties are associated with estimates of the probable magnitude, location, and frequency of occurrence of earthquakes that might affect the site. The available information indicates the potential seismic sources that may affect the site can be grouped into three independent categories: subduction zone events related to sudden slip between the upper surface of the Juan de Fuca plate and the lower surface of the North American plate, subcrustal events related to deformation and volume changes within the subducted mass of the Juan de Fuca plate, and local crustal events associated with movement on shallow, local faults within and adjacent to the Tualatin Basin. Based on our review of currently available information, we have developed generalized design earthquakes for each of these categories. The design earthquakes are characterized by three important properties: size, location relative to the subject site, and the peak horizontal bedrock accelerations produced by the event. In this study, earthquake size is expressed by the moment magnitude (Mw); location is expressed as the closest distance to the fault rupture, measured in kilometers; and peak horizontal bedrock accelerations are expressed in units of gravity (1 g = 32.2 ft/sec2 = 981 cm/sec2). G R U B-2 Subduction Zone Event. The last interplate earthquake on the CSZ occurred in January 1700. Geological studies show that great megathrust earthquakes have occurred repeatedly in the past 7,000 years (Atwater et al., 1995; Clague, 1997; Goldfinger et al., 2003; and Kelsey et al., 2005), and geodetic studies (Hyndman and Wang, 1995; Savage et al., 2000) indicate rate of strain accumulation consistent with the assumption that the CSZ is locked beneath offshore northern California, Oregon, Washington, and southern British Columbia (Fluck et al., 1997; Wang et al., 2001). Numerous geological and geophysical studies suggest the CSZ may be segmented (Hughes and Carr, 1980; Weaver and Michaelson, 1985; Guffanti and Weaver, 1988; Goldfinger, 1994; Kelsey and Bockheim, 1994; Mitchell et al., 1994; Personius, 1995; Nelson and Personius, 1996; Witter, 1999), but the most recent studies suggest that for the last great earthquake in 1700, most of the subduction zone ruptured in a single Mw 9.0 earthquake (Satake et al., 1996; Atwater and Hemphill-Haley, 1997; Clague et al., 2000). Published estimates of the probable maximum size of subduction zone events range from moment magnitude Mw 8.3 to >9.0. Numerous detailed studies of coastal subsidence, tsunamis, and turbidites yield a wide range of recurrence intervals, but the most complete records (>4,000 years) indicate average intervals of 350 to 600 years between great earthquakes on the CSZ (Adams, 1990; Atwater and Hemphill-Haley, 1997; Witter, 1999; Clague et al., 2000; Kelsey et al., 2002; Kelsey et al., 2005; Witter et al., 2003). Tsunami inundation in buried marshes along the Washington and Oregon coast and stratigraphic evidence from the Cascadia margin support these recurrence intervals (Kelsey et al., 2005; Goldfinger et al., 2003). The U.S. Geological Survey (USGS) probabilistic analysis assumes four potential locations for the location of the eastern edge of the earthquake rupture zone, shown on Figure 4B. The 2008 USGS mapping effort indicates three rupture scenarios are assumed to represent these interface events: 1) Mw 9.0±0.2 events that rupture the entire CSZ every 500 years, 2) Mw 8.3 to 8.7 events with rupture zones that occur on segments of the CSZ and occur over the entire length of the CSZ during a period of about 500 years, and 3) Mw 8.0 to 8.2 events that rupture only segments of the CSZ every 500 years (Petersen et al., 2008). The assumed distribution of earthquakes is shown on the Assumed Magnitude-Frequency Distribution, Figure 5B. This distribution assumes the larger Mw 9.0 earthquakes likely occur more often than each of the smaller segmented ruptures, as also indicated by the USGS deaggregation for the site. Therefore, for our deterministic analysis, we have chosen to represent the subduction zone event by a design earthquake of Mw 9.0 at a focal depth of 30 km and rupture distance of 74 km. This corresponds to a sudden rupture of the whole length of the Juan de Fuca-North American plate interface with an assumed rupture zone due west of the site. Based on an average of the attenuation relationships published by Atkinson and Boore (2003), Atkinson and Macias (2009), Zhao et al. (2006), and Abrahamson et al. (2015), a subduction zone earthquake of this size and location would result in a peak horizontal bedrock acceleration of approximately 0.21 g at the site. Subcrustal Event. There is no historical record of significant subcrustal, intraslab earthquakes in Oregon. Although both the Puget Sound and northern California region have experienced many of these earthquakes in historical times, Wong (2005) hypothesizes that due to subduction zone geometry, geophysical conditions, and local geology, Oregon may not be subject to intraslab earthquakes. In the Puget Sound area, these moderate to large earthquakes are deep (40 to 60 km) and over 200 km from the deformation front of the subduction zone. Offshore along the northern California coast, the earthquakes are shallower (up to 40 km) and located along the deformation front. Estimates of the probable size, location, and frequency of subcrustal events in Oregon are generally based on comparisons of the CSZ with active convergent plate margins in other parts of the world and on the historical seismic record for the G ` `O B-3 region surrounding Puget Sound, where significant events known to have occurred within the subducting Juan de Fuca plate have been recorded. Published estimates of the probable maximum size of these events range from moment magnitude Mw 7.0 to 7.5. The 1949, 1965, and 2001 documented subcrustal earthquakes in the Puget Sound area correspond to Mw 7.1, 6.5, and 6.8, respectively. Published information regarding the location and geometry of the subducting zone indicates a focal depth of 50 km is probable (Weaver and Shedlock, 1989). We have chosen to represent the subcrustal event by a design earthquake of magnitude Mw 7.0 at a focal depth of 50 km and a rupture distance of 63 km. Based on the attenuation relationships published by Youngs et al. (1997) and Abrahamson et al. (2015), a subcrustal earthquake of this size and location would result in a peak horizontal bedrock acceleration of approximately 0.17 g at the site. Local Crustal Event. Sudden crustal movements along relatively shallow, local faults in the Portland area, although rare, have been responsible for local crustal earthquakes. The precise relationship between specific earthquakes and individual faults is not well understood since few of the faults in the area are expressed at the ground surface and the foci of the observed earthquakes have not been located with precision. The history of local seismic activity is commonly used as a basis for determining the size and frequency to be expected of local crustal events. Although the historical record of local earthquakes is relatively short (the earliest reported seismic event in the area occurred in 1920), it can serve as a guide for estimating the potential for seismic activity in the area. Based on fault mapping conducted by the USGS (Personius et al., 2003), the inferred location of the Canby-Mollala Fault borders the northeastern corner of the site. However, the USGS does not consider the Canby-Mollala Fault to be an active, contributing source in their Probabilistic Seismic Hazard Analysis (PSHA). Based on our review of the USGS deaggregations for the site (U.S. Geological Survey, 2014), the Portland Hills Fault is the closest crustal fault contributing to the overall seismic hazard at the site. The inferred location of the Portland Hills Fault is approximately 12 km from the site, and the fault has a characteristic earthquake magnitude of Mw = 7. A crustal earthquake of this size would result in a peak horizontal bedrock acceleration of approximately 0.45 g at the site based on an average of the next generation attenuation (NGA) ground-motion relations published by Boore et al. (2014), Campbell and Bozorgnia (2014), and Chiou and Youngs (2014). Summary of Deterministic Earthquake Parameters In summary, three distinctly different types of earthquakes affect seismicity in the project area. Deterministic evaluation of the earthquake sources using recently published attenuation ground-motion relations provides estimates of ground response for each individual earthquake type. Unlike probabilistic estimates, these deterministic estimates are not associated with a relative hazard level or probability of occurrence and simply provide an estimate of the ground-motion parameters for each type of fault at a given distance from the site. The basic parameters of each type of earthquake are as follows: Average Earthquake Attenuation Relationships Rupture Focal Peak Bedrock Peak Bedrock Source for Deterministic Spectra Magnitude,Mw Distance,km Depth,km Acceleration,g Acceleration,g Subduction Zone Atkinson and Boore,2003 9.0 74.0 30 0.14 Atkinson and Macias,2009 9.0 74.0 30 0.18 Zhao et al.,2006 9.0 74.0 30 0.23 0.21 Abrahamson et al.,2015 9.0 74.0 30 0.29 G R D B-4 Average Earthquake Attenuation Relationships Rupture Focal Peak Bedrock Peak Bedrock Source for Deterministic Spectra Magnitude,Mw Distance,km Depth,km Acceleration,g Acceleration,g Subcrustal Youngs et al., 1997 7.0 63.0 50 0.10 Abrahamson et al.,2015 7.0 63.0 50 0.24 0' Local Crustal Campbell and Bozorgnia,2014 7.0 12.0 NA 0.54 Chiou and Youngs,2014 7.0 12.0 NA 0.40 0.45 Boore et al.(2014) 7.0 12.0 NA 0.42 Probabilistic Considerations The probability of an earthquake of a specific magnitude occurring at a given location is commonly expressed by its return period, i.e., the average length of time between successive occurrences of an earthquake of that size or larger at that location. The return period of a design earthquake is calculated once a project design life and some measure of the acceptable risk that the design earthquake might occur or be exceeded are specified. These expected earthquake recurrences are expressed as a probability of exceedance during a given time period or design life. Historically, building codes have adopted an acceptable risk level by identifying ground-acceleration values that meet or exceed a 10% probability of exceedance in 50 years, which corresponds to an earthquake with an expected recurrence interval of 475 years. Previous versions of the IBC developed response spectra based on ground motions associated with the Maximum Considered Earthquake (MCE), which is generally defined as a probabilistic earthquake with a 2% probability of exceedance in 50 years (return period of about 2,500 years) except where subject to deterministic limitations (Leyendecker and Frankel, 2000). The recent 2012 IBC develops response spectra using a Risk-Targeted Maximum Considered Earthquake (MCER), which is defined as the response spectrum expected to achieve a 1% probability of building collapse within a 50-year period. The design-level response spectrum is calculated as two-thirds of the MCER ground motions. Since the MCER earthquake ground motions were developed by the USGS to incorporate the targeted 1% in 50 years risk of structural collapse based on a generic structural fragility, they are different than the ground motions associated with the traditional MCE. Although site response is evaluated based on the MCER, it should be noted that seismic hazards, such as liquefaction and soil strength loss, are evaluated using the Maximum Considered Earthquake Geometric Mean (MCEG) peak ground acceleration (PGA), which is more consistent with the traditional MCE. The 2012 IBC design methodology uses two mapped spectral acceleration parameters, Ss and Si, corresponding to periods of 0.2 and 1.0 sec to develop the MCER earthquake. The Ss and Si parameters for the site located at the approximate latitude and longitude coordinates of 45.4032°N and 122.7691°W are 0.96 and 0.42 g,respectively. Estimated Site Response The effect of a specific seismic event on the site is related to the type and quantity of seismic energy delivered to the bedrock beneath the site by the earthquake and the type and thickness of soil overlying the bedrock at the site. A ground-motion hazard analysis was completed to estimate this site-specific behavior in accordance with Section 21.2 of ASCE 7-10. The ground-motion hazard analysis consisted of four significant components: 1) estimation of bedrock response using recently developed attenuation relationships (deterministic evaluation), 2) estimation of bedrock response using the 2014 USGS-based PSHA (probabilistic evaluation), 3) comparison of the deterministic and probabilistic bedrock-response G ` 'O B-5 • spectra to determine the controlling spectrum, and 4) development of recommended response spectra for the four hazard levels. The following paragraphs describe the details of the ground-motion hazard analysis. To estimate the deterministic bedrock-response spectrum, recently developed attenuation relationships were used to evaluate bedrock ground motions at the site. Based on our review of the USGS deaggregations (U.S. Geological Survey, 2014), crustal seismicity and an event on the CSZ represent the largest contributing sources to the seismic hazard at the site. Considering this, we have chosen to estimate the deterministic bedrock response using 84th-percentile ground motions from the following two earthquake scenarios: 1) a Mw 7.0 crustal earthquake at a distance of 10 km from the site and 2) a Mw 9.0 subduction zone earthquake at a distance of 74 km from the site. The same attenuation relationships outlined in the deterministic section were used to evaluate the crustal and subduction earthquake response. The resulting deterministic bedrock-response spectra are shown on Figure 6B and indicate crustal seismicity controls the hazard at the site. The deterministic MCER bedrock spectrum is taken as the larger of the 84th-percentile ground motions and the deterministic lower limit. The probabilistic bedrock- response spectrum was acquired through the use of the USGS Interactive Deaggregation (U.S. Geological Survey, 2014). The deaggregation was evaluated for a 2% in 50 years probability over a period range of PGA to 5 sec. In accordance with Section 21.2 of ASCE 7-10, the site-specific bedrock MCER response spectrum is taken as the lesser of the probabilistic and deterministic MCER bedrock motions. Figure 6B demonstrates the probabilistic bedrock spectrum is the lesser of the spectra. The site is classified as Site Class D, or a stiff-soil site, based on the average Vs3o of 1,050 ft/sec in accordance with Section 20.3 of ASCE 7-10. Corresponding short- and long-period adjustment factors Fa and Fy of 1.12 and 1.58, respectively, were used to develop the probabilistic Site Class D MCER response spectrum. We recommend using the Site Class D MCER and design response spectra shown on Figure 7B for design of the new structure. Seismic Hazards Liquefaction/Cyclic Softening. The results of our evaluation indicate there is a potential that zones of the interbedded silt and silty sand deposit below the groundwater surface at the site could lose strength or liquefy during a code-based earthquake. Based on our analysis, potentially liquefiable soils are present at a depth of 10 ft below the ground surface and extend to a depth of about 35 ft. Our analysis indicates the potential for 1 to 2 in. of seismically induced settlement, which may occur during the earthquake and after earthquake shaking has ceased. Other Hazards. Based on site topography, the risk of earthquake-induced slope instability and/or lateral spreading is low. The risk of damage by tsunami and/or seiche at the site is absent. The inferred location of the Canby-Mollala Fault borders the northeastern corner of the site (Personius et al., 2003); however,the USGS does not consider the Canby-Mollala Fault to be an active, contributing source in their PSHA. The USGS considers the Portland Hills Fault, located about 12 km northeast of the site, to be the closest crustal fault source contributing to the overall seismic hazard at the site. Unless occurring on a previously unmapped or unknown fault,the risk of fault rupture at the site is low. CIRpB-6 . • CONCLUSIONS The 2012 IBC design methodology uses two spectral response parameters, Ss and S,, corresponding to periods of 0.2 and 1.0 sec to develop the MCER response spectrum. The Ss and Si parameters for the site are 0.96 and 0.42 g, respectively. The results of the ground-motion hazard analysis indicate the 2012 IBC Site Class D spectrum provides an appropriate estimate of the spectral accelerations at the site. We recommend use of the Site Class D design spectrum shown on Figure 7B for design of the new structure at the site. References Abrahamson, N.A., Gregor, N., and Addo, K., 2015, BC hydro ground motion prediction equations for subduction earthquakes, Earthquake Spectra In-Press. Adams, J., 1990, Paleoseismicity of the Cascadia subduction zone: Evidence from turbidites off the Oregon-Washington margin: Tectonics,vol.9,no.4,pp.569-583. Atkinson, G.M.,and Boore, D.M.,2003, Empirical ground motion relations for subduction zone earthquakes and their application to Cascadia and other regions:Seismological Research Letters,vol.93,no.4,pp. 1703-1729. Atkinson,G.M.,and Macias,M.,2009, Predicted ground motions for great interface earthquakes in the Cascadia Subduction Zone: Bulletin of the Seismological Society of America,vol.99,no.3,pp. 1552-1578. Atwater, B.F., and Hemphill-Haley, E., 1997, Recurrence intervals for great earthquakes of the past 3,500 years at northeastern Willapa Bay,Washington: U.S.Geological Survey Professional Paper 1576,p. 108. Atwater, B.F., Nelson, A.R., Clague, J.J., Carver, G.A., Yamaguchi, D.K., Bobrowsky, P.T., Bourgeois, J., Darienzo, M.E., Grant, W.C., Hemphill-Haley, E., Kelsey, H.M.,Jacoby, G.C., Nishenko,S.P., Palmer, S.P., Peterson, C.D.,and Reinhart,M.A., 1995, Summary of coastal geologic evidence for past great earthquakes at the Cascadia subduction zone: Earthquake Spectra. 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Hyndman, R.D.,and Wang, K., 1995,The rupture zone of Cascadia great earthquakes from current deformation and the thermal regime:Journal of Geophysical Research,vol. 100,no.B11,pp.22133-22154. Kelsey, H.M., and Bockheim, J.G., 1994, Coastal landscape evolution as a function of eustasy and surface uplift rate, Cascadia margin,southern Oregon:Geological Society of America Bulletin,vol. 106,pp.840-854. Kelsey, H.M., Nelson,A.R., Hemphill-Haley, E.,and Witter, R.C., 2005,Tsunami history of an Oregon coastal lake reveals a 4600 yr record of great earthquakes on the Cascadia subduction zone:GSA Bulletin,vol. 117,pp. 1009-1032. G A B-7 Kelsey, H.M.,Witter, R.C.,and Hemphill-Haley, E.,2002, Pl.-boundary earthquakes and tsunamis of the past 5500 yr,Sixes River estuary,southern Oregon:Geological Society of America Bulletin,vol. 114,no.3,pp.298-314. Leyendecker, E.V., and Frankel, A.D., 2000, Development of maximum considered earthquake ground motion maps: Earthquake Spectra,vol. 16,no. 1. Madin, I.P.,1990, Earthquake-hazard geology maps of the Portland metropolitan area:Oregon Department of Geology and Mineral Studies Open-File Report 90-02. Mitchell, C.E., Vincent, P., Weldon, R.J. III, and Richards, M.A., 1994, Present-day vertical deformation of the Cascadia margin, Pacific Northwest,United States:Journal of Geophysical Research,vol.99,no.B6,pp.12257-12277. Nelson, A.R., and Personius, S.F., 1996, Great-earthquake potential in Oregon and Washington-An overview of recent coastal geologic studies and their bearing on segmentation of Holocene ruptures, central Cascadia subduction zone, in Rogers, A.M., Walsh,T.J., Kockelman,W.J., and Priest,G.R., eds.,Assessing earthquake hazards and reducing risk in the Pacific Northwest:U.S.Geological Survey Professional Paper 1560,vol.1,pp.91-114. Personius,S.F., 1995, Late Quaternary stream incision and uplift in the forearc of the Cascadia subduction zone,western Oregon: Journal of Geophysical Research,vol.100,no. 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U.S.Geological Survey,2014, Unified hazard tool lookup by latitude,longitude,accessed 04/11/17 from USGS website: https://earthquake.usg,s.gov/hazards/interactive./. Wang, Y., He, J., Dragert, H., and James, T.S., 2001, Three-dimensional viscoelastic interseismic deformation model for the Cascadia subduction zone:Earth,Planets and Space,vol.53,pp.295-306. Weaver,C.S.,and Michaelson,C.A., 1985,Seismicity and volcanism in the Pacific Northwest-Evidence for the segmentation of the Juan de Fuca Pl.:Geophysical Research Letters,vol. 12,no.4,pp.215-218. Weaver, C.S., and Shedlock, K.M., 1989, Potential subduction, probable intraplate and known crustal earthquake source areas in the Cascadia Subduction Zone:U.S.Geological Survey Open-File Report 89-465,pp. 11-26. Witter, R.C., 1999, Late Holocene Paleoseismicity, tsunamis and relative sea-level changes along the south-central Cascadia subduction zone,southern Oregon:University of Oregon, unpublished Ph.D dissertation,pp.178. 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G B-8 • c . 1 Quo! 130°� �1' 126° 122°W � PORTLAND -` - 1 /Ol K, ,a •,.. 52°N Coastal Cascades NORTH AMERICA Coastline dGr `O Q PLATE I Range I " - ‘11-.0 . British Columbia 1 1 t ' r I o4�i /� Q12d.3 t2<.o 123.5 123.0 122.5 1220 121.5 }.a? s. . A LONGITUDE F EXPLOREPLATE Vancouver DISTANCE IN KM Crustal faults L/t o(i1 N2 P1,�ctoria 10. 30. 50. 70. 90. 110. 130. 150. 170 190. 210. 230 250. 270. 290. p � Seattle t0. o o % } '' .e 761 a .., . d ° 10. SO A 30 0 ':,•_ :1l'. o 0 30. D� y o `� • • " North America Plate G Washington o f 0 © A Plate Crustal 50. ¢��/ 4 x Interface °an de seismicity - JUAN DE FUCA C0 F O �c' / PLATE Portlande z 70. 4ca plate - 70. // NI o ti Intraslab , 90. earthquakes - 90. c a) a) 44° e� c.) A 1 to. _ Ito. aq co-a`/ „ Oregon t<oheo 130. 130. 0 PACIFIC PLATE ~ �_ _ _ _ - - 150. 1 t 1 1 t-� tso. o) - - 124.5 124.0 123.5 123.0 122.5 122.0 121.5 I m a GORDA LONGITUDE 0 PLATE California DEFORMATION B) EAST-WEST CROSS-SECTION THROUGH WESTERN OREGON AT THE LATITUDE OF PORTLAND,SHOWING THE SEISMIC N 40° SOURCES CONSIDERED IN THE SITE-SPECIFIC SEISMIC HAZARD STUDY{MODIFIED FROM GEOMATRIX,1995) 0 200 km Mendocino Fault 9C w' A) TECTONIC MAP OF PACIFIC NORTHWEST,SHOWING ORIENTATION AND EXTENT OF CASCADIA SLBDUCTION ZONE(MODIFIED FROM DRAGERT AND OTHERS,1994) G R 0 TECTONIC SETTING SUMMARY . FEB.2018 JOB NO. 5970-F FIG. 113 /` ` �'� -- TeB`� ;_.\` �'�{t�[1 Ie n>�il ,' ci. _�• f. �a .-�- t�,_.: wa. ��m :q��?-_, ., p.� d , ,bi �,-. - 1, TC ) llif 1 �� � \�� � a � ` 1/ r s 1 4>, y c t�f 1 _ J 4< .-.ter Ir uat r er , `l[ \1 1a ssq �' a b a Tss •a� • . 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Thole yt aTK `1 Bea‘,. 11, E �.� ., � �PV. ti. / r.,,•t Tcw s � �: Q •e(� i. . 7' t,S, Irv) _ �' , �'�.... a r ctiE '\a „ Y , \ ,- _. N,. ua'„,w 4!x ' .9 r'1 a ty 7•M o�y,�t ,iituat [ rTf T _ [. ..� 4''� 1 i 117 ems,.^ r :its I {�@ QT3 _ _Qba QTmv )-- / .,�Ttvl �e _ d a amasc x Trb t ..,,, , J""! ,-'� °Is "'`— .. �' V ), y _ �r :7 h` t •a, Reis. ! �. Tf--., '49 "� �', Qls Tl=" , t' � �,.. r :�. '"1 •T'� 4 • • _ T y315 �_ `-/�!� Y"� � ;,r• "s ,'}: j\ ~ ti ' �JJ''" ert° - L ! � f \�s, fiTc..� v `�`„ .q FROM: ? , 17 � J ( .! ' �-.•1^ •��� �.I ,I\J �`1.a,�X✓ 'Tbaa � ,. ` �� �. _ itvrtt, ; Ty T + • I v e ' WALSH,T.J.,KOROSEC,MA.,PHILLIPS,W.M.,LOGAN,R L,AND SCHASSE,H.W., .4t � rse` ! . �d ar I T • ?� 1987,GEOLOGIC MAP Of WASHINGTON-SOUTHWEST QUADRANT;1:250,000: � _ 1 rr i WASHINGTON DIVISION OF GEOLOGY AND EARTH RESOURCES,6M 34 i/ .c§." .�t` Cam. 4 7 I. , r t f • C Tfc Th[ ▪ s�i5 IE 1 j1 .{l f,.� / f,1, ,7 i �7 , c_ WALKER,G.W.,AND MACLEOD,N.S.,1991,GEOLOGIC MAP OF OREGON:U.S. A�4V Ti a- II/ - To <:z5 thi GEOLOGICALSURVEY ' ii ���� ,/ ta-- y _ � H[flhla Butt _ ,I \ '1 'I Ate. ded _. i TI .L, �� t �, ,y �'�`f.�, �f�� /� Ji � �° QTba - � J— Tbaa? . �r;u.�vv� to /'Tfc arc - jf � t .'. is � .�,2` �'� ' ` f+ ,�.i 41 ac+6 i>� .I+•!, ,'�„ `, [ I �.`'^QTbe a0 ,�^t G!S �r. e:, ``/t" r North } oIC /� Tss Ttvm7 � € s 9 Tfc? 1y ;1: MILES a r n, l� •. N ? fR y nnta // v+ (� Tfc 0 10 20 M ".-to � �-;: r_' ,��1 �. 8 -� i6pat-Mtn'3�[ ll i� (f i(:'jJ J ti /� ,� \ ��U G Thaa i T ) ♦rrc o y Trb 0 10 20 KILOMETERS 'ji of > e� . i I Tub Tna a , r tea 1'.• tn� i,t'.I 1 rt: -1 "� ` r f } l 6 t �� S TY 1 � 1 -,k -. ' QIS _ - _ Contact—Approximately located v 1 n _L?—••• Fault— Dashed where inferred;dotted where concealed;queried where doubtful;ball and bar on downthrown side REGIONAL GEOLOGIC MAP A ?A•�• Thrust fault— Dashed where inferred; dotted where concealed;queried where doubtful; sawteeth on upper plate Strike and dip of bed FEB.2018 JOB NO. 5970-F FIG. 2B \ (/ .% " ,— - , Kelso _ MAP EXPLANATION �� . ,'� s; t` �--" i `\!Y�/ \ TIME OF MOST RECENT SURFACE RUPTURE �—STRUCTURE TYPE AND RELATED FEATURES r — Holocene(<10,000 years)or post last glaciation(<15.000 years;15 ka); Normal or high-angle reverse fault `. ( \P - no historic ruptures in Oregon to date T Strike-slip fault Gearhart \ `� ' - Late Quaternary post penultimate glaciation) � Thrust taut - Sld" �At - 1 \- z:„a — Late and middle Quaternary(<750,000 years;750 ka) —(— Anticlinal fold 781 T qs "' Quaternary,undifferentiated(<1,600,000 years;<1.6 Ma)At —I1— Synclinal fold _ _ l� " — Class B structure(age or origin uncertain) Monoclinal fold it �� � Plunge direction of fold SLIP RATE c, ) - f Fault section marker 1 PaL >5 mm/year � r ,: : _ „<: 1.P5.0 mmryear 3Ch 7$1 4 �•,r :�; J . ff" DETAILED STUDY SITES 718 \ - — 0.2-1.0 mm year 2 \ St. I"Ie12Tl$- / r Trench site .pe .i r. - l ■ <D2 mmfyear P 44444✓ 1 i81-2�, Subduction zone study site ' '" ` 1/ 1r TRACE�} � Mostly CULTURAL AND GEOGRAPHIC FEATURES ✓m a .,) .rr�t %� ostycontinuous at map scale ;^ 5B0 Mostly discontinuous at map scale _- Divided highway to !'. r.,a' V' ;`� l f '�• t 1Gi1/ i� v�� Interred or concealed Pnmary or secondary road -� 1 ,t It.i� __—' �- Permanent river or stream /b - i V "''N��� f ����® - --`` Intermittent river or stream ,i_1! = `) Permanent or intermMent lake ancouver 'achr '� Ii .` tr _- t .714 4 T '�; ' � r ■4)1•111 1 v880 ��� 78, 867, Garibaldi �! t ��,; / f FAULT NUMBER NAME OF STRUCTURE Bay City -•: Forest i I�z r0 . �% 714 HELVETIA FAULT rl •Mt. 1 r • 04IMM.Ca�r � �`� �\ �! e r "lri1E1«l��� A 715 BEAVERTON FAULT ��881 Wills' �.,9 14, 41.:i xtiz lvsTa 716 - CANBY-MOLALLAFAULT 882 �., �:. T�a, -,_ a !4 e■••aRa4� a 868� 866 1 . _ �i ) ,•1a� ,879F 717 NEWBERG FAULT TllalyToak 718 /�' EWW1*1 -'\ILA! �` 718 GALES CREEK FAULT ZONE 1;'z �� 111111641111. w• � 1W �� l ,I 719 SALEM-EOLA HILLS HOMOCLINE IlltriaVolt 864 CLACKAMAS RIVER FAULT ZONE f 78 R 6 l f' ` r� •,- • rt l a n d . 867 EAGLE CREEK THRUST FAULT ci�i.7 868 BULL RUN THRUST FAULT 4. ,, �� 872 WALDO HILLS FAULT Area of MOUNT / 873 MOUNT ANGEL FAULT •'ir PI -ortland mapHOOD 874 BOLTON FAULT illir � , �� 875 OATFIELD FAULT McMinnville;:V , -: off, 876 EAST BANK FAULT HIb - it 6 / 877 PORTLAND HILLS FAULT 878 GRANT BUTTE FAULT �� ,r lalla 879 DAMASCUS-TICKLE CREEK FAULT ZONE -.� ^ p - �; , a oAdbum v Sheridan Amity % 1 'k,, \ 880 LACAMAS LAKE FAULT �, �'• f 1 �' 881 TILLAMOOK BAY FAULT ZONE II` �, (�.. IF_ r Silvelt• `.873`.781_g ` FROM: PERSONI US,S.F.,AND OTHERS,2003,MAP OF QUATERNARY FAULTS AND -, ti. : `-` 4• .. - FOLDS IN OREGON,USGS OPEN FILE REPORT OFR-03-095. / \t a4+ �A 7b.i u Dallas, —91 A' ` Salem Independent•-'i A," 872 864 v I ��\� i 0 10 20 MILES r' Monmouth 8I I 884 , •`/\--.--1"ii 871 Stayton i 0 20 40 KILOMETERS Mill City G RO LOCAL FAULT MAP FEB.2018 JOB NO. 5970-F FIG. 3B ' 0.0009 - 0.0008 -- N 0.0007 w 0 0.0006 Q 0 0.0005 0.0004 -- z < 0.0003 0.0002 0.0001 0 7.8 8 8.2 8.4 8.6 8.8 9 9.2 9.4 MOMENT MAGNITUDE Figure 22. Magnitude-frequency distribution of the Cascadia subduction zone. FROM: PETERSEN,M,FRANKEL,A,HARMSEN,S, AND OTHERS,2008,DOCUMENTATION FOR THE 2008 UPDATE OF THE UNITED STATES NATIONAL SEISMIC HAZARD MAPS:US GEOLOGICAL SURVEY,OPEN FILE REPORT 2008-1128 G RO ASSUMED MAGNITUDE-FREQUENCY DISTRIBUTION (CASCADIA SUBDUCTION ZONE) FEB.2018 JOB NO. 5970-F FIG. 58 2.2 Deterministic Bedrock Response Spectrum Based on Portland Hills Fault Hazard 2.0 (84th Percentile Ground Motions) 1.8 I Deterministic Lower Limit Response Spectrum Im 1.6 1.4 Probabilistic Bedrock Response Spectrum Based on 2% in 50-Year Hazard = i (USGS 2014 Deaggregation) 0 1.2 GJ 61 U t U Q 1.0 — U a -/ \ Deterministic Bedrock Response Spectrum Based on Subduction Zone Hazard 0.8 (84th Percentile Ground Motions) \ A 0.6 I A r \\ 0.4 r \ N 0.2 0.0 0 1 2 3 4 Period, T, seconds G RD DETERMINISTIC VS PROBABILISTIC BEDROCK RESPONSE SPECTRA (5%DAMPING) FEB.2018 JOB NO. 5970-F FIG. 6B t 1.2 1.0 0.8 bq , Recommended MCER Response Spectrum i fL li fL u 0.6 I- L N I Recommended Design Response Spectrum 0.4 0.2 0.0 0 1 2 3 4 Period, T, seconds G RD DESIGN AND RECOMMENDED RESPONSE SPECTRA (5%DAMPING) FEB.2018 JOB NO. 5970-F FIG. 7B