It is imperative that landlords of commercial and residential properties alike comply with the various laws governing discrimination in the management, leasing and sale of real property. To most people the word “discrimination” implies an animus towards a person because of his or her membership in a protected class such as a racial or ethnic minority, gender, or sexual orientation. However, anti-discrimination law contemplates a broader set of conduct which also includes activity that is arguably founded upon business judgment which nonetheless impacts a protected class, for e.g. whether the design of a structure accommodates persons with disabilities. Whatever the basis for the alleged discrimination, a landlord’s cost of litigating the lawsuit with the added potential for high jury awards for the victim’s emotional distress (that, most likely, are not covered by insurance), counsel in favor of a thorough analysis of landlord’s compliance with the law.
Whether the aggrieved individual pursues relief through an administrative proceeding or by filing a complaint in state or federal court, the party defending the action could be held liable for damages arising from the victim’s mental anguish or emotional distress and punitive damages. See New York State Div. of Human Rights v. Caprarella, 82 A.D.3d 773, 774-75 (2d Dep’t 2011) (upholding $10,000 award in punitive damages and $7,500 in damages for mental anguish where agency found that landlord could not proffer credible nondiscriminatory reason for turning down a non-white couple); Parris v. Pappas, --- F.Supp.2d ---, 2012 WL 573033 at *6-9 (D. Conn. 2012) (awarding $100,000 in compensatory damages and $150,000 in punitive damages where mobile-home park owner failed to reasonably accommodate plaintiff’s live-in aide and failed to service septic system as required by the lease agreement).
The obligation to provide “reasonable accommodation” to disabled persons is typically the most difficult standard to comply with given the fact-sensitive nature of the issue. Leaving aside the vexing question of what qualifies as a “reasonable accommodation” given a person’s disability, who qualifies as a “disabled person” is a threshold issue that deserves careful attention. Take, for example, the case of Shirley Carper of California who, in 2005, reached a $1 million settlement with the landlord of an apartment development after a five-year legal battle.  The case centered on Ms. Carper’s need for a parking space closer to her unit after she developed a degenerative knee disorder that worsened over time, in part, due to the flight of stairs she had to ascend to reach her apartment from her original parking space. Id. The landlord maintained the position that Ms. Carper was merely demanding accommodation of a “convenience” rather than a disability. Id. After trial, the jury awarded $250,000 in compensatory damages and, during deliberations for punitive damages, the case ultimately settled for $1 million. Id.
Similarly, in the context of commercial leases, both landlords and commercial tenants of “public accommodations” are subject to liability for discrimination against disabled individuals for failure to comply with the ADA’s public access requirements. 42 U.S.C. § 12182(a). The ADA precludes recovery of compensatory and punitive damages, although special damages for emotional distress may be recoverable upon sufficient pleading and proof under Federal Rule of Civil Procedure 9(g), and the court may award reasonable attorney fees under 42 U.S.C. § 12205. Thus, Congress intended that the ADA offers plaintiffs a prospective remedy in the form of injunctive relief, 42 U.S.C. § 2000a-3(a), i.e. that may require landlord and/or tenant to remove “architectural barriers” to the extent such alterations are “readily achievable”. 42 U.S.C. § 12182(b)(2)(A)(iv). Alteration of “architectural barriers” can, no doubt, be costly and the question is who bears this cost?  The ADA simply imposes liability on “any person who owns, leases (or leases to), or operates a place of public accommodation.” 42 U.S.C. § 12182(a). Under the enacting regulations, tenants and landlords may allocate the cost by agreement. 28 C.F.R. § 36.201(b). Nevertheless, merely because a lease agreement may hold tenant solely responsible for complying with all the laws and regulations, a landlord cannot escape liability under the ADA. In other words, a third-party can still hale the landlord into court and obtain injunctive relief requiring the landlord to effectuate structural changes that bring a public accommodation within compliance of the ADA’s access requirements. The landlord can then seek indemnification from the tenant to the extent this is permitted under the lease agreement.
From the above discussion, it is easy to see how landlords of commercial and residential properties can get tripped up and be subject to costly litigation if they are not in compliance with all the relevant anti-discrimination laws. Therefore, nowhere else may the old adage be more apt than in this case, that an ounce of prevention is worth a pound of cure, and knowledgeable counsel should be consulted to ensure compliance and minimize exposure.
 California Dept. Of Fair Emp't & Housing v.2001 California St. Apart., No. 03–423255 (Cal.Sup.Ct.2005), Vol. E1, Issue 1 National Fair Housing Advocate (Dec.2005).
 Liability is imposed upon “any person who owns, leases (or leases to), or operates a place of public accommodation” that discriminates against an individual on the basis of disability. Id. Discrimination includes the failure to remove “architectural barriers” in existing facilities where such removal is “readily achievable.” 42 U.S.C. § 12182(b)(2)(A)(iv).
 Powell v. National Bd. Of Medical Examiners, 364 F.3d 79, 86 (2d Cir. 2004).
 Id.; See Botosan v. Fitzhugh, 13 F.Supp.2d 1047, 1053 (S.D. Cal. 1998) (recognizing that, in an appropriate case, emotional distress can constitute special damages under Fed. R. Civ. P. 9(g)).
 Santiago v. Coco Nail HB, Inc., No. CV 10-3373, 2012 WL 1117961, *2 (E.D.N.Y. 2012) (awarding reasonable counsel fees in ADA public accommodation case).
 Roberts v. Royal Atlantic Corp., 542 F.3d 363, 373 (2d Cir. 2008) (characterizing “readily achievable” as a plausible proposal for barrier removal, the costs of which, facially, do not clearly exceed its benefits).
 Brooks Shopping Centers, LLC v. DCHWWC Restaurant, Inc., 32 Misc.3d 55, *56-7 (N.Y. Sup. Ct. App. Term. July 8, 2011) (reversing trial court and holding that cost of bringing premises into ADA compliance which exceeded $500,000 was not sufficient to demonstrate that removal of barriers was unduly burdensome and new trial was ordered to determine tenant’s ability to render the premises ADA-compliant to the “maximum extent feasible”).
 Yates v. Delano Retail Partners, LLC, No. C 10-3073 CW, 2012 WL 1094444, *2-3 (N.D. Cal. 2012) (recognizing parties’ ability to rely on lease agreements to seek indemnification but holding that indemnifying party not an indispensable party for an action brought under the ADA).


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Grachev, A. A., Persson, O. G., Uttal, T., Akish, E. A., Cox, C. J., Morris, S. M., Fairall, C. W., Stone, R. S., Lesins, G., Makshtas, A. P. and Repina, I. A. (2017) Seasonal and latitudinal variations of surface fluxes at two Arctic terrestrial sites. Climate Dynamics . DOI 10.1007/s00382-017-3983-4.
Thibodeau, B., Bauch, D. and Voss, M. (2017) Nitrogen dynamic in Eurasian coastal Arctic ecosystem: Insight from nitrogen isotope. Global Biogeochemical Cycles, 31 (5). pp. 836-849. DOI 10.1002/2016GB005593.
Osborne, A. H. , Hathorne, E. C. , Schijf, J., Plancherel, Y., Böning, P. and Frank, M. (2017) The potential of sedimentary foraminiferal rare earth element patterns to trace water masses in the past. Geochemistry, Geophysics, Geosystems, 18 (4). pp. 1550-1568. DOI 10.1002/2016GC006782.
Driemel, A., Fahrbach, E., Rohardt, G., Beszczynska-Möller, A., Boetius, A., Budéus, G., Cisewski, B., Engbrodt, R., Gauger, S., Geibert, W., Geprägs, P., Gerdes, D., Gersonde, R., Gordon, A. L., Grobe, H., Hellmer, H. H., Isla, E., Jacobs, S. S., Janout, M., Jokat, W., Klages, M., Kuhn, G., Meincke, J., Ober, S., Østerhus, S., Peterson, R. G., Rabe, B., Rudels, B., Schauer, U., Schröder, M., Schumacher, S., Sieger, R., Sildam, J., Soltwedel, T., Stangeew, E., Stein, M., Strass, V. H., Thiede, J. , Tippenhauer, S., Veth, C., von Appen, W. J., Weirig, M. F., Wisotzki, A., Wolf-Gladrow, D. A. and Kanzow, T. (2017) From pole to pole: 33 years of physical oceanography onboard R/V Polarstern. Earth System Science Data, 9 (1). pp. 211-220. DOI 10.5194/essd-9-211-2017.
Janout, M. A., Hölemann, J., Timokhov, L., Gutjahr, O. and Heinemann, G. (2017) Circulation in the northwest Laptev Sea in the eastern Arctic Ocean: Crossroads between Siberian River water, Atlantic water and polynya-formed dense water. Journal of Geophysical Research: Oceans, 122 (8). pp. 6630-6647. DOI 10.1002/2017JC013159.
Merkulov, V. A., Ashik, I. M. and Timokhov, L. A. (2017) Trends of long-term variability of the sea level at coastal stations of the Arctic Ocean. Problemy Arktiki i Antarktiki (Problems of the Arctic and Antarctic, in Russian) (3). pp. 51-66.
Polyakov, I. V., Pnyushkov, A. V., Alkire, M. B., Ashik, I. M., Baumann, T. M., Carmack, E. C., Goszczko, I., Guthrie, J., Ivanov, V. V., Kanzow, T., Krishfield, R., Kwok, R., Sundfjord, A., Morison, J., Rember, R. and Yulin, A. (2017) Greater role for Atlantic inflows on sea-ice loss in the Eurasian Basin of the Arctic Ocean. Science, 356 (6335). pp. 285-291. DOI 10.1126/science.aai8204.
Selyuzhenok, V., Mahoney, A., Krumpen, T., Castellani, G. and Gerdes, R. (2017) Mechanisms of fast-ice development in the south-eastern Laptev Sea: a case study for winter of 2007/08 and 2009/10. Polar Research, 36 (1411140). DOI 10.1080/17518369.2017.1411140.
Spengler, T., Claud, C. and Heinemann, G. (2017) Polar Low Workshop Summary. Bulletin of the American Meteorological Society, 98 (6). ES139-ES142. DOI 10.1175/BAMS-D-16-0207.1.
Spielhagen, R. (2017) Case Studies of the Role of Arctic Gateways in Plio-Pleistocene Arctic Changes . Polarforschung, 87 (1). pp. 23-31. DOI 10.2312/polarforschung.87.1.23.
Stein, R. (2017) From greenhouse to icehouse: the late Mesozoic-Cenozoic Arctic Ocean sea ice and climate history. Polarforschung, 87 (1). pp. 61-78. DOI 10.2312/polarforschung.87.1.61.
Thibodeau, B. and Bauch, D. (2016) The impact of climatic and atmospheric teleconnections on the brine inventory over the Laptev Sea shelf between 2007 and 2011. Geochemistry, Geophysics, Geosystems, 17 (1). pp. 56-64. DOI 10.1002/2015GC006063.
Jørgensen, L. L., Archambault, P., Blicher, M., Denisenko, N., Gudmundsson, G., Iken, K., Roy, V., Sørensen, J., Anisimova, N., Behe, C., Bluhm, B. A., Denisenko, S., Metcalf, V., Olafsdóttir, S., Schiøtte, T., Tendal, O. S., Ravelo, A. M., Kędra, M. and Piepenburg, D. (2017) Benthos. In: State of the Arctic Marine Biodiversity Report. . Conservation of Arctic Flora and Fauna (CAFF), Akureyri, pp. 85-107.
Hansen, M., Kraan, C. and Piepenburg, D. (2017) Benthic spatial biodiversity patterns in the seas of the Arctic Siberian shelf. [Paper] In: Arctic Change. , 10.12.2017, Quebec, Canada .
Piepenburg, D. and Bergmann, M. (2017) Impacts of increasing human activities and protecting the Arctic marine environment . [Paper] In: Environmental Science and the Work of the Arctic Council. , 10. October 2017, Helsinki, Finnland .
Belter, H. J., Ross, E., Hendricks, S., Janout, M., Ricker, R., Krumpen, T. and Haas, C. (2018) A new time series of sonar-based sea-ice draft records from the Siberian shelf. [Talk] In: POLAR 2018. , 15.06.-26.06.2018, Davos, Switzerland .
Bogorodsky, P. V. and Makshtas, A. P. (2018) Protsessy teplomassoperenosa v deyatel’nom sloe grunta i pripaqzskh l’dakh v rayone stantsinonara „Mys Baranova“(Processes of heat transfer in the active soil layer and in fast ice in the area of the station of "Cape Baranov"). [Talk] In: Seminar No. 2 Proekta „Izmenchivost’ Arkticheskoy Transpolyarnoy Sistemy“ (Seminar No. 2 of the Project "Variability of the Arctic Transpolar System"). , 21.03.2018, St. Petersburg, Russia .
Danshina, A. V. and Ivanov, V. V. (2018) Adaptatsiya modeli NEMO dlya issledovaniya transpolyarnoy sistemy i provedenie testovykh eksperimentov (Adaption of the model NEMO to the study of the transpolar system and realization of test experiments). [Talk] In: Seminar No. 2 Proekta „Izmenchivost’ Arkticheskoy Transpolyarnoy Sistemy“ (Seminar No. 2 of the Project "Variability of the Arctic Transpolar System"). , 21.03.2018, St. Petersburg, Russia .
Ivanov, V. V. (2018) Zadachi proekta i osobennosti ego vypolneniya v 2017 g (Project tasks and their realization in 2017). [Talk] In: Seminar No. 2 Proekta „Izmenchivost’ Arkticheskoy Transpolyarnoy Sistemy“ (Seminar No. 2 of the Project "Variability of the Arctic Transpolar System"). , 21.03.2018, St. Petersburg, Russia .
Makhotin, M. (2018) Okeanograficheskie usloviya i formirovanie vodnxkh mass v more Laptevykh: resul’taty v 2017 g (Oceanographic conditions and water-mass formation in the Laptev Sea: result of 2017). [Talk] In: Seminar No. 2 Proekta „Izmenchivost’ Arkticheskoy Transpolyarnoy Sistemy“ (Seminar No. 2 of the Project "Variability of the Arctic Transpolar System"). , 21.03.2018, St. Petersburg, Russia .
Makshtas, A. P. and Bogorodsky, P. V. (2018) Khzarakteristiki energo- i massoobmena podstilayuscshey gjverkhnosti s atmosferoy po dannym nablyudeniq na dreyfuyushchikh stantsiyakh „Severnyy olyus“ i statsionare „Mys Baranova“ (Characteristics of the energy and mass exchange of the surface layer with the atmosphere according to data collected at the drift station "Severnyy Polyus" and the station of "Cape Baranov"). [Talk] In: Seminar No. 2 Proekta „Izmenchivost’ Arkticheskoy Transpolyarnoy Sistemy“ (Seminar No. 2 of the Project "Variability of the Arctic Transpolar System"). , 21.03.2018, St. Petersburg, Russia .
Martynov, F. M., Povazhnyy, V. V. and Kryukova, I. (2018) Issledovanie pervichnoy produktsii fitoplanktona i taksonomicheskogo sostava zooplanktona, planiruemye v ramkakh proekta CATS v 2018 g (Investigation of primary production of phytoplankon and taxonomic zooplankton composition within the framework of the CATS project in 2018). [Talk] In: Seminar No. 2 Proekta „Izmenchivost’ Arkticheskoy Transpolyarnoy Sistemy“ (Seminar No. 2 of the Project "Variability of the Arctic Transpolar System"). , 21.03.2018, St. Petersburg, Russia .
Novikhin, A. E. (2018) Biogeokhimicheskie tsikly v more Laptevykh: resul'taty rabot za 2017. [Talk] In: Seminar No. 2 Proekta „Izmenchivost’ Arkticheskoy Transpolyarnoy Sistemy“ (Seminar No. 2 of the Project "Variability of the Arctic Transpolar System"). , 21.03.2018, St. Petersburg, Russia .
Ricker, R., Girard-Ardhuin, F., Krumpen, T. and Lique, C. (2018) Satellite-derived sea ice export and its impact on Arctic ice mass balance. [Talk] In: POLAR 2018. , 15.06.-26.06.2018, Davos, Switzerland .
Ricker, R., Hendricks, S., Paul, S., Kaleschke, L. and Tian-Kunze, X. (2018) Recent achievements in deriving sea ice thickness from radar altimetry. [Talk] In: POLAR 2018. , 15.06.-26.06.2018, Davos, Switzerland .
Stein, R., Kolling, H., Kremer, A., Saina, J., Syring, N., Fahl, K., Krawczyk, D. and de Vernal, A. (2018) Reconstruction of Arctic sea ice cover: new insights and questions from biomarker and microfossil proxy records. [Talk] In: 6. International Conference on "Palaeo-Arctic Spatial and Temporal (PAST) Gateways". , 16.04.-20.04.2018, Durham, U.K. .
Stein, R., Kremer, A., Fahl, K. and Niessen, F. (2018) Arctic ice sheets and sea ice extent during MIS 6/MIS 5: a biomarker perspective. [Talk] In: POLAR 2018. , 15.06.-26.06.2018, Davos, Switzerland .
Vyazigina, N., Timokhov, L. A. and Janout, M. (2018) Vliyanie gidrometeorologicheskikh faktorov i solnechnoy aktivnosti na kolebaniya ledobitocti Severo-Evropeyskogo basseina (Influence of hydrometeorological factors and solar activity on the variability of the ice cover of the North European Basin). [Talk] In: Seminar No. 2 Proekta „Izmenchivost’ Arkticheskoy Transpolyarnoy Sistemy“ (Seminar No. 2 of the Project "Variability of the Arctic Transpolar System"). , 21.03.2018, St. Petersburg, Russia .
Willmes, S., Heinemann, G. and Reiser, F. (2018) Spatial and temporal patterns of Arctic and Antarctic sea-ice leads, 2003-2016. [Talk] In: POLAR 2018. , 15.06.-26.06.2018, Davos, Switzerland .
Gutjahr, O. and Heinemann, G. (2017) Impact of the horizontal resolution on the simulation of wind extremes in the Arctic. [Talk] In: Polar CORDEX Meeting 2017. , 18.10.-20.10.2017, Cambridge, United Kingdom .
Heinemann, G. (2017) Task 2C: Modeling of atmosphere/sea-ice/ocean interaction. [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Heinemann, G. and Gutjahr, O. (2017) Analysis of Arctic wind extremes using high-resolution regional climate simulations. [Talk] In: Arctic Science Summit Week 2017. , 31.03.-07.04.2017, Prague, Czech Republic .
Ivanov, V. (2017) Task 1C: Hydrodynamic modeling of changes in sea ice and their impact on ocean dynamics. [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Ivanov, V. and Hölemann, J. A. (2017) Work Package 1: Dynamics of physical processes in Siverian shelf seas and their impact on the transpolar system. [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Janout, M. (2017) Task 1B: The shelf-edge system influence oshelf processes on the Arctic Boundary Current (TRANSDRIFT XXIV and XXV). [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Kohnemann, S. and Heinemann, G. (2017) A model-based climatology of wintertime low-level jets in the Nares Strait. [Talk] In: Arctic Science Summit Week 2017. , 31.03.-07.04.2017, Prague, Czech Republic .
Laukert, G. (2017) Task 3A: Mechanisms controlling the distribution of nutrients in the Laptev Sea (TRANSDRIFT XXIV and XXV). [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Makhotin, M. (2017) Task 1A: Hydrography and water mass formation in the Laptev Sea (TRANSDRIFT XXIV and XXV). [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Makshtas, A. (2017) Task 2A: Processes in the atmospheric boundary layer of the high Arctic (Cape Baranov and historical data). [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Makshtas, A. and Heinemann, G. (2017) Work Package 2: Changes of atmosphere/sea-ice/ocean interactions and their impact on the transpolar system. [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Naidina, O. D. and Bauch, H. A. (2017) Conditions in the Arctic according to the study of microfossils from the postglacial sediments of the Laptev Sea. [Talk] In: 22. International Conference on Marine Geology. , 20.11.-24.11.2017, Moscow, Russia .
Novikhin, A. (2017) Task 3B: Changes of the hydrochemical structure of the Laptev Sea (TRANSDRIFT XXIV and XXV). [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Novikhin, A. (2017) Work Package 3: Biogeochemical cycles in the Laptev Sea. [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Ovsepyan, Y., Taldenkova, E., Chistyakova, N. O., Stepanova, A. Y. and Spielhagen, R. (2017) Late Quaternary paleoclimate reconstructions in the Fram Strate region in relation to Atlantic water inflow to the Arctic. [Talk] In: 22. International Conference on Marine Geology. , 20.11.-24.11.2017, Moscow, Russia .
Piepenburg, D. (2017) Task 4A: Dynamics of pelagic particle processes and benthic functional diversity patterns (TRANSDRIFT XXIV and XXV). [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Piepenburg, D., Breckwoldt, A., Hansen, M. and Holstein, J. (2017) Mine the data: a roadmap towards stakeholder-oriented biodiversity warehousing. [Talk] In: Arctic Change 2017. , 11.12.-15.12.2017, Quebec, Canada .
Povazhnyy, V. (2017) Task 4B: Biological community changes in the Laptev Sea (historical data; TRANSDRIFT XXIV and XXV). [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Povazhnyy, V. and Piepenburg, D. (2017) Work Package 4: Ecological consequences of climate change in the Siberian shelf seas. [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Schirnick, C. (2017) Data management for the CATS Project. [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Spielhagen, R. (2017) Vortrags- und Diskussionsrunde für die Oberstufe: "Klimawandel: Fakt oder Fiktion?". [Public Lecture] In: Aktuelle Wissenschaften: Einblicke in die Arbeit von Meeresgeologen und Klimaforschern. Vortrags- und Diskussionsrunden für Schülerinnen und Schüler rund um das Thema "Arktis". , 15.05.-16.05.2017, Mainz, Germany .
Spielhagen, R. (2017) Vortrags- und Diskussionsrunde für die Unterstufe: "Mit der POLARSTERN zum Nordpol: Forschung am Ende der Welt".. [Public Lecture] In: Aktuelle Wissenschaften: Einblicke in die Arbeit von Meeresgeologen und Klimaforschern. Vortrags- und Diskussionsrunden für Schülerinnen und Schüler rund um das Thema "Arktis". , 15.05.-16.05.2017, Mainz, Germany .
Stein, R. (2017) The long-term Arctic Ocean Climate History: a major challenge for the IODP scientific community. [Talk] In: 22. International Conference on Marine Geology. , 20.11.-24.11.2017, Moscow, Russia .
Taldenkova, E. (2017) Task 5B: Holocene sea-ice variability, water mass distribution and sea-level history in the western Laptev Sea (TRANSDRIFT XXV). [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Timokhov, L. (2017) Task 5A: Long-term variability of the thermohaline state of the Arctic transpolar system (historical data). [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Timokhov, L. (2017) Work Package 5: Long-term and paleoceanographic changes of the Arctic climate system. [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Timokhov, L. and Frolov, I. (2017) The transpolar system of the Arctic Ocean in 2016: state and tendency of changes. [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Willmes, S. (2017) Task 2B: Monitoring of sea-ice properties and production (multi-sensor satellite remote sensing). [Talk] In: Kickoff Workshop of The Changing Arctic Transpolar System – CATS. , 16.05.-18.05.2017, St. Petersburg, Russia .
Heinemann, G., Kohnemann, S. H. E., Schnaase, F., Preußer, A. and Willmes, S. (2018) Multi-method long-term analysis of atmosphere-ocean-ice interactions in Arctic polynyas. [Poster] In: EGU General Assembly 2018. , 08.04.-13.04.2018, Vienna, Austria .
Janout, M., Rabe, B., Randelhoff, A., Hölemann, J., Korhonen, M., Hoppmann, M., Horn, M., Pisarev, S., Nikolopoulos, A., Savy, J. P. and Villacieros, N. (2018) Under-ice turbulence and vertical fluxes in the Arctic Ocean in 2015. [Poster] In: POLAR 2018. , 15.06.-26.06.2018, Davos, Switzerland .
Kolling, H., Stein, R. and Fahl, K. (2018) Pan-Arctic distribution of sea ice biomarkers – a synthesis of old and new surface sediment records. [Poster] In: 6. International Conference on "Palaeo-Arctic Spatial and Temporal (PAST) Gateways". , 16.04.-20.04.2018, Durham, U.K. .
Niessen, F., Stein, R., Schreck, M., Matthiessen, J., Gebhardt, C., Hong, J. K., Polyak, L. and Nam, S. I. (2018) Pleistocene glaciations along the East Siberian continental margin and their extensions into the Arctic Ocean. [Poster] In: 6. International Conference on "Palaeo-Arctic Spatial and Temporal (PAST) Gateways". , 16.04.-20.04.2018, Durham, U.K. .
Preußer, A., Willmes, S., Ohshima, K. I. and Heinemann, G. (2018) Sea-ice production in Arctic polynyas between 2002 and 2017. [Poster] In: POLAR 2018. , 15.06.-26.06.2018, Davos, Switzerland .
Selyuzhenok, V., Krumpen, T., Leppäranta, M. and Gerdes, R. (2018) East Siberian Sea fast ice: linkage with atmospheric forcing. [Poster] In: POLAR 2018. , 15.06.-26.06.2018, Davos, Switzerland .
Drüe, C., Schwarz, P., Heinemann, G. and Makshtas, A. (2017) Investigation of low-level jet events in the Laptev Sea area by ground-based remote sensing. [Poster] In: EMS Annual Meeting & European Conference for Applied Meteorology and Climatology 2017. , 04.09.-08.09.2017, Dublin, Ireland .
Janout, M., Hölemann, J. A., Timokhov, L., Gutjahr, O. and Heinemann, G. (2017) Circulation in the northwest Laptev Sea in the eastern Arctic Ocean: crossroads between Siberian river water, Atlantic water and polynya-formed dense water. [Poster] In: EGU General Assembly 2017. , 23.-28.04.2017, Vienna, Austria .
Kraan, C., Piepenburg, D. and Brey, T. (2017) Quantitative models link declining sea ice to pan-Arctic benthic diversity and ecosystem functioning. [Poster] In: Gordon Research Conference. , 26.03.-31.03.2017, Ventura, USA .
Laukert, G. , Frank, M., Hathorne, E. , Krumpen, T., Werner, K. and Peeken, I. (2017) Tracing the origin of Arctic sea ice and freshwater by neodymium isotopes and rare earth elements. [Poster] In: Goldschmidt Conference 2017. , 13.-18.08.2017, Paris, France .
Meulenbroek, K., Siebert, C., Hathorne, E. C. and Frank, M. (2017) Are barium isotopic signals in planktic foraminifera a proxy for palaeosalinity?. [Poster] In: Goldschmidt Conference 2017. , 13.-18.08.2017, Paris, France .
Preußer, A., Heinemann, G., Oshima, K., Willmes, S. and Paul, S. (2017) Long-term analysis of circumpolar polynya regions and ice production in the Arctic. [Poster] In: Arctic Science Summit Week 2017. , 31.03.-07.04.17, Prague, Czech Republic .
Zentek, R., Heinemann, G. and Kohnemann, S. (2017) Ship-borne wind LIDAR measurements in the Arctic and Antarctic. [Poster] In: EGU General Assembly 2017. , 23.-28.04.2017, Vienna, Austria .
Laukert, G. (2017) Ocean circulation and shelf processes in the Arctic Mediterranean traced by radiogenic neodymium isotopes, rare earth elements and stable oxygen isotopes. (Doctoral thesis/PhD), Christian-Albrechts-Universität, Kiel, III, 167 pp.
Paixao, F. (2017) Verifikation eines lokalen Klimamodels in der Arktis anhand gemessener Oberflächen-Flussdichten. (Bachelor thesis), Trier University, Trier, 34 pp.

matches: V. V. V. V. V. V. V. V. V. V. V. V. V.


Plaintiff United States of America ("United States"), pursuant to Section 2(b) of the Antitrust Procedures and Penalties Act ("APPA" or "Tunney Act"), 15 U.S.C. § 16(b)-(h), files this Competitive Impact Statement relating to the proposed Final Judgment submitted for entry in this civil antitrust proceeding.
On May 11, 2004, Syngenta AG ("Syngenta"), Syngenta Crop Protection AG, AstraZeneca Holdings B.V., AstraZeneca PLC, Koninklijke Vanderhave Groep B.V., and Koninklijke Cooperatie Cosun U.A. entered into an agreement under which Syngenta would purchase all the assets of Advanta B.V. ("Advanta"), a seed company jointly owned by AstraZeneca Holdings B.V. and Koninklijke Vanderhave Groep, B.V. The United States filed a civil antitrust Complaint on August 25, 2004, seeking to enjoin the proposed acquisition. The Complaint alleges that the acquisition likely would substantially lessen competition in the market for sugar beet seeds suitable for growing in the United States, in violation of Section 7 of the Clayton Act. As a result of this loss of competition, prices of sugar beet seeds likely would increase and fewer new or improved sugar beet seed varieties likely would be developed, to the detriment of purchasers of sugar beet seeds, sugar beet processors, and consumers of sugar beet-based products.
At the same time the Complaint was filed, the United States also filed a proposed Final Judgment and Hold Separate Stipulation and Order, which are designed to eliminate the anticompetitive effects of the acquisition. Under the proposed Final Judgment, which is explained more fully below, Syngenta is required divest the worldwide sugar beet seed business of Advanta. Under the terms of the proposed Final Judgment and Hold Separate Stipulation and Order, Syngenta will take certain steps to ensure that Advanta's sugar beet seed business is operated as a competitively independent, economically viable, and ongoing business concern that remains independent and uninfluenced by the consummation of the acquisition, and that competition is maintained during the pendency of the ordered divestiture.
The United States and Defendants have stipulated that the proposed Final Judgment may be entered after compliance with the APPA. Entry of the proposed Final Judgment would terminate this action, except that the Court would retain jurisdiction to construe, modify, or enforce the provisions of the proposed Final Judgment and to punish violations thereof.
Syngenta is a corporation organized and existing under the laws of Switzerland, with its principal offices in Basel, Switzerland. Syngenta is the ultimate parent entity of Syngenta Crop Protection AG. Syngenta is engaged in the development, production, and sale of agricultural products, including pesticides and seeds. Syngenta's total sales in 2003 were approximately $6.6 billion. In 2003, Syngenta's sales of sugar beet seeds in the United States were approximately $10 million; its global sugar beet seed sales were $99 million.
AstraZeneca PLC is a private limited company with its headquarters in London, England. AstraZeneca PLC is the ultimate parent entity of AstraZeneca Holdings B.V. Koninklijke Cooperatie Cosun U.A. is a cooperative with its headquarters in Cosunpark 1, the Netherlands. Koninklijke Cooperatie Cosun U.A. is the ultimate parent entity of Koninklijke Vanderhave Groep B.V.
Advanta is a company incorporated in the Netherlands with its headquartersin Kapelle, the Netherlands. AstraZeneca Holdings B.V. and Koninklijke Vanderhave Groep B.V. each hold 50% of the shares of Advanta. Advanta sells various kinds of agricultural seeds throughout the world, with global sales of 444 million euros ($548 million) in 2003. Advanta sells sugar beet seeds in the United States through its business unit Interstate Seeds. Advanta also markets sugar beet seeds in the United States through collaborations with Holly Hybrids, Seedex, and Croplan. In 2003, Advanta directly and through these collaborations had sugar beet seed sales of about $7 million in the United States.
Syngenta's acquisition of Advanta, as initially agreed to by Defendants on May 11, 2004, would lessen competition substantially in the market for sugar beet seeds suitable for growing in the United States. This acquisition is the subject of the Complaint and proposed Final Judgment filed by the United States on August 25, 2004.
Sugar beet seeds are used by growers to produce sugar beets, which in turn are sold to sugar beet processors, who convert them into sugar for human consumption. Sugar beet growers in the United States purchased $50 million worth of sugar beet seeds in 2003.
Sugar beets are grown under many different climatic and environmental conditions throughout the United States. These different growing regions require sugar beet varieties with different characteristics. A sugar beet seed company identifies desirable traits for each region and breeds those traits into new varieties.
Advanta and Syngenta each have invested extensively in sugar beet seed research and development programs over a number of decades. Syngenta has breeding facilities in Longmont, Colorado and in Western Europe. Advanta also has several breeding facilities, all in Europe. Both develop sugar beet varieties specifically for the unique growing conditions found in various regions of the United States. For example, a sugar beet seed variety that is suitable for cultivation in France is not likely, without further development, to be suitable or attractive to growers in Minnesota or Idaho. The seed companies have not been equally successful in developing seeds for the various growing regions of the United States, and they compete to improve their sales in each region by developing seeds with better disease resistance, yield per acre, and sugar content.
Developing marketable sugar beet seeds can take five to ten years. During this development period, the seed developer will conduct coded registration trials in the region where the beet is intended to be grown. The results of these field trials are used to determine which new varieties will be submitted to sugar beet processors for coded registration trials. Each sugar beet processor in the United States annually conducts coded registration trials to select varieties of sugar beet seeds to recommend to the growers in the processor's growing region. These trials take two to three years to complete. Sugar beet growers typically will select for purchase only seed varieties that have been tested and recommended by the sugar processors to which they intend to market their crops.
Sugar beet seed companies that have processor-approved varieties compete for sales to growers based upon price and characteristics desired by growers, for example, traits that lower production costs, offer higher yield per acre, or provide resistance to diseases and pests prevalent in the growers' geographic region.
Syngenta develops and sells sugar beet seeds in the United States under the brand name Hilleshog. Syngenta accounts for nearly 20% of all the sugar beets seed developed and sold in the United States. Advanta sells sugar beet seeds through its business unit, Interstate Seeds, under the brand name Vanderhave. Sugar beet seeds bred from genetic material developed by Advanta are also sold in the United States by Holly Hybrids and other companies. Advanta-bred sugar beet seeds account for more than 16% of the seeds sold in the United States.
The market in the United States for sugar beet seeds suitable for growing in the United States is highly concentrated. Syngenta and Advanta are two of only three significant firms that develop sugar beet seeds for cultivation in the United States. The market for sugar beet seeds suitable for growing in the United States will become substantially more concentrated if Syngenta acquires Advanta. Syngenta's acquisition of Advanta will lessen competition substantially and make more likely increased prices and a slower pace of innovation.
New entry is not likely to thwart these anticompetitive effects. Successful entry into the sugar beet seed business is difficult, time consuming, and costly. Developing a new sugar beet seed variety can take five to ten years. Completing the trial tests required by a sugar beet processing company for acceptance on the processor's approved list of varieties can take an additional two to three years.
The divestiture requirement of the proposed Final Judgment will eliminate the anticompetitive effects of the acquisition in sugar beet seeds by establishing a new, independent, and economically viable competitor. The proposed Final Judgment requires Syngenta, within 90 days after the filing of the Complaint, or 5 days after notice of the entry of the Final Judgment by the Court, whichever is later, to divest, as a viable ongoing business, Advanta's worldwide sugar beet seed business. These assets include all tangible and intangible assets necessary to run Advanta's worldwide sugar beet seed operations, including research and development facilities, customer lists and contracts, all registered plant breeders rights, and licenses. The United States may extend the period of time available to Syngenta to complete the divestiture up to an additional 120 days.
The proposed Final Judgment requires divestment of Advanta's worldwide sugar beet seed business, including two major research facilities located in Europe that focus on sugar beet seeds: a facility in Rilland-Bath, the Netherlands; and a facility in Tienen, Belgium. At these facilities, Advanta also develops sugar beet seeds with characteristics desirable for production in the United States, such as beets that are resistant to diseases found in the U.S., but not in Europe. Advanta then contracts with a U.S.-based company to grow the varieties of seeds it intends for the U.S. market. Requiring the divestiture of Advanta's worldwide sugar beet seed business, including these European operations, will insure that the acquirer will have the assets necessary to continue to develop sugar beet seeds suitable for growing in the United States, as well as to produce and sell those seeds.
The assets must be divested in such a way as to satisfy the United States in its sole discretion that the operations can and will be operated by the purchaser as a viable, ongoing business that can compete effectively in the relevant market. Defendants must take all reasonable steps necessary to accomplish the divestiture quickly and shall cooperate with prospective purchasers.
In the event that Syngenta does not accomplish the divestiture within the periods prescribed in the proposed Final Judgment, the Final Judgment provides that the Court will appoint a trustee selected by the United States to effect the divestiture. If a trustee is appointed, the proposed Final Judgment provides that Syngenta will pay all costs and expenses of the trustee. The trustee's commission will be structured so as to provide an incentive for the trustee based on the price obtained and the speed with which the divestiture is accomplished. After his or her appointment becomes effective, the trustee will file monthly reports with the Court and the United States setting forth his or her efforts to accomplish the divestiture. At the end of six months, if the divestiture has not been accomplished, the trustee and the United States will make recommendations to the Court, which shall enter such orders as appropriate to carry out the purpose of the trust, including extending the trust or the term of the trustee's appointment.
The divestiture provisions of the proposed Final Judgment will eliminate the anticompetitive effects of the acquisition on the market for sugar beet seeds suitable for growing in the United States.
The United States and Defendants have stipulated that the proposed Final Judgment may be entered by the Court after compliance with the provisions of the APPA, provided that the United States has not withdrawn its consent. The APPA conditions entry upon the Court's determination that the proposed Final Judgment is in the public interest.
The APPA provides a period of at least sixty days preceding the effective date of the proposed Final Judgment within which any person may submit to the United States written comments regarding the proposed Final Judgment. Any person who wishes to comment should do so within sixty days of the date of publication of this Competitive Impact Statement in the Federal Register. All comments received during this period will be considered by the Department of Justice, which remains free to withdraw its consent to the proposed Final Judgment at any time prior to the Court's entry of judgment. The comments and the response of the United States will be filed with the Court and published in the Federal Register.
The United States considered, as an alternative to the proposed Final Judgment, a full trial on the merits against Defendants. The United States could have continued the litigation and sought preliminary and permanent injunctions against Syngenta's acquisition of Advanta. The United States is satisfied, however, that the divestiture of assets described in the proposed Final Judgment will preserve competition for the development, production, and sale of sugar beet seeds suitable for growing in the United States.
(B) the impact of entry of such judgment upon competition in the relevant market or markets, upon the public generally and individuals alleging specific injury from the violations set forth in the complaint including consideration of the public benefit, if any, to be derived from a determination of the issues at trial. 15 U.S.C. § 16(e)(1)(A) & (B). As the United States Court of Appeals for the District of Columbia Circuit has held, the APPA permits a court to consider, among other things, the relationship between the remedy secured and the specific allegations set forth in the government's complaint, whether the decree is sufficiently clear, whether enforcement mechanisms are sufficient, and whether the decree may positively harm third parties. See United States v. Microsoft Corp., 56 F.3d 1448, 1458-62 (D.C. Cir. 1995).
[a]bsent a showing of corrupt failure of the government to discharge its duty, the Court, in making its public interest finding, should ... carefully consider the explanations of the government in the competitive impact statement and its responses to comments in order to determine whether those explanations are reasonable under the circumstances.
United States v. Mid-America Dairymen, Inc., 1977-1 Trade Cas. (CCH) ¶ 61,508, at 71,980 (W.D. Mo. 1977).
The proposed Final Judgment, therefore, should not be reviewed under a standard of whether it is certain to eliminate every anticompetitive effect of a particular practice or whether it mandates certainty of free competition in the future. Court approval of a final judgment requires a standard more flexible and less strict than the standard required for a finding of liability. "[A] proposed decree must be approved even if it falls short of the remedy the court would impose on its own, as long as it falls within the range of acceptability or is 'within the reaches of public interest.'" United States v. AT&T, 552 F. Supp. 131, 151 (D.D.C. 1982) (citations omitted) (quoting Gillette, 406 F. Supp. at 716), aff'd sub nom. Maryland v. United States, 460 U.S. 1001 (1983); see also United States v. Alcan Aluminum Ltd., 605 F.Supp. 619, 622 (W.D. Ky. 1985) (approving the consent decree even though the court would have imposed a greater remedy).
1. See United States v. Gillette Co., 406 F. Supp. 713, 716 (D. Mass. 1975) (recognizing it was not the court's duty to settle; rather, the court must only answer "whether the settlement achieved [was] within the reaches of the public interest"). A "public interest" determination can be made properly on the basis of the Competitive Impact Statement and Response to Comments filed by the Department of Justice pursuant to the APPA. Although the APPA authorizes the use of additional procedures, 15 U.S.C. § 16(f), those procedures are discretionary. A court need not invoke any of them unless it believes that the comments have raised significant issues and that further proceedings would aid the court in resolving those issues. See H.R. Rep. No. 93-1463, 93rd Cong., 2d Sess. 8-9 (1974), reprinted in 1974 U.S.C.C.A.N. 6535, 6538.
2. Cf. BNS, 858 F.2d at 463 (holding that the court's "ultimate authority under the [APPA] is limited to approving or disapproving the consent decree"); Gillette, 406 F. Supp. at 716 (noting that, in this way, the court is constrained to "look at the overall picture not hypercritically, nor with a microscope, but with an artist's reducing glass"). See generally Microsoft, 56 F.3d at 1461 (discussing whether "the remedies [obtained in the decree are] so inconsonant with the allegations charged as to fall outside of the 'reaches of the public interest'").


text:
IEEE TRANSACTIONS ON POWER ELECTRONICS, VOL. 27, NO.
Abstract—This paper describes the design and performance of a 6-kW, full-bridge, bidirectional isolated dc–dc converter using a 20-kHz transformer for a 53.2-V, 2-kWh lithium-ion (Li-ion) battery energy storage system. The dc voltage at the high-voltage side is controlled from 305 to 355 V, as the battery voltage at the lowvoltage side (LVS) varies from 50 to 59 V. The maximal efﬁciency of the dc–dc converter is measured to be 96.0% during battery charging, and 96.9% during battery discharging. Moreover, this paper analyzes the effect of unavoidable dc-bias currents on the magnetic-ﬂux saturation of the transformer. Finally, it provides the dc–dc converter loss breakdown with more focus on the LVS converter. Index Terms—Bidirectional isolated dc–dc converters, dc-bias currents, energy storage systems, lithium-ion (Li-ion) battery.
RC-snubber loss. Insulated-gate bipolar transistor (IGBT) equivalent collector–emitter resistance. Total HVS dc resistance. Total LVS dc resistance. MOSFET on-state resistance. HVS equivalent diode forward resistance. MOSFET lead resistances. HVS net dc voltage. LVS net dc voltage. PC40-core parameters. Maximum ac ﬂux.
LIST OF SYMBOLS Air-gap ﬂux density. Maximum dc-plus-ac ﬂux density. Net magnetomotive force. Magnetizing intensity. High-voltage-side (HVS) net dc-bias current. Low-voltage-side (LVS) instantaneous switching currents at speciﬁc times. LVS net dc-bias current. LVS dc-link capacitor ripple current. LVS magnetizing current. RC-snubber rms current. LVS stray inductance. HVS magnetizing inductance. LVS magnetizing inductance. Transformer turns ratio. Transformer HVS turns number. Transformer LVS turns number. Battery power. DC–DC converter power transfer.
40-A·h Li-ion battery bank. this paper considers the power ﬂow as shown in (1).6-V. and N is the transformer turns ratio. However. ω is the angular switching frequency. 50-Hz ac-side through the ac-link inductor Lac . and the switching-ripple-ﬁlter circuit is represented by LF . CF . Parallel connections of multiple capacitors are accompanied by bulkiness. and RF . increased cost. The circuit conﬁguration is similar to that in . the ac current. The so-called “dc-blocking capacitors” are typically used to prevent the transformer from magneticﬂux saturation. 40-A·h Li-ion battery modules connected in series. 2-kW·h Li-ion battery energy storage system based on the 6-kW full-bridge bidirectional isolated dc–dc converter using a 20-kHz transformer. voltage v2 leads voltage v1 .2 V. It provides experimental and theoretical discussions concerning the effect of dc-bias currents on the magnetic-ﬂux saturation of the high-frequency transformer. The battery bank consists of two 26. available high-frequency capacitors may not meet high-current requirements. L F = 44 μH (0. A lossless capacitor is connected in parallel with each of the IGBTs to achieve zero-voltage switching and to minimize . where L S is the background system impedance (<1%). and may damage the dc–dc converter in the worst case. The dc–dc converter with a symmetrical structure consists of two voltagesource converters that are referred to as bridge 1 and bridge 2 in this paper. EXPERIMENTAL SYSTEM A. this paper aims at demonstrating the performance of the bidirectional isolated dc–dc converter for low-voltage and high-current battery applications. This paper presents the 53. NO.  proposed a dc–dc converter power ﬂow model that takes into account the switching-device voltage drop and dead time. 3. The dc–dc converter is in the charging mode when voltage v1 leads voltage v2 . L is the sum of the transformer leakage inductance Ltrans and the auxiliary inductances LAL and LAH . 27. The dc–dc converter allows bidirectional power transfer by means of controlling the phase-shift angle δ [rad] between square voltages v1 and v2 as follows : PD = |δ| VD1 VB N δ 1− ωL π (1) II. and . and δ is denoted as negative. 1. the 6-kW bidirectional isolated dc–dc converter. Li-ion battery bank of 53. However.2 Ω (3%). . and C F = 150 μF (33%) on a three-phase 200-V. Krismer and Kolar  designed a 100-kHz transformer with a low magnetic-ﬂux density not only to achieve a low core loss but also to provide a large safety margin prior to the saturation ﬂux density. The PWM converter is connected at the point of common coupling to the 200-V. TABLE I CIRCUIT PARAMETERS OF THE BIDIRECTIONAL ISOLATED DC–DC CONVERTER 355 V to keep the dc-voltage ratio of the HVS to LVS close to the transformer turns ratio. and reliability. MARCH 2012 Fig. VOL. reduce efﬁciency.2-V. VB is the amplitude of v2 .2%). 40 A·h connected to the 6-kW bidirectional isolated dc–dc converter. and δ is denoted as positive. The high-voltage dc bus is adjusted between 305 and where VD1 is the amplitude of v1 . 6-kW. However. L A C = 280 μH (1. cost.2 V and the operating voltage range of 50–59 V are determined from a practical point of view. the LVS uses laminated bus bars so that ripple currents can ﬂow into the dc capacitor CD2 that is the combination of electrolytic capacitors and high-frequency ﬁlm capacitors. and the 53.3%). To minimize stray inductances. Klopper and Ferreira  proposed a sensor for measurement of ﬂux below a saturation level. In the discharging mode.2-V. This consideration helps in designing an appropriate air-gap length in high-frequency transformers with different voltage and current ratings. Table I summarizes the parameters of the electrical components in the bidirectional isolated dc–dc converter. and 50-Hz base.1238 IEEE TRANSACTIONS ON POWER ELECTRONICS. Xie et al. and decreased reliability. They cause additional current stress in the switching devices. they did not address any dc-bias current in the transformer. However. R F = 0. The overall loss breakdown of the dc–dc converter compares the loss distribution of the low-voltage high-current converter with that of the high-voltage low-current converter. Bridge 1 consists of four 600-V. Note that both the nominal battery voltage of 53. Each IGBT module contains two devices in series. considering system-level safety. 1 shows the experimental setup consisting of a threephase pulse-width-modulated (PWM) converter. 200-A trench-gate IGBTs (CM200DY-12NF). System Conﬁguration Fig.
and the discharging operation is carried out from an initial voltage of 55. 2.7 V. the sum of the wire bond resistance. Photo of the two Li-ion battery modules used for experiment. Altera’s Max 7000s complex programmable logic device is used to generate eight gate signals for all the gate-drive circuits of the IGBTs in bridge 1 and the MOSFETs in bridge 2. 3–6 were observed through the Tektronix TDS3014B.8 V. 1 TABLE III ADJUSTMENT OF HVS DC-LINK VOLTAGE WITH BATTERY VOLTAGE Fig. vD1 and the battery voltage vB during battery charging and discharging. Each MOSFET module also contains two devices in series. However. The peak current of i2 is 153 A. The high-voltage dc bus is 355 V. turn off overvoltage across the collector–emitter terminals of the IGBT. Table III indicates the relation between the dc voltage at the HVS.: DESIGN AND PERFORMANCE OF A BIDIRECTIONAL ISOLATED DC–DC CONVERTER FOR A BATTERY ENERGY STORAGE SYSTEM 1239 TABLE II DATA OF THE LI-ION BATTERY MODULE USED IN THE EXPERIMENTAL CIRCUIT SHOWN IN FIG. The switching periods of bridges 1 and 2 are the same as 50 μs (20 kHz). Fig. The effect of stray inductance at the LVS is seen as the following change in voltage v2 . Adjustment of vD1 is carried out by changing the reference voltage of the three-phase PWM converter. The time resolution of the controller is 40 ns/bit. Li-Ion Battery Modules Table II presents the speciﬁcations of the Li-ion battery modules of the Li-ion battery bank shown in Fig. III. 3 presents the ac voltage and current waveforms of bridges 1 and 2 when the Li-ion battery is charged and discharged at PB = ±5. EXPERIMENTAL RESULTS This section demonstrates the performance of the designed dc–dc converter in charging and discharging the Li-ion battery bank by adjusting the HVS dc voltage vD1 . From the datasheet by Nihon Inter Electronics Corporation. In Fig. No voltage-balancing circuit is required for the series-connected battery modules. and its speciﬁc power to be 313 W/kg.78 μs. It may also cause MOSFET drain–source overvoltages and increase the switching loss.7 A/μs.36 A in i2 . contact resistance between the source and drain metallization and the silicon. with each of the four MOSFETs.5 mΩ.16 A in i1 . The rate of change of current causes a voltage drop of ∼4 V across the stray inductance in bridge 2.e. Bridge 2 is operated in synchronous rectiﬁcation mode to minimize conduction loss. For the duration of a phase-shift angle of δ = 5. The dc voltage vD1 is controlled with the battery voltage vB in such a way as to minimize the voltage change across the auxiliary inductors and transformer leakage inductor. (2) dt A nonnegligible stray inductance of a few tens of nanoHenries makes it difﬁcult to achieve soft switching at turn off at the LVS.TAN et al.9 V. A small-sized RC snubber is connected in parallel. i. each of which is rated at 26. Each of the modules consists of seven Li-ion battery cells connected in series.9 kW. where di2 + vB . as the battery voltage vB at the LVS varies. the rate of change of i2 (within the dotted lines) is calculated as di2 /dt = −46.6 V and 40 A·h.24 μs is set for each leg in bridges 1 and 2. RDS(ON) . 3(a). v1 leads v2 because the Li-ion battery is charged.. Bridge 2 consists of four 100-V. 0. Control Method The control method is based on an open-loop. feedforward control intended for investigating the basic operating performance of the dc–dc converter in the Li-ion battery energy storage system. The battery charging operation is carried out from an initial voltage of 54.6 nH from (2). Experimental Waveforms Fig. 1. and the battery voltage (the low-voltage dc bus) is 59 V. A. Due to the existence of ﬁnite turn on and turn off times of the IGBTs and MOSFETs. This gives the stray inductance as Lσ 2 = 85. v2 = Lσ 2 . where the nominal voltage of each battery cell is 3.29◦ /bit. C. the on-state resistance. the speciﬁc energy of the Li-ion battery module can be determined to be 63 W·h/kg. a dead time of 1. to reduce its switching loss and to damp out an overvoltage and the resultant ringings. The switching-frequency-based waveforms of Figs. The dc-bias current is 0. The initial voltage is measured at iB = 0. and −4.6 mΩ. 500-A MOSFETs (PDM5001). B. 2 shows the photo of the two Li-ion battery modules used in the experimental setup. From the speciﬁcations given. is as low as 0. and the contact resistance between the metallization and lead frame is not negligible because the total resistance Rwcm l reaches 0.
and iB .2 A/μs. (a) Battery charging at P B = 5. 5. V D 1 = 355 V. For the duration of a phase-shift angle of δ = 7. 3(b).35 μs. Waveforms of v D 1 . MARCH 2012 Fig. and V B = 59 V. v2 leads v1 because the Li-ion battery bank is discharged. and 6. NO.9 nH. Drain–source and gate–source voltages of a leg in bridge 2 at P B = 5.9 kW.5 V across the stray inductance in bridge 2.05 A in i2 . (b) Discharging mode at P B = −5.9 kW. The peak current of i2 is 171 A. v B . and the battery voltage is 50.9 kW (V D 1 = 305 V). This gives the stray inductance as Lσ 2 = 82. The high-voltage dc bus is 305 V.9 kW (V D 1 = 355 V). Experimental waveforms with dc-voltage control at the HVS. the rate of change i2 (within the dotted lines) is calculated as di2 /dt = −42.9 kW. (a) Charging mode at P B = 5. In Fig. .5 V.33 A in i1 . The dc-bias current is 0. 27. Fig.1240 IEEE TRANSACTIONS ON POWER ELECTRONICS. 3. VOL. 4. which is nearly equal to the value from Fig. 3. (b) Battery discharging at P B = −5. The rate of change of current causes a voltage drop of ∼3.
8% between PB = −1. However. especially during battery charging because a higher switching current causes a higher turn off overvoltage.6 kW. This is deduced from the rise time of the MOSFET drain–source voltage. The maximum efﬁciency of the converter is achieved around the onset of zero-voltage switching. 6(b) presents the time-expanded waveforms of vDS and iR C . and the battery power PB is calculated from measurements of vB and iB by using the Hioki 3139 power meter having a moving average function. 4(a) shows the waveforms of the dc voltage at the HVS.TAN et al. and the battery current when the battery is charged at PB = 5. The adjustment of vD1 with the variation of vB can minimize the rate of change of the currents i1 and i2 over the time interval of conduction. Additionally. it is observed from Fig. B. (a) Drain–source voltage and RC-snubber current. At battery charging. The losses in the Li-ion battery. Note that iD2 is 0 during the resonant transition time. this paper will consider the low-side dc-link voltage as the battery voltage vB . 5 presents the time-expanded waveforms of the drain– source and gate–source voltages of a leg in bridge 2 at PB = 5. 4 that the current iB .9 kW. Effects of the RC-snubber on a MOSFET in bridge 2 during battery charging at P B = 5. Note that both functions use the least-mean-square approximation. which is 6% during battery charging and 9% during battery discharging. This means that an amount of current ﬂows in the MOSFET during turn off. and that of the other diagonal MOSFET pair swings from vB to 0. the measured converter efﬁciency peaks at 96% at PB = 1.9 kW. At battery discharging. and that the current results in a nonnegligible turn off switching loss. 1 and Table I. which is almost simultaneous to the rise time of the RC-snubber current. RC snubbers mitigate the overvoltages and ringings across the MOSFET drain–source terminals. and VB = 59 V.9 μs. Fig. which is less than half of the MOSFET peak turn off current. Therefore. iB should contain a 40-kHz component. As bridge 2 is operated in synchronous rectiﬁcation. Fig. Since i2 has a frequency of 20 kHz. During the resonant-transition time. or out of. 7 shows the measured plots of the dc–dc converter efﬁciency and the battery terminal voltage when the battery is charged and discharged between 500 W and 5.9 kW. This stray inductance value is used to design the RC snubbers shown in Fig. 3(a).2 kW and PB = −2. This method contributes to minimizing the peak switching current. (b) Time-expanded waveform of v D S and iR C . the snubber loss due to hard switching or incomplete zero-voltage switching. VD1 = 355 V. On the other hand. the PWM converter.9 kW. Fig. 7 are ﬁtted with an exponential function. measured converter efﬁciency averages at 96. The .9 kW.: DESIGN AND PERFORMANCE OF A BIDIRECTIONAL ISOLATED DC–DC CONVERTER FOR A BATTERY ENERGY STORAGE SYSTEM 1241 Fig. which increases the turn off switching loss. The time taken to charge the snubber capacitor across the MOSFET to vB (resonant-transition time). ﬂowing into. 6. The current ﬂow in the RC snubber produces a snubber loss in bridge 2 that is estimated in Section V-A. Fig. and the cables connecting the battery and PWM converter to the dc–dc converter are not considered. An observable decrease in converter efﬁciency exists when the battery Fig. The ripples in the battery voltage vB are also negligible. and the switching loss in bridges 1 and 2 would dominate the loss at this low-power level. 6(a) shows the drain–source voltage vDS of a MOSFET in bridge 2 and the RC-snubber current iR C of the corresponding snubber circuit during battery charging at PB = 5. Therefore.9 kW. Pdc1 is calculated from measurements of vD1 and iD1 . The accuracies of the measuring instruments are listed in the Appendix. The conduction loss in bridge 1. Converter Efﬁciency Fig. the transformer. whereas the measured battery voltage points are ﬁtted with a ﬁrst-order polynomial function. In bridge 2. the current iB becomes a part of the rectiﬁed current of i2 . and that the ac current of bridge 2 i2 circulates in the four RC snubbers. Fig. A 40-kHz ripple voltage exists at the HVS. the conduction loss in bridge 2 should not be signiﬁcant at a low conduction current. The efﬁciency of the dc–dc converter is less than 92% at low power levels (<1 kW). 4(b) shows those when the battery is discharged at PB = −5. the drain–source voltage of one diagonal MOSFET pair swings from 0 to vB . Power at the HVS. The measured efﬁciency points in Fig. including the settling time of the voltage ringings (parasitic-resonance time) is approximately 0. 6(b) implies that the turn off switching loss in bridge 2 may not be negligible. the Li-ion battery voltage. Fig. indicating that the snubber capacitor is not large enough to minimize the rate of change of the MOSFET drain–source voltage. voltages vGSU and vDSU show that zero-voltage switching at turn on is achieved in the upper MOSFET with a negligible turn on switching loss. a rather ﬂat top ac current is observed in i1 and i2 . the peak current in the RC snubber is 22 A. and the LVS auxiliary inductors. Moreover. The so-called “Miller effect” that lasts for approximately 400 ns is observed from the gate–source voltage of the lower MOSFET vGSL . the Li-ion battery bank have almost none of the 40-kHz ripples.2 kW.
Almost no power ﬂows when the reference phase-shift angle is about −4◦ . The experimental phase-shift angle is the reference angle from the controller. The maximum ac-ﬂux density Bac can be calculated as Bac = VD1 = 0. at δ = 0. When charging. and δ=− π + 2 ωπLPD π2 + 4 VD1 VB N (4) Table IV presents the speciﬁcations of the 20-kHz transformer with an air-gap length of 1 mm. and VB are substituted into (3) or (4). In order to calculate the theoretical phase-shift angle. 3. The curves for the theoretical and experimental points are ﬁtted with the function in (1) using the least-meansquare approximation. 27. The absolute value of the power transfer increases. The magnetizing inductance at the LVS is equivalent to 48. power is between ±3 and ±6 kW. The deviation is more severe when the dc-voltage ratio of the HVS to the LVS is not equal to the transformer turns ratio. Power Transfer Versus Phase-Shift Angle Fig.1242 IEEE TRANSACTIONS ON POWER ELECTRONICS. its LVS-referred peak and rms . where the core material is ferrite PC40. The estimation of loss breakdown is shown in Section V. (a) Battery charging. but it is considered reasonable in making the comparison. Bidirectional power transfer versus phase-shift angle. PD is equal to the power at the HVS Pdc1 and it is denoted as negative. Fig. 7.76 mH/36). VOL. as the phase-shift angle increases in the positive and negative directions. but is 500 W. the power transfer is not zero.  have proposed a more accurate power transfer model for the dc–dc converter. IV. 8. Measured dc–dc converter efﬁciencies and battery voltages. TRANSFORMER OF 20 KHZ This section discusses the dc-bias current phenomenon observed in the transformer. and switching losses. An air-gap length of 1 mm is inserted to prevent the transformer from magnetic-ﬂux saturation. This is the result from an increase in conduction.9 μH (=1. 8 compares the experimental and theoretical power transfer PD versus phase-shift angle δ during battery charging and discharging. A. ohmic. The calculation of the theoretical phase-shift angle is derived from (1) to be δ= π − 2 ωπLPD π2 − 4 VD1 VB N (3) in the calculation. When discharging. MARCH 2012 Fig. C. the theoretical and measured curves are seen to be in good agreement because the dc-voltage ratio of the HVS to the LVS is kept close to the transformer turns ratio. The model is useful for accurate power management when vD1 = N vB . Note that the theoretical calculation is not entirely idealized due to the usage of the experimental values where VD1 = 360 V is the rated voltage at the HVS. the power transfer PD is equal to the battery power PB and it is denoted as positive. and at the battery voltage of 59 V. The magnetizing current is triangular. resulting in a power transfer that deviates from the converter power transfer model in (1). copper. even in the worst case that considers transient conditions and a margin of manufacturing and component tolerances. VD1 . For the charging and discharging modes of operation. Xie et al. The dead time in the converter causes a phase difference between voltages v1 and v2 .104 T 4N1 Ae f (5) during battery discharging. the measured values of PD . Based on the experimental results. which is divided into eight different regions of operation. NO. Maximum AC-Flux Density and Magnetizing Current during battery charging. as the operating current increases to 100 A during battery charging. (b) Battery discharging. and to 120 A during battery discharging at the LVS.
and the equivalent diode forward resistance have a role in reducing the dc-bias currents in the dc–dc converter. 9 shows the theoretical digital control signals for the two diagonal-pair MOSFETs. DC-Bias Currents at the HVS and LVS The bidirectional isolated dc–dc converter produces squarewave ac voltages that are applied to the HVS and LVS of the 20-kHz transformer. The same situation can appear in square-wave voltage v1 at the HVS.16 mΩ. Table V presents the measured dc-resistance values of the passive components and connecting cables in the dc–dc converter. As shown in Fig. The dash-line waveform of v2 represents the ac voltage at the LVS under the ideal condition. the presence of the delay-time differ- Fig.5 V.75 mΩ from the same data sheet. However. The solid-line waveform of v2 shows the following difference caused by the delays: The diagonal-lined voltage-time area is larger than the gray-shaded voltage–time area. 5) unequal turn on and turn off times of the IGBTs and MOSFETs. Based on the calculated dc resistance Rdc2 . 9.TAN et al. the polarity of v2 changes from negative to positive when the gate signal for diagonal-pair 2 changes from high to low. and no transient overvoltage is considered in Fig. Manufacturing and component tolerances make it impossible to predict accurate degrees of mismatches in the IGBTs/MOSFETs and their gate-drive circuits at the stage of system design. resulting in a positive or negative net dc voltage every switching cycle. ence deﬁned by Δt(=t1 − t2 ) yields a net positive dc voltage in the LVS every switching cycle. and it is estimated as 7. A constant current source of Idc = 5 A was connected across the transformer HVS. In this case.3 and 8. possible delays in the turn on and turn off times of the MOSFETs and the output signals from their gate-drive-circuits can defer the polarity change in v2 . and the ensuing square-wave ac voltage v2 at the LVS under both ideal and practical conditions. which excludes the threshold voltage from the iA –vAK curve of the diode. Similarly. and cables. B.05 A at the rated power in the discharging mode. An unequal voltage–time area in each ac voltage would result in a net dc voltage in either side of the transformer. the IGBT equivalent collector–emitter resistance. The total dc resistance at the LVS in the charging and discharging modes Rdc2 is expressed as Rdc2 = 2RDS(ON) + 2Rwcm l + RALo + R2o = 3. Fig. This burdens the switching devices with additional current stress. 2) unequal saturation voltage in the IGBTs. the voltage–time area for one half of a switching cycle is equal to that for the other half of the switching cycle. The dc-bias current in the LVS was measured as 6. (6) The total dc resistance at the HVS in the charging and discharging modes Rdc1 is expressed as Rdc1 = 2RFW D (or 2RCE ) + RAHo + R1o + RCBLo = 42 mΩ(or 33. These include the following: 1) unequal gate-drive circuits. RCE is the equivalent collector–emitter resistance in the charging mode. the delay-time difference Δt is calculated as Vdc2 T = 19 ns (8) Δt = VB where T = 50 μs and VB = 50. other components. . It shows that at times ta and tc . the net dc voltage of v2 Vdc2 is deduced as 19. An optimal design in the air-gap length results in a reasonable magnetizing current in the transformer. A net negative dc voltage is also possible. Note that the on-state resistance across each MOSFET is assumed equal. There are several factors that can cause dc-bias currents in the transformer. 3. whereas Idc = 25 A was connected across the transformer LVS. From Fig. and the terminal dc voltage of each of them was measured. and may damage the dc–dc converter. 4) asymmetry in transient overvoltages across the IGBTs and MOSFETs. Note that the MOSFET on-state resistance. 9. A dc-bias current can cause magnetic-ﬂux saturation that produces high current pulses in the transformer.8 mΩ from the datasheet by Mitsubishi Electric. Hence. 9.1 mV. and it is estimated as 3.9 A. an amount of dc-bias current was observed in the experimental waveforms.9 mΩ) (7) where RFW D is the equivalent diode forward resistance in the discharging mode. which excludes the threshold voltage from the iC –vCE curve of the IGBT.: DESIGN AND PERFORMANCE OF A BIDIRECTIONAL ISOLATED DC–DC CONVERTER FOR A BATTERY ENERGY STORAGE SYSTEM 1243 TABLE IV SPECIFICATIONS OF THE 20-KHZ TRANSFORMER USED IN EXPERIMENT currents are 15. 3) unequal on-state resistance of the MOSFETs. Theoretical MOSFET gate signals and the resultant ac voltage v 2 at the LVS that emphasizes on the time delay during the dead time. the polarity of v2 changes from positive to negative when the gate signal for diagonal-pair 1 changes from high to low. reduces converter efﬁciency. At the time tb . respectively.
This means that the operating condition is almost at the saturation ﬂux density because Fig.35 T at 120 ◦ C. the net magnetomotive force can be expressed as Fnet = N2 I2o − N1 I1o = 24. 10(b) indicates that the change in ﬂux with current is approximately 0. the positive dc-bias currents at both HVS and LVS produce a dc ﬂux that has a slight cancellation effect. Therefore. Each dc-bias current produces a dc ﬂux that has a slight cumulative effect. The net magnetomotive force is expressed as Fnet = −(N2 I2o + N1 I1o ) = −31. Note that VD1 is much higher than Vdc1 . the initial dc-ﬂux density Bdc is 0. Net DC Magnetomotive Force In the discharging mode of Fig. and μ0 (=4π × 10−7 H/m) is the permeability of air. 0.9 A·turns in the charging mode. the calculated delay-time differences are only approximations. (a) Comparison between the transformer with no air gap and four different air-gap lengths. the net dc voltage of v1 Vdc1 is deduced as 13. 10(b).33 A at the rated power in the discharging mode.0 mm at a temperature of 120 ◦ C. 10 shows the relationship of the magnetic ﬂux density and ﬂux with the magnetomotive force of the transformer with no air gap and air-gap lengths of 0. (b) The magnetomotive force is expanded.0. From Fig.9 A·turns. This means that the controller resolution is not responsible for the dc-bias currents. As a result. 3(b). measurements with reasonable accuracy are challenging. Since the delaytime differences are very small. 3. The saturation ﬂux density of the core reduces to 0.5% to 2% of reading). VOL. Design of an Optimal Air-Gap Length In the worst case of the experimental results.3 ns VD1 (9) where VD1 = 305 V. NO. and that measurements of the net dc voltages Vdc1 and Vdc2 cannot be carried out because reasonable accuracy is unachievable as v1 and v2 also include transient overvoltages. 1.241 T for the transformer with no air gap. magnetomotive force and magnetic ﬂux density: F = Hlc + lg Bg μ0 (12) where I1o and I2o are the dc-bias currents at the HVS and LVS. The following assumption is made: Neither fringing effects at the air gap nor leakage ﬂux exists. Based on the dc resistance Rdc1 . and Lm 1 = 1. as shown in (11). MARCH 2012 TABLE V DC-RESISTANCE VALUES OF THE PASSIVE COMPONENTS IN THE DC–DC CONVERTER The dc-bias current in the HVS was measured as 0. D. (11) where H is the magnetizing intensity.9 mV. Considering that the measured values are subjected to the accuracies of the measuring instruments (0.008 mWb/A in the transformer with the air-gap length of 1 mm. and 2. 3(a). The delay-time difference Δt is calculated as Δt = Vdc1 T = 2.1244 IEEE TRANSACTIONS ON POWER ELECTRONICS.008 mWb/A di N2 (13) where N = 6 from Table I.2. Bg is the air-gap ﬂux density. Fig.4 A·turns (10) Fig. Ferrite-core (PC 40) ﬂux density and ﬂux versus magnetomotive force at 120 ◦ C. The dc-bias currents can also be of opposite polarity. C. as shown in the charging mode of Fig. the dc-bias currents at the HVS and LVS accumulate to produce a net magnetomotive force of 31.76 mH and N2 = 6 from Table IV. the delay-time differences in (8) and (9) are calculated based on the measured dc-bias currents in the HVS and LVS and the measured dc resistances that are shown in Table V. This theoretical value is in good agreement with the one calculated from the measured magnetizing inductance value as follows: (Lm 1 /N 2 ) dφ = = 0. The delay-time differences are seen to be less than the controller resolution (40 ns). respectively. 10. 27.5. These curves are derived based on the B–H curve obtained from the datasheet of ferrite PC40 with the help of the following basic relationship between .
the transformer with an air-gap length of 1 mm has an initial dc-ﬂux density of 0.139 T.6 W. Although the method of loss breakdown is the same as that in the previous papers . and k = 150 W · Hz−1 · T−2 · m−3 from the datasheet. V. From the measured waveform of iR C in Fig.4% of the rated current. which is close to the saturation ﬂux density. for various air-gap lengths. respectively. the RC snubbers in the LVS produce an amount of loss.4 W.9 kW and PB = 4. which could increase the copper loss. the inductor core losses in the HVS and LVS at PB = 5. and that for the dc-ﬂux density could come from a margin of manufacturing and component tolerances. LOSS BREAKDOWN This section presents the estimated loss distribution in the dc–dc converter under the following two battery charging conditions: One is at PB = 4.5 to 1 mm. At PB = 4. However. Assuming that the same rms current ﬂows in the other three RC snubbers. the magnetizing inductance.9 kW. Core Loss in Auxiliary Inductors The voltage drop across the high-voltage and LVS inductors are deduced from the rate of change of currents i1 and i2 per switching cycle. For an air-gap length of 0. The effect of dc-bias currents on the transformer would become more severe when SiC-MOSFETs are employed in the HVS. respectively.1 kW.345 T.86 T. in which the nanocrystalline soft-magnetic material with a saturation ﬂux density of 1. which is obtained from the curves in Fig. AND DC-FLUX DENSITY FOR VARIOUS AIR-GAP LENGTHS AT 120 ◦ C the gate-drive circuits could aid in minimizing or even canceling the dc magnetomotive force at the primary and secondary sides with the same voltage and current ratings. the maximum dc-plusac ﬂux density in the possible worst case would be 0. Note that the maximum dc-plus-ac ﬂux density in the possible worst case for the transformer with an air-gap length of 1 mm.9 kW and PB = 4.035 × 2) = 0. B.3 and 37.9 A·turns. VD1 = 340 V.1 kW. The total loss can also be calculated from its energy loss for one switching cycle as follows: 2 PR C = 4CSL VB f. Using the Steinmetz parameters and the improved generalized Steinmetz equation. The dc–dc converter in . β = 2. and IB = 70 A.035 T. . . and Lm 2 is the magnetizing inductance at the LVS.3 T (14) where the doubling factor for the ac-ﬂux density could come from start-up and transient conditions. This would result from using the same IGBTs and gate-drive circuits. However. Manufacturing and component tolerances are unavoidable in any practical bidirectional isolated dc–dc converter.278 < 0.14 A. and calculates the auxiliary-inductor core loss using an improved generalized Steinmetz equation  that is suitable for nonsinusoidal excitation waveforms.6 and 4. the rms magnetizing current would be 14. These RC-snubber losses agree fairly well with those calculated from the measured RC-snubber current. VD1 = 355 V.1 kW. the rms current IR C is 2. Slight component mismatches in bridges 1 and 2 produce a net dc magnetomotive force in the transformer with different voltage and current ratings in the primary and secondary sides.27 A at PB = 5. 6(b). Table VI summarizes the peak and rms magnetizing currents. The peak magnetizing current at the LVS is calculated as N2 φac (15) Im 2 = Lm 2 where φac is the maximum ac ﬂux.2 and 20.8 V. For an air-gap length of 2 mm. PR C = 30. and the dc-ﬂux density that would be present when the net magenetomotive force is 31. the maximum dc-plus-ac ﬂux density reaches 0. the total loss in the four RC snubbers is given by 2 PR C = 4IR C RSL = 34.1 kW are 39. Table VI suggests that the acceptable air-gap length would be in the range of 0. The component mismatches are due to the dc–dc converter employing different switching devices and gate-drive circuits. because the on-state resistance of the SiC-MOSFETs is expected to be one-ﬁfth as low as that of the Si-MOSFETs . Bm ax is given by Bm ax = (0.3 T.5 mm. . leaving a decent margin prior to the saturation ﬂux density. VB = 59 V.: DESIGN AND PERFORMANCE OF A BIDIRECTIONAL ISOLATED DC–DC CONVERTER FOR A BATTERY ENERGY STORAGE SYSTEM 1245 TABLE VI CALCULATED MAGNETIZING INDUCTANCE AND CURRENT AT THE LVS. The maximum dc-plus-ac ﬂux density reaches 0. Finally. The Steinmetz parameters of ferrite PC44 are α = 1. in consideration of the possible worst case. This transformer with no air gap did not experience any magneticﬂux saturation. and  had a toroidal-core transformer with a turns ratio of 1:1. VB = 57. (17) The total loss at PB = 5. However. . A.1 kW. 10.9 kW are 3.33 T. RC-Snubber Loss Zero-voltage switching is achieved in both the HVS and LVS at PB = 5.TAN et al.9 kW. was used and the ac magnetic ﬂux density was designed as 0. inserting an appropriate air gap into the transformer ensures stable operation of the dc–dc converter without causing magnetic-ﬂux saturation. named Finemet.104 × 2) + (0. the following differences exist: This paper takes into account the RC-snubber and dc-capacitor losses at the LVS.2 W. Therefore. the inductor core losses in the HVS and LVS are 1. where good matches in the IGBTs and (16) At PB = 4.6 W. the air-gap length of 1 mm is concluded to be optimal for this system. respectively. and IB = 100A .8 W. since IR C = 2. and the other is at PB = 5.
measurements of the currents iD2 and iCD2 are restricted.0 V. Based on the estimated losses in earlier sections.1246 IEEE TRANSACTIONS ON POWER ELECTRONICS. Inoue and Akagi  showed that the total switching loss for the 350-V bidirectional isolated dc–dc converter using IGBTs was estimated as 90 W at 10 kW. respectively.9 kW. the ESR loss in CD2 is 16 W. transformercore. which include the transformer and auxiliary inductors at the rated power. VOL. the LVS switching loss is 93 W. The analysis of iD2 and iCD2 are based on the following assumptions: Since Fig.9 kW. 3. the rms current of iCD2 can be calculated from Fig. it can be assumed that the ripple current ﬂows into the dc capacitors CD2 .1 kW are estimated as 334 W (424 W − 90 W) and 145 W (225 W − 80 W). Therefore. Assuming that the switching loss is proportional to the power transferred. where iD2 can be analyzed from the charging operation.9 μs.1 kW or PB = 5. The total loss in the switching devices.9 kW. inductor-core. V D 1 = 355 V. As specially designed laminated bus bars are used at the LVS.4 A so that the ESR loss in CD2 is 5 W. Relationship between the dc–dc converter loss and switching frequency at P B = 4. VB = 59.52 μs. MARCH 2012 Fig. and IB = 100 A. and at P B = 5. Therefore. the measured dc– dc converter loss and efﬁciency are 424 W and 93.1 kW. respectively. which include conduction and switching . The experimental data were ﬁtted with a ﬁrst-order polynomial and extrapolated to zero switching frequency. Because the waveform of iB contains almost no 40-kHz ripple. 13 shows the estimated overall loss breakdown at battery charging powers of 4.9 kW.8 V. the LVS switching loss is 200 W.8 V. 11(a) shows the experimental waveforms of i2 and iB at PB = 5. V D 1 = 340 V. The total ohmic loss in the dc–dc converter is deduced to be in a range of 4–10 W at PB = 4. VD1 = 355 V. the currents I21 and I22 are measured to be 99 and 73 A. the switching loss in bridge 1 is 27 W at PB = 5. 14 presents the loss distribution in bridges 1 and 2. and V B = 59 V.2 mΩ. and iCD2 is the ripple current ﬂowing in the dc capacitors at the LVS. and switching loss. E. At PB = 5. 12 shows the experimental results relating the dc–dc converter loss with the switching frequency between 10 and 20 kHz for battery powers of 5. The ripple current is obtained as iCD2 = iD2 − iB . Observed and simpliﬁed theoretical current waveforms at P B = 5. At PB = 5. including the conduction. Therefore. HVS switching loss. The estimated loss. Estimated Loss Distribution and Considerations Fig.9 kW. Hence. VB = 57. The ﬁgure illustrates that the switching loss of 200 W in bridge 2 is the largest portion (47%) of the loss in the dc–dc converter. the calculation for the switching currents can be found in  and . They include LVS snubber loss. only the dc and fundamental components are considered in the simpliﬁed waveforms. VD1 = 340 V.9 kW.1 and 5. DC-Capacitor Loss at the LVS Fig. Fig. Fig.1 kW. 6(c) shows that the resonant transition time of i2 is less than 0. The estimated loss excluding the LVS switching loss is 132 W. this time in which iD2 is 0 is negligible with respect to its fundamental frequency (40 kHz). The contactpoint ohmic losses can be estimated to be in the range of 2–5 W. Fig. and IB = 70 A. Hence.1 kW. and the magnetic components. the measured dc–dc converter loss and efﬁciency are 225 W and 94. 27. copper. Referring to the datasheets. 11(b) to be 71.9 and 4.9 kW. C. 11. At PB = 4.9 kW and PB = 4.1 kW and δ = 3. The ohmic loss is considered to be mainly contributed by the contact points and laminated bus bars at the LVS.3%. is 224 W.1 A. 11(b) shows the simpliﬁed waveforms of iD2 and iCD2 . and ohmic loss. and V B = 57. the equivalent series resistance (ESR) of the ﬁlm and electrolytic capacitors can be approximated to be 3. At PB = 4. In addition. The frequency-dependent losses at PB = 5. the switching loss is the most dominant of the frequency-dependent loss. NO.9 kW.1 kW. HVS Switching Loss and Ohmic Loss Fig.8%. the rms current of iCD2 can be calculated to be 39. While the values of I21 and I22 in this section are obtained from experiments. (a) Observed waveforms of i2 and iD 2 . ESR. and 18 W at PB = 4. transformer-core and inductorcore losses. (b) Simpliﬁed waveforms of iD 2 and iC D 2 . Note that I21 and I22 are deﬁned as the instantaneous switching currents of bridge 2 at speciﬁc times. 12. RC-snubber. D.
vol.” Proc. no. 1–8. F. Inoue and H. Kheraluwala. and ID 2 = 70 A. 2007. 1744–1756. Mar. and Y. V D 1 = 355 V. respectively. Liu.05% of full scale. V D 2 = 59 V. and M. Available: http://www.TAN et al. D. 13. V D 1 = 355 V. 63–73. “A bidirectional isolated dc-dc converter as a core circuit of the next-generation medium-voltage power conversion system. Nov. M. One of the best methods of improving the efﬁciency of the dc–dc converter is to operate it at a lower switching frequency.. 6. 2. at a switching frequency of 10 kHz. Appl. However. the dc–dc converter can . B. pp. IEEE. A. pp. no. and that at P B = 5. this class of MOSFET is more optimal at lower switching frequencies to minimize the switching loss. Ind. Ind. pp. Estimated loss breakdown at P B = 4. The results have veriﬁed the proper operation of the Li-ion battery energy storage system. Gascoigne. A. Kheraluwala. 12. M. Crow. 1294–1301. APPENDIX The 9602 ac/dc clamp-on meters of Hioki 3139 have the following speciﬁcations: 1) Voltage meter a) DC-voltage measurement range: 0–600 V. W. REFERENCES  New Energy and Industrial Technology Development Organization (NEDO). CONCLUSION This paper has presented the experimental results from the combination of a 53. The transformer with an air-gap length of 1 mm has been shown experimentally to be robust against magnetic-ﬂux saturation. 22. 2001. Ribeiro. and E. The MOSFET employed in bridge 2 has a very low on-state resistance (0.  R.. 28.  S. and IB = 100 A. Estimated loss distribution in bridges 1 and 2. V D 2 = 59 V. Johnson. L. and the magnetic components at P B = 5.9 kW.2% of full scale. For example.9 kW. D. V B = 57. 12 indicates that the efﬁciency of the 6-kW bidirectional isolated dc–dc converter at the rated power can be improved when the switching frequency of the dc–dc converter is reduced. (2008). “A threephase soft-switched high-power-density dc/dc converter for high power applications. K. “Performance characterization of a high-power dual active bridge dc-todc converter.” IEEE Trans. K. “Energy storage systems for advanced power applications.2-V. no. Fig. Appl. D. “Advancement of energy storage devices and applications in electrical power system. Power Electron. Feb. Smith. The dc-current ranges for the 9278 universal clamp-on current transducer at the HVS and LVS are 0–50 and 0–200 A. Discussions focusing on magneticﬂux saturation due to unavoidable dc-bias currents at the highvoltage and LVSs have been carried out. M. 1991. A. b) accuracy: ±0. From the estimation of loss distribution in the dc–dc converter.” IEEE Trans. Fig.  M. Jul. this method is accompanied by acoustic noise generation and a bulky transformer.1 kW. is 71% of the loss in the dc–dc converter at the rated power. Divan. vol. a large portion of the loss at the rated power is caused by the turn off switching loss at the LVS.2% of full scale. 89. Global warming counter measures: Japanese technologies for energy savings/GHG (greenhouse gases) emissions reduction (Revised ed.” IEEE Trans. Akagi. As a result of a tradeoff.: DESIGN AND PERFORMANCE OF A BIDIRECTIONAL ISOLATED DC–DC CONVERTER FOR A BATTERY ENERGY STORAGE SYSTEM 1247 achieve an efﬁciency as high as 97% at a switching frequency of 5 kHz at the expense of generating acoustic noise and requiring a bulky transformer. General Meeting. and IB = 100 A./Dec. Fig. C. the very low onstate resistance compromises the switching loss in the MOSFET. vol. P. Divan. VI. W. vol. no. Baumann. Sen.” in Proc. R.. 2008. H. De Doncker. pp.5 mΩ). 1992. Their accuracies are ±0. 1.  P.go. 40-A·h Li-ion battery bank with a single-phase full-bridge bidirectional isolated dc–dc converter. 535–542. [Online].). Kroposki. 14. the overall efﬁciency of the dc–dc converter can be improved by 2–3% at the rated power. The extrapolation based on the experimental results shows that at the rated power. 2) Current meter a) DC-current range: 0–500 A (depending on the rating of the clamp-on current transducer). The bidirectional isolated dc–dc converter exhibits high efﬁciency in the low-voltage and high-current operation. Dec. b) accuracy: ±0. V D 1 = 340 V. H. Therefore. 27. IEEE Power Energy Soc.8 V.nedo. losses. and B.jp  S. pp. Arsoy. even in the worst cases.
M. L. Xie. 771–781. Malaysia. He is the recipient of the 2001 IEEE William E. Klopper and J. the 2004 IEEE Industry Applications Society Outstanding Achievement Award. “Transformercoupled multiport ZVS bidirectional dc-dc converter with wide input range. “An Isolated Three-Port Bidirectional DC-DC Converter With Decoupled Power Flow Management. Kolar. 3.” IEEE Trans.. Int.. Tokyo.. in 1976 and 1979. 2002. Power Electron.Eng. Okayama University. “A sensor for balancing ﬂux in converters with a high-frequency transformer link. degree in international development engineering. both from Tokyo Institute of Technology. IEEE Power Electron. Nov.) degree from the University of Shefﬁeld. high-frequency resonant inverters for induction heating and corona discharge treatment processes. pp. Y.S.Eng. she has been a Senior Lecturer in the Department of Electrical Power Engineering. S. no. M. Power Electron. Zhao. Tarui. “A 6-kW. 2003. Oomori. Int.D. Abe.  D. 23. Sun. Jun. 2. Tan is a Graduate Member of the Institution of Engineers Malaysia (IEM) and a Member of the Institution of Engineering and Technology (IET). Conf. Fujihira. degrees from the Tokyo Institute of Technology. Walter and W. “Hybrid modulation for dual-activebridge bidirectional converter with extended power range for ultracapacitor application. Conf.” IEEE Trans.. vol. vol. Kotsopoulos. He has made presentations many times as a Keynote or Invited Speaker internationally. 2008.D. Imaizumi. 25. W. Weischedel and G.. respectively. “Voltage balancing of a 320-V. Available: http://home. 422–429. vol. Tokyo Institute of Technology. “A symmetry correcting pulsewidth modulator for power conditioning applications. Japan. C. Kobayashi. active and passive electromagnetic interference (EMI) ﬁlters. The total citation index for all his papers in Google Scholar is more than 15. vol. MARCH 2012  J.” IEEE Trans.” in Proc. and utility applications of power electronics. Ind.  S. He has received ﬁve IEEE Transactions Prize Paper Awards and nine IEEE Conference Prize Paper Awards. 774–779. no. U. R. Mar. Japan. 21.. His current research interests include power conversion systems. 2008. no. Jul. Zhu. motor drives.” in Proc. (2007). T. Since October 2010. respectively. vol. 2443–2453.  K. “A novel soft-commutating isolated boost full-bridge ZVS-PWM dc-dc converter for bidirectional high power applications. 1434–1442. 46–52. “Accurate power loss model derivation of a high-current dual active bridge converter for an automotive application. and the M. 2006. no. [Online].. 3. Mar. pp. pp. Westerman. Kola. 2755–2765.1248 IEEE TRANSACTIONS ON POWER ELECTRONICS. degree from Tokyo Institute of Technology. 23. Ind. Japan. Madison. Kinouchi. and M. Spec. he was a Visiting Scientist at Massachusetts Institute of Technology (MIT). pp. He was elected as a Distinguished Lecturer of the IEEE Power Electronics and Industry Applications Societies for 1998–1999. in 2010. Nov. no.gsyuasa. W. 2008. Ferreira. Her current research interests include power conversion systems and bidirectional isolated dc–dc converters. 6. pp.” in Proc.2 kV 4H-SiC MOSFETs with the very low on-state resistance of 5 mΩcm2 .  H. Yuasa. Duarte. Akagi was the President of the IEEE Power Electronics Society for 2007–2008. no. He received the B.. Hendrix. 3. Venkatachalam. M. Power Electron.” in Proc. the M. Since January 2000. Freudenberg. “Successful development of 1. and Ph. Jun. L. and S. in 1987.  C. VOL. Tan. degree in electrical and electronic engineering. She received the B. pp. 2-kWh lithium-ion battery energy storage system using a bidirectional isolated dc-dc converter. (PESC). Appl. “Middle-frequency isolation transformer design issues for the high-voltage dc-dc converter.  L. Power Electron. 22. pp.K.  N. Jun. D. IEEE Power Electron. 318–322. “Power ﬂow characterization of a bidirectional galvanically isolated high-power dc/dc converter over a wide operating range. Krismer and J.  H. Tokyo. Power Electron. J. Japan. IEEE Workshop Comput. 27–32. Vinnikov. Laugis. Abdallah. 12-F electric double-layer capacitor bank combined with a 10kW bidirectional isolated dc-dc converter. 1930–1936. he was a Professor in the Department of Electrical Engineering. M. on August 19. where he became an Associate Professor in the Department of Electrical Engineering. pp. NO. From March to August 1996.S. and then at MIT.  S. and J. Appl. vol. Hirofumi Akagi (M’87–SM’94–F’96) was born in Okayama. Tacca. Miura.S. Newell Power Electronics Award. self-commutated back-to-back (BTB) systems. degree from the Nagoya Institute of Technology. and the Ph.” IEEE Trans. Inoue. M. 23. and T. and the M. In 1979. no. 36–41. Ind. Takahiro Abe was born in Ibaraki. Jun. pp. in 1974. Conf. S.html  N. no.” IEEE Trans. Akagi. 2008. and the 2008 IEEE Richard H. no. pp. S. 2007. 27.S. and H. Sep.” in Proc.” IEEE Trans. T. Tokyo. Khambadkone. Tokyo. Power Electron. and ﬂexible ac transmission systems (FACTS) devices. vol. He has authored and coauthored more than 100 IEEE Transactions papers and two invited papers published in the Proceedings of the IEEE in 2001 and 2004.-I. R. L.  Y. Japan. pp. Cambridge. he was a Visiting Professor at the University of Wisconsin. Power Electron. and H. Mar. “A bidirectional dc-dc converter for an energy storage system with galvanic isolation. Tan. Ind. for ten months. He was involved in the research on bidirectional isolated dc–dc converters at Tokyo Institute of Technology. Spec. A. Akagi.000. W. and I Galkin. May/Jun.  G. all in electrical engineering. 1997. degree from Universiti Tenaga Nasional. . 2006. pp. May/Jun. 54–66. such as active ﬁlters. Kajang. as an Assistant Professor. in 2002. “High-power galvanically isolated dc-dc converter topology for future automobiles. Inoue and H.” IEEE Trans.  H.” IEEE Trans. Dr. Symp. A.gyps. Nagoya. 1. Y. 5. S. Zhou and A. all in electrical engineering. 881–891. 2010.  F. 1–4. “Accurate prediction of ferrite core loss with nonsinusoidal waveforms using only Steinmetz parameters. From 1991 to 1999. 2299–2306. no. J. 33.. Industrial Use Lithium-Ion Battery (LIM40/LIM80 Series) (in Japanese). in 2009 and 2011. Tao.. Sullivan. 1951. 45. pp. T. A. 4. pp. Dr. Japan. Watanabe. Okayama. Kaufmann Technical Field Award. Nadia Mei Lin Tan (S’07–M’10) was born in Kuala Lumpur. (IPEC). During 1987. he joined Nagaoka University of Technology. 1. and H. Power Semicond. Universiti Tenaga Nasional.com/products/li/lim40. in 2007. Power Electron. 2. Jan. vol. He is currently working for JFE Steel Corporation./Aug. vol. 2010. Electron. 57. (Hons.” IEEE Trans.  N. Appl. K. pp. vol.. A.” IEEE Trans. J. He received the B. 1973.. 3. and is currently the Senior Past President. he has been a Professor in the Department of Electrical and Electronic engineering. 2010. (PESC). 2009. Round. 6. Nakao. Jun. Akagi. IA-9. Shefﬁeld. Malaysia. Japan. vol. De Doncker. Devices Ics (ISPSD).

matches: V. V. V. V. V. V. V. V. V. V. V. V. V. V. V. V. V. V. V. V. V. V. V. V. V. V. V. V.


the Supreme Administrative Court of 31 August 2011, ref. no.
judgment of the Supreme Administrative Court of 31 August 2011, ref.
of 8 February 2008, ref. no. 450 811 075/428, are annulled.
judicial and other legal protection under Art. 36 of the Charter.
Railways, again as an engineer.
Social Security Administration, pursuant to § 29 let. a) of Act no.
to the total period of insurance in all member states.
Constitutional Court’s case law (in particular, judgments file no. II.
par. 2 of the Charter.
Justice of the European Union (the “ECJ”) for a preliminary ruling.
pension insurer, and for all ordinary courts.
22 June 2011, C-399/09, according to the cited judgments file no. Pl.
the European Union decision of 22 June 2011, C-399/09.
Supreme Administrative Court did not find any arguments for applying it.
and with its intentions. In this regard he refers to judgment file no.
Republic’s entry into the European Union.
ensue from it for a supplementary payment to an old age pension.
judgment of the Supreme Administrative Court of 31 August 2011, file no.
discrimination under Art. 1 a Art. 3 par. 1 of the Charter.
25/2010-146, and 3 Ads 130/2008-204).
judges of the relevant panel and of the parties to the proceeding.
ref. no. Pl. ÚS 5/12-1.
greater detail, a hearing in this matter was waived.
accordance with the legal rules of the Czech Republic.
remaining (Case of 26 January 1999, Terhoeve, C-18/95,  ECR I-345, paragraph 57, and the case law cited).
two requirements should be deprived of such a payment.
50/04, Pl. ÚS 66/04, Pl. ÚS 19/08, and Pl. ÚS 29/09.
functioning, and object are directly connected to Community law.
identity and exceeding the scope of the entrusted competences.
judgment of 22 June 2011, C-399/09 on the present case.
Parliament and Council Regulation (EC) No 629/2006 of 5 April 2006.
date of dissolution o the Czech and Slovak Federal Republic.
par. 1 of the Charter (see file no. II. ÚS 405/02, III. ÚS 252/04, IV.
the Constitutional Court merely stated that “[i]nsofar as Act no.
abandoning the principle audiatur et altera pars.
fundamental rights and freedoms involves several components (file no.
and application of substantive law.
directly applicable provisions of international treaties.
constitutionality of the relationship of § 61 of Act no. 155/1995 Coll.
Republic and the Slovak Republic on Social Security, published as no.
issue of application of the Agreement in its decisions file no.
of 1 January 1993 by the dissolution of the joint Czechoslovak state.
which Act no. 428/2011 Coll. was published, was distributed.
benefit are not governed by any legislation. Thus, § 106a of Act no.
prohibition on payment of social benefits that is not governed by law.
reason for the law ceases to exist, the law itself ceases to exist).
concerning § 106a of Act no. 155/1995 Coll., as amended by Act no.
which the decisions contested by the constitutional complaint arose.
Instruction: This judgment cannot be appealed.
which I refer in full.
I do not consider correct the conclusion that European law, i.e.
cases analogous to the complainant’s is an overreaching by an EU body.
Annex III remain applicable [Art. 7 par. 2 let. c)]. The content of Art.
became EU law, and is a provision that is binding for all member states.
discriminate against nationals of other member states.
receives contradictory instructions, which is it bound to observe.