茨城大学重点研究「知的で持続可能な社会基盤および防災セキュリティ技術研究創出事業」茨城大学工学部附属防災セキュリティ技術教育研究センター 2 0 1 5年度報告書茨城大学重点研究プロジェクト「知的で持続可能な社会基盤および防災セキュリティ技術創出事業」平成27年度報告書刊行にあたってプロジェクト代表本研究課題は，平成23年4月に課題募集プロポーザルとその審査により茨城大学重点研究課題として認定され、工学部、教育学部および茨城大学センター教員から構成される異分野研究者の集う場として活動が始まりました.本年度は活動５年目を迎えることとなりました.この報告書では，学術誌論文をはじめとする研究成果に加えて，参加メンバーの学術企画の開催や参加，受賞例についても整理しましたが，「防災セキュリティ技術」という学際領域に類するテーマが示すように，昨年度に引き続き，多種多様な研究活動とその成果を収めてきたことがわかりました。 2015年を廟みますと，9月には我々茨城大学が所在する茨城県内で甚大な自然災害が発生しました. 「平成27年9月関東・東北豪雨」に伴う洪水被害です.茨城県西部を流下する鬼怒川が氾濫したことにより，多くの住宅や公共施設，そして耕作地が被害を受け，人命にまで影響が及ぶ事態となりました.大規模な豪雨の影響は,河川・湖沼の氾濫だけに留まらず，広く流域全体としての自然・社会環境への影響もあり，定常時から社会インフラをセンシングしておくことの重要性を再認識する１年となりました．今後，本教育研究センターでは，最終年の研究取りまとめに向けて，主として「統合化」をキーワードとした研究展開を進めるプランを持っております.今後、センターで得られた成果がより多くの社会基盤維持管理の分野で利用または応用展開が進むよう，メンバー一同，努力を続けて行きたいと考えております.末筆とはなりますが，茨城大学重点研究課題として採択頂き，茨城大学を代表する研究課題の一つとして諸方面の応援と援助を頂きました茨城大学に心から感謝申し上げますとともに，必ずしも十分でなかった研究交流にも関わらず，本誌に示す多大なる研究成果を上げている参加メンバーに心から敬意と謝意を表します.今後の研究活動への努力をお約束し、関連する皆様に感謝を申し上げますとともに，ここに平成27年度の研究成果を報告させて頂きます.平成28年３月吉日プロジェクト代表茨城大学工学部附属防災セキュリティ技術教育研究センター （2015年度）プロジェクト参加教員

1.活動概要
− 1 −
2.研究報告
1. Ibrahim, A., Wu, Z., Fahmy, M., and Kamal, D, Experimental study on cyclic response of concrete bridge columns reinforced by steel and basalt FRP reinforcements, ASCE Journal of Composites for Construction, ISSN: 1943-5614, (2015-10). ー 6 ー 2. Hiroshi Noguchi, Yasuhiro Ohtaki, Masaru Kamada, Design and Practice of File Backup System Taking Advantage of Remotely Distributed Campuses, Network-Based Information Systems (NBiS), pp. 694 - 697, (2015-9) — 2 5 — 3.山崎正稔,石内鉄平,桑原祐史, EOS-Terra/ASTERを用いたマングローブ城抽出に関する研究一ミャンマー沿岸域を対象として一,日本沿岸域学会誌, Vol.28, No.3, pp.73-82, (2015-12) — 2 9 —
3.プロジェクト業績
活動実績− 3 9 −業績一覧— 4  3 — 1.活動概要

1. 技術･研究開発分野
担当者氏名:県智深Email : zhishen.wu.prof@vc.ibaraki.ac.jp ・計画名:社会基盤を対象とした維持管理および防災/減災へ向けたセンサ技術・空間情報応用研究実施予定時期:平成27年4月1日～平成28年3月31日･防災･環境教育の実施と地域還元のためのツール開発実施予定時期:平成27年4月1日～平成28年3月31日•システムとしての防災・減災技術の確立検討（企業・自治体・学術連携）実施予定時期:平成27年4月1日～平成28年3月31日
1. 実施内容:
•センサ技術の開発光ファイバセンサ・カーボンファイバセンサ・RFIDタグ・UAV・防災/気象用センサICタグ(水位・加速度・傾斜・降雨・大気関連） •センサデ一タの安定した通信/解析技術の開発と省電力化推進マルチホップ通信による耐災害メッセージングシステム津波を受ける構造物の外力評価•センサ技術の現地実証の推進と課題抽出超小型衛星を用いた情報利用技術の検討(継続課題)老朽化したインフラ構造物の災害リスクを考慮した雑持管理計画(継続課題) ・UAVを活用した新たな空間情報の防災システムへの活用と超小型衛星との連携
・地元企業・学校等との学術連携実現・他大学・企業・自治体連携での研究資金獲得・シンポジウム開催（県内・都内） ・自治体との各種連絡会議推進（茨城県・日立市・ひたちなか市等） ・国際共同研究を実施（アメリカ、英国、イタリア、中国、韓国等） ・大型プロジェクトを申請2. 実施体制(注外部の人も含む)責任者:呉智深(センター長)メンバ:齋藤修,鎌田賢,沼尾達弥,今井洋,羽淵裕真,渋沢進,湊淳,横田浩久,原田陸郎,武田茂樹,外岡秀行,桑原祐史,車谷麻緒,石田智行(茨城高専・東北大学・東京大学・青山学院大学の教員の正式参加を調整中） 3.実施における課題:特になし. 4. 実施結果(年度末に記載) (結果) ・茨城県・水戸市連携による、内水氾濫監視システムの実証実験推進（H27年3月に設置準備完了。H28年4月以降に機器設置実施予定） ・茨城県との連携によるUAVのダム長寿命化への応用検討（H27年10月に第一回撮影実施） •茨城県との連携による常陸大宮地区での傾斜計設置。地盤災害への対応検討実証実験（H27年１月より実施） ・茨城県との連携による常陸大宮、引田橋における加速度センサ設置による橋梁長寿命化対応実証実験（H26年6月より継続中） ・株式会社 KSK 構造診断研究所との連携による新潟県、国道18号妙高大橋における光ファイバセンサ設置による橋梁健全性評価手法の実証実験（H27年5月より実施） ・株式会社KSK構造診断研究所との連携による川根大橋におけるカーボンファイバセンサ設置による橋梁振動モニタリング実証実験（H27年11月より実施） ・中国の蘇通長江大橋（橋梁管理局と協力）でのインフラ構造物の早期損傷検知実験（H27年11月より実施）八戸工大：橋梁長寿命化研究での連携青山学院大学: UAV飛行におけるプロトコル確立研究連携東京大学・長岡技術大学：橋梁長寿命化についての情報交換会ひたちなか市立外野小学校・前渡小学校での環境・情報特別授業実施（H27年9月）宮城県名取市関上地区のUAVによる震災被害地アーカイブ映像撮影　宮城教育大学との連携宮城県女川町・石巻市連携による地盤沈下調査（来年度継続）・UAVによる震災アーカイブ映像撮影日本地球惑星連合(JPGU) 2015 :セッション「UAVが拓く新しい世界」コンビーナとしてUAVの活用を推進。茨城県生活環境部・土木学会茨城支部・㈱新星コンサルタント等においてUAV操作指導・安全教育を実施。常総市水害（９月１０日）のUAVによる調査・記録・分析に協力（国交省委託事業の一環） 「建設フェスタ2014」(2014年11月9日開催）においてUAV技術展示・成果PRを実施(課題）　特になし6.その他(参考資料、報告書など)
2.人材育成
担当者氏名: 呉智深1.計画名・実施予定時期Email :  zhishen.wu.prof@vc.ibaraki.ac.jp計画名:「センサシステム開発-情報ネットワーク-データ解析法」に関する授業の学部～社会人への展開実施予定時期:平成27年4月1日～平成28年3月31日2. 実施内容:下記の学内授業実施および学生国際会議企画を行い、研究で得た最新情報を学部学生ヘフィードバックし、また、社会人ドクターの教育を通じて高度化する. ①学部学生に対する教育通信理論,ソフトウェア実現,空間情報工学②大学院博士前期課程学生に対する教育構造物の数値解析法,衛星画像および地理情報の解析法,社会基盤情報処理特論③社会人（専門技術者向）に対する教育土木学会茨城会で行っているイブニングセミナ(地域の技術者対象)に防災・センシング関連で講座を出す④その他:小中高校への環境。防災教育の提供3. 実施体制(注外部の人も含む)責任者:呉智深(センター長)メンバ:鎌田賢,齋藤修,沼尾達弥,今井洋,羽淵裕真,渋沢進,湊淳,桑原祐史,横田浩久,原田隆郎,武田茂樹,外岡秀行,車谷麻緒,石田智行
4.実施における課題:  特になし
5. 実施結果(年度末に記載) (結果) ・茨城県「いばらき近未来実証推進事業」プロジェクト推進委員会委員（アドバイザー）として、UAVを含むロボット実証試験に安全管理・実証方法についてアドバイスを行っている（齋藤、桑原、鎌田）。 (課題）特になし6.その他(参考資料、報告書など)
3.資金獲得
担当者氏名：呉智深1.計画名・実施予定時期Email : zhishen.wu.prof@vc.ibaraki.ac.jp計画名:大型プロジェクトの獲得.実施予定時期:平成27年4月1日～平成28年3月31日. 2. 実施内容:科研費はセンターメンバー個々人がチャレンジすることとし,それ以外の大型プロジェクト(国土交通省,経済産業省,総務省)に,共同研究等で関係のある産官学学外機関との連携を取り,チヤレンジする.昨年度に不採択となった概算要求書をベースとして生かす。 3. 実施体制(注外部の人も含む)責任者吳智深(センター長)メンバ:鎌田賢,齋藤修,沼尾達弥,今井洋,羽淵裕真,渋沢進,湊淳,桑原祐史,模田浩久,原田隆郎,武田茂樹,外岡秀行,車谷麻緒,石田智行4. 実施における課題:大型プロジェクトの申請に関する作業時間をより確保する必要がある。 5. 実施結果(年度末に記載) (結果) ・平成26年度募集SCAT研究費助成獲得　青山学院大学戸辺教授との連名「複数異種UAV間協調動作プロトコルの開発と広域環境観測への応用」複数のUAVを協調動作させる。防災への応用が期待できる。(H27年度から3年) (課題)特になし6.その他(参考資料、報告書など) 2.研究報告(H27年度参加教員発表の代表的な学術論文集)
Arafa M. A. Ibrahim ¹; Zhishen Wu, M.ASCE ²; Mohamed F. M. Fahmy, Ph.D. ³; and Doaa Kamaı ⁴ Abstract: This paper presents the seismic performance of concrete bridge columns reinforced with both steel and fiber-reinforced polymer (FRP). A bond-based parametric experimental study was conducted on five FRP steel–reinforced concrete (FSRC) bridge columns—using basalt FRP (BFRP) bars and BFRP sheets—and two steel-reinforced concrete (SRC) bridge columns served as references to investigate the fundamental characteristics of the proposed reinforcement. The investigated bond parameters included the texture of the FRP bars (smooth and ribbed), diameter of the FRP bars, location of the FRP bars with respect to the steel bars, and application of external FRP confinement. All columns were tested under the combined effect of constant axial load and reversed cyclic loading. The experimental results showed that the FSRC bridge column could realize the existence of a stable postyield stiffness (hardening behavior) as well as a reasonable displacement ductility of up to 10 before encountering strength degradation. Moreover, the FRP bars added for column longitudinal reinforcement did not have a substantial impact on the column elastic stiffness. The bond condition of the FRP bars to the surrounding concrete had pronounced effects on the column failure mode, postyield stiffness, residual displacement, and ductility; thus, it could be adopted as a design parameter. DOI: 10.1061/(ASCE)CC.1943-5614.0006614. © 2015 American Society of Civil Engineers. Author keywords: Basalt fiber-reinforced polymer (FRP); Bridge column; Ductility; Damage control; Bond; Postyield stiffness.
Introduction
Ductility, energy dissipation, and strength demand are important seismic performance indices in seismic design of modern concrete bridge columns (Kawashima et al. 1998 ). In addition, seismic per- formance of bridges during and after an earthquake is considerably dependent on the importance level of the bridge. For instance, col- umns of essential bridges should open quickly to all traffic (quickly come back to the original function) after a massive earthquake (AASHTO 2011 ). Several studies such as Christopoulos et al. (2003 ) and Pettinga et al. ( 2007 ) addressed the importance of the existence of structure postyield stiffness, i.e., ability of structure to continue carrying load after yielding. Actually, the existence of postyield stiffness was considered a seismic performance index (Christopoulos and Pampanin 2004) because the structure residual inclination (permanent deformation) and damage level are dependent on the slope of the postyield stiffness. Other studies (e.g., Priestley et al. 1996; Wu et al. 2009) emphasized that the uncontrollable damage of conventional steel-reinforced concrete (SRC) structures results from the elastoplastic characteristics of ordinary steel bars. To guarantee the existence of postyield stiffness and control damage of a structure, Pettinga et al. ( 2007 ) suggested coupling the main structural system with a more flexible secondary system, in which the secondary system is designed to undergo elastic deformations while the main system undergoes inelastic deformations. During the last two decades, several experimental and analytical investigations have been conducted with the aim of control of the postyield stiffness and mitigation of the postearthquake damage and residual displacement of concrete bridge columns. In this di- rection, Iemura et al. ( 2006 ) incorporated unbonded high-strength tendons in bridge piers to realize a structure with stable postyield stiffness. To enhance the self-centering capability (i.e., the ability of the structure to return to its original configuration after seismic events) of bridge piers, Sakai et al. (2006) and Zatar and Mutsuyoshi (2002) used posttensioned tendons. To avoid the need for mechanical anchors, Davis et al. ( 2012 ) used unbonded preten- sioned strands. Other studies (e.g., Trono et al. 2015; Dawood et al. 2012 ; ElGawady and Sha'lan 2011 ; Chou and Chen 2006 ) realized the self-centering of concrete bridge piers by adopting posttensioned segmental systems. However, this system has a low energy dissipa- tion capacity compared with conventional monolithic columns (Nikbakht et al. 2013 ). Therefore, studies by ElGawady and Sha’lan ( 2011 ), Marriott et al. ( 2009 ), Currie et al. ( 2009 ), and others sup- plied the system with internal and/or external energy dissipaters. Saiidi and Wang ( 2006 ), Saiidi et al. ( 2009 ), and Tazarv and Saiidi ( 2013 ) used shape memory alloy as longitudinal reinforce- ment of concrete bridge columns to mitigate residual displacement and plastic hinge damage. However, the high cost, lack of clear understanding of thermomechanical processing, and difficulty in machining of shape memory alloys would bar its expanded use (Desroches and Smith 2003 ; Varela and Saiidi 2014).
04015062-1
In recent years, numerous research findings evidenced the success of external confinement of concrete columns using fiber-reinforced polymer (FRP) jackets to improve shear, bond, and ductile behavior of concrete columns (Seible et al. 1997; Xiao et al. 1999; Hamad et al. 2004; Haroun and Elsanadedy 2005; Wu et al. 2006; Choi et al. 2013). Zhu et al. ( 2006 ) encased a single- segment posttensioned column in a concrete-filled FRP polymer tube to enhance its ductility, and ElGawady et al. ( 2010 ) used the same procedure for segmental concrete columns. The columns reached high drift levels with no strength degradation. Despite the effectiveness of FRP confinement to increase the concrete compres- sion strength and restrain the lateral expansions on the core concrete, Fahmy et al. ( 2010a ) concluded that concrete confine- ment is not the technique that directly controls the column postyield stiffness and residual displacement. Wu et al. (2009, 2010a) proposed a steel-FRP composite bar (SFCB) as a technique that can be applied in the field of civil en- gineering to combine the advantageous mechanical and physical properties of both steel and FRP composites. By reinforcing RC bridge columns with SFCB, it would be suitable to control their postyield stiffness and residual displacement. Furthermore, the postyield stiffness depended on the properties of the fibers used, whereby columns reinforced with steel-basalt fiber composite bars exhibited a larger drift capacity before the rupture of the basalt fi- bers than those reinforced with steel–carbon fīber composite bars (Fahmy et al. 2010b). Although scale model columns reinforced with this type of composite bars exhibit a favorable seismic perfor- mance with damage-controllable states after yielding, the produc- tion of comparable composite bars for practical application with larger inner steel cores may require an excessive amount of outer longitudinal fiber material, which would complicate the production process; thus, using many composite bars with a smaller inner steel core would be necessary. In addition, the bond performance be- tween composite bars and concrete must be further improved. Figs. 1(a and b) present the schematic reinforcement of two conventional SRC structures; namely SRC1 and SRC2. Design of SRC1 and SRC2 was assumed to satisfy the strength demands by moderate and strong earthquake, respectively. Therefore, addi- tional increase in the concrete dimensions and/or steel reinforce- ment is necessary in SRC2, comparing with that of SRC1. This paper proposes coupling the system of the SRC1 with additional FRP reinforcement [Fig. 1(c)]. Fig. 2 presents a detailed mechani- cal model of the expected performance of the proposed FRP-steel reinforced concrete (FSRC) system in comparison with the two SRC systems. This model is a modified form of the proposed behavior of FRP-RC damage-controllable structure defined by Fahmy et al. ( 2009 ), in which further developments meeting the requirements of modern codes for new structures are considered. The FSRC system is proposed for the life safety performance ob- jective because the structure would be able to withstand moderate earthquakes with elastic performance, and the demand strength by strong earthquake could be ensured by the existence of a stable postyield stiffness, K 2 (Fig. 2 ). Prior to yielding of steel reinforce- ment, FSRC exhibits initial elastic stiffness, K 1 , similar to that of SRC1, but less than that of SRC2. Comparing with SRC systems, FSRC system is characterized by a stability zone of the peak strength, along the line M-S [zero stiffness (K 3 )], whereby the structure system demonstrates a desirable ductile performance be- fore degradation of strength. Failure of the FSRC system is defined when the contribution of FRP bars to the lateral resistance is com- pletely lost (at point F). From the viewpoint of postearthquake re- coverability, because the FSRC structure contains elastic members and has positive postyield stiffness, both damage level and residual deformations could be controlled. Recently, basalt FRPs (BFRPs) have emerged as a promising alternative to conventional FRPs in reinforcing concrete structures (El Refai et al. 2015). BFRPs show advantageous characteristics in mechanical, chemical, and high ratio of performance to cost in comparison to other FRPs. For instance, BFRP has a higher strength and modulus, a similar cost, and more chemical stability compared with E-glass FRP; a wider range of working tempera- tures and much lower cost than carbon FRP (CFRP); and greater than five times the strength and approximately one-third the density of commonly used low-carbon steel bars (Wu et al. 2010b; Sim et al. 2005; Palmieri et al. 2009). Because of the aforementioned characteristics, BFRP was nominated for application in the FSRC system in this study.
Research Significance
To control postyield performance of concrete bridge columns, this study proposed combining steel and FRP reinforcement for con- crete structures. Bond between FRP bars and the surrounding con- crete was adopted as a key parameter that affects the structural behavior of the proposed column. The objectives of this research were threefold: (1) assessing the cyclic response of RC bridge col- umns reinforced by steel and FRP reinforcements; (2) examining Fig. 1. Schematic reinforcement: (a) SRC1 structure; (b) SRC2 structure; (c) proposed FSRC structure 04015062-2 from ascelibrary.org by University of Waterloo on 10/13/15. Copyright ASCE. For personal use only; all the effect of bond conditions between FRP bars and concrete on the structural performance of the proposed column; and (3) investigat- ing replacement of some steel confinement reinforcement with external FRP jacket. To achieve these objectives, a total of seven one-sixth scale columns were constructed and tested under the effect of combined constant axial load and reversed cyclic load.
Experimental Program
An experimental investigation on concrete bridge columns rein- forced with both FRP and steel reinforcement was conducted. One column (CS-4%) was designed as a reference for an RC bridge column that can safely resist strong earthquakes, and another col- umn (CS-2%) was considered to resist nearly half the lateral force of the first column. A complete design was also provided for an FSRC concrete column with the same concrete parameters and steel reinforcement of CS-2% in addition to FRP reinforcement. The subsequent section describes the design parameters of both the steel and FRP reinforcements.
Design of Column Specimens
General Design Parameters
A cantilever bridge column with an overall height of 1.0 m, a cross section of 200 × 200 mm, and a distance from the column base to the point of the lateral load application of 850 mm, yielding an aspect ratio of 4.25, was proposed for this experimental study. These dimensions approximately represent a one-sixth scale of a typical bridge column and were selected based on the available fa- cility capacity of the laboratory, where the tests were carried out. A target compressive strength ( f ζ ) of 30 MPa was applied to all © ASCE led from ascelibrary.org by University of Waterloo on 10/13/15. Copyright ASCE. For personal use only; a examined columns. The yield strength (f v ) and ultimate strength of the 13-mm-diameter deformed steel bars used as longitudinal reinforcements were 375 and 560 MPa, respectively; and the cor- responding values for the 6-mm-diameter steel transverse reinforce- ments were 400 and 625 MPa, respectively (Table 1). BFRP bars and BFRP sheet with the mechanical material properties shown in Table 1 were used to investigate the proposed FSRC design. Before applying the proposed cyclic loading that will be explained later, an axial load was applied on all column specimens. Typically, the axial load owing to the superstructure dead weight in bridge columns varies between 5 and 10% of the capacity of the columns. However, in the study of Park ( 1996 ), the weight of the bridge superstructure. was estimated from the dimensions of the superstructure and col- umns to cause an axial compression load ratio at the bottom of col- umn specimens of approximately 3%. Other laboratory studies (e.g., Kawashima et al. 1998; Zatar and Mutsuyoshi 2002) assumed an axial compression stress of 1 Mpa to represent the dead weight
sectional area of each bar and based on the manufacturer, the basalt fiber content was 60% of the cross-sectional area.
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J. Compos. Constr. of a highway bridge's superstructure in Japan. Therefore, in this study, a constant axial load of approximately 40 kN (corresponding to 3.3% of the nominal capacity of the column) was applied to each column specimen, inducing an axial compression stress of 1 MPa. Using the aforementioned data, a detailed design of the tested spec- imens is described next.
Longitudinal  Reinforcement
AASHTO ( 2012 ) specifies that the area of the longitudinal reinforcement ( A t ) of columns located in a high-seismic-hazard area is to be no less than 0.01 or more than 0.04 times the gross cross-sectional area A g (i.e., 0.01 ≤ ρ 1 ≤ 0.04, where ρ 1 is the steel reinforcement ratio and equal to A t / A q ). In this study, the first steel- reinforced column (CS-4%) was reinforced with 12 steel bars of 13-mm diameter (i.e., ρι = ρ tmax = 4%, where ρ tmax is the maxi- mum steel reinforcement ratio), whereas the second column (CS-2%) used six longitudinal steel bars of 13-mm diameter (i.e., ρ 1 = 2%). The theoretical lateral strengths of these two sam- ples were 67.0 and 36 kN, respectively, and both values were de- fined using the AASHTO ( 2012 ) rectangular stress block for concrete in compression, which has a mean stress of 0.85 f ′ , an ultimate concrete compression strain of 0.003, and a steel stress of f v for the longitudinal steel bars. The FSRC column was de- signed to have the same steel reinforcement as column CS-2%, as stated previously. For this column, to reach a strength compa- rable to that of column CS-4%, additional FRP bars were added. Assuming full contribution of the used BFRP bars (i.e., design strain equals the ultimate tensile strain), four BFRP bars of 10-mm diam- eter or six BFRP bars of 8-mm diameter would theoretically increase the column strength to the desired value. By summing the area of steel and BFRP bars, the resultant longitudinal reinforcement ratio, ρ 1 , became 2.8%. To prevent problems of plastic hinge relocation, anchorage failure, or other failure modes, the BFRP bars in this study were extended to a height of 700 mm from the column base and embedded in the column footing to a depth of 300 mm.
Transverse Reinforcement
In this experimental program, the amount of transverse reinforce- ment was controlled by the following requirements: (1) confine the plastic hinge regions and prevent early buckling of the longitudinal reinforcement; (2) provide sufficient shear strength; and (3) prevent buckling of FRP bars and improve their bond behavior. To guar- antee the aforementioned requirements, the transverse reinforce- ments in the form of internal steel stirrups and external FRP wrapping were designed as described next. According to AASHTO ( 2012 ), 6-mm-diameter steel stirrups with a spacing of 25 mm should be provided to the plastic hinge region of the columns. Because this amount of transverse reinforce- ment is independent on the demanded ductility and may result in congestion of reinforcing cages and create concreting problems, 50% of the inner transverse steel reinforcement was replaced with an external FRP jacket. Based on analytical and experimental re- sults, Wehbe et al. ( 1999 ) proposed a design expression relating the amount of confining reinforcement to attainable displacement ductility as shown in the following:
where A sh = area of transverse reinforcement in each of the trans- verse directions; h c = dimension of the concrete core of the section measured perpendicular to the direction of the hoop bars to the outside of the perimeter hoop; S h = center-to-center vertical spacing of the hoops; A g = gross area of the column cross section; f cn = 27.6 MPa; f sn = 414 MPa; f vl = yield stress of longitudinal steel reinforcement; f ξ = concrete compressive strength; P = axial load; and μΔ = displacement ductility factor. According to Eq. ( 1 ) and using transverse reinforcement of 6-mm-diameter stirrups with a spacing of 50 mm, the resultant μΔ was approximately 8. Referring to the relationship between the curvature and displacement ductility factors of Park and Paulay ( 1975 ) and using a reasonable value of the expected plastic hinge length ( L p ), the equivalent curvature ductility, fみ, of approximately 11 could be obtained. To determine the length of the plastic hinge region, Eq. (2) proposed by Paulay and Priestley ( 1992 ) was firstly investigated to give a length of 175 mm
where L = height of the column to contra-flexure point; and d b and f γ l = diameter and yield stress of longitudinal bars, respectively. How- ever, it was reported by Sheikh and Khoury (1997) that L p is approx- imately equal to the cross section size of the confined concrete columns. Therefore, the value of L p was assumed to be 200 mm. To determine the required thickness of the BFRP jacket in the plastic hinge region, the demanded ductility for the proposed struc- ture was first identified based on the available literature. Sheikh and Khoury (1997) suggested values of μ δ ≥ 16, 16 ≥ μ δ ≥ 8, and μ ♂ δ 8 for concrete columns with high ductility, moderate duc- tility, and low ductility, respectively. The New Zealand Society for Earthquake Engineering (NZSEE 2006) recommended values of μ δ ≥ 20 and μ δ ≥ 10 for ductile and limited ductile columns, re- spectively. Wehbe et al. ( 1999 ) used values for μ Δ ≥ 10 for high levels of ductility. In this study, to assess the structural performance of the proposed FSRC columns up to high levels of ductility, a value of μ α = 20 (μΔ > 10) was assumed. Consequently, a BFRP jacket should be applied to the plastic hinge of the columns to in- crease μ ς by a value equal to 9. Therefore, the developed procedure. of Sheikh and Li (2007) [Eq. (3)] was adopted to determine the required external FRP jackets
where t F = total thickness of FRP jacket; f F u = ultimate tensile strength of FRP jacket; β = 0.25 for square columns; h = dimension normal to the loading direction of rectangular columns; k p is the nominal axial load ratio; and μ ò ,in = μ ò = μ ò ,org , in which μ φ and μ φ ,org are the curvature ductility factors after and before FRP jacketing, respectively. Hence, by assuming a material strength reduction factor of the BFRP jacket equal to 0.65 (CSA 2012 ), wrapping the plastic hinge region of the columns with a 0.666-mm-thick BFRP jacket was sufficient to achieve the demanded ductility. At high levels of ductility, in the end regions of concrete columns, for small values of axial compression, the contribution of concrete to resist shear forces is very small ( AASHTO 2012 ). Consequently, by providing the columns with only the internal steel stirrups, the shear capacity of the columns was approximately the same as the theoretical flexural strength of the columns with the highest reinforcement ratio. Therefore, to ensure a dominant flexu- ral failure mode, only half of the calculated BFRP jacket thickness, which was required for plastic hinge confinement, was reasonably extended to a length of 1.5 times the cross-section size of the col- umns (Seible et al. 1995) (i.e., a 0.333-mm-thick BFRP jacket was extended up to 300 mm height above the column base). Outside the end regions, the contributions from both the transverse steel © ASCE 04015062-4 reinforcement and the concrete to resist shear forces were enough to prevent any expected shear failure. Ultimately, for FSRC columns with BFRP bars placed in the concrete cover, outside of the steel stirrups, the 0.333-mm-thick BFRP jackets were extended up to a height of 600 mm above the column base. However, because the sensitive design of FRP jacketing for bond purposes is out of the scope of the present study and may be the focus of other works, it is not pursued further in this work.
Test Specimens and Experimental Parameters
Seven different column units were tested to meet the objectives of this study. All column units had a deep heavily reinforced concrete base of 1.0 × 0.5 × 0.5 m (length × width × depth), which simulated a rigid foundation for the tested column. To avoid any unexpected local failure at the loading region, the transverse steel reinforcements were spaced at 30 mm in the highest 300-mm portion of the column units. All columns were attached to a strong steel floor using four vertical high-strength steel rods. The geometry and instrumentation of a typical column unit are shown in Fig. 3. A detailed description of the test specimens follows and is supported by Figs. 3–5 and Table 2: . Specimen CS-2% [Fig. 4(a)] served as a reference specimen for the steel RC columns. In this column, the longitudinal steel reinforcement consisted of six 13-mm-diameter bars (i.e., ρ 1 = 0.02), and the transverse reinforcement consisted of 6-mm-diameter internal closed stirrups spaced every 50 mm. Specimen CS-2%-J [Fig. 4(b)] was reinforced with the same steel reinforcement as specimen CS-2%. In addition, a BFRP jacket with 0.333-mm thickness was applied to the lowest 300 mm of the column (i.e., L j1 = 300 mm) and another jacket of 0.333-mm thickness was added to only the lowest 200 mm of the column portion (i.e., L j2 = 200 mm). In addition to the steel reinforcement details of column CS-2%, specimen CSF-2.8%-IS-D10 [Fig. 4(c)] was reinforced with two 10-mm-diameter BFRP bars placed on each of the two opposite sides of the column (those with the highest tension/ compression) at the same place as the longitudinal steel bars. The surface texture of the BFRP bars contained small prefab- ricated indentations (factory product) [Fig. 5(b)]. Specimen CSF-2.8%-IS-D10-J [Fig. 4(d)] was similar to speci- men CSF-2.8%-IS-D10 but wrapped at the lowest region with the same BFRP jacketing as in column CS-2%-J. Specimen CSF-2.8%-IR-D10-J [Fig. 4(d)] was the same as specimen CSF-2.8%-IS-D10-J, but the surface texture of the BFRP bars was spirally roughened with BFRP strips in two perpendicular directions [Fig. 5(c)]. Specimen CSF-2.8%-ES-D10-J [Fig. 4(e)] was the same as spe- cimen CSF-2.8%-IS-D10-J, but the BFRP bars were placed out- side the steel stirrups, i.e., in the concrete cover. In addition, a BFRP jacket of 0.333-mm thickness was applied first to the low- est 600 mm of the column portion (i.e., L j1 = 600 mm) and then another jacket of 0.333-mm thickness was added to only the lowest 200 mm of the column portion (i.e., L i2 = 200 mm). Specimen CSF-2.8%-ES-D8-J [Fig. 4(f)] was reinforced with both steel and BFRP reinforcement and wrapped with the BFRP jacketing as in specimen CSF-2.8%-ES-D10-J, but the FRP re- inforcement consisted of three 8-mm-diameter BFRP bars placed external to the steel stirrups on each of the two opposite sides of the column (those with the highest tension/compression) in the concrete cover. The surface texture of the BFRP bars contained small indentations [Fig. 5(a)]. The total reinforcement ratio in this column was approximately 2.8% (i.e., ρ 1 = 0.028).
Loading and Instrumentation
All columns were subjected to a constant axial load of 40 kN and several excursions of lateral cyclic loading using a dynamic actua- tor with a capacity of 700 kN. The reversed cyclic loading sequence was determined based on the column displacement at the yielding load ( Δy ), which was numerically defined for the reference column and kept the same for all specimens. The lateral loading sequence Fig. 3. Typical test unit (dimensions are in mm) and instrumentation of a test specimen 04015062-5 from ascelibrary.org by University of Waterloo on 10/13/15. Copyright ASCE. For personal use only; all om ascelibrary org by University of Waterloo on 10/13/15. Copyright ASCE. For personal use only; started with two cycles of 0.5Δy followed by two cycles of Δy and then three cycles each of 2Δy, 3Δy, 4Δy, 6Δy, 8Δy, and 10Δy, until failure. A linear variable differential transformer (LVDT) was used during testing to record the horizontal displacement of the tested columns. The axial strain histories for both the steel and FRP reinforcements were recorded during the test by using a set of 5-mm-long strain gauges, arranged as shown in Fig. 6. Although efforts were made to keep the axial load constant during the exper- imental tests, laterally displacing the columns resulted in some variations in the axial load, particularly at high levels of lateral displacement. The actual applied axial load was in the range of 40–100 kN during the tests. All installed instrumentations are shown in Fig. 3 .
Experimental Results and Discussion
General Observations and Hysteretic Curves
In this section, the results of each column specimen are individually discussed with reference to its hysteretic response (V-δ curve) and failure mode. Fig. 7 shows the lateral load versus the column drift ratio for all tested columns, and Fig. 8 shows the final failure model of these columns. The deformation capacity of the columns is ex- pressed as the member lateral drift, which is defined as the ratio of the lateral displacement at the point of the load application of each column to the effective height of the column (850 mm). The lon- gitudinal strains of the FRP bars are also examined, where the lat- eral load-FRP bar strain hysteresis loops were recorded by the FRP bar's strain gauge in each column (Fig. 9 ). Table 3 summarizes the characteristic values and experimental findings in both the positive and negative loading directions together with the observed failure. mode of each specimen, and Table 4 contains the average character- istic values of both directions. The terms V cr and V v represent the cracking load and steel yielding load, respectively, and the terms δ cr and δ v represent the corresponding lateral displacements. The cracking load and corresponding displacement were approximately defined from the first turning points of the load-displacement curves when the plastic hinge zone was covered with an FRP jacket. The yielding loads and displacements were obtained from the results of the strain gauges attached to the longitudinal steel bars. The term V P represents the peak load, and it is characterized by two displacement values δρ 1 and δρ 2 , in which δρ 1 corresponds to the peak load and δ P 2 corresponds to the end of the plateau zone, if any. The terms V u and δ u represent the ultimate load and its cor- responding displacement, respectively, and they were defined for SRC specimens at the drift level at which the load capacity decreased to 80% of the peak load. For the other columns (FSRC columns), the ultimate loads and displacements were defined at the degradation of the peak load to the peak load of the SRC col- umn, column CS-2%. Values of the displacement ductility factor, μ , at different characteristic points were superimposed on the hys- teretic responses (Figs. 7 and 9), where μ = δ/δ , and δ = lateral
04015062-6

Note: f ρ = actual concrete compressive strength on the day of testing; L j 1 = length of the first 0.333-mm-thick FRP jacket above the column base; ρ = total reinforcement ratio of the steel and FRP bars. displacement at the load application point. Moreover, a complete description of the failures noted during the loading of all columns was included in the hysteretic curves using marked footnotes. The next sections provide a detailed discussion of the behavior of all tested columns.
Columns Reinforced with Steel Bars
For the control specimen CS-2%, flexural cracks first occurred near the column base in the first loading cycle at a drift level of approx- imately 0.1%, corresponding to a lateral load of 9.2 kN. While dis- placing the specimen in both loading directions, new horizontal and slightly inclined cracks formed and propagated with further loading and distributed in the lowest 300 mm of the column in both loading directions. The first yielding of the steel bars was observed at a drift level of 0.65%, corresponding to a lateral load of 26.0 kN [Fig. 7(a)]. Afterward, a stable hardening cyclic response appeared and continued up to a drift level of 3.5% ( μ = 5.3), corresponding to an average peak load of 37.5 kN in both the push and pull directions. Following this drift level, a peak-loading horizontal © ASCE 04015062-7 from ascelibrary.org by University of Waterloo on 10/13/15. Copyright ASCE. For personal use only; all plateau was formed with the appearance of significant wide cracks in the plastic hinge regions up to a drift level of 5.9% ( μ = 8.8). A bulking of the concrete covers accompanied with a smooth degra- dation of the cyclic response after this level took place up to a drift level of 7% ( μ = 10.3). Beyond this level, a complete spalling and crushing of the concrete cover within a height of approximately 200 mm above the column footing occurred. After the spalling of the concrete cover, a serious buckling of the longitudinal reinforcement [Fig. 8(a)] was observed, causing a sudden drop in the lateral load and terminating the test. For specimen CS-2%-J, the behavior was significantly affected by wrapping the plastic hinge region with the FRP jacket [Figs. 7(b) and 8(b)]. The use of the BFRP jacket prevented the crack propa- gation from being observed, but the load-displacement curve dem- onstrated that the first turning point, indicating the first cracking, was at a drift level and corresponding lateral load nearly the same as those of the control specimen. The first yielding of the steel reinforcement took place at a drift level of 0.64%, corresponding to a lateral load of 27.2 kN (i.e., V y = 1.05V yc , where V yc is the yielding load of column CS-2%). After the yielding of the steel bars, the column was able to continue carrying a load in a stable manner up to a drift level of 5.9% ( μ = 8.9), corresponding to a peak lateral load of 43.4 kN (i.e., V P = 1.15V PC , where V PC is the average peak load of column CS-2%). Beyond this drift, a gradual bulging of the FRP jacket accompanied by a smooth gradual loss of strength took place up to an average lateral drift of 9.4% ( μ = 14.3), corresponding to a lateral load of 40.8 kN. Upon increasing the lateral displacement beyond this drift level, the bulging of the FRP jacket was accompanied by a local buckling of the steel bars within a height of approximately 100 mm above. the column footing, followed by the rupture of some steel bars [Fig. 8(b)].
Columns Reinforced with Both Steel and FRP
Reinforcements
Before the yielding of the steel bars, the observed behavior of all FSRC columns was slightly affected by the contribution of the FRP reinforcement. This could be because of the small contribution of the FRP bars to both the column strength and deformation, and this was identified from the average longitudinal strain records of BFRP bars located at both loading sides, which ranged from 8.5 to 11% of the uniaxial rupture strain of the BFRP bars (i.e., ε fy = 0.085 to 0.11 ε r , where ε f y is the average strain recorded in the FRP bars at the yielding of the steel bars, and ε r is the rupture strain of the BFRP bar). This small contribution is owing to the small elastic stiffness ratio between FRP and steel (i.e., AfEf/AsEs = 0.095, where A and E are the gross cross-sectional area and elastic modulus, respectively, and subscripts S and f denote the steel and FRP bars, respectively). Beyond the yielding of the steel bars, the contribution of the FRP reinforcement became significant and controllable, where a hardening zone with a clear positive postyield stiffness was realized. Compared with the control specimen, all specimens reinforced with both steel and FRP bars achieved con- siderably higher lateral strength (i.e., 30-85% higher than that of column CS-2%). In contrast to the control column, the failure modes of all FRP-reinforced specimens were never attributed to the buckling of the internal reinforcement because a significant por- tion of the total force in the compression zone was carried by the FRP bars. However, buckling of the longitudinal internal bars occurred abruptly after bond or rupture failure of the FRP bars. The primary observations of the behavior of such columns are subsequently described in detail.
Columns Reinforced with 10-mm-Diameter BFRP Bars
For column CSF-2.8%-IS-D10, small cracks were first observed during the first loading cycle near the column base. With further © ASCE 04015062-8 J. Compos. Constr. ed from ascelibrary.org by University of Waterloo on 10/13/15. Copyright ASCE. For personal use only; all aded from ascelibrary.org by University of Waterloo on 10/13/15. Copyright ASCE. For personal use only; all rights rese loadings, these cracks increased and distributed within the lowest 300 mm of the column height. The applied strain gauges recorded a first yielding strain of the steel bars at a drift level of 0.74%, corresponding to a lateral load of 27.6 kN. Fig. 7(c) shows that beyond the yielding of the steel bars, the column continued carrying loads with a stable postyield stiffness up to a lateral drift of 3.5% ( μ = 4.2), corresponding to a peak load of 48 kN (V P = 1.28V PC ) in one loading direction. In the other loading di- rection, a peak lateral load of 50.5 kN ( V P = 1.35V PC ) was reached at a drift of 4.7% ( μ = 5.6). Within this hardening zone, a consid- erable propagation of cracks appeared and further widened. After reaching the peak load, the concrete cover at the column foot- ing interface and above it started to bulk out, causing a plateau at the peak load level. Along this plateau, concrete crushing and © ASCE 04015062-9 J. Compos. Constr. spalling became significant, particularly within the first 200 mm near the column footing. A popping sound was heard during the loading of the column to a lateral drift of 7% ( μ = 8.5), at which the measured strain ε f of the BFRP bars was 0.47 ε r , indicating the first local bond failure between the FRP bars and surrounding con- crete at the column-footing interface. At this drift level, a 20% de- crease in the load capacity occurred, confirming the effect of slippage between the FRP bars and surrounding concrete. An addi- tional degradation in the column strength of 5% up to the comple- tion of the first loading cycle to a lateral drift of 8.2% ( μ = 9.9) was also observed. A popping sound was heard again when loading to the second and third cycles of the same lateral drift, together with separation of the concrete from the footing, which clearly indicated another local bond failure of the FRP bars. The strain of the FRP bars at this drift level was approximately 54% of the rupture uni- axial strain (i.e., ε f y = 0.54 ε r ) [Fig. 9(a)]. However, as a result of the applied cyclic loading, the partial rupture of the BFRP bars was observed within the first 100 mm above the column base. Conse- quently, another 40% decrease in the carried load occurred. Hence, all the stresses in the longitudinal direction were carried by the steel bars, resulting in a sudden local buckling of some of them and subsequently their rupture. Fig. 8(c) shows the state of column CSF-2.8%-IS-D10 just before the rupture of the steel bars. Figs. 7(d) , 8(d) , and 9(b) show that the observed trim of the behavior of specimen CSF-2.8%-IS-D10-J was similar to that of specimen CSF-2.8%-IS-D10 up to a high loading level. The first yielding of the steel bars was recorded at a lateral load of 28.1 kN, corresponding to a drift of 0.6%. The column achieved a peak load of 56.75 kN (V P = 1.51V PC ) at a drift level of 5.9% ( μ = 6.8) in one loading direction and a peak load of 59 kN ( V P = 1.57V PC ) at a drift of 4.7% ( μ = 5.4) in the opposite loading direction. Beyond the peak load, a stability in this load was observed up to a drift level of 7% ( μ = 8.1), at which point the first bond failure of the FRP bars occurred, causing an average decrease of 15% in the lateral load. Beyond this deformation level, the column maintained the remaining lateral load up to a drift level of 8.2% ( μ = 9.5), at which point a complete debonding of the FRP bars occurred, accompa- nied by another drop in the lateral load by approximately 20%. As a consequence, the steel bars became vulnerable to high stresses, resulting in local buckling and subsequent rupture of some © ASCE
04015062-10
J. Compos. Constr. aded from ascelibrary org by University of Waterloo on 10/13/15. Copyright ASCE. 1
of them, causing a termination of the column test [Fig. 8(d)]. The maximum achieved strain in the FRP bars before the bonding failure was approximately 56% of the rupture strain [Fig. 9(b)]. The response of column CSF-2.8%-ES-D10-J shows the effect of placing FRP bars out of the transverse reinforcement in the con- crete cover [Fig. 7(f)]. Compared with CSF-2.8%-IS-D10-J that was internally reinforced with FRP bars, a slight increase in the carried lateral load could be achieved. This increase was attributed to the increase in the effective depth of the FRP bars (i.e., the dis- tance from the fibers of the maximum compression strain to the center of the FRP bars located on the tensile side). As shown in Fig. 7(f) and summarized in Table 3 , the first yielding of steel bars was recorded at a drift level of 0.67%, corresponding to a lateral load of 33.2 kN. Beyond the yielding of the steel bars, the column showed a gradual increase in the carried lateral load in both loading directions up to drift levels of +4.7% ( μ = 6.7) and −5.9% ( μ = 8.6), corresponding to lateral loads of +63kN (V p = 1.68V pc ) and −60kN (V p = 1.6V pc ), respectively. After- ward, local bond failure of the FRP bars occurred as the bond stress between the FRP bars and surrounding concrete reached the bond strength. Two successive drops of the lateral load were observed in both loading directions, where the lateral load in the two loading directions at a drift level of 9.4% ( μ = 10.3) reached values of +37.5 and −38.3 kN. Before the debonding of the FRP bars, the recorded strain was approximately 61% of the rupture strain [Fig. 9(d)]. The presence of the FRP bars outside the closed stirrups resulted in a pronounced rupture of the external fibers of the BFRP bars, particularly after the bulging of the BFRP jacket [Fig. 8(f)]. Specimen CSF-2.8%-IR-D10 examined the effect of roughen- ing the texture of the BFRP bars on the bond between the FRP bars and concrete. Fig. 7(e) and Table 3 show that the first yielding of steel bars was at a lateral drift of 0.69%, corresponding to a lateral load of 28.1 kN. Beyond the yielding of the steel bars, a hardening © ASCE zone of this column was realized in a manner quite similar to that in column CSF-2.8%-IS-D10, with smooth FRP bars, up to a drift level of 3.5%. After this lateral drift ratio, the hysteretic loops showed a continuous increase in column strength but with a smaller positive stiffness in both loading directions up to the completion of the first loading cycle of the lateral drift of 8.2% ( μ = 11.3), cor- responding to lateral loads in the positive and negative directions of 60 kN (V P = 1.6V PC ) and 69 kN (V P = 1.84V PC ), respectively. Displacing the column in the two directions with the additional two loading cycles to the same lateral drift resulted in an approx- imately 25% loss in the achieved peak lateral strength. This drop was primarily attributable to a complete rupture of the FRP bars at 50 mm above the column base [Fig. 8(e)]. At this loading stage, a loud sound was heard, and a bulge formed in the BFRP jacket. Dur- ing the loading process, the strain gauges located at the section of maximum moment failed early, and only the strain gauge located at 150-mm height continued until the rupture of the BFRP bars. The maximum attained axial strain of the FRP bars recorded by this strain gauge was approximately 71% of the uniaxial rupture strain [Fig. 9(c)]. Following the rupture of the FRP bars, a sudden in- crease in the stresses were experienced by the steel bars, causing some of them to rupture.
Columns Reinforced with 8-mm-Diameter BFRP Bars
The observed behavior of specimen CSF-2.8%-ES-D8-J was highly similar to that of the column reinforced with rough BFRP bars; this could be attributed to good bond conditions between the FRP bars and surrounding concrete in both cases. As shown in Fig. 7(g) and Table 3 , the first yielding of the steel bars was re- corded at a drift level of 0.65%, corresponding to a lateral load of 31.9 kN. Beyond the yielding of the steel bars, this column achieved the largest lateral resistance among all tested columns, 04015062-11 J. Compos. Constr. led from ascelibrary.org by University of Waterloo on 10/13/15. Copyright ASCE. For personal use only; all where at a drift of 7% ( μ = 10.3), the maximum attained lateral strengths in the positive and negative directions were 68 kN (V p = 1.81V pc ) and 71 kN (V p = 1.89V pc ), respectively. At this drift level, a loud sound was heard, indicating the probability of rupture of one or more FRP bars, followed by an approximately 15% decrease in the peak load. The column was then able to main- tain its achieved strength up to a lateral drift of 8.2%, at which point a second loud sound was heard, indicating the probability of rup- ture of the additional FRP bars, accompanied with another 25% decrease in the lateral load, after the completion of the second load- ing cycle at this drift level. Unfortunately, the results of the third loading cycle at this drift level were lost during the transfer of the data from the logger. The continued displacement of the column led to a rupture of some steel bars. The results of this column indicated that the FRP bars fulfilled nearly 94% of its rupture strain before failure [Fig. 9(e)]. Moreover, after removing the FRP jacket at the lowest part of the column, the decreases in the lateral load were confirmed to result from the rupture of the FRP bars [Fig. 8(g)].
Envelope Responses
To investigate the effect of each tested parameter individually, the average skeleton curves of push and pull loading directions of all specimens are shown in five separate figures, where the effect of each parameter can be investigated in one figure [Fig. 10 ]. Refer- ring to Fig. 2, the effectiveness of the proposed column reinforce- ment should be evaluated in light of the targeted mechanical load-displacement model. To this end, the evaluation process is presented in this section in terms of initial stiffness, postyield stiff- ness, stiffness degradation, and ductility measurements [Table 5]. To draw firm conclusions, the response of the steel RC column (CS- 4%), discussed earlier, was numerically simulated using the Open- Sees software (Mazzoni et al. 2009 ). The experimentally applied cyclic loading regime was adopted in the numerical simulation to predict the behavior of column CS-2% for comparison with the ex- perimental results [Fig. 10(a)] and then to predict the performance of column CS-4%. Only the envelope response of the hysteretic curve is presented in this study for comparison with the FSRC columns.
Initial Stiffness
The initial stiffness or elastic stiffness ( K 1 ) is an important seismic performance measure (index) that can be evaluated as K 1 = V v/δv. Higher values of this index would lead to a shorter vibration period of the structure and would generally increase the earthquake forces received (Saiidi et al. 2009). However, as stated by ElGawady and Sha'lan ( 2011 ), during a real earthquake scenario, the stiffness re- duction is anticipated to lengthen the natural period of vibration of the structure, leading to a potential reduction in the seismic demand forces that is associated with larger displacement demands; there- fore, depending on the design approach, there would be a trade-off between the displacement and force demands. K 1 was determined for all tested specimens to investigate the effect of adding FRP reinforcement to the steel reinforcement (Table 5). As shown in Fig. 10 and summarized in Table 5, adding FRP composites to the steel reinforcement resulted in an insignificant change in the initial stiffness values. Although wrapping the plastic hinge region with an FRP jacket resulted in a significant enhance- ment in the column confinement, the increase in the initial stiffness was extremely small: K 1 of column CS-2%-J = 1.06; K 1 of col- umn CS-2%, and K 1 of column CSF-2.8%-IS-D10-J = 1.07; and K 1 of column CS-2%-J= 1.06 times that of column CS-2% and K 1 of column CSF-2.8%-IS-D10-J = 1.07 times that of column CSF-2.8%-IS-D10. Placing additional FRP bars at the same place as the steel bars had no clear effect on the initial stiffness compared with the reference column CS-2%, whereas adding FRP bars of 8- or 10-mm diameter in the concrete cover (outside the transverse reinforcement) resulted in an approximately 20% increase in the initial stiffness. Moreover, the bond condition between the FRP bars and concrete showed no considerable effect on the initial stiff- ness: K 1 of column CSF-2.8%-IR-D10-J ~ K 1 of column CSF- 2.8%-IS-D10-J; and K 1 of column CSF-2.8%-ES-D8-J ~ K 1 of column CSF-2.8%-ES-D10-J. In contrast, Fig. 10(b) and Table 5 show that the initial elastic stiffness of the numerically investigated column CS-4% was more than twice that of the control specimen. That is, increasing the column strength to withstand a strong earth- quake using additional steel reinforcement would greatly affect the elastic stiffness and in turn its vibration period, whereas adding FRP bars to an RC structure would be a reasonable solution to avoid any increase in the imposed seismic force on the structure.
Postyield Stiffness
In this section, the postyielding stage of the tested columns up to the column peak strength was evaluated using two postyield stiffness indices. The first index was the ratio between the column postyield stiffness and elastic stiffness ( k = k , / k 1 %), where the starting point of the postyield stiffness of all tested columns was defined as the theoretical strength of the control specimen CS- 2% defined previously using AASHTO ( 2012 ), whereas the end- point was that corresponding to the maximum achieved lateral strength of the considered specimen (i.e., at a lateral load of V P and the corresponding lateral displacement δ P 1 ) (Fig. 2 ). By defin- ing the postyield stiffness using the aforementioned start and endpoints, the curvature in the hardening zone of the load- displacement curves was idealized to a straight line. The second measure was the column displacement ductility, corresponding to the endpoint of the hardening zone (i.e., μρ 1 = δρ 1 /δρ y ). The two indices were calculated for all specimens and are summarized in Table 5. A detailed explanation concerning the efficiency of the proposed FSRC column design to achieve the desired level of lateral strength (the strength of column CS-4% = 67 kN) as well as the effect of the investigated parameters on the postyield stiffness behavior is as described next. Fig. 10(a) and Table 5 illustrate that control specimen CS-2% continuously increased in lateral strength with the increase in the applied displacement up to a ductility of 5.3 (i.e., μρ 1 = 5.3). The achieved postyield stiffness ratio corresponding to this final duc- tility of the hardening zone was 7.8%. Through wrapping the plastic hinge zone with an FRP jacket, as in column CS-2%-J, the column showed an ability to receive a greater lateral load up to μ P 1 of 9.0, but the postyield stiffness ratio decreased to 6.6%. A substantial decrease in the column postyield stiffness ratio of approximately 52% was observed when increasing the amount of steel reinforcement, as in column CS-4%. In contrast, a signifi- cant increase in the postyield stiffness ratio was observed when adding FRP bars (Fig. 10). Fig. 10(b) illustrates that column CSF-2.8%-IS-D10 had a value of k more than twice that of column CS-2%, although the two columns share nearly the same value of µ P 1 . Changing the location of the FRP bars from the place of the steel fibers (column CSF-2.8%-IS-D10-J) to outside the steel stir- rups (in the concrete covers, such as in column CSF-2.8%-ES-D10- J) resulted in a 30% increase in the value of μ P 1 and a 17% decrease in the value of k [Fig. 10(c)]. Furthermore, by adding cross ribs to the surface of the FRP bars, as in column CSF-2.8%-IR-D10-J, the value of k decreased from 17.3 to 12%, although μρ 1 shifted from 5.8 to 11.3 [Fig. 10(d)]. This gives a high responsibility to the sur- face texture of the FRP bars to control the length of the hardening 04015062-12 ed from ascelibrary.org by University of Waterloo on 10/13/15. Copyright ASCE. For personal use only; a zone. Finally, the postyield stiffness represented by the two indices could be increased by using FRP bars with smaller cross-sectional areas. For instance, columns CSF-2.8%-ES-D10-J and CSF-2.8%- ES-D8-J, with the same FRP-to-steel stiffness ratio, had similar val- ues of k ; whereas μ ρ 1 shifted from 7.7 to 10.4 when decreasing the diameter to 8 mm. This confirmed the previous observation regard- ing the ability of the bond conditions of FRP to concrete to control the ductility at the endpoint of the achieved postyield stiffness. The previous discussion demonstrated that by adding FRP bars with a good bonding condition to concrete (8-mm-diameter bars or rough bars), the column could achieve the demanded lateral strength at high ductility levels through a stable postyield stiffness. Compared with these cases, using FRP bars with a weaker bond to concrete resulted in an unfavorable ductility and lateral strength at the end © ASCE 04015062-13 J. Compos. Constr.
of the postyield stiffness; however, it results in a larger postyield stiffness ratio.
Stiffness Degradation
Because of its importance for nonlinear modeling of structures (ElGawady and Sha'lan 2011 ), stiffness degradation of the tested specimens is compared in this section. Fig. 10(g) shows the nor- malized secant stiffness [i.e., secant stiffness normalized by the ini- tial stiffness (NSS)] versus the corresponding drift ratios for every specimen. As shown in the figure, the NSS values of the specimen CS-4% at all drift levels were clearly small and its secant stiffness degradation was very fast compared with those of the other spec- imens. Similarly, the NSS values of the specimens CS-2% and CS- 2%-J were smaller and their secant stiffness degradations were faster than those of the FSRC specimens. Among all the FSRC specimens, the secant stiffness of the specimens reinforced with smooth BFRP bars located in the outside of the stirrups degraded faster than others. The reasons for the aforementioned remarks are attributed to the elastic and postyield behavior and the failure modes of the columns, where the steel RC columns were charac- terized by cover spalling, concrete crushing, and bar buckling be- fore failure. For the FRP-steel columns, the bond behavior of the BFRP bars was the key parameter controlling the stiffness degradation.
Column Ductility
Aside from the displacement ductility corresponding to the end- point of the postyield stiffness, as explained in the previous section, three other indices were also used in this study to evaluate the duc- tility behavior and investigate the effect of the tested parameters. The first is the peak strength stability factor, which measures the ability of a structure to undergo large displacements after achieving the peak lateral strength and before entering the degra- dation zone [i.e., SI = ( δρ 2 − δρ 1 )/ δρ ]. The second is the degrada- tion stiffness factor, which defines the ability of a structure to reach its ultimate load through a gradual degradation path [i.e., K 4 = (V P − V u )/(δu − δ P2 )]. Through this definition, a smaller value of this index indicates a stable degradation behavior and vice versa. The last ductility measure is the displacement ductility factor at the ultimate lateral strength ( μ π = δ μ /δ γ ). Referring to the proposed mechanical model (Fig. 2 ), Table 5 summarizes the ductility indices for all tested columns. For the control specimen, after reaching its peak lateral load at a μ ρ 1 of 5.3, it achieved a stability index of SI = 3.5. This means that the column maintained its peak lateral strength up to a displacement ductility of 8.8 before entering the degradation zone. Following the stability zone, the column reached its ultimate strength at an ulti- mate displacement ductility, μ π C , of 10.3 through a degradation stiffness factor, K 4C , of 0.86. A significant enhancement in the column ductility could be achieved by wrapping the plastic hinge zone of the steel RC column with an FRP jacket. Although no clear stability zone was formed in column CS-2%-J [i.e., μ p 1 ≠ ( μ p 1 + SI) of column CS-2%], the column fulfilled its ultimate strength (i.e., V u = 0.8V p ) at a dis- placement ductility of 14.7 (i.e., μ n = 1.43 μ n c ) through a degra- dation factor of 0.27 (i.e., K 4 = 0.3K 4C ). In the absence of the FRP jacket, no pronounced enhancement in the ductility indices could be achieved by adding internal smooth FRP bars to the lon- gitudinal reinforcement. Moreover, the weak bond between the FRP bars and surrounding concrete resulted in a relatively sharp degradation curve before reaching the ultimate strength. For in- stance, the ductility indices SI, K4, and μ u of specimen CSF- 2.8%-IS-D10 were 0.97, 1.3, and 0.96 times the corresponding values of specimen CS-2%, respectively. Reviewing the ductility indices of specimens CSF-2.8%-IS-D10-J (which had smooth FRP bars) and CSF-2.8%-IR-D10-J (which had roughened FRP bars) indicated that the ductility behaviors of the two columns were quite different. Whereas column CSF-2.8%-IS-D10-J had ductility indices SI, K4, and μ μ of 2.1, 1.4, and 9.9, respectively, the cor- responding values of column CSF-2.8%-IR-D10-J were 0, 4.14, and 12.3, respectively. This indicates the benefit of roughening the FRP bars to enhance the ultimate ductility of the column (i.e., a 25% increase in μ μ ); however, the column suddenly failed upon reaching its peak strength as a result of FRP rupture. Regard- ing the effect of the location of the FRP bars on the ductility mea- surements, the results showed that placing the FRP bars in the concrete covers instead of in the concrete core caused the column to begin to degrade once it achieved its peak strength, without showing stability under the peak load. Moreover, the column could reach its ultimate strength at a greater displacement ductility through a smoother degradation manner. This was clearly observed in the behavior of column CSF-2.8%-ES-D10-J compared with col- umn CSF-2.8%-IS-D10-J, where the ductility indices K4 and μ ,, of column CSF-2.8%-ES-D10-J were 0.5 and 1.4 times those of col- umn CSF-2.8%-IS-D10-J, respectively. The reason for this might be attributed to the increase in the accumulated stresses in the FRP bars when they were installed at a larger effective depth, which in turn causes an increase in the bond stresses between the FRP bars and concrete. Reinforcing the column with 8-mm-diameter FRP bars (column CSF-2.8%-ES-D8-J) caused sudden drops in the lateral strength, similar to the case in the specimen reinforced with roughened FRP bars (column CSF-2.8%-IR-D10-J). However, the degradation curve was gradual as a result of using three 8-mm-diameter bars that ruptured at two different drift levels. Although columns CSF-2.8%-ES-D8-J and CSF-2.8%-ES-D10-J share the same ultimate ductility factor, the degradation stiffness factor, K4, of column CSF-2.8%-ES-D8-J was nearly three times that of column CSF-2.8%-ES-D10-J. Thus, more experimental in- vestigations on the effect of the bar surface texture (bond condi- tions) on the ductility indices, particularly the degradation © ASCE 04015062-14 behavior of columns reinforced with steel and FRP bars, are required.
Residual Displacement
Residual displacement was defined as the displacement of zero- crossing at unloading on the hysteresis loop from the maximum displacement (Fahmy et al. 2010b). In this study, the residual dis- placements for the first cycle of each imposed lateral displacement level were determined for all tested specimens (ElGawady and Sha'lan 2011 ). Thus, the drift ratio versus normalized residual dis- placement of all tested columns was drawn as shown in Fig. 11. The normalized residual displacement was defined as the ratio of the residual displacement of each column at a drift level to that of the control column at the same drift level (i.e., r = δ r / δ r c ). As the value of r for a column decreases, the column becomes more recoverable and reparable. Fig. 11 illustrates that although the value of r up to a drift of 1.7% (onset of FRP participation) was less than one for all columns, no clear stable trend could be observed. Start- ing from this drift ratio, a stable trend for the value of r could be observed. As expected, wrapping the plastic hinge region using an FRP jacket, as in column CS-2%-J, could not control the residual displacement of the column, where the value of r ranged from 0.95 to 1.02 up to a drift of 8%, at which the control column failed. In contrast, reinforcing the columns with FRP bars clearly decreased the residual displacement. For instance, the average value of the normalized residual displacement of columns CSF-2.8%-IS-D10, CSF-2.8%-IS-D10-J, and CSF-2.8%-ES-D10-J, reinforced with FRP bars having smooth surface textures, was approximately 0.8. The columns maintained this normalized value up to a drift level just before entering the degradation zone of each column as a result of bond failure. Following this drift level, the FRP bars could not continue controlling the residual displacement until the end of the tests. In contrast, the impact on the residual displacement became more considerable when enhancing the bond between the FRP bars and surrounding concrete. This trend can be observed in column CSF-2.8%-IR-D10-J (which had FRP bars with a rough texture) and column CSF-2.8%-ES-D8-J (which had FRP bars with a smaller diameter), in which both columns had an average value of r = 0.7. The columns continued controlling the residual displace- ment with nearly the same value of r until the FRP bars ruptured, at which the FRP bars lost their role in controlling the column performance. As a concluding remark, among all of the tested col- umns, CSF-2.8%-ES-D8-J had the best seismic performance in terms of the smallest residual displacement. This performance was caused by the high postyield stiffness, in the plastic zone.
Dissipated Energy and Damping Ratio
In recent decades, two indices have been defined to describe the ability of a structure to dissipate earthquake energy and hence sur- vive major seismic events. The first index is the cumulative dissi- pated energy, which is computed by summing up the areas enclosed by the hysteretic loops in the lateral load-displacement relation- ships of the structure up to failure. The other index is the viscous damping ratio, which reflects the damage level attained during in- elastic excursions. In this section, the cumulative dissipated energy, E , for all tested specimens was recorded and plotted for each drift level to further evaluate the effectiveness of the proposed reinforce- ment and investigate the effect of each tested parameter (Fig. 12 ). In the same manner, the damping ratio versus the displacement ductility were calculated and plotted for all specimens (Fig. 13 ).
Dissipated Energy
After calculating the energy dissipated by each loading cycle, the cumulative dissipated energy up to each drift level was determined, and the effect of each parameter was addressed (Fig. 12 ). Fig. 12 illustrates that the investigated parameters had only a minor impact on the cumulative dissipated energy. Up to a drift of 6%, all col- umns shared almost the same cumulative dissipated energy of ap- proximately 27 kN · m. The effects of the investigated parameters were apparent beyond this drift level. After a drift of 7%, no pro- nounced increase in the cumulated dissipated energy was observed in the control specimen, CS-2%, as a result of the cover spalling and bar buckling that caused the failure of the specimen. By wrap- ping the plastic hinge with an FRP jacket, as in column CS-2%-J, the column continued dissipating more energy up to a drift of 9.4%, at which E was equal to 1.7 E C , where E C is the total cumulative dissipated energy of the control specimen up to failure. Then, the rate of increase was small owing to the FRP sheet bulging, which led to the failure of that column. Fig. 12 also illustrates that although adding FRP bars to the steel reinforcement caused a sub- stantial increase in the postyield stiffness ratio, the column was able to dissipate more energy up to failure. The aforementioned obser- vations suggest that the efficiency of reinforcing the columns with both steel and FRP bars on the amount of dissipated energy may be controlled by other variables. Thus, further investigations are still needed to determine the best configuration and best surface treat- ment to optimize the dissipated energy.
Equivalent Viscous Damping Ratio
The equivalent viscous damping, ζ i , for the first cycle of all loading sets was calculated by (Jacobsen 1930)
where E j and E si refer to the dissipated energy and elastic energy in the cycle i , respectively. Fig. 13 shows the effect of each testing parameter on the rela- tionship between the displacement ductility, μ , and calculated damping ratio, ζ %. Fig. 13 illustrates that up to a displacement duc- tility of 2, only slight differences in the damping ratio, ζ, were ob- served in all columns, where at this ductility, the average value of ζ was approximately 12%. There was a small decrease in the ζ for specimen CS-2%-J compared to the control specimen CS-2%, par- ticularly after a displacement ductility of 9, at which the concrete cover of column CS-2% spalled [Fig. 13(a)]. A comparison of the results of column CSF-2.8%-IS-D10 with the control column [Fig. 13(b)] indicates that the decrease in ζ was more evident © ASCE 04015062-15 from ascelibrary.org by University of Waterloo on 10/13/15. Copyright ASCE. For personal use only; all up to a displacement ductility of 9, at which both the columns started to degrade; the average loss in the value of ζ in this column was approximately 20%. Moreover, the comparison of the results of columns CSF-2.8%-IS-D10-J and CSF-2.8%-IR-D10-J with that of column CS-2%-J [Fig. 13(d)] illustrates that although reinforc- ing the column with internal smooth FRP bars resulted in a minor decrease in ζ up to the appearance of bond slip failure at a ductility of 8%, this decrease was more evident when using roughened FRP bars (i.e., 33% decrease). For both columns, the value of ζ ap- proaching the failure was closer to that of specimen CS-2%-J. The average value of ζ for I-S-10-J and I-R-10-J up to failure was nearly 86 and 75% of that of specimen C-J, respectively. By placing the FRP bars externally instead of internally, as in col- umn CSF-2.8%-ES-D10-J, the value of ζ became lower, in which the average value of ζ for specimen CSF-2.8%-ES-D10-J was al- most 88% of that of specimen CSF-2.8%-IS-D10-J [Fig. 13(c)]. Finally, Fig. 13(e) illustrates that using FRP bars with 8-mm diam- eter caused a greater decrease in the value of ζ compared with the counterpart 10-mm-diameter BFRP bars. In other words, the de- crease in the damping ratio was more pronounced when using FRP bars with a stronger bond to the surrounding concrete. This discussion demonstrates that although using FRP bars as the main reinforcement helped the column achieve higher levels of postyield stiffness without causing any loss in the column ductility and dis- sipated energy, the decrease in the damping ratio could be accept- able to achieve the aim of damage-controlling structures.
Which FSRC Column Could Successfully Achieve a
Ductile-Recoverable Performance?
In the light of the targeted structural performance of the proposed FSRC structure, two columns could achieve the demanded ductile- recoverable performance. The first is CSF-2.8%-IR-D10-J (which had FRP bars with rough textures and was wrapped with an FRP jacket), and the second is column CSF-2.8%-ES-D8-J (which had FRP bars with a smaller diameter and was wrapped with an FRP © ASCE
04015062-16
J. Compos. Constr. aded from ascelibrary.org by University of Waterloo on 10/13/15. Copyright ASCE. For personal use only; a jacket). Both columns had an average value for the postyield stiff- ness ratio of 12.6%, an average displacement ductility before load degradation of 10.85, an average ultimate ductility of 12.85, and an average residual displacement of 0.7 times that of the steel RC col- umns. This performance was caused by the good bonding between the FRP bars and surrounding concrete.
Summary and Conclusions
In this study, an FSRC structure was proposed as a high- seismic-performance structure. Experimental tests on the effect of constant axial load and several cyclic loadings were conducted on seven RC bridge columns, where two columns simulated the performance of SRC bridge columns and the others showed the response of the proposed FSRC columns. The roles of several bond-based parameters, such as the diameter of the FRP bars, tex- ture of the FRP bars, location of the added FRP bars, and external © ASCE
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J. Compos. Constr. aded from ascelibrary.org by University of Waterloo on 10/13/15. Copyright ASCE. For personal use only; a confinement using an FRP jacket, were examined through the ex- perimental program. The following conclusions could be drawn from this study: • The proposed FRP-steel structural system ensured the existence of a stable postyield stiffness and realized the predesigned de- formability capacity. With a proper design of the proposed FRP- steel reinforcement system, bridge columns could withstand a strong  earthquake. Adding 0.8% FRP reinforcement to 2% main steel reinforce- . ment of concrete bridge column as well as replacement of 50% of the internal steel confining the plastic hinge with exter- nal BFRP jacket guaranteed a comparable lateral strength to that of a bridge column with 4% longitudinal steel reinforcement. The FSRC column fulfilled a displacement ductility of up to 10 before encountering strength degradation. The elastic stiffness of the FRP steel–reinforced column was the same as that of the counterpart steel-reinforced column; however, doubling the main reinforcement ratio of the steel- reinforced column was accompanied by a considerable increase in column elastic stiffness and a decrease in the postyield stiff- ness ratio. Within the range of the tested specimens, the average residual displacement of the proposed FSRC columns ranged from 70 to 85% that of the corresponding steel RC column. The mitigation of the residual deformations was dependent on the bond condi- tions between the FRP bars and the surrounding concrete. The favorable mitigation could be when BFRP bars with rough sur- face texture were used. Both the surface texture configuration and the diameter of the cross section of the FRP bars significantly influence the seismic performance and failure mode of FSRC bridge columns. Using 10-mm-diameter BFRP bars with a rough surface texture or 8-mm-diameter bars with small indentations resulted in a rupture of the BFRP bars that was accompanied by a sharp brittle failure after achieving a high strength level with reason- able ductility. Using BFRP bars with a smooth surface texture of 10-mm diameter caused lower and more stable peak strength at a smaller ductility with a smoother degradation. Location of BFRP bars in the column cross section with respect to the steel reinforcement had a marginal effect on both the column postyield stiffness ratio and column ductility before strength degradation. However, a considerable increase in the elastic stiffness and elimination of the stability plateau at the peak strength before strength degradation was observed when FRP bars were placed in the concrete cover rather than at the same fibers of steel reinforcement. Adding BFRP bars to steel RC columns maintained the cumu- lative dissipated energy and did not significantly decrease the damping ratio of the columns. . Future research should be directed toward providing a better un- derstanding of the behavior of such FRP-steel reinforcements with an emphasis on the bond condition's effect on the postpeak stability, residual displacement, and degradation stiffness. Other design parameters, including the FRP-to-steel stiffness ratio and transverse FRP reinforcement ratio, should also be examined. By monitoring the degree of damage and the condition of both steel and FRP reinforcement, the possibility of replacing FRP elements and damaged concrete to restore the original structural function is also a point of interest for future research. AASHTO. (2011). AASHTO guide specifications for LRFD seismic bridge design, 2nd Ed., Washington, DC. AASHTO. (2012). AASHTO LRFD bridge design specifications , Washington, DC. 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Hiroshi Noguchi and  Yasuhiro  Ohtaki Center for Information Technology, Ibaraki University Hitachi, Ibaraki 316-8511, Japan Email: {noguchi, y.ohtaki}@mx.ibaraki.ac.jp Abstract—We present a file backup system designed for a university on the basis of its experiences with the 2011 east Japan earthquake. This system has two levels of security: (1) The web site for public relations and the e-mail systems for communications that will work continuously even in the case of commercial power failure. (2) The personal computer systems that will be down during power failure but will restart working without any loss of data as soon as the power supply is back. The former has been implemented by employing a private cloud computing platform and a public one in combination. To the latter, the same cloud-computing approach was not applicable because it causes latency and extra cost of communications in the normal time to keep the frequently accessed data in a cloud storage. Fortunately, this particular university has three campuses so far away that one may be destroyed by a single disaster but the other two will survive. Taking advantage of the remote campuses, we designed and implemented a file storage system that keeps the original copy on-premise in a campus and its backup copy in one of the other two campuses. Its operations have shown that the communication traffic among the campuses increased only by half in order to keep the data safe against natural disasters. Keywords-disaster robustness; business continuity planning; distributed data backup The 2011 east Japan earthquake hit the north east part of Japan on March 11th including the three campuses of Ibaraki University [1] in Fig. 1. The buildings in Hitachi campus were severely damaged, in some cases, deep into their structural frameworks. Those in Mito and Ami campuses were also strongly shaken but less damaged for they were located relatively away from the epicenters. The server computers were so firmly fixed to the basement that they did not fall. But they were shaken so badly that we thought the hard disk drives must have been destroyed. After the main electricity got back in several days, we were relieved and even surprised to know that the disks were not damaged at all. Probably the arms in the drives were already retracted in response to the preceding smaller vibrations while the huge main shakes were traveling from the distant epicenters. Even a smaller earthquake can destroy data on the Masaru Kamada Department of Computer and Information Sciences, Ibaraki University Hitachi, Ibaraki 316-8511, Japan Email: m.kamada@mx.ibaraki.ac.jp Figure 1. Three major campuses of Ibaraki University This map shows the geographic outline of the islands and does not indicate the territory of any specific country. disk drives if it takes place just beneath the server room. Data loss can give much more serious damages to the business of universities than the physical damages to the buildings. It is the difficult time in disasters when we really need the means of public relation and communications. But the information systems of the university were completely out of operation for days due to the power failure. The main electricity were not available for three days in Hitachi and two days in Mito and Ami. We really needed to announce what to do with the entrance examination scheduled on the next day of the earthquake. It was obvious that we could not do it as scheduled. But we had to announce how many days it would be put off or if it should be cancelled. We set up an emergency web site [2] on a cloud service [3] to post announcements to the students and the employees. The temporary IP address is registered to the DNS so that the accesses to the old regular site be directed to the emergency site. This emergency site helped people get informed of the announcements from the university for a few days until the regular web site came back into operation. Then we suffered from another problem that the wrongly cached DNS records in different internet service providers kept pointing the emergency site. The heroic emergency site became a ghost web site that confused people by showing outdated information persistently. On the basis of those bitter experiences, we have re- designed and implemented the information system of our university. This system has two levels of security: (1) The web site for public relations and the e-mail systems for communications that will work continuously even in the case of power failure. (2) The personal computer systems that will be down during power failure but will restart working without any loss of data as soon as the power supply is back. This university has three major remote campuses as shown in Fig. 1. The center for information technology is located in Hitachi campus where the computers have been historically used much more actively than the other two campuses. That is the reason why the connection to the Internet is going through the firewall system in Hitachi as shown in Fig. 2. The campuses are connected in the star-shaped topology via the wide-area ethernet at the speed of 1Gbps with the interfaces of L3 switches. Each campus has a file server that stores the home directory of all the users in the campus. The home directory is mounted on the PC when the user logs on and also stores mails to be referred to by the mail server for the students there. The file server stores the systems and the data for the virtual servers on the real server computers managed by VMware. The campuses have been often struck by power failure for a short time due to local thunder storms. Each campus has one of the IdP server systems so that the campus can be indifferent to power failure of other campuses. From time to time, we had power failure at Hitachi campus where the mail server for the staff and the main web servers are hosted. But the black out did not last more than half an hour. That was strong enough against small and local disasters. But the story was different in the case of huge disaster in the wide area. The earthquake severely damaged Hitachi campus but fortunately claimed no casualties. We lost water supply for more than 10 days. However, in the aspect of communica- tions, the most difficult situation was caused by the power failure for three days. The main web server and the mail server for the staff were completely down when we really needed to post announcement and discuss recovery plans to cope with the huge disaster. The emergency web site [2] set up on a cloud service [3] was barely fulfilling the public relations. But it took some time for this new web site to start operating even though people needed information as soon as possible to know what was going on. Besides, after the regular web site came back into operation, the emergency site had suddenly become a ghost web site that confused people by showing outdated information persistently due to the wrongly cached DNS records in different internet service providers. It is the only and best policy to keep the regular web server working throughout the disaster. The same is true of the mail servers. We have never heard of temporary mail services that can substitute existing mail services of the regular use. They are useless because we do not have a means to identify the new mail addresses of students and colleagues. Loss of data is better than loss of lives, but it is worse than damaged buildings. It can take an enormous amount of time or may be even impossible to reproduce the lost data. We learnt that an information system really strong against mega disasters needs the following two features: (1) The web site for public relations and the e-mail sys- tems for communications that will work continuously without interruption even in the case of power failure. (2) The personal computer systems that will be down during power failure but will restart working without any loss of data as soon as the power supply is back. The former can be implemented simply by employing the recent private or public cloud systems. The latter requires an idea. If we employ the cloud service to keep the data safe in preparation for possible mega disasters, the data exchange between the local PC and the remote cloud causes latency and traffic in the normal time. The users should suffer from slow response of the PC due to the latency in data retrieval from the remote storage. Frequent data exchanges between the PC and the remote storage should require wide bandwidth to increase the communications cost or slow down the data transfers. In order to store the data in a safe place while keeping the data readily accessible in the on-premise file server, we can exploit the three remotely distributed campuses connected by the 1Gbps wide-area ethernet. The local file server stores the data for the local users in each campus as it has been practiced normally. Let its duplicate copy be stored in a fellow file server in another campus. Then, at least one copy of the data will survive strong earthquakes just beneath the file server even if it would not give time to retract the disk arms. The web and mail services can be interrupted for a short time in the normal time for people usually accept regular server maintenance. On the contrary, those services should not be interrupted in the time of disaster since the latest information is crucial for the people to make decisions in emergency. The web site can work continuously without interruption only if the server stays physically healthy and the power is uninterrupted even in the case of commercial power failure. There are roughly two kinds of solutions by means of private and public cloud systems operated at some data centers. The main web server of Ibaraki university has been moved into a private cloud system set up in a containerized data center built in one of the campuses. It is designed to survive the strongest earthquakes in the history of Japan and supplied with main electricity by the dedicated generator. The DNS servers are extended to include the secondary DNS service hosted in a public cloud service. We are planning to move the web server further to a public cloud service. The mail system has emigrated to the cloud service Mi- crosoft Office 365. The IdP servers are extended to include a slave server also in a public cloud service. That is the simplest and affordable solution by virtue of the academic discount. Each campus has a file server Model VNX5300 from EMC [4] that stores the data to be preserved. They include the work space of students, electronic contents for the classes, and documents for class management. They are crucial for continuation of teaching activities through small or middle scale disasters or their quick recovery after huge disasters. If we send the frequently updated huge data to a remote backup server outside the university in anticipation of possi- ble mega disasters, the data transfer from the local server to the remote server causes heavy traffic in the network in the normal time. However, the fear of losing important data has been driving several Japanese universities to start a coalition to keep the data of one another [5] [6] despite the extra cost for communications and storage. In the case of Ibaraki university where we have been suffering from the overhead cost for managing remote cam- puses, we have a good set of three campuses geographically apart from one another. We have only to synchronize the original copy of the data with its duplicate copy in another campus. We can keep them in the plain format without any encryption that is mandatory for the coalition of several different universities. The backup data is a simple duplicate copy of the original one. A system image of the virtual machine and its data partition in the backup can migrate onto a different real server to start working even in the case of physical damage on the original server. Figure 3 shows a schematic diagram of the three file servers. Each disk system is partitioned into the four regions: original copy, local duplicate copy, remote duplicate copy, and local generational archive. The local duplicate copy is just for increasing the robust- ness against malfunction of the disk systems. The file server VNX5300 offers the capability of updating the local dupli- cate copy immediately when the original copy is modified. The remote duplicate copy in another campus is also updated by the same mechanism except that the updates take place not immediately but every 10 minutes. This time interval of 10 minutes is a parameter that can be tuned in accordance with the real situation. A too short interval may increase the base load traffic. A too long interval may cause a burst traffic to transfer piled up updates at a time. The local generational archive stores an ordinary incre- mental backup data from which users can reverse their file operation. No data are encrypted since everything is under the control of the same institute. Nor are they even compressed. Backup data in the same plain format as the original are advantageous to save computational load for encryption and compression, and troubles for decryption and extraction in case the backup copy is needed in emergency. The traffic observed before and after the introduction of new backup system is plotted in Fig. 4. In the period of the measurement, the total volume of data was around 2.5T bytes. Variation from day to day of the week depends mainly on the change in the timetable of the classes. The traffic has increased to roughly 1.5 times but stays less than 1M Bytes per second. For the wide-area ethernet connecting the campuses at the maximum speed of 1G bits per second, the new traffic rate is a moderate good figure not too low or high. Through the bitter experiences during the disaster in March 2011, we have redesigned the information system for Ibaraki university. The web site and e-mail systems were moved to the private and public cloud services so that they will never be interrupted again. The file servers are made secure by keeping a duplicate copy in another campus which is located more than 30 km away. Taking advantage of the remote campuses, we designed and implemented a file storage system that keeps the original copy on-premise in a campus and its backup copy in one of the other two campuses. Its operations have shown that the communication traffic among the campuses increased only by half in order to keep the data safe against natural disasters. [1] Ibaraki University, http://www.ibaraki.ac.jp/ [2] VIOPS Virtualized Infrastructure Operators group, http://www.viops.jp/2011/03/post-12.html, March 2011. [3] Cloudpack, http://cloudpack.jp/ VNX5300, [4] EMC http://www.emc-storage.co.uk/emc-vnx- 5300-emc-vnx5300-vnx5300-storage [5] K. Nishimura: Implementation and evaluation of secure out- sourcing scheme for cloud storage services using secret sharing scheme, The 8th International Workshop on Security (IWSEC) 2013), November 2013. [6] Y. Kumagai et al.: A distributed file management system based on identity federation, IPSJ Technical Report , Internet and Operation Technology(IOT), Vol.2012-IOT-18, No.8, June 2012. [7] M. Iwano et al.: Demonstration of non-stop backup system for business continuity plan, IEICE Technical Report , Vol.112, No.22, 111-116, May 2012.公開文EOS-Terra/ASTERを用いたマングローブ域抽出に関する研究—ミャンマー沿岸域を対象として一A Study on extraction of mangrove area in Myanmar coastal zone by using EOS-Terra/ASTER —Focused on the Myanmar coastal zone—山崎　正稔*・石内　鉄平**・桑原　祐史*** Masatoshi YAMAZAKI and Teppei ISHIUCHI and Yuji KUWAHARA要旨：近年，地球温暖化による海面上昇，高潮，台風やサイクロンによる被害の増加が懸念されている.これらの影響による災害の鍼災を目的とした適応策としてマングローブが注目されているが，過伐採や植林などにより急速な分布城の変化が問題となっている。モニタリングによって分布域変化の把握が課題となっているが，分布城の広域性や多様な生物が生育する潮間帯に生育するため，急速な変化を広域に波り捉えることの難しさが問題となっている．本研究では，この問題点と水分量に対して特長的な反応をする短波長赤外の衛星データに着目し，桑原ら（2010）の手法で課題とされていた沿岸域の地形特性（山地等）への対応とその適応効果について考察した．結論として，メコンデルタなどを対象とした桑原らの手法を改良し，沿岸城の山地に対応したマングロープ城抽出手順を提案，その適用効果を検証し，新たな手順を提案した.キーワード:マングローブ,湖間帯,可視近赤外,短波長赤外, EOS-Terra/ASTER ,と淡水の入り混じる環境で生育することから，強序論1.風や高潮から海岸線を守る海岸防災林の役割を持1.1研究の背景ち，水産資源の涵築の場としての機能も備わって近年，地球温暖化による海面上昇，高潮，台風いる。また，資金調達が難しい発展途上国にとっやサイクロンによる被害の増加が懸念されている.て，自然物であるマングロープは防波堤建設などミャンマー連邦共和国（以下、ミャンマーと表記に比べ、コスト面でも安価であり有効的な海岸保する）では，平成20年5月，サイクロン・ナルギ護や侵食対策になりえる。ミャンマー環境保全森スにより多くの死傷者や被害が出たリ. その原因林省では技術者の育成を進め，洪水農業に加え，のひとつが防風・防波の役割を果たしていたマングローブ林の激減である²² マングローブは，近マングローブ等を指標とした環境問題への取り組みに対する衛星利用の重要性を認識している ²².隣に生活する人々に生活に欠かせない様々な恵みまた、現在行われているマングロープ分布域を確を提供してきたか.また,マングローブ林は海水◆学生会員茨城大学大学院理工学研究科都市システム工学専攻. ■■明石工業商等専門学校都市システム工学科*** 正会員茨城大学広域水圈環境科学教育研究センター認する手段として，マングロープアトラスによる確認，各種被覆情報などの地理情報，さらに，衛星利用の三つが主である。このような中，マングローブ分布域作成を目的とした場合，広域性・同時性・周期性に優れた衛星データの利用が有効である。既往研究における報告では，LANDSAT/TMの短波長赤外バンドを利用し陸域植生の区分を行った後，マングローブ樹種の区分に対し近赤外と短波長赤外バンドを用いることが有効であるとの報告がある(*),また, ASTERデータを用い, NDVI （正規化植生指標）から植物の活性度の高い領域を抽出し，可視近赤外の各バンドの値(DN : Digital Number,以下, DN値とする）から分布を検討した例もあるか.以上を統合すると，衛星利用に注目が集まるなか，将来的に全球を対象として，高い更新頻度に即した簡易なマングローブ域の抽出法の提案が望まれている。日本研究では，気候変動による海面上昇や高潮，台風，サイクロンによる被害を推定し，適応策提案に結びつけるために必要な地理情報として，術星データからミヤンマー沿岸部のマングローブ分布域を作成することを目的とした。その際に，マングロープ分布域の作成方法は，桑原らの方法を基本とした。ミャンマー西部沿岸域では低地と山地が隣接し混在する特徴がある。したがって、地形を考慮に入れた手順の改良を行うことで他地域への汎用性を検討することにした．具体的には，ミャンマーの地域性を考慮した潮間帯の抽出方法の改良とおよび低地と山地が混在する沿岸域の対応を工夫することを目的とした．以上の目的を踏まえた研究の流れを図−1に示す. ープの植生が見られるとされるイーヤワディ川デルタ地帯(以下,イーヤワディデルタと表記する)とする。イーヤワディデルタのマングローブ分布図が現存する。マングローブアトラス(2010)を図−2に示すり、この図を，衛星データからの抽出結果の良否を判断する基準データとした.研究対象地域は，ミャンマーにおいてマングロEOS-Terra/ASTERを用いたマングローブ枕抽出に関する研究本研究では解析データとしてASTERデータを使用する. ASTERはEOS-Terra衛星に搭載されている光学センサである．解像度はバンドによって異なるが，而解像度のものは15mであり，本論でのマングロープ域の精度は約1/50,000程度の縮尺を目途として議論し，広範囲に生育するマングロープ林などの抽出や解析への対応を目指す.また,マングローブのような蒸発散品の多い地物の抽出に適する短波長赤外（SW）を計測している点に注目した．ただし，ASTERはセンサ故障のため，現在,SWの計測はされていないが，2014年8月に打上げに成功したWorldView-3では,解像度約4mとASTERより高い空間分解能でのSW観測が開始されているため，本研究の成果はそちらのデータを用いた研究への応用展開も望める.表-1に使用データの一覧を示す. ASTERデータの取得時期は，雲の影響を考慮し，乾季とされる10月中旬から２月中旬を中心に７時期のデータを用いた.陸域地形データとしてSRTM?,水深データとしてETOPOI®,時間毎の潮位データとしてUHSLC hourlydata",海岸線マスクデータとしてSWBD101を使用した.  GISデータであるSRTM, ETOPO1, SWBDの比較を表−2に示す.これらのデータとASTERデータの各分解能の違いについて,オーバーレイした後に出力するデータを90mとして解像度を統一することとした. NDVIとは，マルチバンドデータから求められる植生の有無・多少・活性度を示す指標のひとつである.  ASTERデータより算出した対象地域のND
※SWBDはベクタ形式のため解像度なし表−3　各被羅におけるNDVI
VIから耕作地とマングローブ域,それぞれ5箇所ずつをトレーニングエリアとして選定し、NDVIを利用してマングロープ域の抽出が可能か検証した．各被疫の指標値の分布を表−3に示す．表−3より，マングロープ域周辺に分布する耕作地とマングローブ城の指標値は類似した値であり，值の分布が被察間で重複しているため，NDVIのみによってマングロープ城を抽出することは困難であり，抽出プロセスに工夫が必要であることが確認された. NDVIのみを利用した方法では，マングロープB3：バンド3  n：バンド5・6・7 (C,は-1～+1の値をとり，+1に近いほどマングローブの分布密度や活性度が高いことを示す）域を抽出することが困難であるため，桑原らによって，メコンデルタ等を対象に提案された指標,マングロープ抽出指標 C, (式(1)）を利用した.事前に，マングローブ域と耕作地の地表面反射率を比較した結果，両者の反射率に大きな違いがあることが確認できた．具体的には，マングローブ域と耕作地から各5地点を選出し，DN値を求めた．結果を図−3に示す．クロロフィルの影響によりマングローブ域のDN値が最も高くなるパンド3（近赤外）,マングローブの蒸発散に反応しマングロープ域と耕作地が安定して分類でき，かつ,バンド3 とのデータレンジ差が大きいバンド5・ 6・7（短波長赤外）を利用できることがわかる.
以上から、近赤外域と短波長赤外域のDN値を用い,式(1)に示す比演算式により計算するマングローブ抽出指標：C,の適応がミャンマーにおいて可能であることが確認できた.水域から設抽出を取り除くため，水域データの作成を行った。一般的に，近赤外より長波長の光は水に強く吸収される特性があるため，近赤外より長波長のセンサを用いて得られた画像では陸域と水域の区別がつきやすい.そこで,本研究では, EOS-Terra/ASTERを用いたマングロープ域抽出に関する研究細かい江線近傍の地形特徴を抽出することを考慮に入れ，短波長赤外より解像度の高い近赤外データ（15m）を用いて水域の抽出を行った。処理に関しては，溯間帯を削り過ぎずに領域を除くことができる衛星データの閥債を試行により検討し(表-4),二値化処理により水域を抽出した.水域,マングロープ域および陸域がオーバーレイされている結果を図−4に示す（処理前：1,809,797pixel,処理後: 718,335pixel).なお，関値はマングロープ域と他の土地被覆の境界のDN値から取得し， ASTERデータの取得時期ごとに分けて決定した.図−4より，この処理において主として海域や湾部近傍の水域における誤抽出はほぼ除去できたが、内陸部の水田・湖沼部といった湿地部近傍にもマングローブ城が点在しており，陸域においてデー夕処理での対応が必要となった.この処理では，マングロープの生育条件に合致しないと思われる陸域の誤抽出域の除去を目的とした。そのため，マングロープの生育域である潮間帯の特徴に注目した。潮間帯は，植生，土壌，水域が混在する。そこで，NDVI・NDWI（正規化水指標）・NDSI（正規化土壌指標）を利用し、潮間帯におけるマングローブ分布域の特性を限定した。マングローブ分布域，耕作地の土地被覆ごとに各５か所トレーニングエリアとしてサンプルを抽出した結果，マングローブ分布域が示す範囲は0.25 ≤ NDVI ≤ 0.60, 0.35 ≤ NDWI ≤ 0.70, -0.85  ≤  NDSI  ≤  -0.50 であった．本研究では，この条件を同時に満たす領域をマングローブ生育の条件とし、満たさない領域を削除した。抽出結果を図−5に示す（処理前：718,335pixel,処理後： 451,763pixel)次に，海岸線からの地形形状に注目し，禍間帯図−5生育条件による処理を行った処理結果(図−2中STEP2)そのものを特定した.マングローブは，海水と淡水の入り混じる環境である潮間帯に生育する。そこで，海水が陸方向に浸入する限界線を推定することで抽出結果を生成できると考えた．対象地域の禍位データ（UHSLC hourlydata）から，扱高潮位とASTERデータ撮影時の潮位を算出し,また,水深はETOPOIより算出した．その際に，図−6に示す地形勾配法が必要となる。勾配は，沿岸に5本の測線を発生させ, ETOPOLより計算した結果0.011度となった.  ETOPOIは約2kmメッシュの水深データであるため，推定した勾配は概略値であることは否めない。しかし、デルタ地帯の地形が平坦（内陸100kmまで標高差5m以内）であるため，概ね妥当と考えた．この点について分解能が約2kmのETOPOLを用い，境界を数値で識論することには限界があるが，海外を対象とした解析において全球を同一基準で整備したデータは数少なく，傾斜面での解析精度向上社合後の課題としたい。以上の方法により，作成した水域データから内陸側に4,600m以内の範囲をマングローブ生育領域（図−6中/ :湖間帯)とした．他の領域を削除した結果を図−7に示す（処理前： 451,763pixel,処理後: 236,208pixel) .次に，マングロープ生育の条件を満たさない山地にも誤抽出が見られるため，SRTMを用いて作成した陸域の傾斜角から認抽出の除去を行う。平地と山地を区別する際の関値は農学と建築の分野の定義から決定した．まず，農地における最大傾斜角は放牧牛が傾斜地を自由に歩行できる範囲が斜面傾斜角6度であるため6～12度であるリリ。また,建築の分野においても傾斜角が6～8度以上になると安全性が失われ危険を伴うため，傾斜地用の基礎処理を行った後に構造物の建設を行う。よって,本研究では平地と山地の閾値を6度とした.傾斜角6度以上の誤抽出城を取り除いた結果を図−8に示す（処理前：236,208pixel,処理後： 207,566pixel) .本研究で使用した衛星データの分解能は15m～ 90mである。このため，90m×90mの矩形領域内の被覆が混在している場合に必ずしもマングローブ域でない地点が抽出される可能性がある.このため本研究では，マングローブ城抽出を目的とした研究であることから，領域としては小さすぎる弧立した領域の除去が必要であるため，遥別と疑集処理率'を行った．延集と適別処理により分類画像のノイズを除去することがでる。選別処理によって関値のサイズにしたがって孤立したビクセル群を除去することができ，凝集処理によって分類後の類似性の高い隣接領域を接合して既存領域の空間的干渉性の向上が期待できる。遥別の関値を9Pixel,延集の関値を3×3Pixelとし,その処理結果を図−9に示す.図−7 潮間帯を抽出する処理結果(図−2中 STEP3)図−8 陸域傾斜角で抽出した山地を除いた処理結果（図−2中 STEP4）図−9 選別・凝集処理結果EOS-Terra/ASTERを用いたマングロープ成抽出に関する研究マングローブ域抽出結果（2002-2004）とマングローブアトラス(2010), Google Earth(2013)との比較両像を図−10に示す．また，図−10における領域①と領域②の拡大図を図−11，図−12に示す．データ取得時期に差異はあるが，抽出処理結果はマングローブアトラスと概ね同等の抽出結果が得られたことが分かる。しかし，山地に対して平地と同条件での抽出処理を行ったため，既存研究のメコンデルタでは確認できなかった山地における頭抽出がみられた。よって，山地に対して傾斜角を考廊に入れた抽出処理が必要となった。また，イーヤワディデルタは干満差が約5mもあり9,平地では非常に平坦な土地が続くため7,内陸の河川まで塩水であることを条件として潮間帯を抽出し,内陸部の誤抽出域除去を行った結果，マングローブ域の抽出に精度向上が見られた。また，水坂の通出や生育条件の導入，そして適別と議集といった個別のデータ処理を組み合わせた本研究のマングローブ抽出プロセスは数値や関値に依存するため,試行検討による最適値の導出が重要であった. ①式(1)で計算するマングローブ抽出指標に加え，西部沿岸域における低地と山地への処理プロセスを加えることによって，ミャンマーにおいても高精度にマングローブ域を抽出できることが確認できた．しかし，この手法は地域毎に関値を試行検討し、最適値を導出する必要がある．この点について本研究では，空中写真や標高データ，土地利用図などを総合的に考慮し，関値決定の方法を支援し，抽出精度の向上に繋げることができた。関値選定を支援するデータを精査することにより適用効果と精度の向上を検討した. ②イーヤワディデルタは平地と山地が入り混じる複雑な地形を持つため，平地と山地で同じ生成法による抽出を行った場合，山地に誤抽出が多く見られた。その結果，陸域勾配を考虚し，山地の誤抽出を削除することで精度の向上が見られた。 ③イーヤワディデルタは内陸の支川部にも多くのマングロープが生育していることを考慮し、このエリアに対し、陸域の標高データと潮位から塩水となる海域エリアを，河口から内陸の河川まで探索させることで河川が内陸部まで入り組む潮間帯の変化に適応できることが確認された.凝集と選別により分類画像のノイズ除去をすることができる。通常はまず，遥別処理を行い閾値のサイズ（ピクセル数）に従って孤立したピクセル群を除去し、次に，疑集処理を行って分類後の類似性の高い隣接領域を結合して既存のクラスの空間的コヒーレンシーを向上させることができる. 1）ジャパン・プラットフォーム,ミャンマー・サイクロン被害者支援: http://www.japan platform.org/programs/myanmar-cyclone/ , 2015.8 2）丸紅エアロスペース株式会社：インフラ・システム輸出促進調査等（ミャンマーにおける小型衛星インフラ・システム輸出調査）調査報告告, 2012. 3）海津正倫,平井幸弘:海面上昇とアジアの海法, 2001 4）財団法人資源環境観測解析センター：資源・環境リモートセンシング実用シリーズ⑤地球親測データの利用(2), pp. 53-67, 2005. 5)  Y Kotera,M Ochi,M hato : Generation of seismic base mapping using satellite image in the southern deltaic area, people' s republic of Bangladesh,  Proceedings of the 18th Asian Conference on Remote Sensing, 1997. Mark Spalding, Mami Kainuma, Lorna Colli 6) ns: World Mangrove Atlas, pp. 118-121, 2010 7)アメリカ地質調査所(USGS) : http://dds.c r.usgs.gov/srtm/, 2015.2 NOAA ETOPOl Global Relief Model, https:// 8) www.ngdc.noaa.gov/mgg/global/global.htm l, 2015.2 9)  UHSLC : http://uhslc.soest.hawaii.edu/, 2015.2 10)アメリカ地質調査所(USGS) :https://dds.cr .usgs.gov/srtm/version2_1/SWBD/, 2015.2 11）井出保行，小島誠，林治雄：傾斜放牧草地の地形と草地管理： 1. 裸地と排義の分布，日本草地学会誌Vol. 44 (3) ,  pp. 208–214,  1998 12）桑原祐史,田中健太:マングローブ分布域の抽出精度向上に関する研究,地球環境研究論EOS-Terra/ASTERを用いたマングローブ城抽出に関する研究文集Vol. 18,  pp. 71-79,  2010 13）桑原祐史,田中健太:メコンデルタを対象としたマングローブ分布域の推定方法に関する研究，地球環境研究論文集Vol.17, pp. 69-75, 2009 14) Mark Spalding,Francois Blasco and Colin Field : World Mangrove Atlas, pp. 75,11997 15）大野勝弘：ミャンマー・エーヤワディーデルタにおけるマングローブ生態系の在地的管理に関する研究, 2007 16）田村正行：短波長赤外バンドとディジタル標高モデルによるマングローブ林の抽出,日本リモートセンシング学会誌Vol.32, pp. 221-231,  2012 17）竹内渉,安岡善文:衛星リモートセンシングデータを用いた正規化植生,土壌,水指数の開発,写真測量とリモートセンシングVol. 43(6), pp. 7-19, 2004 18) IPCC第5次報告書19) EXELIS : ENVIチュートリアルマルチスベクトル分類茨城大学大学院理工学研究科都市システム工学専攻（茨城県日立市中成沢町4-12-1),平成3年生まれ,平成27年3月茨城大学工学部都市システム工学科卒,同年4月茨城大学大学院理工学研究科都市システム工学専攻入学，士木学会学生会員. E-mail:15nm820y@vc. ibaraki. ac. jp明石工業高等専門学校都市システム工学科(兵庫県明石市魚住町西岡679-3), 2001年表城大学工学部都市システム工学科卒業，2003年同大学院修士課程修了。2008年同大学院博士課程修了，2008年博士（工学），同年茨城大学産学官連携イノベーション創成機構事常勤研究員, 2010年明石工業高等専門学校都市システム工学科助教,講師を経て，2014年より同校准教授，現在に至る。土木学会・日本都市計商学会・日本造国学会・日本リモートセンシング学会・環境情報科学センター・日本協社のまちづくり学会・日本測量協会、各会員. E-mail : ishiuchieakashi.ac.jp茨城大学広域水圈環境科学教育研究センター (茨城県日立市中成沢町4-12-1), 1992年東京理科大学理工学部士木工学科卒業。1995年茨城大学工学部都市システム工学科助手，講師，准教授，を経て，2015年茨城大学広域水園環境科学教育研究センター教授(センター長)，現在に至る。博士（工学)。日本沿岸域学会，土木学会，日本リモートセンシング学会，地理情報システム学会，環境情報科学セングー,地盤工学会,各会員. E-mail : yuji.kuwahara.rseve.ibaraki.ac.jp A Study on extraction of mangrove area :خin Myanmar coastal zone by using EOS-Terra/ASTER —Focused on the Myanmar coastal zone— Masatoshi YAMAZAKI and Teppei ISHIUCHI and Yuji KUWAHARA ABSTRACT : The objective of this study is to propose the mangrove distribution map in Myanmar coastal zone by using satellite remote sensing data based on the estimation method of mangrove area (Kuwahara et al. 2010). Global warming and sea-level rise are projected to affect seriously on low-land area. Therefore, recently, the mangrove forest is important to decrease the damage which is received from storm surge and sea-level rise. However, because of the mangrove forest distribution area decreases, it is thought that the monitoring becomes important to confirm the area.   So we adopted this method to mangrove area extraction and verified the versatility of its method. KEYWORDS : Mangrove, Intertidal zone, Visible and near infrared, Short wavelength infrared, EOS-Terra/ASTER 3.プロジェクト業績
【都市インフラの強靱化技術開発】炭素繊維やパサルト繊維による構造補強材の高性能化により、インフラ構造物の耐震性やResilience (自己回復性)向上手法を開発し、都市インフラの強靱化技術に関する取り組みを行っており、産学連携を検討している。 （呉教授） 【都市インフラモニタリングや知能化による高度化センサ技術の開発】 1）光ファイバセンサ方式とカーボンファイバセンサ方式による長寿命・高性能化された分布型領域センサの実製作・性能究明および実用性を実験室レベルで検討した。 2）領域センサにより得られるひずみ分布の直接利用により、構造物の早期損傷検知に対する各種指標を検討した。 3）開発したセンサ、装置、検知アルゴリズムを基に、早期損傷検知システムを開発し、茨城県内のRC桁橋、新潟県内の既損傷PC箱桁橋、東北新幹線高架橋、中国の蘇通大橋（１ ０００ｍ級長大橋）など既設構造物のリアルタイム健全性評価システムの検証を進めている。 4）高度化された光ファイバ分布センシング技術による広域地圏環境(河川堤防・地すべり)防災システムの構築に関する産官学連携プロジェクトを開始し、基礎検討を行っている。 (呉教授) 【センサデータの安定した通信/解析技術の開発と省電力化推進】 1)大規模災害により、竜源インフラが完全消失した状況において、市民にどのように、情報を伝達するか、また市民の安否情報をどのように収集するかについて研究を進めている。内部電源を必要としないUHF帯RFIDを利用している。 （武田教授） 【老朽化したインフラ構造物の災害リスクを考慮した維持管理計画】 1）橋梁の（異常監視）ヘルスモニタリングシステムの実証実験を常陸大宮市の引田橋での継続的に実施した。（齋藤特命教授,原田准教授，鎌田教授） 2)ひたちなか海浜鉄道湊線での企業との傾斜計の実証実験を推進.本件は10月に茨城県土木部常陸大宮土木事務所との連携で常陸大宮地区の地すべり危険地域へ３基設置.現在データ収集中（齋藤特命教授） 3)構造物の耐久性設計に資する高度化シミュレーション技術を開発している(車谷准教授) 【UAVを活用した新たな空間情報の防災システムへの活用と超小型衛星との連携】 1)茨城県生活安全部とUAVによる安全監視・不法投棄防止システムの実現の連携検討を開始した。 （齋藤特命教授） 2)宮城教育大学とのUAVを用いた空間環境情報可視化と津波災害地域の空撮を行った。(齋藤特命教授） 3)宮城県石巻市、女川町での空撮による災害地域の土地利用変化、震災遺構の映像保存プロジェクトをスタートさせた。 （齋藤特命教授） 4）茨城県土木部河川課との連携で、茨城県が管理するダムの長寿命化のための映像撮影を10月より開始。 （齋藤特命教授） 【アウトリーチとしての防災・減災、それらを含む環境教育の実施】 1）日本リモートセンシング学会との共催で、守谷市内小学校で防災における空間情報応用に関する講演を実施した。（桑原教授） 2)ひたちなか市外野小学校、前渡小学校にて環境・情報教育を実施した。 (齋藤特命教授) 3)（社)次世代センサ協議会第44回センサ＆アクチュエータ技術シンポジウムにてUAVと防災について講演した。（齋藤特命教授） 1）防災情報を配信する方法には、インターネットやＦＭ放送、拡声器などを利用する方法があるが一長一短がある。そこで追加手段の１つとして交通信号機を配信局として用いる方法を研究している。特に、光点滅（ＬＥＤのオンオフ、人間の目ではそのオンオフは感知できない）による光ワイヤレス通信の高信頼化について検討している。 2）大規模震災では、集中制御が基となる携帯網・ワイヤレス通信網では情報共有は難しいまた、信号機等の停電により大規模な車両渋滞が生じると考えられる。その渋滞車両を積極的に通信局として利用するネットワーク網を構築する研究を進めている。ＭＡＣプロトコルとして、変形２進カウントダウン法を考案している。（羽渕裕真教授） 【地元企業・学校等との学術連携実現】 1）湊線の鉄道運行支援システム構築について、連携研究を実現（ひたちなか海浜鉄道、福山C）した。（斎藤特命教授，桑原教授） 1）齋藤　修：ひたちなか市外野小学校等における理科教育の支援の一環として、理科・環境・情報等の総合的教育を行った（2015年4月～2016年3月）。 2)鎌田　賢：第38回全国高等学校総合文化祭「いばらき総文祭2014」コンピュータ部門のプログラミングコンテストにおいて、予選・本選（平成26年7月28日(月) 29日(火) )の審査委員を務めた。本選ではデモ展示も行った。 1）東北大学と共同で申請した科研費（挑戦的萌芽）「地盤中の間隙水の挙動調査」採択（東北大学風間基樹教授，茨城大学安原一哉名誉教授，鎌田賢教授）（齋藤修特命教授の企画・アレンジによる）において、地中埋め込み型の土砂内水流・水圧センサと無線モジユールを開発した。 2)シンポジウム「SICE計測部門セミナー都市のスマートセンシング」を企画して開催した。（齋藤特命教授） http://rcl.it.aoyama.ac.jp/sice-sss/sice_seminar_20140917.html 【自治体との各種連絡会議推進（茨城県・日立市・ひたちなか市等）】 1）茨城県商工労働部、土木部、企画部、生活安全部との定期情報交換（齋藤特命教授）（鎌田教授，桑原教授は随時参加） 2）日立市土木部、生活環境部、市長、副市長との定期情報交換（齋藤特命教授）（鎌田教授，桑原教授は随時参加） 3)ひたちなか市長との情報交換会1回／年度（齋藤特命教授）（鎌田教授，桑原教授は随時参加） 【国際共同研究を実施（アメリカ、英国、イタリア、中国、韓国等）】 1)科研S-8 (安原名誉教授)におけるメコン川流域とメコンデルタの気候変動影響と災害事情調査(UAVによる調査を含む)に参画した。成果は2014年8月30日NHK-TV　NHKスペシャルにて放映された。 2)センシング技術による地下鉄防災システムの高度化に関する共同研究を英国のケンプリッジ大学、長大橋の長寿命化に関する研究を中国の東南大学、光ファイバ技術による高速道路橋の長期モニタリングに関してアメリカＦＨＷＡ（連邦道路管理局）橋梁の長期性能検討プロジェクトチームなどと共同研究により推し進めている。（呉教授他） 1）５名の社会人博士課程入学者をリクルートした2013年度につづき、2014年度には3名の社会人博士課程入学者をリクルートした。（齋藤特命教授）センター教員の連携による指導体制を構築している。 2)当センター所属院生を中心にして、茨城大学学生国際会議（１１月１５日１６日：水戸キャンパス）の企画・運営を行った。（呉教授、沼尾教授、鎌田教授、湊教授、桑原教授、外岡教授） 3)羽渕研究室学部4年生高柳翔太君の研究成果発表「ＬＥＤのオンオフ信号による光ワイヤレス通信の研究成果」が電子情報通信学会東京支部学生会学生奨励賞を受賞した。 (全発表数数：２２９、受賞者数：２１） 4)羽渕研究室博士前期課程2年生ライサチン君の国際会議発表「Proposal of Turbo- Coded Differential OOK for Optical IM/DD System」がRISP International Workshop on Nonlinear Circuits, Communications and Signal Processing (NCSP'15)にてStudent Paper Awardを受賞した。
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Tomoyuki Ishida, Kaoru Sugita, Noriki Uchida, Yoshitaka Shibata, "Proposal of Information Acquisition Method utilizing CCN in a time of Large Scale Natural Disaster", Proc. of the 5th International Workshop on Multimedia, Web and Virtual Reality Technologies and Applications (MWVRTA2015), 2015年11月. 21) Hiroshi Noguchi, Yasuhiro Ohtaki, Masaru Kamada, Masanori Itaba, Satoru Ozawa, "Study of future strategy ICT center – Better quality of services in the case of Ibaraki University –", Proc. of the International Symposium on ICT in Development 2015 (ISICTD2015), 2015年11月. 22) Kazuhiro Takahagi, Tomoyuki Ishida, Noriki Uchida, Yoshitaka Shibata, "Proposal of the Common Infrastructure System for Real-Time Disaster Information Transmission", Proc. of the 30th IEEE International Conference on Advanced Information Networking and Applications Workshops, 2016年3月. 23) Kaoru Sugita, Ken Nishimura, Tomoyuki Ishida, "Optimizing Media Quality Parameters to Reduce Energy Consumption for Emergency Communication Service", Proc. of the 30th IEEE International Conference on Advanced Information Networking and Applications Workshops, 2016年3月. 24) Shota Takayanagi, Hiromasa Habuchi, Yusuke Takamaru, Yusuke Kozawa, "Data-Rate-Enhanced Optical-Wireless Code-Shift-Keying System Having Two-Stage Demodulation", Proc. of NCSP2016, 2016年3月. 25) Mitsuru Saotome, Yusuke Kozawa, Yohtaro Umeda, Hiromasa Habuchi、 "Differencial-OOK System for Underwater Visible light Communications", Proc. of NCSP2016, 2016年3月. 26) Atsuhiro Takahashi, Hiromasa Habuchi, Yusuke Kozawa, "Rigorous Symbol Error Rate Analysis of Hierarchical MPPM-CNK through Optical Wireless Channel", Proc. of NCSP2016, 2016年3月. 27) Yuto Matsuda, Hiromasa Habuchi, Yusuke Kozawa, "Impact of Improved On-Off-Keyed Turbo-code System in Optical Wireless Channel", Proc. of NCSP2016, 2016年3月. 1)鎌田和樹，渡邉俊哉，渋沢進，  “深度画像センサを用いた拮抗体操支援システムの改良と評価”，情報処理学会第163回ヒューマンコンピュータインタラクション研究会，2015年5月. 2)高橋貴大，羽渕裕真，小澤佑介，“光無線MPPM-CNKにおける競合計数器型フレーム同期法の一検討”，電子情報通信学会ＷＢＳ研究会，2015年5月. 3)高柳翔太，羽渕裕真，“拡張擬直交M 系列を用いる光無線コードシフトキーイングの一検討”，電子情報通信学会ＷＢＳ研究会，2015年5月. 4)松田優人，ライサチン，小澤佑介，羽渕裕真，“光無線通信における差動符号化OOK を用いるパンクチャードターボ符号システムの一検討”，電子情報通信学会ＷＢＳ研究会， 2015年5月. 5)外岡秀行，塩澤美咲，加藤篤，加藤恵理，片山晴善，酒井理人，加藤創史，“ALOS-2/CIRCの代替校正・相互校正の初期結果”，日本リモートセンシング学会第58 回学術講演会論文集, 2015年6月. 6)田川博務,外岡秀行, “ASTER火災検出アルゴリズムの安定化のための一手法(その２)”,日本リモートセンシング学会第58回学術講演会論文集，2015年6月. 7)加藤篤，外岡秀行，“雲を含む熱赤外画像マッチングにおける位相限定相関法のロバスト性評価”，日本リモートセンシング学会第58回学術講演会論文集，2015年6月. 8)黒木幹，伊藤将司，内田智昭，石田智行，米倉達広，“情報収集ツールを活用した市民協働のインフラ維持管理システムの運用に関する研究”,第51回土木計画学研究発表会(春大会）, 2015年6月. 9)安藤翔，石田智行，“相互関連性を考慮した拡張現実型ジオメディアの研究”，電子情報通信学会第30 回サイバーワールド研究会，2015年6月. 10)江藤拓也，羽渕裕真，橋浦康一郎，“一般化拡張プライム符号を用いるリレーネットワークの性能評価”，電子情報通信学会ＷＢＳ研究会，2015年7月. 11)高柳翔太，羽渕裕真，小澤佑介，“2段階復調を行う光無線コードシフトキーイングの誤り率性能評価”，電子情報通信学会ソサイエティ大会，2015年9月. 12)伊形俊介，村田直也，小澤佑介，樸田洋太郎，羽渕裕真，“デジタル制御型カラーシフトキーイング方式におけるターゲットカラー制御法の一検討”，電子情報通信学会ＷＢＳ研究会, 2015年10月. 13)奥村淳平，村田直也，小澤佑介，櫟田洋太郎，羽渕裕真，“4色LEDを用いた一般化空間変調方式の一検討”，電子情報通信学会ＷＢＳ研究会，2015年10月. 14)高橋貴大,羽渕裕真,小澤佑介, “競合計数器型フレーム同期法を用いる光無MPPM-CNKにおけるシンボル誤り率特性”，電子情報通信学会ＷＢＳ研究会，2015年10月. 15)松田優人，羽渕裕真，小澤佑介，“非均一電力割当法を用いる光無線ターボ符号化DOOKの一検討”，電子情報通信学会ＷＢＳ研究会，2015年10月. 16)野口宏,大瀧保広,鎌田賢, “オンプレミスの電子メールサーバからOffice365への移行”,大学ICT推進協議会2015年度年次大会, 2015年12月. 17)森卓哉，小澤佑介，株田洋太郎，羽渕裕真，“高速水中可視光通信のための複数色LEDを用いた畳み込み符号化OOK方式の一検討”，電子情報通信学会ワイドバンド研究会, 2015年12月. 1) 車谷麻緒:2014年度日本計算工学会論文奨励賞,「破壊シミュレーションのための構造要素を用いた離散体解析手法」 1.探択された競争的資金等の外部資金1)竜気通信誉及財団研究助成金,「消失通信路に耐性のある光無線ターボ符号システムに関する研究」 1,000,000円，2015年度，研究代表者：羽渕裕真2)受託研究  JAXA,「地球観測用小型赤外カメラ(CIRC)に関する校正検証」、平成27年度，研究代表者：外岡秀行3)受託研究宇宙システム開発利用推進機構,「ASTER画像の信頼性・利用性の向上に関する研究」、平成27年度，研究代表者：外岡秀行． 4)民間企業との共同研究,「スマートグラスの処理、チューニング方法等の助言」,研究代表者:鎌田管5)民間企業との共同研究,「路面状態センシング方式と放射分光応用に関する研究」,研究代表者：外岡秀行6)民間企業との共同研究,「路面状態センシング方式と放射分光応用に関する研究」,研究代表者：外岡秀行7)地域法人との共同研究,「訪日外国人観光客を対象としたARスマホアプリ(iOS版)の研究開発」、研究代表者：石田智行8)地域法人との共同研究,「訪日外国人観光客を対象としたARスマホアプリ(Android版)の研究開発」，研究代表者：石田智行1）科研費基盤（B）  （研究課題番号15H02693）大規模災害時の劣悪通信環境で繋がる次世代ネバー・ダイ・ネットワークとその応用代表柴田義孝（岩手県立大学） 分担石田智行（茨城大学）研究期間　2015年4月1日～2018年3月31日研究費総額 7,150,000円2）科研費挑戦的萌芽研究 （研究課題番号：26630215）課題名地盤中の水の挙動の調査を格段に進展できるワイヤレスマルチセンサの開発への挑戦代表風間基樹（東北大学） 分担 安原一哉，鎌田賢（茨城大学）研究期間　2014年4月1日～2017年3月31日研究費総額 4, 915, 000円3）科研費基盤研究（C）  （研究課題番号：26420409）課題名  可変張力つき2変数スプラインの導出とその画像補間への応用代表 鎌田賢（茨城大学）研究期間　2014年4月1日～2017年3月31日研究費総額　1, 700, 000円4） 科研費基盤研究(C)課題名センサネットワークと知識ベースを用いた高齢者見守リシステムの研究代表澁澤進，H26年度澁澤分550千円研究期間　2012年4月1日～2014年3月31日研究費総額　4, 800, 000円5） 科研費基盤研究(C)課題名 ITS のための光／竜波融合型通信の高度化代表羽渕裕真，H26年度 1100千円研究期間　2012年4月1日～2015年3月31日研究費総額　4, 200, 000円茨城大学重点研究「知的で持続可能な社会基盤および防災セキュリティ技術研究創出事業」茨城大学工学部附属防災セキュリティ技術教育研究センター 2 0 1 5年度報告書発行日平成28年4月
茨城大学  工学部  都市システム工学科教授　呉　智深〒316-8511日立市中成沢町4-12-1 Tel: 0294-38-5179  Fax: 0294-38-5268 ※禁無断転載茨城大学重点研究http://www.ibaraki.ac.jp/generalinfo/activity/researching/juuten/茨城大学工学部附属教育研究センター http://www.eng.ibaraki.ac.jp/research/centers/index.html防災セキュリティ技術教育研究センター http://www. ibaraki. ac. jp/research/centers/disaster/index. html