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HomeMy WebLinkAboutTechnical Review Meeting-Sutton-Glennallen Intertie 1993State of Alaska Walter J. Hickel, Governor Alaska Energy Authority A Public Corporation June 21, 1993 To: DISTRIBUTION a ao AA {V From: Richard Emerman f pot Alaska Energy Authority | ~+ Subject: Technical Review Meeting -- Sutton-Glennallen Intertie This is to confirm that a technical review meeting on design issues associated with the Sutton-Glennallen intertie will be held at the Alaska Energy Authority on July 6, 1993 at 1:00 p.m. The purpose is to review and discuss the design criteria and design features proposed thus far by R.W. Beck & Associates, the feasibility study contractor that has been working on this for the Energy Authority over the last several months. The meeting is expected to last all afternoon. Attached is material developed by R.W. Beck on these issues. Your review of this material and your participation at the meeting would be much appreciated. DISTRIBUTION: Dave Eberle, Alaska Energy Authority Afzal Khan, Alaska Energy Authority Brent Petrie, Alaska Energy Authority Stan Sieczkowski, Alaska Energy Authority Remy Williams, Alaska Energy Authority Dora Gropp, Chugach Electric Association Mike Massin, Chugach Electric Association Mike Easley, Copper Valley Electric Association Clayton Hurless, Copper Valley Electric Association Power Engineers, Inc., c/o Copper Valley Electric Association Jim Hall, Matanuska Electric Association Bob Mau, Matanuska Electric Association Bob Ylvisaker, Matanuska Electric Association Paul Dorvel, R.W. Beck & Associates PO. Box 190869 701 EastTudor Road Anchorage, Alaska 99519-0869 (907) 561-7877 Fax: (907) 561-8584 Alaska Energy Authority Copper Valley Intertie Some Points for Discussion Design and Construction Element Previous Design Loadings 1", Failure Seq Conductor 556 Dove ACSR Structure Type Hframe Self-supp Insulators Polymer Foundation Types DE, pipe pile Construction Zones 4 Right-of-Way 100 ft Contingency 10% other as needed (clearances..... ) H= high degree of disagreement,high impact RWB 1-2"/wind 605 Teal 30/19 Xframe guyed Polymer HP, DE 4 150 ft 20% min M= medium degree of disagreement,medium impact L= low degree of or no disagreement,low impact Degree except Vpost Draft Report System Configuration Options PED 6/6/93 There are several basic options for constructing an intertie to CVEA from MEA. These options are discussed below. 1. Voltage Selection for the Intertie The MEA 115-kV and CVEA 138-kV systems are voltage incompatible and an autotransformer is required to mate the systems. The auto could be located at either end of the intertie, but higher losses and maybe overall higher lifecycle costs for a 115 kV line make it preferable to locate the auto on the MEA system and operate an intertie at 138 kV. This voltage level is adequate to serve the assumed long term power needs of the CVEA system, but it is not adequate to serve as a link to other systems beyond CVEA as might be desirable in the context of regional network planning. 2. Location of the Autotransformer on MEA System The 115/138 kV autotransformer could be located logically in any one of three places: at the existing MEA O'Neill Substation in Sutton, at a new substation fed from the 115 kV O'Neill Tap Line, or at the O'Neill Tap point. Location of the autotransformer in the existing O'Neill Substation has few advantages other than existing right-of-way and easement. Considerable expansion of the substation would be required while maintaining service to Sutton. To avoid working the substation hot a mobile substation could be used to maintain service. However, since the existing substation is located in a residential zone between two existing homes off Jonesville Road, selecting a getaway route for an Intertie is very restricted and could engender increased public opposition. This option was selected for study in [1]. The Mat-Su Borough Planning Commission proposed, and we support, use of a new substation located on a borough parcel about 0.7 miles west of O'Neill adjacent to the O'Neill 115 kV Tap Line. The autotransformer could also be placed at the O'Neill Tap point. This would require the uprating of the tap line to 138 kV operation and replacement of certain equipment at the O'Neill Substation. Increasing voltage level could also require modification of easement agreements and lead to further public opposition. MEA has considered conversion to 138 kV. 3) Improved Operational Switching Flexibility The O'Neill Substation is fed radially via a 16-mile, 115 kV line from an unbreakered three-way tap Teeland-Eklutna line. Currently any fault between Teeland, Eklutna and O'Neill will cut off service to O'Neill. A 3-breaker, 3-position ring bus switching configuration is desirable to improve system reliability and continuity of service. It can be used to control CVEA backfeed to the Sutton service area and to sever the two systems to limit fault impacts. Two options exist for siting the ring-bus: at the new substation in Sutton or at O'Neill Tap, replacing the existing three-way disconnects. If the ring bus is placed at the new substation in Sutton and the O'Neill Tap point is not changed, any fault between Teeland and Eklutna or on the O'Neill Tap Line, will sever MEA service both to an Intertie and O'Neill Substation . This may, however, place a 1.5 MW Sutton load on the CVEA system which could then experience underfrequency problems as its generators attempt to pick up the CVEA system load. The breakers at the new substation could be relayed to drop the O'Neill Substation load [verify practical]. The possibility of CVEA backfeeding Sutton is discussed in Appendix [PTI]. In any case, reliability of the Railbelt feeds to an Intertie are affected by faults in any direction from the tap point. If the ring bus is placed at the O'Neill Tap point, a fault on either side of the tap toward Teeland or Eklutna would leave intact the other feed to the O'Neill Tap Line and an Intertie. A fault on the O'Neill Tap Line would still sever an Intertie and O'Neill Substation. However, some increased operation flexibility is achieved for the MEA system and reliability is increased marginally for an Intertie. All lines involved have been highly reliable [VERIFY WITH JIM HALL]. In this scenario, the new substation in Sutton would be a single bus configuration. In Phase | we have assumed a ring-bus at the new substation in Sutton. In Phase 2 we recommend that the location of the ring bus at the O'Neill Tap be addressed in detail. 4. Future Regional Grid Concerns The link from MEA to CVEA has been considered as part of an eventual second intertie with GVEA through Delta Junction. These regional plans would logically include consideration of an upgrade of the MEA 115-kV system and extension of the existing 230-kV system to at least O'Neill Tap. We have not given any consideration in the feasibility study to the sizing or rating of intertie equipment for the eventual purpose of relocating to other parts of the grid. [For a discussion of possible region-wide ramifications see Appendix ____[ PTI].] In summary then, in Phase | we select a system based on a new substation in Sutton with a 3- position ring bus and a 138 kV Intertie voltage, with the O'Neill Tap remaining unchanged. Other options such as a double circuit line from a 4-position ring bus at O'Neill Tap were not considered. Draft Basis of Feasibility Design Transmission Lines Only June 19, 1993 A: Background A Sutton to Glennallen transmission line has been studied previously on its own to serve the Copper Valley system [1], as part of the Railbelt Intertie[2], or as a link from the Railbelt to Cordova [6]. For the latter two studies the proposed line was rated at 230 kV and was intended to interconnect the Golden Valley Electric Association system or the City of Cordova to the Railbelt. The most recent study [1], prepared for CVEA by Power Engineers ("authors") covers the Sutton to Glennallen 138-kV Intertie along the same basic route as the present feasibility study. It also represents the most thorough discussion to date of design criteria in concert with the current operating and maintenance desires of CVEA. In [2] the capital cost of the complete second Railbelt Intertie from Palmer through Glennallen to Delta Junction was estimated at $156 million in 1989 dollars and was subsequently judged to be too expensive to pursue at the time of the study. The scope of the present feasibility study requires only consideration of a 138 kV transmission line to serve the needs of Copper Valley Electric Association without regard for possible future use as a link in a regional network despite demonstrations that a second Railbelt Intertie would have definite regional electrical system performance advantages [2]. A line designed for 138 kV operation and with a conductor sized only to serve a 40 MW projected Copper Valley load (i.e. 556 ACSR Dove or similar) would rule out effective upgrading to serve as a link in a wider network without essentially completely reconstructing the line. Further studies would be required to verify that 138 kV is or is not suitable for a second intertie to GVEA, although assuming 605 ACSR Teal conductor, preliminary computations for just the Sutton-Glennallen link indicated losses on the order of 7% for a 90 MW transfer with an 18 % voltage drop [see Appendix ___, PTI Report]. No computations were made to predict the impact a 90 MW flow would have on operating temperature. In agreement with the Energy Authority we have taken [1] as a reasonable starting point for the basis of feasiblity design. To supplement our firm's experience, we have discussed the Intertie with the engineering and operations staff of CVEA, MEA, GVEA and AEA nd other transmission design engineers familiar with Alaskan conditions. In the following section we review the assumptions and criteria of this study and formulate revised criteria where deemed appropriate or to suit different route situations. B. Project Location The Sutton to Glennallen 138-kV Intertie will originate in Sutton at a new substation located approximately 0.7 miles west of the O'Neill Substation on a parcel owned by the Matanuska-Susistna Borough. The line will terminate and interconnect with the CVEA system at the Pump Station No. 11 Substation, currently owned by AEA but operated and maintained by CVEA , located adjacent to the Alyeska Pipeline in Glennallen about one mile south of Glenn Highway. The Intertie alternative routes will principally follow combinations of the Matanuska River Valley, Boulder Creek, Caribou Creek, Hicks Creek, Squaw Creek, and Alfred Creek before crossing into the Copper River Basin where the line will run generally parallel to the Glenn Highway. Refer to maps | through 11 in Section for the routes. For most of the routes study segments, elevations are 500 to 2500 feet. In the mountain passes, elevations will increase to nearly 4500 feet and will require consideration of more severe loading criteria. The Intertie would be approximately 135 miles in length and would be located at latitude 62 degrees north and longitude 145.5 to 148.5 degrees east. Cc. Transmission Line 1. Electric Loading The Intertie will be designed for the transfer of 40 MW at 0.90 lagging power factor (receiving) at 138 kV. The basic conductor for the Intertie was designated by others [1] as 556 kcmil 26/7 ACSR (codename "Dove"). The selection of 556 Dove was based largely on CVEA stocking considerations. However, alternative conductors with strength and sag advantages should be considered for the potential of significant project cost benefits. Alternative conductors include 556 Dove with an extra- high strength steel core, standard 605 and 636 kemil sizes of ACSR with 30/19 stranding (i.e. higher steel content stranding), 605 kcmil size of steel-supported aluminum conductor (SSAC ™ ), and Dove ACSR/SD self-damping conductor. We performed sag tension computations and compared sags for several different conductors and recommend the use of 605 kcmil 30/19 ACSR, code name Teal. Computations for 556 Dove, based on a pure radial transfer of 40 MW, 0.90 power factor lagging (receiving end conditions), indicate about 3.8% power losses and a voltage drop of about 5.8%, which are acceptable for the feasibility study. For the same loading conditions, losses for 605 Teal are 3.6%. System studies will provide more definitive values for losses and system performance under steady state heavy and light loading plus transient behavior. See Appendix _____ [ PTI report]. 2 Ampacity The authors of [1] propose developing ampacity ratings and maximum operating temperature based on (1) maximum MVA transfer rather than arbitrary standard values such as 75 or 100 degrees C; (2) maximum ambient temperatures 90 F (32 C) summer full sun and 50 F (10 C) winter with no solar input, i.e. cloudy; and (3) windspeed at 2 feet per second (fps), coefficient of absorptivity of 0.55 and coefficient of emissivity of 0.5. They predicted maximum operating temperatures of about 110 F in summer and less than 60 F in winter for 40 MVA loading. Recognizing that CVEA is a winter- peaking utility, the authors state that future consideration could be given to reducing the MVA basis for the summer ampacity rating, but that their recommendation is to use 110 F as the maximum design conductor operating temperature for clearance determination. We agree that actual maximum MVA should be the basis for determining maximum operating temperature in accordance with the NESC. We have verified that the ambient temperature assumptions are reasonable. The value of windspeed of 2 fps (2.9 mph) is a commonly-used and accepted number in the industry and we have no reason to question its use in [1], except that a meteorological study could be requested to recommend an average value of windspeed which might be different and serve to reduced operating temperatures. The use of 110 F as a maximum operating temperature is not in compliance with NESC Rule 232 which requires a minimum of 120 F for the application of Rule 232. Actual maximum operating temperature is determined from maximum loading and ambient conditions and is used to verify that a selection greater than 120 F is not required. For some of their length, most line route segments follow valleys which may tend to channel wind parallel to the line with reduced cooling effect[3]. The assumptions of 0.55 for solar absorbtivity and 0.5 for infrared emissivity are typical of new, unoxidized aluminum conductors. The line will be in service many years and values for a typical blackened conductor, 0.95 for solar absorptivity and 0.91 for infrared emissivity, would appear appropriate for a check. We used alternative software to predict maximum operating temperatures. These are given in Table 4-1. Table 4-1 Conductor Operating Temperatures '“ Winter 10C Summer 32C Summer 32C Summer 32C Wind E-W Wind E-W Wind N-S Wind E-W Aged Conductor Teal 605 ACSR? 24.5 C (76F) 55.8 C (132F) 45.5 C (114F) 61.0 C (142F) Dove 556 ACSR’ 24.9 C (77F) 56.1 C (133F) 45.5 C (114F) 61.1 C (142F) 37 No. 9 AW® 35.3 C (96F) 67.0 C (153F) 48.9 C (120F) 69.6 C (157F) Notes 1. Line current is 170 amps, latitude 62 N, longitude 147 E, summer date July 4 time 15:00 PST ,winter date January 1 time 13:00 PST. 2 Elevation is 2500 feet above sea level, windspeed 2 feet per second. 3. Elevation is 4500 feet above sea level, windspeed 3 feet per second. 4. Coefficient of solar absorbtivity is 0.55 and coefficient of infrared emissivity is 0.50, except for aged conductor which uses 0.95 and 0.91 respectively. Based on the results in Table 4-1, we recommend and will assume for the feasibility study a maximum operating temperature of 150 F (66 C) for 556 Dove ACSR or 605 Teal ACSR. For reference we also include temperatures for a 37 no. 9 Alumoweld conductor which might be used in the Chitna Pass area.. The maximum operating temperature for this conductor will be chosen as 160 F (71 C). 3; Weather Data Weather data is critical for formulating reasonable physical loading criteria. Most critical is information on ambient temperatures, windspeed and direction, ice,snow and frost accumulation and densities, snow pack, and isokeraunic level (i.e. thunderstorm activity) for lightning protection design. Weather data for the Intertie corridor is very limited. This situation is complicated by microclimates which can create very severe and special loading conditions. Major sources of basic weather data measurements are the Alaska Climate Summary (ACS), the Western Regional Climate Center (WRCC, Reno, Nevada), and the National Climate Data Center (Asheville, North Carolina). In addition, reference [3] provides expert meteorological analysis of limited weather data for the Palmer-Glennallen corridor in the Matanuska River Valley and parallel to the Glenn Highway and develops detailed recommendations for extreme loading conditions /See Part 4 below]. Firsthand observations by local utility workers and line designers are also invaluable for selecting extreme loading conditions. Minimum and maximum ambient temperature data were obtained for weather stations at Palmer, Sutton, Gunsight Mountain, Snowshoe Lake, Eureka, Little Nelchina Road Camp, Tahneta Pass, Glennallen, and Gulkana from the WRCC. This basic information was used to evaluate the appropriateness of conductor operating temperatures and ambient conditions under different loadings presented in [1]. a. Extreme Windspeed The authors state that they were directed to assume an extreme wind speed of 100 mph by CVEA. Recorded values of 70-75 mph at Glennallen in 1992 and 115 mph in Palmer in 1979 were cited. The consensus opinion of people familiar with the area, including the Anchorage National Weather Service, was that the line route can be a "very windy place." The ACS provides no wind data of significance. One fastest mile sample reading was obtained for Gulkana and converted to 50-year and 100-year recurrent interval wind speeds or 91 mph and 100 mph respectively. The authors conclude that an extreme 100 mph windspeed is reasonable and should be used. They further suggest that reductions in the design wind speed for the shielding effects of trees and terrain can be considered in final design. An extreme windspeed of 100 mph is generally in agreement with meteorological analysis of the corridor [3], which estimated a maximum one-minute 75-year windspeed of 89 mph and maximum 5-second 75-year gust windspeed of 121 mph. These maxima were predicted for a portion of the corridor near Sheep Mountain and Lions Head. The wind channeling in this area is high and similar loads could be expected in backcountry valley routes. Weighted averages of the 75-year windspeeds for the entire length of the corridor were calculated to be 75 mph for the one-minute windspeed and 103 mph for the 5-second gust windspeed. We will use an extreme windspeed of 100 mph applied to horizontal spans and a gust windspeed of 120 mph applied to the structure only for most of the Intertie, Loading Zones 1 and 2. Shielding effects of trees and terrain may be possible for transverse winds in certain locations and this is considered in detail in [3]; in the absence of a detailed meteorological analysis for the present routes, however, shielding effects will not be considered in this study. Higher, channeled winds may be experienced when crossing major streams and creeks. Longitudinal winds would be largely unshielded but their loading effect on wires is also reduced by the shallow impact angle. We will use an extreme windspeed of 125 mph applied to horizontal spans and a gust windspeed of 150 mph applied to the structure only for the backcountry routes (Loading Zone 3) which cross passes at up to 4500 feet (e.g. Chitna Pass). Extreme winds of 100-175 mph have been used for design of several high, exposed line segments in Alaska, e.g. 175 mph for Glennallen-Valdez line in Thompson Pass and 200 mph Snettisham [verify]. The dependence of the wind pressure coefficient on temperature and elevation [4, Table 2.1-1] can be taken into account to reduce wind loading in final design where appropriate. b. Ambient Temperature Temperature records for Gulkana, Snowshoe Lake, Tahneta Pass, and Sutton, are listed [1]. The Gulkana station recording period is the longest (1942-1987) and shows the most extreme ambient temperature values (-65 F record low and 91 F record high). Several ambient temperatures are of interest: maximum summer temperature for determining maximum operating temperature, extreme minimum temperature for studying uplift situations, ambient temperatures during extreme loading conditions, and the annual average minimum temperature (AAMT) used with tension limits to control aeolian vibration. We agree with the selection of maximum ambient temperature of 90 F in summer for computation of maximum conductor temperature. Since CVEA is a winter-peaking utility it is unlikely that the maximum operating temperature will be reached thus giving the Intertie substantial thermal margin. The authors recommend a single cold temperature of -60 F for the entire project. This recommendation is based on the fact that the cold temperature tension limits will not control sag-tension behavior in view of the extreme ice loadings, except in very short spans where tensions at -60F should be checked against limits. The selection of -60 F as the extreme minimum cold condition is appropriate for the line and will, as the authors state, affect only structure locations and types in uplift situations, e.g. in the bottom of valleys and toes of slopes. We agree with the use of -60 F. An AAMT is used essentially for aeolian vibration tension limits and the authors later choose -25 C (-13 F), based on a 10% coldest temperature criterion. This value appears appropriate for the western portion of the line but not for the backcountry routes and eastern portion based on estimated minimum temperatures in Table 4-2. For this feasibility study we will use an AAMT of -32 C (-25 F) for the ACSR conductors. Note that SSAC™ and ACSR/SD conductor types have inherent self-damping characteristics which do not require application of the AAMT limits. Table 4-2 Average of Annual Minimum of Monthly Average of Daily Minimum Temperatures (deg F) Highest/Lowest Station Years Annual Minimum Mean Std Dev Mean-2SD Sutton 1978-1992 16.4/-11.3 2.8 7.8 -12.8 Palmer 1961-1992 12.5/-11.4 -0.6 6.3 -13.2 Gunsight 1966-1974 -20.9/-40.4 -29.7 5.7 -41.1 Eureka 1957-1968 10.5/-2.1 43 4.1 -3.9 Snowshoe 1963-1992 -9.7/-35.8 -23.1 6.2 -35.5 Glennallen 1965-1992 -1.68/-37.7 -23.2 9.1 -41.4 Gulkana 1961-1992 -0.3/-36.5 -19.7 84 -36.5 c. Snow Ground Cover The authors cite maximum snow accumulation data from the ACS for Gulkana (48"), Snowshoe Lake (36"), Tahneta Pass (48"), and Sutton (36"). A design value of 48 inches of snow ground cover was chosen for the determination of NESC clearances to grade. The authors prudently mention the need to consider greater snow cover or clearances in areas of high snow machine usage. We agree with the selection of 48 inches snow cover as the basis for the feasibility study for non-backcountry areas. We recommend and will assume a snow cover of 60 inches in backcountry areas. d. Ice and Snow Accumulation The authors state that the ACS does not tabulate ice accumulation information and that local experience is the best guide for selecting this design parameter. The selection of ice and snow loading criteria for transmission line design in Alaska has been the subject of much discussion and study over the past few years. Engaging meteorological consultants to develop recommended extreme loadings has been common in recent transmission line design and failure analysis work in Alaska. We recommend that a meteorological consultant be hired to develop such recommended loadings preferably prior to completion of the feasibility study and at least prior to final design. Very large build-ups of snow and underlying ice have been observed on several lines including the Healy-Willow 345-kV Transmission Line, the Glennallen-Valdez 138-kV Transmission Line, and the Tyee Lake 138-kV Transmission Line in Southeast Alaska. Each of these lines has experienced excessive sags due to accumulated ice and snow, but no outright cascading failures due to ice build-up. On the Tyee Lake 7ransmission Line 556 Dove section across Vank and Woronkofski Islands, excessive sags have led to conductor-snow/ground contact and low-level ground faults [4]. On the Healy-Willow line with 2-954 ACSR conductors, contact with 15' tall trees and 7 severely reduced highway clearances have been experienced [8]; ice has also been observed to bridge the subconductors at spacer locations. CVEA personnel reported that 5"-6" radial snow at unknown density over 1" radial ice were observed on the CVEA Glennallen-Valdez line just north of Thompson Pass with no failures. Extreme diameters of heavy ice and snow accumulation of up to 18 inches and typically 4-12 inches have commonly been observed on Alaskan lines but efforts to measure accumulation have been only partially successful. MEA designs its distribution and transmission lines to NESC Heavy and reportedly has encountered no failures[1]. CVEA also designs its distribution lines for NESC Heavy and has also experienced no failures to date. In [1] a value of 1" of radial ice is selected for design extreme ice loading, reasoning that it is better to accept the risk of greater ice loading and consequent failures than the certainty of increased capital costs for improved reliability. This is a trade-off made for virtually all lines designed for areas where extreme loading is unknown or where the cost to withstand infrequent but severe loading (e.g. hurricanes, tornadoes) would be exhorbitant. A 1" radial ice extreme loading is consistent with criteria selected for several lines in Alaska, including loading zone II for the Glennallen-Valdez line extending from Pump Station No. 11 south to mile 70. The reliability of this line segment has been excellent in its 8-9 year service life. We note that the section is parallel to prevailing winds in the Copper River basin, whereas an Intertie would be parallel to the prevailing winds. The loading imparted by snow and ice depends on their radial thickness, extent and density. The four basic categories of ice include glaze ice at 57 pcf, rime ice at 18 to 56 pcf, wet snow at 18 to 50 pef, and hoarfrost at less than 18 pef. Each of these categories is formed under different circumstances of air temperature, wind speed, water content and liquid drop size as described in [3,5]. The formation of ice in its various combinations on the conductor is further complicated by the impact of resistive heating by the electrical load current, but estimating this effect is not practical. Reference [5] cites recommended procedures for assuming ice loadings based on maximum measurement or observations in the absence of statistical data. Specifically one suggested procedure would assume a mean ice radial thickness of 0.60 times the maximum observed thickness with a standard deviation of 0.40 times the mean [5].. Table 4-la below gives several examples of how this methodology would apply to observations. Assuming a Gaussian distribution the one-standard deviation radius would have an annual probability of occurrence of 15.9%, corresponding to approximately a 6-year mean recurrence interval, while a two-standard deviation radius would have a 2.2% annual probability, corresponding to a 50-year recurrence interval. A probability of occurrence for a given project lifetime can be computed and is shown in Table 4-1b. Table 4-la Extreme Ice Assumptions in the Absence of Statistical Data (All values of radial ice in inches) Maximum Observed Estimated Calculated One Std Two-Std Ice Radius Mean Std Dev Radius Radius (inches) (inches) (inches) (inches) (inches) 0.50 0.30 0.12 0.42 0.54 1.00 0.60 0.24 0.84 1.08 1.50 0.90 0.36 1.26 1.62 2.00 1.20 0.48 1.68 2.16 2.50 1.50 0.60 2.10 2.70 3.00 1.80 0.72 2.52 3.24 Table 4-1b Probability of Occurrence of 1-SD and 2-SD Ice Loadings Project Probability Probability Lifetime of 1 SD Ice of 2 SD Ice Years J 0.578542 0.108463 10 0.822373 0.205161 15 0.925138 0.291372 20 0.968449 0.368231 25 0.986702 0.436755 30 0.994396 0.497846 35 0.997638 0.552311 40 0.999005 0.600869 45 0.99958 0.644159 50 0.999823 0.682755 For this feasibility study we will assume ice and snow loadings as given in Table 4-4, Assumed Design Data. 4. Meteorological Research, Inc. Technical Report [3] Reference [3] contains a detailed analysis of limited available weather information and develops specific extreme loading criteria for the Palmer-Glennallen corridor, basically following the Glenn Highway. One of the principal authors of [3] was contacted during the present study to discuss the appropriateness of the findings in [3] to the route alternatives under consideration in this study. He advised that the findings in [3] are not valid for the route alternatives under study but that they are still probably valid for the routes studied in 1982. Despite its limitation to the 1982 routes, the findings and recommendations of this report are discussed in this section. Two climatic zones are identified. The western portion lies in a transition zone through the Matanuska River Valley, characterized by a mixture of cool, moist weather and cold, dry weather. The eastern portion lies in the continental zone through the Copper River Valley, characterized by cold, dry weather. 9 A tabulation for twelve weather stations indicates maximum ambient temperatures of 90 F for Palmer and Glennallen, 91 F for Gulkana, and 85 F for Eureka and Snowshoe Lake. This corroborates the selection of 90 F for maximum ambient temperature. A tabulation for the same twelve weather stations indicates minimum ambient temperatures of -60 to -65 F in the Copper River Valley and -35 to -44 F from Palmer to Eureka. This corroborates the selection of -60 F for extreme cold temperature. A tabulation for the same twelve weather stations indicates a mean annual snowfall of 65 to 68 inches on the western end of the Matanuska Valley, 41 to 53 inches between Snowshoe Lake and Gulkana in the Copper River Valley, and 117 inches in Eureka. This would seem to corroborate the selection of 48 inches of ground snow cover for clearance checks. However, the value of nearly 10 ft of total annual snowfall in Eureka indicates a higher value of snow cover would be appropriate for the areas of Eureka and backcountry routes. We will use a snow cover of 60 inches for clearance checks in these areas. Analysis of maximum one-minute hourly winds (e.g. extreme wind values) at 30 ft above ground showed 50-year return period windspeeds for Gulkana of 59 mph, Sheep Mountain 69 mph and Palmer 74 mph. The direction of extreme winds was generally NNE or SSE for Gulkana, S and SSE for Snowshoe Lake, and NE or ENE for Sheep Mountain and Palmer respectively. Extreme wind occurrences are distributed evenly from fall to spring for Gulkana but are somewhat concentrated in the winter months November to March for Sheep Mountain, Palmer and Snowshoe Lake. This data indicates that extreme winds will generally be transverse to the Intertie route east of Eureka and longitudinal to the route in the valleys west of Eureka, and that extreme winds are very likely to occur in winter. MRI suggests that a reduction of windspeed up to 20% in the "lee of stands (of spruce)" is possible and that larger reductions of 20-50% are possible in denser, taller forest stands. The value of 20% should be used, according to MRI, where the transmission line cleared right-of-way is parallel to the prevailing winds. The MRI report discusses span factors which reflect the span coverage of winds of different durations. For instance, gusts will not typically affect more than 100 feet of span while steady one- minute winds will affect up to 1000 feet. Mountainous terrain is recognized as creating special wind turbulence. The report culminates with tables giving 25-, 50- and 75-year return period values for one- minute wind speeds, 5-second gust speeds, mixed icing loads, glaze icing loads and rime icing loads by proposed line segment. The windspeed values are adjusted for exposure, elevation and for a common height above ground of 30 ft. Total transverse wind load on iced conductors is included. Wet snow occurrences were cited as being too infrequent and short to result in any significant accumulation. We reverse-computed assumptions for coincident wind loading embedded in the 50- year and 75-year tables and derived Tables 4-3a and 4-3b which summarize the MRI recommended loadings for the 1982 route. Table 4-3a Reduction of MRI 50-year Data for Intertie Study Criteria Loading condition West East Maximum Select Extreme Wind 74 59 85 102 One-minute mph Gust 101 82 115 138 5-sec mph Mixed Ice Radial inches 0.75 1.00 1.00 1.00 Weight lb/ft 0.68 1.05 1.05 1.05 Wind mph 5 10 10 10 Rime Ice Radial inches NA NA 1.50 1.50 Weight lb/ft NA NA 1.98 1.98 Wind mph NA NA 40 40 Glaze Ice Radial inches 0.25. 0.25 0.25 0.25 Weight |b/ft 0.36 0.36 0.36 0.36 Wind mph 15 1S 18 18 Table 4-3b Reduction of MRI 75-year Data for Intertie Study Criteria Loading condition West East Maximum Select Extreme Wind 78 61 89 107 One-minute mph % 1.2! 73 Gust 106 84 121 145 5-sec mph 1.2 101 Mixed Ice Radial inches 1.00 1.50 1.50 1.50 Weight lb/ft 105 1.98 1.98 2.00 Wind mph 7 12 12 12 Rime Ice Radial inches NA NA 1.75 1.75 Weight |b/ft NA NA 2.55 2.55 Wind mph NA NA 50 50 Glaze Ice Equivalent NA NA 0.573 12 0.893 48 0.25 18 Glaze Ice Equivalent NA NA 0.893 14 1.055 60 10 11 Glaze Ice Radial inches 0.3 0.3 0.3 0.3 0.3 Weight lb/ft 0.45 045 0.45 0.45 Wind mph 20 20 23 23; 23 Notes 1. Weight of ice loadings is weight of ice only over Dove 556 ACSR conductor. 2. All ice loadings considered at 0 F. 3. Rime and mixed ice at 0.4 g/ce or 25 pef an glaze ice at 0.9 g/ce or 56 pef. MRI comments that mixed icing events (i.e. rime and hoarfrost) are the result of light winds and dense fog conditions and that significant accumulations can be expected at elevations below 2000 feet in the Matanuska River Valley. In the Copper River Valley significant rime ice accumulations are expected at elevations above 3300 feet, which would also apply to backcountry areas in our opinion, and that significant mixed icing events will occur in the basin from Tolsona to Glennallen, at elevations below 2500 feet. No recommendations are given for assumed ambient temperatures during ice and wind loading. We will assume 0 F for all extreme ice and extreme combined ice and wind loadings. MRI recommended low windspeeds for icing events based on the line being generally parallel to prevailing winds in the Matanuska River Valley and based on weather records in the Copper River Valley. Observed extreme loadings in the region have apparently exceeded MRI findings. 5. Weather Load Cases The authors [1] recommend five loading cases namely (1) NESC Heavy Loading Zone with 1/2" radial ice, 4 psf wind, at 0 F; (2) NESC Extreme Wind 12.5 psf (70 mph) at 14 F; (3) extreme wind 26 psf (100 mph) at 14 F; (4) combined extreme loadings (a) 1" radial ice at 57 pcf, no wind, 0 F, west of longitude 147 35 (approximately at Eureka)and (b) 1" radial ice at 57 pcef, 4 psf (40 mph) wind, 0 F, east of longitude 147 35; and (5) construction load equal to 700 Ibs vertical and 2000 Ibs longitudinal applied at any one or all conductor attachment points. Each load case is treated separately below. Case (1) is the standard NESC Heavy loading zone and is required. NESC Grade B overload factors will apply to loads calculated for these conditions. Case (2) is the NESC extreme wind condition based on ASCE 7 wind maps giving the fastest- mile at 33 feet above ground level and for a 50-year return period, which corresponds to a 0.02 probability of occurrence in any one year. The NESC specifies an ambient temperature of 60 F for this condition, while the authors opted to use 14 F to reflect conditions more suitable to the Intertie. This temperature approximates the statistical mean low for most observation sites for the months of April and October, and the mean high for the months of February or March and November. This gives a reasonable expectation that the wind value will occur at or above 14 F. NESC specifies an 12 overload factor of 1.0 for steel structures and foundations for this load condition. Case (3) is an additional extreme wind case using a base wind speed 100 mph (26 psf). The authors discuss the possible shielding effect of trees and terrain features as a means by which to apply a reasoned reduction of extreme wind speed along the line to reduce costs. They correctly note that increasing windspeed from 70 mph to 100 mph will double the wind load from 13 psf to 26 psf, and that this would have a major impact on initial costs. The authors do not specify overload capacity factors for this case explicitly. They imply that NESC OCFs would apply. We recommend and will use a 1.10 OCF for all loads in this case. Detailed rules for determining the wind speed reduction are proposed for consideration in final design, but are not considered for this study. The authors state that the use of these rules may be invalid given the uncertainty of the tree cover over the lifetime of the Intertie due to fire, harvest or, we might add, disease; a major bark beetle infestation threatens the black spruce forests east of Eureka. We do not recommend reduction of the 100 mph extreme wind load in the knowledge that many portions of the line will be in area where wind channeling can take place and in light of the concerns of the authors raised above. Furthermore, portions of the study route segments are perpendicular to the direction of Matanuska Valley and its prevailing winds. It is not practical to fine tune wind loading at this level of study. Case (4) is the extreme combined ice and wind loading case. It is the most difficult to estimate and the most fraught with risk if underestimated. A 1" radial ice load at 0 F is proposed by the authors with 4 psf (40 mph) wind load east of Eureka and without wind west of Eureka. They recognize the observations of 5" - 6" diameter rime ice accumulation in the Glennallen area. This level of rime ice at a low end density of 20 pcf would equate to a radial glaze ice of 3.5" at 57 pef, and would, as the authors state, control design of the line to a considerable degree by generally requiring stronger conductor, support assemblies, structures and foundations or shorter spans. The authors do not specify overload capacity factors for this case explicitly. They imply that NESC OCFs would apply. We recommend and will use a 1.00 OCF for all loads in this case. We agree with the approach to divide the line into different loading zones but would expand their number to three to accommodate the backcountry, high elevation route alternatives. We further agree with the selection of 0 F as the temperature, noting that even though ice formation will initially take place typically at temperatures above 25 F it is highly likely that a cold front could pass through on the tails of the icing event [5,pg 66]. Considering the weather analysis in [3], the assumptions of [1], and discussions with knowledgeable transmission line design engineers in Alaska, we recommend and will assume combined extreme ice and wind loadings as given in Table 4-4 Case (5) is a construction load which we believe is adequate for Loading Zones | and 2. For Loading Zone 3, however, we will increase the loads as shown in Table 4-4. The authors do not specify overload capacity factors for this case explicitly. They imply that NESC OCFs would apply. We recommend and will use a 2.00 OCF for all loads in this case. Table44.xls LOADING CONDITION NESC Heavy OCF wind 2.50 OCF vert 1.50 OCF tension 1.65 Extreme Ice No Wind OCFs 1.10 Extreme Wind No Ice OCFs 1.10 Extreme Combined Wind and Ice OCFs 1.0 Construction OCFs 2.00 Ambient Temperatures Elevation Spans Load Calcs Air Gap Structure Dims LOADING.XLS Table 4-4 Assumed Study Design Data Parameter Radial Ice Wind speed Wind PSF Temperature Radial Ice Temperature Wind speed Wind PSF Gust Gust Temperature Radial Ice Radial Snow Snow Density Ice Equiv Wind speed Wind PSF Wind Equiv Temperature Vertical Longitudinal AAMT Maximum Mimimum Maximum Ruling 1.2 Horizontal 1.4 Vertical Average No Wind Mod Wind High Wind Loading Units Zone 1 in 0.5 mph 40 lb/sf 4 deg F 0 in 1 deg F 0 mph 100 lb/sf 26 mph 120 lb/sf 37 deg F 20 in 1 in 2 pcf 30 in 2.24 mph 20 lb/sf 1 lb/sf 1.28 deg F 30 Ibs 700 lbs 2000 deg F -25 deg F 90 deg F -60 ft 2500 ft 1000 ft 1200 ft 1400 ft 900 in 54 in 36 in 20 Page 1 Loading Zone 2 0.5 40 4 oO 1.5 0 100 26 120 37 10 = 20 1.9 40 5.86 20 700 2000 -25 90 -60 3500 1200 1440 1680 1100 54 36 20 Loading Zone 3 0.5 40 4 0 2 0 125 40 150 58 10 Np NA 75 14 14 20 2000 3000 -25 90 5000 1100 1320 1540 1000 60 40 24 Maximums from Reference [3] 50-Year 75-Year 85 89 18.5 20.2 115 121 33.9 37.5 NA NA 1.5 1.75 0 0 Oo Oo 0.89 1.06 40 50 4 6.4 5.93 9.33 NA NA Ref [1] 0.5 40 oF = 100 26 14 =-0O00oO2 opt 700 2000 13 Table 4-4 summarizes the loading conditions we will assume for the feasibility study as well as other design assumptions. Insert Table 4-4 here. 6. Conductor Sag-Tension Design Conductor sag tension limiting conditions are summarized in Table 4-5. Table 4-5 Conductor Tension Limits Percentage of URS Loading Condition Reference [1] Feasibility Study NESC Heavy 60 %! 60%! NESC Extreme Wind No Limit No Limit Extreme Wind, No Ice 70 % 70% Extreme Ice, No Wind No Limit 70 % Combined Extreme Loading 60 % 80 % Initial Unloaded Tension 60 % (AI only) 33 % (-25F) Final Unloaded Tension 25 % 25 % (-25F) Notes 1. NESC Rule 261 required limits. 2. Based on tension/mass ratio of 7400 ft; temperature in [1] is -25C which equates to - 13F, not 13F as used in [1]. 3; At -40F, approximates NESC 35% limit for ACSR Dove. We note that [1] contains no condition of -40 F in sag tension runs as indicated in this section of their report; we have added this condition to sag tension runs in this study. Also unloaded tension limits close to maximums required by NESC but at much lower temperatures than the NESC 60 F are used in [1]. Our standard practice is to use reduced unloaded, initial tension limits on the order of 20 % ultimate rated strength (URS) at the AAMT without vibration dampers; this is in line with REA 62-1 [9] recommendations. However, this will lead to increased sags and shorter spans with an attendant cost penalty. NESC stipulates that the tension not exceed 35% URS under initial conditions or 25% URS under final conditions at 60 F. REA [9] and Alcoa recommend about the same limits, 33% initial and 25% final, but at 0 F for the NESC Heavy zone. The line can generally be divided into two segments where prevailing winds are generally parallel (west of 14 Eureka) or perpendicular (east of Eureka) to the alignment. For Loading Zones | and 3, REA-recommended limits at the AAMT will be used, while for Loading Zone 2, where steady and strong northerly or southerly winds can be expected, tension limits of 25% initial and 20% final will be used. Vibration dampers would be used as needed, depending on the final conductor chosen. Dove ACSR sag tension runs for Loading Zones 1,2 and 3 are included for ruling spans 800- 1200 feet. Maximum final sag for clearances at 1/2 inch radial ice and 30F or 150 F are quite high leading quickly to the need for structure heights over 80 feet to reach target span lengths on the project. We recognize that the selection of Dove was largely based on its existing use on CVEA lines. However, we believe that alternative conductors should be examined if they hold promise for reducing project cost while meeting the physical loading criteria established. Stocking requirements are a comparatively minor cost and inconvenience when viewed in light of line reliability. Alternative conductors were briefly evaluated as discussed below. Tables 4-6a to 4-6c give a comparison of the sags and typical structure height advantages for various loading conditions and conductors. 1. Dove ACSR with an Extra High Strength (EHS) Steel Core Standard ACSR conductor is typically fabricated using high strength steel core. It is possible to fabricate Dove ACSR with an extra-high strength steel core. Special compression fittings would probably be required to be compatible with the increased strength of the core /we are verifying this].Because the Dove/EHS would be identical in appearance to the standard Dove used elsewhere on the system, there is a risk that the standard Dove conductor or fittings could be inadvertently installed in lieu of the Dove/EHS conductor or fittings for repairs. Sag tension obtained data from ALCAN for Dove/EHS shows a 7 ft smaller sag than standard Dove for an average span of 833 ft in a 1000 ft ruling span section in Loading Zone 1 and a 10 ft smaller nominal structure height. Given the high time labor premium for working above 80 feet this sag savings could more than make up for the increased cost of structures. It is worth pointing out that any final line optimization run should account for this high time labor premium. 2. 605 kcmil ACSR 30/19 stranding, Code Name "Teal" Teal ACSR has a nominally larger overall diameter of 0.994 inches compared to Dove ACSR at 0.927 inches. The higher content of steel in Teal ACSR gives it markedly improved sag characteristics over Dove ACSR for the same loading conditions. In Tables 4-6a thru 4-6c the maximum sags for Dove and Teal for Loading Zones 1,2 and 3 are given along with nominal structure height based on direct embedment foundations. Teal allows the use of 10-15 ft shorter structures than Dove. Teal ACSR will allow spanning out to 1000 ft in a 1200 ft ruling with an 80 ft structure; Dove ACSR on an 80 ft typical structure would be limited to spans just over 850 ft. In addition to the structure height savings, there would be lower power losses on the line. The disadvantages to using Teal are (1) new requirements for spare materials and (2) higher tensions and marginally higher design wind and ice loads. Additional sag benefits could be obtained with an EHS core. 5. Steel-supported aluminum conductor, SSAC", Teal stranding 605 kcmil SSAC™ uses a standard steel core (high strength or extra high strength) with soft temper (alloy 1350- O temper) aluminum outer strands, forcing the steel core to take up most or all of the conductor mechanical load throughout the loading regime. It appears essentially the same as its standard ACSR counterpart. An SSAC conductor has a high degree of self-damping capability which is enhanced by prestretching; 15 the conductor also shows a high degree of resistance to vibration fatigue. Since the aluminum wires are fully annealed (i.e. soft temper) there is no reduction in conductor rated strength over time when operated at emergency current levels, although there are high temperature limits that must be observed to avoid damage to core coverings. The soft temper of SSAC aluminum wires gives it 63% conductivity compared to the standard EC-H19 temper of 61% conductivity based on copper conductivity 100%. Indications are that standard conductor fittings for the comparable ACSR conductor can be used successfully with the SSAC™ conductor. Tensions for the same sags as standard ACSR will be lower. The disadvantages of Teal SSAC are (1) new requirements for spare materials, (2) higher cost for conductor and (3) a higher cost to install if prestretching is specified. Additional sag benefits can be obtained by specifying an EHS steel core. Sag tension data for Teal SSAC was obtained from CABLEC. Table 4-6a for Loading Zone 1 shows that Teal SSAC has a 0-5 ft [verify with CABLEC] structure height advantage- and marginally higher tensions- compared to Dove ACSR. 4. Alumoweld Conductor Alumoweld strand has been frequently used in Alaska where severe loadings are expected (e.g. Glennallen-Valdez 138-kV line in Thompson Pass uses 19 No. 5 AW, Tyee Lake - Petersburg 138-kV line at high elevations uses 37 No. 8 AW, etc.). .CWEA has experienced a major avalanche failure in Thompson Pass when the alumoweld conductor did not break and took out several structures which were outside the avalanche zone. The recurrence of this type of incident can be countered by carefully routing around possible landslide chutes, placing sacrificial structures and low strength conductors/assemblies in the chute where unavoidable, and installing strong structures either side of the chutes with comparatively weak deadend strings which would fail before cascading could occur. Alumoweld conductor shows much less sag than ACSR or SSAC™ conductors for the same loading conditions. The disadvantages of alumoweld include (1) much higher tensions, (2) higher design loadings and (3) greater power losses. Alumoweld conductor would only be considered for use in Loading Zone 3 and spans over 2000 ft if encountered. Based on the above discussion and relative merits of the conductor type considered,we will assume for the feasibility study the use of 605 Teal ACSR in Loading Zones 1,2 and 3. The scope of the feasibility study does not permit detailed conductor optimization. We recommend that such an optimization be performed in the preliminary engineering phase of a final project. % Failure Containment and Sequence A decision to implement a failure containment design is based on the premise that it is too costly and perhaps unreliable to design the line against any failures for any extreme loading conditions. Recognizing the possibility of severe extreme loading conditions in Alaska, failure containment is a primary concern of all experienced line designers in Alaska. Several different approaches have been taken ranging from guyed X structures designed to have their guy yokes yield under high unbalanced longitudinal load (e.g. Tyee-Wrangell) to the frequent application of longitudinally strong structures (e.g. Swan Lake ). 16 The authors discuss at length a design philosophy to contain failures and achieve a specified failure sequence based on a CVEA-stated preference that "loadings that develop high conductor tenions should result in conductor failure before structure failure." The basis for this preference is the widespread failure that occurred due to an avalanche in Thompson Ridge Pass in 1991[verify date]; it is reported that the high strength 19 No. 5 Alumoweld conductor (URS 73350 Ibs) used in the pass did not break during the avalanche and tore down several structures which otherwise might not have failed. We agree with the authors point that coordinating conductor and structure failure is difficult to implement with confidence. We further agree with their rephrasing of the design requirements : "... the failure of an element on the line should not be allowed to propagate into the failure of other elements such that the cost of the failure is disproportionate to the probablility of its occurring." Implementing this philosphy is made difficult by a lack of knowledge of the probability of extreme loading conditions. The failure containment philosophy in [1], i.e. based on the premise of accepting limited failure and containing failures by providing periodic longitudinally and tranvsersely strong structures, is commonly used and comparatively easily implemented. The authors discuss the nature of failures due to vertical, transverse and longitudinal loading and the practicality of limiting those failures by conventional, cost-effective means. Some key elements of their discussion are summarized below. The authors correctly state that relying on the conductor as the weak link in the system is not reliable due to the nature of ACSR stress-strain at high tensions and that other elements of the system would have to be substantially overdesigned to accommodate this weak link. We agree with the conclusion that depending on the conductor to fail before structures is not a practical or desirable solution. A failure containment approach is discussed which accepts a premise of limited failures and containment of those failures by periodic and strategically placed strong structures. To limit failures due to primarily vertical overloading a fuse link concept is proposed. The link would be located at the line end of suspension insulator strings on tangent and light angle structures only, and would allow the conductor to separate before damaging the structure. The failure sequence does depend on the probability of failures of the different components, and testing to verify or establish the probability distribution is recommended. Limiting failures due to transverse loading is more unpredictable. Failure of angle structures would indeed feed comparatively large amounts of slack into adjacent spans, leading to large longitudinal unbalance and probable cascading failures. These structures must be made strong. The variability of transverse loading along the line, e.g. wind span variation, makes it very difficult to implement a fuse link concept. The authors recommend accepting limited failures of tangent structures under extreme transverse wind loading, while containing those failures by periodic logitudinally and transversely strong structures. This is a common, practical approach requiring full deadend capability in containment structures. Avalanches are a major concern, especially in backcountry routes. Where practical unavoidable avalanche chutes should be spanned and where not practical structures should be made "sacrificial." Strong containment structures would be used either side of known or suspected chutes or with deadend fuse links. This is a practical approach. Where spanning avalanche chutes with the expectation of avoiding the forces of moving rock or snow, high frontal avalanche winds should nevertheless be considered for the span. The use of fuse links, propagation damping and containment are discussed to limit failures due to unbalanced 17 longitudinal loads. Using fuse links will require structures with high longitudinal strength to withstand the dynamic release of conductor tension from one side of the structure and the residual longitudinal unbalance due to the remaining conductor which may or may not drop. This method should be used to prevent the conductor from exceeding its serviceable, working limit. However, if the fuse link is designed to act at or before the working limit and if the working limit is applied to reasonably probable loading conditions, nuisance failures can be expected. Propagation damping relies basically on insulator swing and structure deflection to relieve longitudinal unbalance loads. This feature is inherent in a line with suspension strings and unguyed H-frame or guyed X-frame construction. The level of damping may be fairly well predicted for given loading cases and structure load-deflection data; typically, longitudinal unbalance due to the extreme case of a broken conductor will die out after 5-10 spans in a tangent run. Containment relies on strong deadend structures to withstand longitudinal loads. This is the most reliable method of limiting failures due to unbalanced longitudinal loads. In this feasibility study we will not attempt to account for a detailed failure sequence philosophy such as presented in [1]. We will include failure containment by applying rules for the placement of deadend structures as discussed in the following section. 8. Comments on Proposed Failure Sequence Design The proposed failure sequence, while reasonable in general, has the following shortcomings. First, the failure sequence specifies that conductor attachments and insulator string assemblies will be designed to withstand all load cases combined V, T and L loads with NESC load factor 2.00 ( actually a strength factor of 50%, Rule 277). This factored load is then used as the basis for calculating loads on all other structure components. Application of NESC load factors to extreme loads will distort the strength requirements for the insulator string and structure components. While this imparts strength to the line it can also be costly in materials. It can also be unreliable in predicting a failure sequence. Rather than relate all loads to NESC values, we recommend for final design a load factor reliability design approach in which ultimate loads are used in conjunction with component or line importance factors and the ultimate strengths of system components [5] to establish line performance. Design loads would be compared to NESC Rule 252 loads to verify compliance. Second, the rules for applying longitudinal loads are not clear, especially regarding crossarm and pole strength. Very strong, full deadend structures are apparently designed to cases 1,3 and 4a. Intermediate strength structures meeting 2 and 4b and standard tangents with reduced longitudinal pole strength are mentioned. The rule states that "a series of structures designed with an (longitudinal) OCF as low as 0.50 may meet the longitudinal strengths by insertion of a structure with poles meeting the...requirements of 1 and 4a (full deadends)...(which) shall be inserted into the line so that no more than 10 consecutive structures .... that do not meet ... point 4b (intermediate strength structure) exist." This implies that the use of intermediate strength structures is all that is required. We would recommend that this be clarified in the final design. For the feasibility study we will not apply a detailed failure sequence or compute design loads for achieving an order of strength. We will determine maximum loadings for each loading zone and type of structure and apply rules for placement of longitudinally strong structures, i.e. full deadends, at intervals along the line as follows: (a) deadend spans longer than 1500 feet and (b) deadend at least every 10000 feet. In final design adjacent span ratios 18 and a span length limit in terms of ruling or average span would be used as rules for applying deadends in addition to (a) and (b) above and the requirement for deadend capacity at heavy angles. 10. Electrical Clearances The authors cite NESC requirements for electrical clearances to grade, Rule 232. Under winter clearance conditions, to account for snow cover, surveying tolerances and construction variances, 5 feet is added to clearances over land and driveways; 2 feet is added for other categories of grade. Under summer clearance conditions 2 feet is added to clearances. These margins assume 4 feet of snow cover where applicable and the requisite NESC adder for voltage. A minor adjustment for elevations greater than 3300 feet as required by NESC for some portions of backcountry routes was made for this feasibility study. We have made some adjustments to the clearances in [1] for this study as shown in Table 4-7. Table 4-7 Electrical Clearances (ft) NESC Net Winter Net Winter Table 232-1 Low Elevation High Elevation Nature of Crossing Clearance’ Clearance™* Clearance*? Railway 26.5 30.6 NA Major Road 18.5 22.6 NA Minor Road 18.5 26.6 32.8 Driveway Land 18.5 26.6 32.8 Accessible to Vehicles Land 14.5 22.6 28.8 Inaccessible to Vehicles Water Bodies 17.0 21.1 27.2 No Boating Notes 1. Includes NESC reference, electrical and mechanical components of clearance. 2. Includes voltage adder of 2.1 feet and 2.0 feet for survey and construction variances. 3) Includes adder for elevation at 3% times the voltage adder per 1000 feet in excess of 3300 feet, assuming 5300 feet maximum elevation. 4. Includes 4.0 feet snow cover. 5. Includes 5.0 feet snow cover and 5.0 feet margin to account for structure deflection. 19 We will determine the above-clearances under maximum sag conditions either at NESC Heavy ice loading ,i.e. 0.5 radial incles of ice at 32F, or maximum operating temperature, 150 F. On studies for the Bradley Lake transmission line maximum vehicle height on snow-covered terrain was determined to be about 12 ft (Sno-Cat with antenna). Added to the 5 ft snow cover this results in a reference height of 17 ft, compared to the standard NESC reference height of 14 ft derived from state regulations limiting vehicle heights. In the clearances in Table 4-7 we have applied the NESC reference height on top of the snow cover, giving an extra margin of 2 ft clearance. This extra margin is desirable in part to account for increased sag due to structure and insulator deflection. The substantial amount of snowmobiling in backcountry areas in Loading Zone 2 makes it very important to consider clearances under extreme loading conditions. The situation of very high sags in remote areas of the Tyee Lake 138-kV line (i.e. on Vank and Woronkofski Islands) cannot be tolerated on the Intertie because of the snowmobile activity. Means to provide extra clearance include shortening spans, reducing the longitudinal flexibility of structures, raising structure height, using inverted V insulator strings, and perhaps other methods. In the case of actual contact with the snow, ground fault current may be severely limited by high resistivity snow and too low to trip the line on the CVEA end. Methods to deliberately short the line to ground with interset structures or underbuilt ground wires could be used but are considered much less desirable than providing extra clearance margins. Where practical, the line should be routed away from established snowmobile routes, recognizing though that the right-of-way itself will be an attractive new route for some snowmobilers. To account for the increase in sag under extreme ice loading consider the case of Teal ACSR, Loading Zone 2, and a 1000 ft span. The sag under 1.5 inches radial ice at 0 F (37.6 ft) is 12.9 ft greater than the sag at 0.5 inches radial ice at 32 F (24.7 ft) under which Table 4-7 clearances are applied. The basic clearance to grade for high country clearances was 27.8 ft in Table 4-7 before considering structure deflection. The clearance to 5 ft snow would be 9.8 ft without structure deflection. A very small total deflection will yield a much larger increase in sag. For example, an increase in slack due to structure or insulator deflection of about | ft will lead to an increase of about 5 ft in sag. A 1 ft limit on deflection is stringent. To account for deflection, we would not recommend less than 10 ft clearance to the 5 ft snow cover under extreme ice and therefore have added 5 ft to the basic clearance margin in Table 4-7. NESC controls other design values such as working clearances, climbing space, clearances to adjacent structures, clearances over wires carried on different supporting structures,etc. Reference [1] is not correct in implying that NESC controls phase-phase clearances; no value is specified for line to ground voltages greater than 50 kV. 11. Structure Families The authors selected as the basic structure for most of the line (i.e. Construction Zones 1,3 and 4) an unbraced Hframe structure fabricated using standard self-weathering steel poles based on wood pole class equivalents, i.e. the Meyer LD series. For Construction Zone 2 a single steel pole tangent structure was selected. We agree with the selection of self-weathering steel pole structures; they have been used 20 extensively in Alaska with apparent success. Steel offers significant strength to weight ratio and erection benefits when compared to wood. Given current volatile market conditions, the supply of wood poles in required lengths and classes at reasonable cost is not at all assured. In addition, the variability of wood fiber requires the application of significantly higher NESC overload capacity factors which would further complicate the failure sequence design. The use of standard steel structures however, may lead to unacceptably high structure deflection for loads specified in the failure sequence section. Limits on structure deflection would be desirable as part of the overall study of clearances and failure sequences. The selection of H-frame, 3-pole and single pole structures for a structure family is acceptable although we have concerns about using the standard LD series of poles without extreme ice loading deflection limits. We prefer and will assume for all construction zones the use of tangent guyed X- frame structures, commonly used in Alaska and which offer advantages in flexibility, inherent longitudinal capacity and perhaps in long term maintenance. We will assume guyed 3-pole structures for angles and deadends. Both the X-frame and the H-frame structure will permit long span construction and reduce the number of structure sites/foundations. The unbraced H-frame will require moment foundations, making it necessary to size piles with comparable section modulus to that of the steel poles; the X- frame by its nature has foundation design controlled typically by compression/uplift forces. To the extent that frost jacking can be overcome by design, bracing could increase the span capability of the H-frame and allow use of foundations controlled by axial loading. The X-frame structure is forgiving in jacking situations, tilting until outboard piles can be lengthened and redriven. Redriving X-frame Hpiles would be easier than redriving pipe pile under H-frame poles. The X-frame structure is also lighter than a comparable self-supporting H-frame and perhaps more cheaply installed by helicopter. The X-frame does require driving four piles at the foundations and guy anchor points but the use of Hpile foundations and anchors allow foundation installation to be performed quite a bit ahead of tower erection; use of direct embedded H-frames would require augering and shoring holes with a caisson, then covering up until the H-frame is erected. It is expected that installation of the X-frames would go much more quickly than the H-frames, even with the guy attachments, due to the additional difficulty of leveling H-frame structures and backfilling holes properly /verify/. 12. Foundations, Guys and Anchors The authors propose direct embedment of structures in granular soils and erection on driven pipe piles for structures in permafrost and muskeg locations. Pipe piles were selected in [1] due to their omnidirectional strength properties, required to match the strength of H-frame poles. H-piles were eliminated because of their typical application with the weak axis resisting longitudinal loads. The authors recommend study of other pile options with the promise of large cost savings. Rock may be encountered in Construction Zone | and we will assume rock-anchored foundation types for the feasibility study. A typical pile length of 20 is assumed in [1]. For the feasibility study, we will assume that X-frame structures and 3-pole structures in muskeg are supported on Hpile foundations. Assuming an active layer of 3-5 ft in permafrost areas and applying a rule of thumb of permafrost penetration of two times the active layer and adding a 4 ft stick-up, the length of pile of 20 ft will be assumed in this study. We will assume that guy anchors are 20 ft driven Hpiles in muskeg, 15 ft driven Hpiles in 21 granular soils or 8 ft rock anchors. Hpile is assumed as 10x57/verify/. 13. Insulator Assemblies Reference [1] uses polymer insulators for all assemblies. We agree with this selection and will adopt for the feasibility study. X-frame structures use outboard I-strings and a center V-string. We would recommend consideration of inverted V assemblies to limit conductor slack increase. For the feasibility study, insulator maximum working load (i.e. equal to the routine test load "RTL" of polymeric insulators) will not be exceeded under NESC Heavy loading conditions. Because of concern over long term extreme loading effects on the strength rating of polymer insulators and the underlying assumption that the importance of this line increases with time, we will assume that under any extreme loading condition the insulators will not be loaded to more than 125% of the RTL. 14. Right-of-Way Width Right-of-way width provides for clear construction of a line, containment of energized conductors under wind conditions (e.g. blowout), access for maintenance, and removal of encroaching vegetation and danger trees. Conductors must not extend beyond right-of-way limits under wind conditions for safety reasons. This requires that wider right-of-way be acquired for longer spans. Recently electric and magnetic field limits have played a role in determining right-of-way widths and public concern over possible health effects has made magnetic fields an issue in the design and siting of an Intertic. However, no field limits are used to determine right-of-way width for an Intertie and no State regulations limit fields at this time. We have accepted the line routing criterion that magnetic fields in occupied structures should not be measurably greater than they would be without an Intertie, i.e. practically meaning the increase should be calculated at less than 0.2 milligauss. Reference [1] assumes a 100 ft right-of-way. Right-of-way width is determined by the need to contain the line and its conductors within the right-of-way under moderate and extreme wind conditions. It is our practice to determine right-of-way width based on conductor blowout at 60 mph wind (6PSF) maintaining NESC Rule 234 clearances to the edge of right-of-way or under extreme wind conditions with a margin of 1-2 ft. Structure deflection must also be considered. Based on computations of sag and blowout for Teal ACSR, a 100 ft right-of-way would severely limit span lengths. A /50-175 ft right-of-way will accommodate 1200 ft spans in a 1000 ft ruling span section and is judged the minimum desirable for the span lengths being considered for an Intertie. Typical right-of-way widths on crosscountry transmission lines at 115-138 kV are on the order of 150-180 ft (e.g. Tyee Lake, Swan Lake), with wide ranges from less than 100 ft to 300 ft to accommodate special circumstances. Wider rights-of-way widths are often dictated by the need to clear numerous danger trees. The Intertie will occupy several different types of terrain and forest lands and the final right-of-way width can be tailored for each section of line. In Construction Zone 1 with tall stands of cottonwood, the right-of-way can be expected to be wider than 175 ft due to danger trees. An average /50 ft right-of-way will accommodate the range of variation that can be expected. The right-of-way will be clear cut. Timber will be cut and stacked for use by nearby residents or land owners, lopped and scattered, or burned. It has been suggested that stable, stunted black spruce forest east of Eureka could remain. This is not desirable because of the increasing threat of fire due to the bark beetle infestation and conductor damage that would result; also this would lead to a more visually prominent line. 22 15. Construction Zones Reference [1] divides the line into four construction zones to account for different design or construction conditions. Data below refers to the alignment chosen for study in [1]. Construction Zone allocations for different route segments in the present study are indicated on the summary Table 15-1. Zone | :from Sutton to Mile 100 approximately where it heads up Pinocle Creek, distance 40.34 miles, elevation 500-2500 feet, Hframe construction, good land access, year round construction possible, no permafrost but rock and earth foundations, dense forest (100%)of birch, cottonwood, white spruce; Zone 2 :from Mile 100 to Tahneta Pass, distance 24.38 miles, elevation up to 3200 feet, single pole construction, helicopter construction, no permafrost but rock and earth foundations, lightly forested (50%)spruce and cottonwood, non-winter construction only; Zone 3 :from Tahneta Pass to Moose Lake, distance 47.81 miles, elevation 2000-3000 feet, extensive permafrost (80% assumed), good access, Hframe construction pile foundations mostly, transition to all-winter construction in Zone 4, sparsely forested (25%)mostly black spruce; and Zone 4 :from Moose Lake to Glennallen, distance 21.74 miles, elevation 1600-2000 feet, 100% permafrost, Hframe construction pile foundations, 100% forest cover with 30-40 ft black spruce, good access, winter construction only. The division of the line into these zones for estimating purposes and project planning is reasonable and will be adopted generally for this feasibility study with modification for use with other structure types. ‘Alaska Enorgy Authority Copper Valley Intortie Feasibility Study Comperison of Conductor Sags For 32.6 ft clearance X frames 6/11/93 PED Sage in Foot DOVE —TEAL DOVE MRISO MRI 50 MRI 75 1.3 Mex Seg Mex Seg Delta. =~ Max Sag Average Ruling Span 667 800 19.88 17.14 2.74 19.97 750 900 23.46 20.48 2.98 26.23 833 1000 27.40 [24.22 3.18 33.34 917 1100 31.65 | 26.26 | 3.39 41.28 1000 1200 37.68 | 32.57 5.31 50.04 1083 1300 44.99 |_37.15 7.84 59.64 Average span based on RS = Lavg-2/3(Lmex-Lavg) Boxed sags are under 0.5 inches radial ice at 32 F, no wind. Unboxed sags ere at 150 F for Dove and Teal, 160F for Alumoweld. Inoulator L 5h Clesrence 32.8 ft Top Pole VA Embed Add On Embed % 0% 667 800 19.88 17.14 2.74 19.97 Sag for Avg Span 13.81 11.90 1.90 13.87 Length Exposed 62.61 50.70 52.67 Total Length 52.61 50.70 52.67 Nominal 56 50 55 750 900 23.46 20.48 2.98 26.23 Sag for Avg Sp 16.29 © 14.22 2.07 18.22 Length Exposed 55.09 53.02 57.02 Total Length 55.09 53.02 57.02 Nominal 55 55 60 833 1000 27.4 24.22 3.18 33.34 Sag for Avg Span 19.03 16.82 2.21 23.15 Length Exposed 57.83 55.62 61.95 Total Length 57.83 56.62 61.95 Nominal 60 85 65 917 1100 31.65 28.26 3.39 41.28 Sag for Avg Spen 21.98 19.63 2.35 28.67 Length Exposed 60.78 = 58.43 67.47 Total Length 60.78 58.43 67.47 Nominal 60 60 70 1000 1200 37.88 92.67 5.31 50.04 Sag for Avg Span 26.31 22.62 3.69 34.75 Length Exposed 65.11 61.42 73.55 Total Length 65.11 61.42 73.85 Nominal 65 65 78 1083 1300 44.99 37.15 7.84 59.64 Sag for Avg Span 31.24 25.80 5.44 41.42 Length Exposed 70.04 64.60 80.22 Total Length 70.08 64.60 80.22 Nomina! 70 65 80 XFSAGHT.XLS. TEAL DOVE —_—TEAL DOVE —TEAL MRI 75 (\21XTM 121. XTM \22XTM 122 XTM Mex Seg Delta. © Max Sag Max Seg Delta. = Max Sag Max Sag 16.98 3.03 33.97 21.89 | 1238 [ 2687 22.72 20.63 5.60 | 45.59 29.52 | 16.07 | 34.94 27.56 24.65 8.69 88.75 39.40 | 19.35 | 45.28 32.78 29.64 | 11.64 | 73.48 50.55 | 22.93 | 56.88 38.41 36.44 | 13.60 | #961 62.92 | 26.89 | 69.75 48.03 43.92 | 15.72 [107.60 76.50 _} 31.30 |_g3.89 58.74 16.94 3.03 33.97 21.59 1238 26.87 22.72 11.76 210 23.69 14.99 8.60 18.66 15.78 50.56 62.39 3.79 57.46 54.8 50.56 62.39 63.79 57.46 54.58 50 65 55 60 55 20.63 5.6 45.59 29.52, 16.07 34.94 27.56 14.33 3.89 31.66 © 20.50 11.16 24.26 19.14 83.13 70.46 59.30 63.06 57.94 53.13 70.46 59.30 63.06 57.94 55 70 60 65 60 24.65 8.69 58.75 39.4 19.35 45.28 32.78 17.12 6.03 4080 27.36 1344 31.44 22.76 55.92 79.60 66.16 70.24 61.56 55.92 79.60 66.16 70.24 61.56 55 80 70 70 65 29.64 11.64 = 73.48 = 60.55 22.93 56.88 = 38.41 20.58 8.08 51.03 36.10 © '15.92 39.50 26.67 59.38 89.83 73.90 78.30 65.47 59.38 89.83 73.90 78.30 65.47 60 90 75 80 65 36.44 13.6 89.81 62.92 26.89 «69.75 48.03 25.31 9.44 62.37 43.69 «18.67, «48.44 (33.35 64.11 101.17 82.49 87.24 = 72.15 64.11 101.17 82.49 87.24 72.15 65 105 85 90 78 43.92 15.72 107.8 76.5 31.3 63.89 58.74 30.50 1092 74.86 «= 53.13 21.74 = 88.26 = 40.79 69.30 113.66 91.93 97.06 79.59 69.30 113.66 91.93 97.06 79.69 70 115 95 100 80 Page 1 Delta 4.15 7.38 12.50 18.47 21.72 25.18 4.18 2.88 7.38 5.13 12.5 8.68 18.47 12.83 21.72 15.08 25.15 17.47 DOVE \23 XTM Max Sag 36.09 47.58 61.68 76.39 93.28 160.45 36.09 25.06 63.86 63.86 65 47.58 33.04 71.84 71.84 78 61.68 42.83 81.63 81.63 85 76.39 53.05 91.85 91.85 95 93.28 64.78 103.58 103.58 105 160.45 111.42 150.22 180.22 150 TEAL 23 x™™ Max Seg 24.03 32.52 42.19 53.00 65.31 116.43 24.03 16.69 55.49 55.49 55 32.52 22.58 61.38 61.38 65 42.19 29.30 68.10 68.10 70 53.00 36.81 75.61 75.61 75 65.31 45.35 84.15 84.15 85 116.43 80.85 119.65 119.65 120 Delta 12.06 15.06 19.49 23.39 27.97 44.02 12.08 8.38 15,06 10.46 19.49 13.53 23.39 16.24 27.97 19.42 44.02 30.57 37 No9AW DOVE (23 XTM Mex Seg Info Only 10.67 12.92 15.33 17.88 20.58 23.42 10.67 7.41 46.21 46.21 60 12.92 8.97 47.77 47.77 50 15.33 10.65 49.45 49.45 50 17.88 12.42 51.22 61.22 55 20.58 14.29 53.00 53.09 85 23.42 16.26 55.06 55.06 55 PEEnst Max Seg 19,68 23.48 27.52 33.39 40.37 48,02 19.68 13.67 82.47 52.47 55 23.48 16.31 65.11 55.11 65 27.82 19.11 57.91 57.91 60 33.39 23.19 61.99 61.99 65 40.37 28.03 66.83 66.83 70 48.02 33.35 72.18 72.15 78 TEAL PEEast Mex Seg 16.99 20.45 24.22 28.26 32.87 37.18 16.99 11.80 50.60 50.60 20.45 14.20 53.00 53.00 55 24.22 16.82 55.62 85.62 55 28.26 19.63 58.43 58.43 32.57 22.62 61.42 61.42 65 37.18 25.80 64.60 64.60 65 Delta 2.69 3.03 3.30 5.13 7.80 10.87 2.69 1.87 3.03 2.10 3.3 2.29 5.13 3.66 7.8 §.42 10.87 7.85 DOVE PEWest Max Sag 19.8 23.41 27.44 32.05 38.64 45.98 19.8 13.75 52.56 52.55 55 23.41 26.83 65.63 65.63 65 27.44 19.08 57.86 57.86 32.05 22.26 61.06 61.06 65 38.64 26.83 65.63 65.63 65 45.98 31.93 70.73 70.73 70 TEAL PEWest Max Seg 17.08 20.45 24.22 28.26 32.87 37.18 17.08 11.86 50.66 50.66 20.45 22.62 61.42 61.42 65 24.22 16.82 85.62 95.62 85 28.26 19.63 88.43 88.43 60 32.57 22.62 61.42 61.42 65 37.15 28.80 64.60 64,60 6s Dette 2.72 2.96 3.22 3.79 6.07 8.83 2.96 4.22 3.22 2.24 3.79 2.63 6.07 4,22 8.63 6.13 Alaska Enorgy Authority Copper Valley intortie Feasibility Study Comparison of Conductor Sags For DE 3-Pole STr 6/11/93 PED Sage in Foot DOVE TEAL Dove MRI5O — MRI5O MRI 75 1.3 Mex Seg Max Seg Delta = Max Sag Average Ruling Span 667 800 19.88 17.14 2.74 19.97 750 900 23.46 20.48 2.98 26.23 833 1000 27.40 [24.22 3.18 33.34 917 1100 31.65 | 28.26 3.39 41.28 1000 1200 37.68 | 32.57 5.31 50.04 1083 1300 44.99 |_37.15 7.84 59.64 Average span based on RS = Lavg-2/3(Lmax-Lavg) under 0.5 inches radial ice at 32 F, no wind. 160F for Alumoweld. 5 ft 32.8 ft if 3h 10% 667 800 19.88 17.14 2.74 19.97 Sag for Avg Span 13.81 11.90 1.90 13.87 Length Exposed 52.61 0.70 52.67 Total Length 61.78 59.67 61.85 Nominal 65 60 65 750 900 23.46 20.48 2.98 26.23 Sag for Avg Span 16.29 14.22 2.07 18.22 Length Exposed 55.09 53.02 7.02 Total Length 64.55 62.25 66.68 Nominal 65 65 70 833 1000 27.4 24,223.18 33.34 Sag for Avg Span 19.03 16.82 2.21 23.15 Length Exposed 57.83 56.62 61.95 Total Length 67.59 65.13 72.17 Nominal 70 65 78 917 1100 31.65 28.26 3.39 41.28 Sag for Avg Span 21.98 19.63 2.38 28.67 Length Exposed 60.78 68.43 67.47 Total Length 70.87 68.25 78.30 Nominal 70 70 80 1000 1200 37.68 32.57 5.31 50.04 Seg for Avg Span 26.31 22.62 3.69 34.75 Length Exposed 65.11 61.42 73.55 Total Length 75.67 71.58 85.06 Nominal 75 75 85 1083 1300 44.99 37.15 7.84 59.64 Sag for Avg Span 31.24 25.80 5.44 41.42 Length Exposed 70.04 64.60 80.22 Total Length 81.16 75.11 92.46 Nominal 85 75 95 SAGCOMP.XLS TEAL DOVE TEAL bove TEAL MRI 75 \21.XTM 121. XTM 22 XTM = 122 XTM Mex Seg Delta. = Max Seg Max Seg Delta, = Max Sag Max Sag Delta 16.94 3.03 33.97 21.69 | 1238 [ 2687 22.72 4.15 20.63 5.60 45.59 20.52 | 16.07 | 34.94 27.56 7.38 24.65 8.69 88.75 3940 | 19.35 | 45.28 32.78 | 12.50 29.64 11.64 | 73.48 60.65 | 2293 | S668 38.41 18.47 36.44 13.60 | 89.81 62.92 | 26.89 | 69.75 48.03 | 21.72 43.92 | 15.72 [107.80 76.50} 31.30 |_s389 58.74 | 25.15 16.94 3.03 33.97 21.69 12.38 26.87 22.72 4.15 11.76 2.10 23.59 14.99 8.60 18.66 15.78 2.88 50.56 62.39 63.79 57.46 54.58 59.52 72.66 63.10 67.18 63.98 60 78 65 70 65 20.63 5.6 45.59 29.52 16.07 «34.94 27.56 7.38 14.33 3.89 31.66 = 20.80 11.16 24.26 19.14 5.13 53.13 70.46 59.30 63.06 57.94 62.36 81.62 69.22 73.40 67.71 65 85 70 75 70 24.65 8.69 58.75 39.4 19.35 45.28 32.78 12.6 17.12 6.03 40.80 27.36 13.44 31.44 ~—22.76 8.68 55.92 79.60 66.16 70.24 61.56 65.46 91.78 76.85 81.38 n.74 65 95 80 85 78 29.64 11.64 73.48 6085 22.93 56.88 36.41 18.47 20.58 8.08 51.03 36.10 18.92 39.50 26.67 12.83 59.38 89.83 73.90 78.30 65.47 69.31 103.14 85.45 90.33 76.08 70 105 85 90 80 36.44 13.6 89.81 62.92 26.89 69.75 48.03 = 21.72 28.31 9.44 62.37 43.69 18.67 48.44 «33.35 ‘15.08 64.11 101.17 82.49 87.24 72.15 74.56 116.74 94.99 100.26 63.50 75 115 95 100 85 43.92 15.72 107.8 76.5 31.3 83.89 88.74 25.18 30.50 10.92 74.86 §3.13 21.74 $8.26 40.79 17.47 69.30 113.66 91.93 97.06 79.59 80.33 129.62 105.47 111.17 91.77 80 130 105 115, 95 Page 1 DOVE TEAL 123 XTMLZ3 XTM. Mex Sag Max Seg 36.09 24.03 47.88 32.52 61.68 42.19 76.39 63.00 93.28 65.31 160.45 116.43 36.09 24,03 25.06 16.69 63.86 55.49 74.29 64.99 78 65 47.58 32.52 33.04 22.68 71.84 61.38 83.16 71.54 85 75 61.68 42.19 42.83 29.30 81.63 68.10 94.04 79.00 95 80 76.39 3.00 53.05 36.81 91.85 75.61 105.39 87.34 105 90 93.28 65.31 64.78 45.35 103.68 84.15 118.42 96.84 120 100 160.45 116.43 111.42 80.85 150.22 119.65 170.25 136.28 170 140 Delta 12.08 15.06 19.49 23.39 27.97 44,02 12.06 8.38 15.08 10.48 19.49 13.53 23.39 16.24 27.97 19.42 44.02 30.57 37 No9AW DOVE (23 XT™M. Mex Sag Info Only 10.67 12.92 15.33 17.88 20.58 23.42 10.67 7.41 46.21 54.68 55 12.92 8.97 47.77 56.41 60 15.33 10.65 49.45 58.27 17.88 12.42 51.22 60.24 20.58 14.29 53.00 62.32 65 23.42 16.26 55.06 64.52 65 PEEset Mox Sag 19.68 23.48 27.52 33.39 40.37 48.02 19.68 13.67 82.47 61.63 65 23.48 16.31 86.11 64.56 65 27.52 19.11 67.91 67.68 70 33,39 23.19 61.99 72.21 78 40.37 28.03 66.83 77.59 48.02 33.35 72.15 83.50 85 TEAL PEEast Max Sag 16.99 20.45 24.22 28.26 32.67 37.18 16.99 11.80 50.60 59.55 20.45 14.20 53.00 62.22 65 24.22 16.82 55.62 65.13 65 28.26 19.63 58.43 68.25 70 32.87 22.62 61.42 71.68 8 37.15 28.80 64.60 78.11 78 Delte 2.69 3.03 3.30 5.13 7.80 10.87 2.69 1.87 3.03 2.10 3.3 2.29 6.13 3.56 7.8 6.42 10.87 7.85 Dove PEWest Mex Seg 19.8 23.41 27.44 32.05 38.64 45.98 19.8 13.75 52.55 61.72 65 23.41 26.83 65.63 76.26 80 27.44 19.06 57.86 67.62 70 32.05 22.26 61.06 nA 78 38.64 26.83 65.63 76.26 45.98 31.93 70.73 81.92 TEAL PEWest Max Sag 17.08 20.45 24.22 28.26 32.57 37.15 17.08 11.86 50.66 59.62 60 20.45 22.62 61.42 71.58 75 24.22 16.82 55.62 65.13 65 28.26 19.63 58.43 68.25 70 32.57 22.62 61.42 71.58 78 37.15 28.80 64,60 75.11 78 Delta 2.72 2.96 3.22 3.79 6.07 8.83 2.96 4.22 3.22 2.24 3.79 2.63 6.07 4.22 8.83 6.13 COPPER VALLEY INTERTIE FEASIBILITY STUDY ALASKA ENERGY AUTHORITY SEGMENT SUMMARY All Length Minimum | Maximum Loading Construction MRI Route AltA Route Alt B Route Alt C Segments miles Elevation Elevation Zone Zone Pls Segment Segment Segment 1-2 5.56 655 1100 1 1 10-11 10-11 4.25 2500 2600 3 2 na 10-15 12.61 2600 4300 3 2 na 11-12 4.74 2500 2900 3 2 na 12-13 3.14 2700 3000 3 3 29-30 13-14 3.3 3200 4000 3 3 30-31 14-15 1.91 3400 3700 3 3 30-31 14-16 6.2 3200 3500 2 3 30-31 15-16 4.57 3200 4400 3 3 na 15-17 6.26] 3500 3800 3 3 na 16-18 5.89 2950 3200 2 3 32-33 17-19 9.06] — 2600 3900 2 3 na 18-21 13.75 2350 3800 3 3 36-37? 19-20 12.13 2300 3365 2 3 37-38? 2-3 11.36 1100 1900 1 1 10-11 2-31 6.58 800 1125 1 1 10-11__| 20-22 4.99 2250 2725 2 3 38-39 [21-23 5.19 2200 2317 2 3 38-39 22-26 13.56] 2400 3107 Zz 3 39-46 23-24 6.02| 2317 2660__| 2 3 39-46 24-25 4.88 2660 3000 2 3 39-46 25-26 2.27 2400 2850__| 2 3 39-46 26-27 7.97 2000 2200 2 4 39-46 27-28 6.44 1720 2100 2 4 39-46 27-29 9.05 1682 2172 2 4 39-46 28-29 2.6 1682 1720 2 4 39-46 29-30 7.32 1400 1682 Z 4 39-46 3-4 6.68 1100 2200 1 1 12-13-14 31-3 5.89 1100 1900 1 1 10-11 4-5 7.22 2200 3000 1 1 14-15 4-7 13.53 2200 3400 1 1 na 5-6 9.27 2900 3000 1 2 14-15-16 6-8 8.73 2830 3500 3 2 na 6-9 4.71 2200 3400 1 2 17-20 12.52 4900 3 2 F 2800 3 2 9-11 7.37 3 2 21-26? Total Lengths 134.03 133.00 135.62 # of segments 16 21 23 SEGMENTS.XLS 1 of 1 6/18/93