Heavy-Duty-, On- und Off-Highway-Motoren 2019: Zukünftige Herausforderungen 14. Internationale MTZ-Fachtagung Großmotoren [1. Aufl.] 9783658313708, 9783658313715

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Proceedings

Wolfgang Siebenpfeiffer Hrsg.

Heavy-Duty-, On- und Off-HighwayMotoren 2019 Zukünftige Herausforderungen 14. Internationale MTZ-Fachtagung Großmotoren

Proceedings

Ein stetig steigender Fundus an Informationen ist heute notwendig, um die immer komplexer werdende Technik heutiger Kraftfahrzeuge zu verstehen. Funktionen, Arbeitsweise, Komponenten und Systeme entwickeln sich rasant. In immer schnelleren Zyklen verbreitet sich aktuelles Wissen gerade aus Konferenzen, Tagungen und Symposien in die Fachwelt. Den raschen Zugriff auf diese Informationen bietet diese Reihe Proceedings, die sich zur Aufgabe gestellt hat, das zum Verständnis topaktueller Technik rund um das Automobil erforderliche spezielle Wissen in der Systematik aus Konferenzen und Tagungen zusammen zu stellen und als Buch in Springer.com wie auch elektronisch in Springer Link und Springer Professional bereit zu stellen. Die Reihe wendet sich an Fahrzeug- und Motoreningenieure sowie Studierende, die aktuelles Fachwissen im Zusammenhang mit Fragestellungen ihres Arbeitsfeldes suchen. Professoren und Dozenten an Universitäten und Hochschulen mit Schwerpunkt Kraftfahrzeug- und Motorentechnik finden hier die Zusammenstellung von Veranstaltungen, die sie selber nicht besuchen konnten. Gutachtern, Forschern und Entwicklungsingenieuren in der Automobil- und Zulieferindustrie sowie Dienstleistern können die Proceedings wertvolle Antworten auf topaktuelle Fragen geben. Today, a steadily growing store of information is called for in order to understand the increasingly complex technologies used in modern automobiles. Functions, modes of operation, components and systems are rapidly evolving, while at the same time the latest expertise is disseminated directly from conferences, congresses and symposia to the professional world in ever-faster cycles. This series of proceedings offers rapid access to this information, gathering the specific knowledge needed to keep up with cutting-edge advances in automotive technologies, employing the same systematic approach used at conferences and congresses and presenting it in print (available at Springer.com) and electronic (at Springer Link and Springer Professional) formats. The series addresses the needs of automotive engineers, motor design engineers and students looking for the latest expertise in connection with key questions in their field, while professors and instructors working in the areas of automotive and motor design engineering will also find summaries of industry events they weren’t able to attend. The proceedings also offer valuable answers to the topical questions that concern assessors, researchers and developmental engineers in the automotive and supplier industry, as well as service providers.

Weitere Bände in der Reihe http://www.springer.com/series/13360

Wolfgang Siebenpfeiffer (Hrsg.)

Heavy-Duty-, On- und Off-Highway-Motoren 2019 Zukünftige Herausforderungen 14. Internationale MTZ-Fachtagung Großmotoren

Hrsg. Wolfgang Siebenpfeiffer Vieweg Verlag Chefredakteur ATZ/MTZ Stuttgart, Deutschland

ISSN 2198-7440  (electronic) ISSN 2198-7432 Proceedings ISBN 978-3-658-31371-5  (eBook) ISBN 978-3-658-31370-8 https://doi.org/10.1007/978-3-658-31371-5 Die Deutsche Nationalbibliothek verzeichnet diese Publikation in der Deutschen National­ bibliografie; detaillierte bibliografische Daten sind im Internet über http://dnb.d-nb.de abrufbar. © Springer Fachmedien Wiesbaden GmbH, ein Teil von Springer Nature 2020 Das Werk einschließlich aller seiner Teile ist urheberrechtlich geschützt. Jede Verwertung, die nicht ausdrücklich vom Urheberrechtsgesetz zugelassen ist, bedarf der vorherigen Zustimmung des Verlags. Das gilt insbesondere für Vervielfältigungen, Bearbeitungen, Übersetzungen, Mikroverfilmungen und die Einspeicherung und Verarbeitung in elektronischen Systemen. Die Wiedergabe von allgemein beschreibenden Bezeichnungen, Marken, Unternehmensnamen etc. in diesem Werk bedeutet nicht, dass diese frei durch jedermann benutzt werden dürfen. Die Berechtigung zur Benutzung unterliegt, auch ohne gesonderten Hinweis hierzu, den Regeln des Markenrechts. Die Rechte des jeweiligen Zeicheninhabers sind zu beachten. Der Verlag, die Autoren und die Herausgeber gehen davon aus, dass die Angaben und Informa­ tionen in diesem Werk zum Zeitpunkt der Veröffentlichung vollständig und korrekt sind. Weder der Verlag, noch die Autoren oder die Herausgeber übernehmen, ausdrücklich oder implizit, Gewähr für den Inhalt des Werkes, etwaige Fehler oder Äußerungen. Der Verlag bleibt im Hinblick auf geografische Zuordnungen und Gebietsbezeichnungen in veröffentlichten Karten und Institutionsadressen neutral. Springer Vieweg ist ein Imprint der eingetragenen Gesellschaft Springer Fachmedien Wiesbaden GmbH und ist ein Teil von Springer Nature. Die Anschrift der Gesellschaft ist: Abraham-Lincoln-Str. 46, 65189 Wiesbaden, Germany

Vorwort

Für die Hersteller von Großmotoren für mobile, stationäre und maritime Anwendungen sowie deren Zulieferer sind die ständig schärfer werdenden Abgasschadstoffgrenzwerte bei Dieselmotoren eine große Herausforderung. So haben sich beispielsweise die EU-Staaten darauf geeinigt, die CO2-Grenzwerte für Lkw und Busse bis 2030 im Vergleich zu 2019 um 30% zu verringern. „Herausforderung CO2-Grenzwerte“ lautet demzufolge auch das diesjährige Motto der jährlich stattfindenden MTZ-Fachtagung „Heavy-Duty-, On- und ­Off-Highway-Motoren“. Die Schwerpunkte der am 26. und 27. November 2019 in Friedrichshafen stattfindenden Konferenz liegen auf neuen Motoren und Konzepten, der Verbesserung des Grundmotors und der Schadstoffreduzierung. Eine begleitende Fachausstellung sowie die Besichtigung des MTU-Werks von ­Rolls-Royce runden das Programm ab. Nutzen Sie die Gelegenheit, sich über neueste Lösungen zu informieren, Ihr Netzwerk zu erweitern und wertvolle Kontakte zu knüpfen. Ich freue mich auf Ihre Teilnahme an der Tagung. Für den Wissenschaftlichen Beirat Dr. Alexander Heintzel

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Editorial

The increasingly stringent exhaust emission limits for diesel engines represent a major challenge for the manufacturers of large engines for mobile, stationary and maritime applications. One example of this is the agreement by the EU member states to reduce by 30% the CO2 thresholds for trucks and buses from 2019 levels by 2030. For this reason the theme of this year’s annual MTZ Conference “Heavy-Duty, On- and Off-Highway Engines” is “The challenge of CO2 limits”. The key topics at the conference, which takes place on 26 and 27 November 2019 in Friedrichshafen, are new engines and concepts, improving the basic engine and reducing pollution. The event also includes a trade exhibition and a visit to the RollsRoyce MTU plant. Don’t miss this opportunity to find out about the latest solutions, expand your network and make valuable contacts. I look forward to welcoming you to the conference. On behalf of the Scientific Advisory Board Dr. Alexander Heintzel

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Inhaltsverzeichnis

The role of synthetic fuels in an integrated energy system Dr. Achim Schaadt, Robert Szolak, Prof. Christopher Hebling, Florian Rümmele, Max Julius Hadrich, Mohamed Ouda und Bernd Danckert DEUTZ engine portfolio below 56 kW to meet EU Stage V NRMM emission standard Christian Opitz, Christoph Klein, Hartmut Sieverding und Heiner Bülte Compact and powerful: the new 9-liter diesel engine from MAN for off-highway applications Tobias Herrmann, Vanessa Simon, Markus Fuchs, Marc Winterhoff und Reinhard Lämmermann CatVap® – a new heating measure for exhaust aftertreatment system Robert Szolak, Bernd Danckert, Dr. Alexander Susdorf, Paul Beutel, Katharina Pautsch, Christian Ewert, Florian Rümmele, Anand Kakadiya und Dr. Achim Schaadt A novel low-cost aftertreatment solution for lean‑burn gas engines Matthew Keenan, Jacques Nicole und Ben Rogers CFD simulation of particle deposition in exhaust gas treatment systems. Dorian Holtz, Conrad Gierow, Robert Bank, Dirk Kadau und Flavio Soppelsa

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Inhaltsverzeichnis

Variably honed cylinder liners, iron-based cast pistons and variably coated piston rings as PCU system for friction loss and TCO reduction Dr.-Ing. Daniel Hrdina, Dipl.-Ing. Marco Maurizi, Bartek ­Lemm, Dipl.-Ing. Hakan Kahraman und Dipl.-Ing. Guilherme Soares de Faria Parameter study of the appearance and allocation of small oil aerosol particles at the piston, piston ring and cylinder liner surfaces in the engine blow-by and the evaluation of countermeasures Magnus Lukas Lorenz und Prof. Dr. Thomas Koch Application of virtual sensors for stress-related design and ­operation-specific lifetime prognosis Martin Diesch, Dr.-Ing. Thomas Bubolz, Dr.-Ing. Martin Dazer, Dipl.-Ing. Kevin Lucan und Prof. Bernd Bertsche Model-based injector deposit detection Michael Hinrichs, Prof. Dr. Rolf Isermann und Prof. Dr. Peter Pickel Potential and challenges of multiple injection strategies for maritime fuels in large engines Benjamin Stengel, Ibrahim Najar, Fabian Pinkert, Egon Hassel und Bert Buchholz The recuperated split-cycle engine as a sustainable heavy-duty solution. Nick Owen, Prof. Robert Morgan und Prof. Andrew Atkins Potential of low pressure EGR in combination with electric turbocharging for heavy-duty applications Harsh Sankhla, Bartosch Jagodzinski, Sascha Schönfeld, Markus Schönen, Martin Müther und Peter Heuser Modern hybrid propulsion systems for rail and marine applications: environmental and customer benefits through optimized system integration of proven diesel technology with latest electrical innovation. Martin Urban Achieving the proposed EU heavy-duty truck 2030 CO2 legislation Ahmed Meza, Andy Skipton-Carter, Andrew Auld, Nicholas Hasselbach, Önder Bulut, Pascal Revereault und William Missions

Inhaltsverzeichnis

Reducing CO2 emissions in heavy-duty spark ignited engines for electric power using alternative fuels Paul S. Wang, Niko Landin, Michael Bardell, Patrick Seiler, Jas Singh, David Ginter und David T. Montgomery Is liquefied methane the heavy-duty fuel of the future? Max Kofod, Fenna Sleeswijk, Paul Bosma, Maurice van Erp und Bruno Goncalves Tagungsbericht Marc Ziegler

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Autorenverzeichnis

Prof. Andrew Atkins  Ricardo UK Ltd, Shoreham-by-Sea, UK Andrew Auld  Ricardo UK Ltd, Shoreham-by-Sea, UK Robert Bank  FVTR GmbH, Rostock, Deutschland Michael Bardell  Caterpillar Inc, Peoria, USA Prof. Bernd Bertsche  University of Stuttgart, Stuttgart, Deutschland Paul Beutel  Fraunhofer Institute ISE, Freiburg, Deutschland Paul Bosma  Shell Global Solutions Deutschland GmbH, Hamburg, Deutschland Dr.-Ing. Thomas Bubolz  MTU Friedrichshafen GmbH, Friedrichshafen, Deutschland Bert Buchholz  Rostock University, Rostock, Deutschland Heiner Bülte  Deutz AG, Köln, Deutschland Önder Bulut  Ricardo UK Ltd, Shoreham-by-Sea, UK Bernd Danckert  ICCL, Integrated Consulting Company Ltd, Κύπρος, Zypern Dr.-Ing. Martin Dazer  University of Stuttgart, Stuttgart, Deutschland Martin Diesch  University of Stuttgart, Stuttgart, Deutschland Christian Ewert  Fraunhofer Institute ISE, Freiburg, Deutschland Markus Fuchs  MAN Truck & Bus SE, München, Deutschland Conrad Gierow  FVTR GmbH, Rostock, Deutschland

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Autorenverzeichnis

David Ginter  Caterpillar Inc, Peoria, USA Bruno Goncalves  Shell Global Solutions International B.V., The Hague, Niederlande Max Julius Hadrich  Fraunhofer Institute for Solar Energy Systems ISE, Freiburg, Deutschland Egon Hassel  Rostock University, Rostock, Deutschland Nicholas Hasselbach  Ricardo UK Ltd, Shoreham-by-Sea, UK Prof. Christopher Hebling  Fraunhofer Institute for Solar Energy Systems ISE, Freiburg, Deutschland Tobias Herrmann  MAN Truck & Bus SE, München, Deutschland Peter Heuser  FEV Group GmbH, Aachen, Deutschland Michael Hinrichs European Technology Innovation Center, John Deere, Kaiserslautern, Deutschland Dorian Holtz  FVTR GmbH, Rostock, Deutschland Dr.-Ing. Daniel Hrdina  Mahle GmbH, Stuttgart, Deutschland Prof. Dr. Rolf Isermann  TU Darmstadt, Darmstadt, Deutschland Bartosch Jagodzinski  FEV Europe GmbH, Aachen, Deutschland Dirk Kadau  Winterthur Gas & Diesel Ltd, Winterthur, Schweiz Dipl.-Ing. Hakan Kahraman  Mahle GmbH, Stuttgart, Deutschland Anand Kakadiya  Fraunhofer Institute ISE, Freiburg, Deutschland Matthew Keenan  Ricardo UK, Shoreham-by-Sea, UK Christoph Klein  Deutz AG, Köln, Deutschland Prof. Dr. Thomas Koch KIT Karlsruhe Institute of Technology, Karlsruhe, Deutschland Max Kofod  Shell Global Solutions Deutschland GmbH, Hamburg, Deutschland Reinhard Lämmermann  MAN Truck & Bus SE, München, Deutschland Niko Landin  Caterpillar Inc, Peoria, USA

Autorenverzeichnis

XV

Bartek Lemm  Mahle GmbH, Stuttgart, Deutschland Magnus Lukas Lorenz  Daimler AG, Stuttgart, Deutschland Dipl.-Ing. Kevin Lucan  University of Stuttgart, Stuttgart, Deutschland Dipl.-Ing. Marco Maurizi  Mahle GmbH, Stuttgart, Deutschland Ahmed Meza  Ricardo UK Ltd, Shoreham-by-Sea, UK William Missions  Ricardo UK Ltd, Shoreham-by-Sea, UK David T. Montgomery  Caterpillar Inc, Peoria, USA Prof. Robert Morgan  University of Brighton, Brighton, UK Martin Müther  FEV Group GmbH, Aachen, Deutschland Ibrahim Najar  Rostock University, Rostock, Deutschland Jacques Nicole  Ricardo Inc., Santa Clara, USA Christian Opitz  Deutz AG, Köln, Deutschland Mohamed Ouda  Fraunhofer Institute for Solar Energy Systems ISE, Freiburg, Deutschland Nick Owen  Dolphin N2 Limited, Shoreham-by-Sea, UK Katharina Pautsch  Fraunhofer Institute ISE, Freiburg, Deutschland Prof. Dr. Peter Pickel  European Technology Innovation Center, John Deere, Kaiserslautern, Deutschland Fabian Pinkert Forschungszentrum für Verbrennungsmotoren und Thermodynamik Rostock GmbH, Rostock, Deutschland Pascal Revereault  Ricardo UK Ltd, Shoreham-by-Sea, UK Ben Rogers  Ricardo UK, Shoreham-by-Sea, UK Florian Rümmele  Fraunhofer Institute for Solar Energy Systems ISE, Freiburg, Deutschland Harsh Sankhla  FEV Europe GmbH, Aachen, Deutschland Dr. Achim Schaadt Fraunhofer Institute for Solar Energy Systems ISE, Freiburg, Deutschland

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Autorenverzeichnis

Markus Schönen  FEV Europe GmbH, Aachen, Deutschland Sascha Schönfeld  FEV Europe GmbH, Aachen, Deutschland Patrick Seiler  Caterpillar Inc, Peoria, USA Hartmut Sieverding  Deutz AG, Köln, Deutschland Vanessa Simon  MAN Truck & Bus SE, München, Deutschland Jas Singh  Caterpillar Inc, Peoria, USA Andy Skipton-Carter  Ricardo UK Ltd, Shoreham-by-Sea, UK Fenna Sleeswijk Shell Global Solutions Deutschland GmbH, Hamburg, Deutschland Dipl.-Ing. Guilherme Soares de Faria  Mahle GmbH, Stuttgart, Deutschland Flavio Soppelsa  Winterthur Gas & Diesel Ltd, Winterthur, Schweiz Benjamin Stengel  Rostock University, Rostock, Deutschland Dr. Alexander Susdorf  Fraunhofer Institute ISE, Freiburg, Deutschland Robert Szolak Fraunhofer Institute for Solar Energy Systems ISE, Freiburg, Deutschland Martin Urban  MTU Friedrichshafen GmbH, Friedrichshafen, Deutschland Maurice van Erp Shell Global Solutions Deutschland GmbH, Hamburg, Deutschland Paul S. Wang  Caterpillar Inc, Mossville, USA Marc Winterhoff  MAN Truck & Bus SE, München, Deutschland Marc Ziegler Wiesbaden, Deutschland

The role of synthetic fuels in an integrated energy system Dr. Achim Schaadt1, Robert Szolak1, Prof. Christopher Hebling1, Florian Rümmele1, Max Julius Hadrich1, Mohamed Ouda1 and Bernd Danckert2 1

Fraunhofer Institute for Solar Energy Systems ISE, Heidenhofstraße 2, 79110 Freiburg, Germany 2 ICCL, Integrated Consulting Company Ltd., Cyprus

Abstract. The manufactures of internal combustion engines face currently a lot of challenges. Among these are the required drastic emission reductions of carbon dioxide (CO2) as well as harmful substances the most prominent. As a consequence of these harsh market conditions solutions that could cope with this framework are urgently needed. In this context the state of the art of synthetic fuels as an essential part of the energy system is discussed and how they could offer solutions. In addition to the technology developments at Fraunhofer ISE providing innovative solutions in clean mobility such as the CatVap® technology and more efficient synthesis pathways to oxymethylene dimethyl ether (OME) the energy system analysis is one important research topic. One goal of these energy systems simulations is to find the cost optimized pathway of the transformation process. The simulation results are used as a basis for investigations to optimize national / regional energy systems with consideration of all energy carriers and consumption sectors taken into account the CO2 reduction goals. The results clearly show that hydrogen will play a significant role and water electrolysis will be the key technology in order to reduce the transformation costs, to stabilize the grid and to connect the renewable energy power sector with the other sectors (transport, industry, buildings) for defossilization. This green hydrogen can be further converted together with CO2 via Power to X (PtX) concepts into synthetic fuels such as methanol, dimethyl ether (DME) and OME offering both CO2 reductions and reduction of pollutants. PtX will be installed globally in those regions of the world with favourable site conditions for renewable energies (PV, wind). As a consequence this will result in the establishment of a global trading system for renewable energy. Keywords: Synthetics fuels, CO2 reduction, green hydrogen, Power to X

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Introduction

As a result of the growing commitment to renewable energies worldwide, the share of electricity generated from fluctuating sources like solar and wind is steadily increasing. Therefore, new grid-supportive measures are needed to match demand and supply and thus to ensure reliable and efficient operation. © Springer Fachmedien Wiesbaden GmbH, ein Teil von Springer Nature 2020 W. Siebenpfeiffer (Hrsg.), Heavy-Duty-, On- und Off-Highway-Motoren 2019, Proceedings, https://doi.org/10.1007/978-3-658-31371-5_1

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Fig. 1 shows the power generation in Germany each in May 2010 and 2018. The fast development in terms of increased share of photovoltaics (PV) and wind power production is obvious. During this period of time of just eight years the annual share of renewable energy more than doubled from 19.1% to 40.3%. The increase in both PV and wind power production is followed by a decrease of power production based on nuclear and hard coal. Nuclear power plants will be decommissioned completely in Germany in three years.

Fig. 1. Power production mix in Germany in May 2010 and 2018 according to the energy charts (www.energy-charts.de).

The Fraunhofer Institute for Solar Energy Systems ISE with about 1,300 employees is the largest institute for solar research in Europe and has been one of the leading R&Dservice providers in the field of solar energy for the last 30 years. Fraunhofer ISE is working on creating the technological foundations for supplying energy efficiently and on an environmentally sound basis in industrialized, threshold and developing countries. With its research focusing on energy conversion, energy efficiency, energy distribution and energy storage, it contributes to a broad spectrum of renewable energy technologies. In this context Fraunhofer ISE offers also various solutions in mobility such as PV vehicle integration, battery and fuel cell systems, thermal management, power electronics and synthetic fuels.

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Power to X

The energy system analysis is one important research topic at Fraunhofer ISE as this offers the possibility to better understand the interactions of this complex system and to analyze the possible transformation pathways in terms of costs assuming that all CO2 reduction goals are fulfilled. In order to carry out such energy system simulations the cross-sectoral model platform REMod-D has been developed for the German energy system [1]. In this model all relevant energy sources, converters, storage systems and consumption sectors are implemented. The goal was to find the overall minimum of system costs taken into account the security of supply and the CO2 reduction goals.

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The simulations clearly showed that the transition of energy systems is technically and economically feasible. After the energy transition the costs for the operation and maintenance of a renewable energy system will be comparable to our current energy system. Wind and solar power will become dominant. In order to achieve the national and international CO2 reduction targets, these high penetration rates of renewable energy capacities are mandatory and even more. Although Germany achieved a high share of renewable energies in the electricity sector the CO2 reduction goals in the mobility sector are significantly missed (Fig. 2).

Fig. 2. Power-to-X technologies offer the chance to transform renewable energies in other sectors

From the REMod-D simulations it became clear that the coupling of sectors by Power to X technologies is a crucial requisite for a reliable and sustainable energy system in order to transform renewable energy into other sectors. Water electrolysis is the key element to link these different sectors. Chemical energy storage based on electricity conversion via water electrolysis to hydrogen (H2) and oxygen (O2) followed by the conversion of H2 to different energy carriers of interest is essential for long term energy storage, clean mobility solutions and a green chemical and process industry. Furthermore, electrolyzer systems in the distribution grid can be valuable loads for frequency stabilization. In addition to water electrolysis other low carbon intensive hydrogen production methods such as methane pyrolysis might also be needed and applied. Since electrolysis couples directly the electric world with the sectors based on chemical energy carrier it will be the dominating hydrogen production technology in future. In various REMod-D energy system scenarios a ramp up of electrolysis capacity in the range of 50-80 GW was calculated for Germany in the year 2050. Beyond Germany renewable hydrogen production will start in regions with production costs of electricity below 3 cent/kWh and full load hours of more than 4,000 per year [2]. There are many regions in the world such as Australia and Chile which offer these favorable conditions. As a consequence, a global trading system of renewable hydrogen and products, which can be synthesized from hydrogen such as methanol and ammonia, will be established. The techno-economic potential of renewable energies in Germany is currently estimated to be in the range of 1,000 TWh/a. As the primary energy demand is expected to

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be 2,000-2,500 TWh/a in Germany there has to be also in future an import of chemical energy carriers, but these will be more and more based on renewable energy. There is currently much debate on the most appropriate energy carrier to facilitate the hydrogen transport. Liquid organic hydrogen carrier (LOHC), ammonia, liquefied natural gas (LNG), liquefied hydrogen and methanol are currently discussed and investigated at Fraunhofer ISE as a means for hydrogen transport. In the division hydrogen technologies at Fraunhofer ISE, the whole H2 value chain is being considered. The research activities involve the electrochemical H2 production, thermochemical H2 valorization via synthesis, reforming and pyrolysis technologies and the direct utilization of H2 in fuel cells with focus on the mobility sector (Fig. 3). To summarize Fraunhofer ISE optimizes technologies allowing the deployment of the “Matchmaker” H2 in sustainable energy systems.

Fig. 3. Overview of Power to X technologies investigated at Fraunhofer ISE

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Synthetic Fuels Based on Methanol

In terms of fuel synthesis at Fraunhofer ISE the research activities include the development of innovative processes to convert H2 to designer molecules fulfilling the criteria of future fuels and energy carriers defined by [3], namely 1. CO2 “quasi” neutrality 2. sustainability with regard to unlimited availability 3. as low environmental/ecological impact as possible 4. economic efficiency and 5. functionality and best possible integrity with existing technologies. In the department “Thermochemical Processes” (TCP) at Fraunhofer the aim is to increase the energy efficiency of thermochemical processes, to reduce CO2 and exhaust emissions and thus to help the energy transition to success with innovative process engineering solutions. In this context, TCP conducts research on the more sustainable production of different fuels and chemicals from different feedstocks (CO2, syngas,

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exhaust gas from steel industry, etc.). In addition, the techno-economic-ecological aspects of the PtX processes are deeply analyzed. The outcome of the assessment of different fuels showed that methanol and the products that are based on methanol such as dimethyl ether (DME) and oxymethylene dimethyl ether (OME) are very attractive. In this paper the utilization of methanol and OME, which can be produced from methanol, are discussed in terms of the criteria that has been provided before. Methanol (denoted here on as MeOH) is a bulk chemical that is currently produced in large amounts of more than 90 MMT (million metric tons) per year globally. The global demand for MeOH is estimated by market researchers to grow steadily with a compound annual growth rate of 5.5% at least until 2027. MeOH production is predicted to increase from 90 MMT in 2017 to 135 MMT in 2027 [4]. Whilst state-of-theart processes produce MeOH almost exclusively from fossil fuels, MeOH can also be synthesized very selectively via the direct hydrogenation of waste/captured CO2. Possible CO2 sources are bio-based processes (fermentation processes, etc.), steel mills, waste incineration plants, cement and ammonia industry or ultimately from air using direct air capture (DAC), with H2 being provided by renewable energy powered water electrolysis. As these huge amounts of methanol are already available and traded globally the necessary infrastructure is available too. In addition, as it is a liquid, methanol can be handled very easily stored, transported, etc. Furthermore, about 40% of the current methanol production is already used in the fuel market for biodiesel production and the production of methyl tertiary butylether (MTBE), etc. MeOH is an interesting oxygenate fuel for internal combustion engines and works well in blends up to M100 (i.e. 100% MeOH, 0% gasoline). Different fleet tests have been carried out around the world [5]. Some issues such as corrosion on piping or seals have been identified and therefore corrosion additives are required in several standards regulating MeOH blends. Toxicity remains an issue with MeOH in general, causing blindness or death when ingested, but both the toxic and environmental effects of MeOH remain less harmful compared to gasoline [6]. MeOH has recently gained attention as a marine fuel because of the tightening of emissions regulations for ships. Especially inside the International Maritime Organization (IMO) sulphur emission control area (SECA), MeOH can be an interesting option in order to reduce the sulphur and soot emissions. Stena Lines (Sweden) has retrofitted a ferry to run mainly on MeOH [7]. The European Union’s Horizon 2020 FReSMe project utilizes steel mill gases for the production of methanol. The produced MeOH from this project will be used by Stena Lines in the aforementioned retrofitted ferry [8]. In Germany, the project Carbon2Chem®, funded by the Federal Ministry of Education and Research (BMBF), is also evaluating process routes and plant designs in order to convert steel mill gases into chemicals, e.g. MeOH [9] or ammonia. Steel mill gases are rich in CO2, carbon monoxide (CO) and H2 and currently used only for electricity and heat generation in a combined heat and power plant. Carbon Recycling International (CRI) from Iceland has built a plant using renewable energy to produce hydrogen and convert it with geothermal CO2 to methanol. The

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“George Olah plant”, producing around 5,000 MT a-1 is the “world’s largest CO2 Methanol plant” [10]. The CO2 emissions can be reduced by 90% compared to fossil MeOH, which has been certified by ISCC. However, this plant is operated at a constant load, which wouldn’t typically be possible in other renewable scenarios. Dynamic operation of methanol synthesis processes is not a state of the art technology and research is ongoing to find out to which extent the load and other process conditions can be varied. Assuming low cost renewable electricity of around 2.5 ct€ per kWh and a high capacity factor, which are possible in many regions of the world such as Australia, would allow renewable energy methanol production costs of less than 500 € per t MeOH [11]. This is still higher than the current prices for MeOH (ca. 280 €/t) but political pressure and technology development will make the Power to X route more and more attractive. MeOH derivatives like MTBE have been used as octane boost additives for many years and fatty acid methyl ester (FAME, biodiesel) has been blended in with diesel fuel. In addition, the methanol to gasoline route attracts again the interest of the industry as an interesting option and is therefore investigated within the project C3-Mobility funded by the German Federal Ministry of Economic Affairs and Energy (BMWi) [12]. Also the oxygenated fuels based on MeOH i.e. DME and OME are gaining increasing interest as diesel fuel alternatives. OMEn with the constitutional formula CH3–O(– CH2–O)n–CH3 are non-hazardous to human health or the environment, have a high cetane number and can be used in existing infrastructure. The chemical-kinetic analysis showed that the desirable suppression of acetylene formation is attributed to the missing C–C bonds of OME. Since OME show this ideal fuel properties Fraunhofer ISE develops and evaluates new and more efficient OME synthesis routes based on MeOH with the potential to reduce the production costs significantly. The analysis of Fraunhofer ISE showed that at a production capacity of 1 Mt per annum OME3-5 production costs of 571 US$ per t can be achieved [13]. In tests applying a heavy-duty single cylinder engine at the Technical University (TU) of Munich, OME1 as well as OME3-6 showed a strong reduction of particle emissions compared to regular diesel [14]. In even more recent tests the TU Munich confirmed this low soot formation in OME combustion, which is maintained even at high exhaust gas recirculation rates, enabling ultra-low NOx emission operation. As further work Pélerin identified sealing materials that are robust against OME and suggested designs for OME injection systems that are less complex in terms of fewer requirements regarding maximum injection pressure and multiple injection capability [15]. Currently, the influence of OME on the exhaust gas aftertreatment system is analyzed in detail at Fraunhofer ISE especially focusing cold and dynamic engine operation conditions. OME diesel blends have also already successfully been tested in an Euro 6d car with shares of 7% and 15% OME in correspondence to the Real Driving Emissions (RDE) procedure of the European Union. Though even with blending OME particle emissions can be reduced significantly, a diesel particulate filter will be indispensable in the future because of the particle number emission limit. Additional research concerning tailored OME engines, their aftertreatment systems and potentials is carried out within the BMBF project NAMOSYN [16]. In a comparative well-to-wheel life cycle assessment of OME3-5 synfuel production via the power-

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to-liquid pathway has been found out that for the best case the greenhouse gas emissions can be reduced by up to 86% [17].

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Conclusion

Fraunhofer ISE in Freiburg has gained CO2 neutral with its experience in green hydrogen technologies and the fuels of the future resulting from them. It has the best prerequisites to develop the internal combustion engine into a technology of the future. The bulk chemical methanol serves as a fuel, fuel additive and energy carrier for the production of tailor-made fuels with ideal suitability for modern combustion engines. Overall methanol and its derivatives show very interesting properties and a high potential to play a significant role within the future market of green fuels made from renewable energies. This is also true with the new CatVap® technology, a new fuel processing technology that is able to deliver the required thermal energy highly dynamically, for a highly efficient mode of operation of a modern exhaust aftertreatment system that will also be needed for synthetic fuels (see the lecture CatVap® – New Heating Measure for Exhaust Aftertreatment Systems). This highly efficient heating and exhaust system temperature control system allows to develop very low-consumption internal combustion engines of the most varied combustion processes to operate with almost all available fuels, especially future fuels (in particular obtained by Power to X).

References 1. Palzer, A.; Henning, H.-M.: A comprehensive model for the German electricity and heat sector in a future energy system with a dominant contribution from renewable energy technologies – Part I: Methodology. Renewable and Sustainable Energy Reviews 30, 1003-1018 (2014). 2. Hebling, C.; Ragwitz, M.; Fleiter, T.; Groos, U.; Härle, D.; Held, A.; Jahn, M.; Müller, N.; Pfeifer, T.; Plötz, P.; Ranzmeyer, O.; Schaadt, A.; Sensfuß, F.; Smolinka, T.; Wietschel, M. (2019) Eine Wasserstoff-Roadmap für Deutschland. Fraunhofer-Institut für Systemund Innovationsforschung ISI, Karlsruhe; Fraunhofer-Institut für Solare Energiesysteme ISE, Freiburg unter Beteiligung von Fraunhofer-Institut für Mikrostruktur von Werkstoffen und Systemen IMWS, Halle; Fraunhofer-Institut für Keramische Technologien und Systeme IKTS, Dresden. 3. Gaukel, K.; Pélerin, D.; Härtl, M.; Wachtmeister, G.; Burger, J.; Maus, W.; Jacob, E. (2016) The Fuel OME2: An Example to Pave the Way to Emission-Neutral Vehicles with Internal Combustion Engine, 37. Internationales Wiener Motorensymposium, Conference Proceedings, 193-223. 4. Berggren, M. (2017) Methanol to Energy – Challenges and Opportunities, 4th MMSA Methanol Technology and Policy Commercial Congress, 5-7 December 2017, Frankfurt. 5. European Automobile Manufacturers Association (ACEA), Methanol as a Gasoline Blending Component, ACEA Position Paper, 2015.

8 6. Methanol Institute, Methanol Use In Gasoline: Blending, Storage and Handling of Gasoline Containing Methanol, Methanol Blending Technical Product Bulletin, Singapore, Washington. 7. International Maritime Organization (IMO) and DNV GL, Methanol as marine fuel: Environmental benefits, technology readiness, and economic feasibility: Use of methanol as fuel, 2016. 8. http://www.fresme.eu/index.php#NEWS, last accessed 31 October 2019. 9. Nestler, F.; Krüger, M.; Full, J.; Hadrich, M.J.; White, R.W.; Schaadt, A. (2018) Methanol Synthesis – Industrial Challenges within a Changing Raw Material Landscape, Chem. Ing. Tech., 90, 10, 1-11. 10. https://www.carbonrecycling.is/mefco2-project, last accessed 31 October 2019. 11. Hank, C.; Gelpke, S.; Schnabl, A.; White, R.J.; Full, J.; Wiebe, N.; Smolinka, T.; Schaadt, A.; Henning, H.-M.; Hebling, C. (2018) Economics & carbon dioxide avoidance cost of methanol production based on renewable hydrogen and recycled carbon dioxide – powerto-methanol, Sustainable Energy Fuels, 2, 1244. 12. http://www.c3-mobility.de/, last accessed 31 October 2019. 13. Ouda, M.; Mantei, F.K.; Elmehlawy, M.; White, R.J.; Klein, H.; Fateen, S.-E. K (2018) Describing oxymethylene ether synthesis based on the application of non-stoichiomsetric Gibbs minimization, React. Chem. Eng., 3, 277. 14. Härtl, M.; Gaukel, K.; Pélerin, D.; Wachtmeister, G. (2017) Oxymethylene Ether as Potentially CO2-neutral Fuel for Clean Diesel Engines Part 1: Engine Testing, MTZ worldwide, 2, 52-58. 15. Pélerin, D.; Gaukel, K.; Härtl, M.; Jacob, J.; Wachtmeister, G. (2020) Potentials to simplify the engine system using the alternative diesel fuels oxymethylene ether OME1 and OME3−6 on a heavy-duty engine, Fuel 259, 116231. 16. Rösel G.; Avolio, G.; Grimm, J.; Maiwald, O.; Kastner, O.; Brück, R. (2018) Diesel – efuel blends for simultaneous reduction of real driving NOx and CO2 emissions, International Engine Congress. 17. Hank, C.; Lazar, L; Mantei, F.K.; Ouda, M.; White, R.J.; Smolinka, T.; Schaadt, A.; Hebling, C.; Henning, H.-M. (2019) Comparative well-to-wheel life cycle assessment of OME3-5 synfuel production via the power-to-liquid pathway, Sustainable Energy Fuels, 3, 3219.

DEUTZ engine portfolio below 56 kW to meet EU Stage V NRMM emission standard Christian Opitz, Christoph Klein, Hartmut Sieverding, Heiner Bülte DEUTZ AG, 51149 Cologne, Germany Abstract. With the Stage V Emission regulation of the European Union (EU) a new limit for the particulate emission including particulate number is introduced for diesel engines from 19 to 56 kW. Thereby new control strategies regarding particulates are required in this engine category. This article describes the DEUTZ engine platform concept for the engine category from 19 to 56 kW with a diesel particulate filter (DPF) as a standard technology to fulfill the EU Stage V emission regulation. The technical requirements of engines for non-road mobile machinery are described as well as the derived hardware concept of the engine family including the exhaust aftertreatment system. The introduction of a new ECU platform is used by DEUTZ for a new software concept with an airpath model based on a combination of physical engine models and data based numerical models to ensure the best performance and robustness under all ambient conditions. Furthermore, DEUTZ developed a new engine emission model including a DoE tool to reduce the calibration effort and improve the DPF model quality at the same time. The article also gives an insight on development methodology with an outlook on the virtual calibration using the DEUTZ offline simulation tool “xQtec”. Keywords: Stage V, Diesel, Offline Calibration

1

Introduction and project specification

The European emission legislation Stage V for diesel engine-powered non-road mobile machinery in the power output range of 19 to 56 kW is in force for the European market since 2019. The essential changes compared to the previous EU Stage IV legislation are the introduction of a particulate number limit of 1x1012 as well as a further reduction of the maximum permitted particulate mass to 15 mg/kWh. In addition, the requirements for the diagnosis concept has increased and an in-service monitoring system has been implemented. The aim is that adherence to the emissions directive shall be monitored under real-world operating conditions over the engine’s entire service life. Initially, monitoring will be limited to reporting; thereafter mandatory conformity factors are to be introduced based on the database. The DEUTZ diesel engine portfolio for the power output range of 19 to 56 kW includes the engine series TD/TCD 2.2, D/TD/TCD 2.9 and TD/TCD 3.6. The TD/TCD 2.2 engine series is a three-cylinder derivative of the successful 2.9. These engine series with their output parameters maximum torque and power are shown in © Springer Fachmedien Wiesbaden GmbH, ein Teil von Springer Nature 2020 W. Siebenpfeiffer (Hrsg.), Heavy-Duty-, On- und Off-Highway-Motoren 2019, Proceedings, https://doi.org/10.1007/978-3-658-31371-5_2

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Fig. 1. Each engine series is available as a supercharged variant without a charge air cooler (TD) and as a supercharged variant with charge air cooler as well as, in some cases also as a naturally aspirated engine. The variants without charge air cooler reduce the required installation space and the application and integration effort of the equipment manufacturer.

Fig. 1. DEUTZ diesel engine portfolio below 56 kW

These engine series are installed for example in agricultural, construction and material handling machinery. There is, consequently, a great variety of applications, power output derivatives and different load profiles, involving both high and low capacity utilization. Fig. 2 shows exemplarily a selection of different load profiles. The required particulate mass limitation applicable to the Stage IV emissions standard could be met mostly through engine design measures combined with an oxidation catalytic converter. A soot particulate filter was optionally available for DEUTZ engines smaller than 56kW when used in operating environments with enhanced emissions requirements. Due to the introduction of a limit on the number of particulates, a particulate filter is mandatory for all Stage V engines between 19 and 56 kW as standard technology. Development objectives were to keep the installation space for the engine, the auxiliaries and the exhaust-gas aftertreatment system for the different applications as small as possible and enabling the Stage V engines to be used in the existing machinery concepts. The article first describes the changes made to the engine concept and to the exhaustgas aftertreatment system. The new software concepts for the air path control and the models for the exhaust aftertreatment are also presented. Due to the challenge to bring the great variety of applications to the market within a short development period,

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DEUTZ developed a modular approach for dataset calibration and used virtual calibration for some calibrations tasks to simultaneously shorten the development time and increasing robustness.

Fig. 2. TCD2.9 load profiles for different applications and load utilization rates

2

Hardware design of the base engine and exhaust aftertreatment

The base engine concept was adopted from the respective engines of the emission standard EU Stage IV [1], [2]. All engine series are equipped with a common rail injection system, exhaust gas recirculation and a wastegate turbocharger. A cooled exhaust gas recirculation system is used to reduce nitrogen oxide emissions within the engine. In comparison to the previous generations, the latest generation of Bosch MDG1 control units is employed; which includes the AMU hardware acceleration module, thereby enabling data-based models to be calculated on a separate micro-controller (see the sections on software). In order to ensure passive regeneration (CRT effect) of the particulate filter, a throttle valve is used for thermal management in the fresh air in front of the intake manifold. An EGR sensor is used to calculate the exhaust gas recirculation mass flow. In order to satisfy the high requirements for engine charge model accuracy, additionally to a pressure and temperature sensor in the intake manifold, a pressure sensor is employed in the exhaust manifold downstream of the engine. Fig. 3 provides an overview of the system together with the relevant new features.

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Fig. 3. System overview for EU Stage V concept with the TCD2.9 as an example

By a range of installation variants for the different applications DEUTZ provides customers with a great degree of freedom regarding the positioning of the air inlet into the engine (flywheel side or belt side) and of the turbocharger. This allows the customer to make the best use of the installation space. To reduce the calibration effort of the dataset for the different variants, the EGR sections and mixers of the variants were designed with coupled 1D/3D CFD calculations to be as thermodynamically neutral as possible beforehand. Thereby besides the EGR-distribution to the different cylinders, the focus was also on temperature layering in the intake manifold in order to avoid impacting the charge air calculation. Fig. 4 shows the simulation results for the CO2 mass fraction in the intake manifold for two different EGR induction variants in the horizontal and vertical layer.

Fig. 4. Comparison of EGR induction for two inlet variants

The challenge when designing the exhaust-gas aftertreatment system was to show how the additional functional component (particulate filter) could be integrated with the least effect on the available installation space. Based on former excellent experience of passively regenerated DPF systems, DEUTZ will continue to rely on this design for Stage V. Since these engines are also used in customer applications with low-load

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working cycles, e.g. in material handling applications, DEUTZ engines fitted with DPF are equipped with a graduated thermal management strategy which supports soot burnoff. Since such measures cannot be achieved with a completely neutral effect on fuel consumption and equipment availability, a further important design objective was minimizing the regeneration intervals required. The requirement for a small installation space on the one hand and the highest possible soot load limit and ash capacity on the other can only be resolved by using a more advanced DPF technology. DEUTZ has therefore decided to use an asymmetric cell geometry with a cellularity of 300 cpsi and a 9 mil wall thickness. Compared with technologies which were used for Stage IV, this offers a 30 % greater ash storage capacity for the same installation space. This is achieved by increasing the volume of the inlet channels compared to the outlet channels. Conversely, if the accumulated ash time requirement remains the same as before, this gain can be directly used to reduce the installation space needed.

Fig. 5. Depiction of an asymmetric cell geometry

At the same time, by using aluminum titanate which is thermally more robust than cordierite, it was possible to increase the permissible soot load limits by up to 200 %. This enables an installation space reduction even when soot raw emissions remain the same. Moreover, the regeneration intervals between two active regeneration periods can be extended for extremely low-load applications. A direct comparison of the installation space between Stage IV and Stage V engines shows therefore a slight increase of approx. 20 % compared to the Stage IV system without DPF. Compared with the DPF variants available in Stage IV, it was possible to reduce the installation space by 20 – 40 % depending on the engine series. Fig. 6 shows the qualitative size comparison between the three exhaust-gas aftertreatment systems.

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Fig. 6. Size comparison of Stage IV and Stage V exhaust-gas aftertreatment systems

3

Software design for air path and exhaust gas aftertreatment

For an optimal adaption of the DEUTZ engine concepts to customer requirements, many of the software functions are developed inhouse by DEUTZ. In the following the new concept for controlling the air path, the emissions model and the exhaust-gas aftertreatment component models are described. A highly-dynamic control of the air path is essential for an optimal engine response, a high performance, low emissions and low fuel consumption. At the same time, regeneration strategies ensure that, even with unfavorable load profiles, the necessary temperature window for the CRT effect can be achieved. With the introduction of Stage V engines and making use of the opportunities offered by the Bosch MDG1 control unit, DEUTZ, in conjunction with Robert Bosch GmbH, introduced a new model-based air path control concept. The actual fresh air mass flow and EGR mass flow values are calculated by physical models plus two numerical four-dimensional models. One model calculates the volumetric efficiency of the cylinder head. The cylinder head mass flow is calculated using the physical relationship between engine speed and density in the inlet manifold. The second model is a correction model for the EGR mass flow sensor. Both models are implemented synchronously with the power stroke on the control unit’s separate ‘AMU’ processor [3]. To improve robustness when affected by environmental conditions such as altitude, heat and cold and so as to reduce the effort and time involved in calibrating, a hybrid numerical model taking of measurement data from the engine test bench and engine model data (GT-Power) into account is used in the volumetric efficiency model. This achieves a high degree of accuracy when modeling the fresh air mass even under extreme conditions such as operating at high altitudes. Fig. 7 shows the quality of the hybrid model compared with a map-based 2D model and a numerical 4D model using engine test bench data. Besides altitude (ambient pressure), the quality of the model is affected by numerous parameters such as exhaust gas backpressure (soot and ash load of the particulate filter) or the ambient temperature. When

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using numerical models there is a risk of entering the extrapolation zone where the model quality is severely reduced. By applying the hybrid approach, DEUTZ is able to represent most of the operating range with a high degree of model accuracy already at the start of a project. Fig. 7 shows the accuracy of three model approaches under changed boundary conditions. Due to the model-based control strategy and the improved calculation of actual air path values emissions could be reduced compared to predecessor engine series. In particular, in highly dynamic cycles more rapid control response by the new software has reduced nitrogen oxide emissions by approx. 10 %, while also reducing soot emissions. Fig. 8 illustrates soot and nitrogen oxide emissions in a representative high-dynamic customer cycle.

Fig. 7. Model quality of the air path under standard and high-altitude operating conditions

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Fig. 8. Dynamic soot and NOX reduction by the model-based air path during a customer cycle

Knowing the current extent of the soot load in the diesel particulate filter as precisely as possible is of great interest in two respects. On the one hand, the exhaust gas backpressure must be limited and the system and its components need to be protected against thermal damage by preventing an excessive build-up of soot. Otherwise, the exothermic reaction when burning off a critical mass of soot could result in the destruction of the filter material or other components. In order to avoid these critical load conditions, at certain load thresholds the NO2-based soot burn-off is specifically supported by engine control strategies in the medium temperature range. If necessary, the soot is oxidized at high temperatures during an active regeneration period. This escalation-based regeneration strategy functions all the more efficiently regarding fuel consumption and equipment availability the more the engine management system knows in real-time about soot build-up and burn-off behavior. Generally, this is achieved by continuously balancing the soot input taken from the engine emissions model and the current rate of soot burn-off. An integrator thus supplies live information about the soot load state which can, in turn, be compared against the thresholds for triggering the measures outlined above. It is therefore necessary to model the engine’s raw emissions as precisely as possible. To achieve this, DEUTZ has developed its own tool chain for the Stage V engine gen-

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eration which contains all aspects from test plan creation, model generation, implementation in the control unit and validation (see the section on methodology). It consists of six-dimensional, data-based, individual models, which are optimized for the particular engine mode and the emissions being modeled. Compared with Stage IV, this has produced a significant improvement in the accuracy of the soot model. Fig. 10 shows the achieved model accuracy exemplarily for the certification cycle. The model error for nitrogen oxide lies within the range of 2 %; for soot emissions the error is less than 10 %. The soot burn-off model is made up of various sub-models, all of which take a physical modeling approach. Temperature models are integrated for both the oxidizing catalytic converter and the DPF, which resolve heat transfer and storage locally. The oxidation reaction rates for NO to NO2 in the DOC and for soot with NO2 and oxygen in the DPF are formulated as Arrhenius equations. This approach, which largely abandons the use of mapping models, provides a direct interpretability, on the one hand, and scalability of the models both between different conditions and to other hardware variants on the other. Fig. 9 schematically illustrates the structure of the exhaust-gas aftertreatment system models. This physical forecast model is assisted by regular comparison with and correction against a sensor-based backpressure model of the DPF. This prevents any load model drift caused by the summation of slight model errors as soon as the sensor value enters ranges with a high degree of reliability. Fig. 10 also illustrates the accuracy of the overall model using the certification cycle as an example.

Fig. 9. Schematic overview of exhaust gas aftertreatment models

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Fig. 10. Accuracy of the engine emission and burn-off models

4

Calibration and validation methodology

To improve efficiency while developing datasets and at the same time improving the robustness of those datasets, DEUTZ employs two approaches. On the one hand a modular approach is used for dataset development and on the other hand DEUTZ has developed its own platform tool for offline simulation. Introducing the new generation of control units and improvements to software functions has added more complexity while at the same time reducing development times. To improve the transferability of existing sub function datasets to other engine series and variants, DEUTZ employs a modular approach. The parameters are divided up into different stages. Within any stage, the parameters for the particular engine concept are identical. For example, the Stage 2 parameters are identical for all engines with the DOC/DPF plus exhaust gas recirculation aftertreatment variant. These parameters may for example be air path control or injection configuration parameters and include approx. 30 % of the combustion-relevant parameters. This reduces the effort for the dataset calibration for each engine variant and therefore shortens the development time. At the same time, the validation depth for the relevant calibration is increased, which improves both the quality and the robustness of the datasets. Fig. 11 shows the modular approach and the distribution of the parameters to different stages as well as their distribution by percentage.

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Fig. 11. DEUTZ modular approach for dataset development

The DEUTZ offline simulation program ‘xQtec’ is used throughout the entire development process [4]. During the Stage V project, the tool was used for function development, for calibrating and for validating datasets. As a platform the tool enables different models and control unit software to be linked and simulation tasks to be carried out. The currently supported modules include e.g. GT-Power engine models, control unit software as C code (Simulink + Target-Link, etc.), Bosch control unit software and models (e.g. ASCMO, MATLAB/ Simulink). It is possible for just those functions which are required to be simulated. This allows the fastest possible simulation speed to be achieved, for diagnostic functions this can for example be up to 2,000 times faster than in real time. All current measurement file formats are supported (e.g. csv, mdf, etc.). Fig. 12 illustrates schematically the options offered by the tool. With this tool it is for example possible to calibrate physical models of the engine control unit very efficient or to carry out robustness analysis for critical operating ranges and cycles.

Fig. 12. DEUTZ simulation platform ‘xQtec’ in the development process

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For example, during calibrating and validating the emissions and soot burn-off model, application-specific customer cycles are first measured directly in the target application without any additional instrumentation work being necessary. Important target parameters, especially those for emissions and also parameters needed for engine control unit calculations, can then be determined by reproducing the customer cycles on an engine test bench equipped with additional measuring technology. The offline simulation offers the option of simulating the dynamically measured process output parameters via the process input parameters. In this context, xQtec provides the software-in-the-loop tool which can be supplemented by parameter optimizing algorithms and hardware-in-theloop components. The design of the soot loading model already offers excellent calibratability and scalability due to its physical approach. Because of the ability to simulate and optimize all the relevant cycles offline, the result is a set of parameters which is suitable for a range of applications and which can be transferred to the engine control unit. The extent of the validation work required on the engine test bench and also later on the customer application can be significantly reduced by offline optimization. This is particularly beneficial in the market environment for industrial engines where extensive testing is often uneconomical due to small production numbers and great model variance. The development methodology based on the V process (see Fig. 13) has resulted in greatly reduced development times and has improved accuracy and robustness. Furthermore, as a consequence, DEUTZ is in a position to rapidly forecast the expected regeneration intervals also for as yet untested applications operated by new customers.

Fig. 13. V-process for calibrating and validating exhaust gas aftertreatment models

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Summary

This article describes the DEUTZ portfolio of 19 to 56 kW power output range engines which meet the Stage V emissions legislation. Compared with Stage IV, a particulate filter has been introduced as standard technology for the 2.2, 2.9 and 3.6 engine series. In order to be able to integrate this additional component without adverse implications for installation space and equipment availability, DEUTZ has introduced an asymmetrical cell geometry and thermally more robust aluminum titanate substrates. This has reduced the installation space by 40 % compared with the optional DPF variants available for Stage IV. Compared with the Stage IV systems without a DPF, the installation space requirement increased by only 20 %, making it possible to install the Stage V engines in existing machinery designs. In addition to this, DEUTZ has introduced model-based software functions for air path control and exhaust gas aftertreatment. This combination of physical and numerical models has produced great accuracy even under different circumstances such as high-altitude operating conditions. To shorten development times and at the same time improve robustness, DEUTZ now established its proprietary simulation platform xQtec. xQtec enables high-speed modular simulation of application-specific cycles based on engine models, control unit functions, numerical models and measurement data. 60 % of the work involved in calibrating and validating exhaust gas aftertreatment models can now be carried out offline. Consequently, a very significant validation depth can be realized despite the huge range of different applications.

References 1. Köhne, M.: New Off-Road Engine Generation for Tier 4 from DEUTZ, in: ATZ Offhighway Special Edition (2010), pp. 8-17 2. Pister, M., Eisenhauer, F.; Martini, T., Ortjohann, T.: DEUTZ TCD 3.6 Four-Cylinder InLine Engine for Agricultural Machinery Applications, in: ATZ Heavy Duty Edition 3/2017 3. Nork, B., Diener, R., AMU-Based Functions on Engine ECUs, in: International Conference on Calibration Methods and Automotive Data Analytics 2019. 4. Broll, P., Weyers, J., Qriqra, A.: Off-line Simulation of the Means of Controlling Diesel Engines with Exhaust Gas Aftertreatment, in: 19th MTZ Simulation and Test Conference, 2017

Compact and powerful: the new 9-liter diesel engine from MAN for off-highway applications Tobias Herrmann, Vanessa Simon, Markus Fuchs, Marc Winterhoff, Reinhard Lämmermann MAN Truck & Bus SE – External Engines Abstract. In 2019, MAN launched a newly developed engine with a displacement of 9 liters. In addition to traditional applications for trucks and coaches, the industrial version has been presented at the same time. For its first application in the off-highway segment, the engine is used in agricultural tractors, where SCR-only combustion and exhaust gas aftertreatment have been completely revised and adapted. The low-speed concept and excellent engine dynamics were of particular importance. A large number of field cycles were analyzed during testing and the test program was adapted to the tough conditions in the tractor. Moreover, this report presents activities and results for recording the tractor’s Real Driving Emission in a wide range of applications. Keywords: Validation, performance, & emissions, in-service monitoring (ISM).

1

Introduction

A few years ago, MAN decided to start developing the new D15 engine platform, bridging the gap between the 7-liter and 12-liter displacement class. During the development process, the main objectives included a low power-to-weight ratio and compact dimensions. The straight-six power unit delivers between 205 and 324 kW (279 and 440 HP) and its highest power variant achieves a maximum torque of 1970 Nm between 1150 and 1300 rpm (see Fig. 1, on the left). This makes the MAN D1556 the lightest off-road engine in its displacement class, with a dry weight of just 860 kg. Even at low speeds, it delivers an impressively high torque. In its first application as an industrial engine, the new D1556 engine is used in tractors where, due to space restrictions and high utilization, it has a maximum output of 305 kW. In order to serve the global market, the engine has been certified for emission stages EU Stage V, EPA Tier 4 Final, and EU Stage IIIa for Low Regulated Countries. The following figure shows the current state of the power curve in tractors. A competitive comparison (see Fig. 1, on the right) clearly shows that the D15 is in a leading position – particularly at maximum torque.

© Springer Fachmedien Wiesbaden GmbH, ein Teil von Springer Nature 2020 W. Siebenpfeiffer (Hrsg.), Heavy-Duty-, On- und Off-Highway-Motoren 2019, Proceedings, https://doi.org/10.1007/978-3-658-31371-5_3

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Fig. 1. Power curve, as well as specific output/effective mean pressure

Furthermore, great emphasis was placed on the modularity of the 9-liter engine system. The following figure shows the different variants in trucks, coaches, natural gas and tractor engines.

Fig. 2. CAD representation of the D15 as truck, coach, natural gas and tractor variant

The variants differ, for example, in fan attachments and belt drives, cylinder head covers, and oil sumps made of plastic and aluminum, as well as turbo chargers with wastegate or Variable Turbine Geometry (VTG). In contrast to its diesel counterpart, the E1856 natural gas variant features cooled exhaust gas recirculation (EGR). The coach variant can be supplied with a crankshaft starter alternator as an option. The following section describes in detail the development (see the following table for technical data) and integration of the D15 into tractors.

3 Table 1. Technical data – tractor engine

Engine type Exhaust gas status

D1556LE5xx EU Stage V, EPA Tier 4 Final, Downgrade UN-ECE R96, H (complies with EU Stage IIIA)

Number of cylinders/arrangement/valve s per cylinder Bore/stroke [mm] Displacement [cm³] Compression [-] Injection/ignition pressure [bar] Weight (dry) [kg] Rated power [kW] at speed [rpm] Max. torque [Nm] at speed [rpm]

6/inline/4 115/145 9.0 20:1 2,500/230 Approx. 860 (without exhaust gas aftertreatment), approx. 960 (with exhaust gas aftertreatment) 217 239 261 283 305 1,700 1,550 1,650 1,750 1,850 1,970 1,100– 1,100– 1,100– 1,150– 1,200– 1,300 1,350 1,400 1,400 1,350

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Combustion development

2.1

SCR-only concept

The emission requirements of EU Stage V and EPA-T4F are fulfilled with SCR-only technology – in other words, without the use of external exhaust gas recirculation (EGR). In addition to the mechanical advantages such as reducing the weight and complexity of the system and relief of the cooling system, there are also considerable advantages for the combustion design. The SCR-only concept makes it possible to optimize the combustion efficiency and thus adjust it to the best possible fuel consumption. Due to the optimum combustion process, particle emissions are at an extremely low level, which in turn has a positive effect on the service life of the particulate filters and minimizes the ingress of soot into the oil. In addition to the advantages of high consumption efficiency and low particulate emissions, the physical properties of a highly efficient combustion process also result in high engine-out NOX emissions. This problem was countered by an adapted combustion chamber geometry that is specifically designed for low nitrogen oxide emissions (see Section 2.2). In conjunction with an optimized AdBlue fluid dosing strategy and a high-performance SCR system, MAN has succeeded in combining optimum fuel consumption with low AdBlue fluid consumption.

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2.2

Combustion chamber

MAN conducted numerous studies on combustion chamber designs. A large number of different piston recess geometries combined with different injection nozzle variants were studied both in simulations and on combustion test benches, resulting in an entirely redesigned piston recess that is ideal for the SCR-only concept and enables both low fuel consumption and low engine-out emissions. The piston recess has been designed so that the flame is directed towards the beam splitter on the edge of the piston, splitting the flame into two parts. The lower part of the flame is deflected in the piston recess and runs along the piston towards the piston center. The upper part of the flame is first directed towards the head of the piston via an incline in order for it to spread radially in a further step. This delays the combustion at the incline which means that the peak temperatures are lowered and NOX emissions are reduced (see Fig. 3).

Fig. 3. Flame front, combustion chamber temperatures, and NOx emissions in the combustion chamber at 5%, 50%, and 90% combustion progress

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2.3

Injection system

The engine is equipped with state-of-the-art common rail injection technology. The injectors feature leakage-free control valves which have a greater hydraulic efficiency than conventional injectors with gap leakage. This reduces the fuel consumption by approx. 1%. Significantly improved hydraulics also increase the accuracy of the injection quantity. Even the smallest pilot quantities can be displayed steadily over the entire service life, which has a positive effect on emissions and combustion noise. A 12-hole injection nozzle is used together with the redesigned piston recess geometry. The maximum system pressure is 2,200 bar. 2.4

Turbocharging and engine dynamics

Off-road applications place high demands on torque characteristics and engine dynamics. The deployment profiles range from light, semi-static operation on harvesters, for example, to heavy alternating operation on wheel loaders. Torque and dynamics should accordingly be available over a wide band – even under different climatic conditions and at high altitudes. With the turbo charger with variable turbine geometry (VTG) used, high air masses are provided over the entire engine characteristics range. Even at low engine speeds and small injection quantities, sufficient charging pressure is provided to ensure a fast, transient load response. Load jumps at different engine speeds were run on a highly dynamic engine test bench and the relevant combustion parameters were calibrated in several optimization loops, resulting in dynamic power development over the entire speed range.

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Exhaust gas aftertreatment

In the case of tractors, the available installation space of the previous machine model had to be maintained. The previous model’s engine featured exhaust gas recirculation and had a lower max. torque with a lower displacement. The previous model’s engine featured exhaust gas recirculation, low torque, and displacement. The development team was therefore faced with the challenge of improving the uniform flow distribution via the SCR at the one hand side and to optimize the AdBlue fluid preparation with significantly higher exhaust-gas mass flows at higher NOX conversion rates on the other side. At the same time the maximum exhaust-gas back pressure had to been taken into account. 3.1

Particulate filters

The diesel oxidation catalyst (DOC) and diesel particulate filter (DPF) geometries were set by the size of the engine compartment lid situated above them. A substrate design with asymmetric channel geometry was implemented to maximize the ash cleaning intervals of the DPF. The field and endurance tests resulted in an ash cleaning interval of at least 8,000 operating hours. The soot loads in the DPF proved uncritical during

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the entire test. Due to the high engine-out NOX emission level and the correspondingly high NOX/particulate ratio, it is usually not necessary to regenerate the DPF over and above a fixed regeneration interval of 1000 hours, even in the case of low-load applications. However, in order to validate the quality of the DPF loading model, a vehicle was tested in an extremely low-load application on the MAN plant premises (see Section 4.1). This unilateral and extremely low-load machine application produced DPF regeneration intervals of approx. 250 hours. 3.2

SCR catalytic converter

The previous model’s dimensions were also adopted for the exhaust piping and AdBlue line. In the previous vehicle version, the SCR was already dimensioned to 13 inches – the largest standard diameter available on the market for cordierite substrates. It was only by extending the SCR substrates that the higher requirements for exhaust-gas mass flow and NOX conversion were taken into account. The uniform flow distribution at the SCR inlet had to be increased to a uniformity index of at least 0.98. In order to fulfill the requirements regarding uniform distribution and exhaust-gas back pressure, various optimization loops were calculated during a simulation test at different operating points. During validation, the target variant was checked for uniform flow and NH3 distribution on a hot-gas test bench to confirm the simulation results.

Fig. 4. Variants of flow optimization

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4

Testing

4.1

Functional test

The functional tests are limited to the parts for use in tractors which differ from the parts for use in on-highway engines. Amongst others, the following parts were subjected to extensive tests: the load on the viscous damper, the alternator, and the air compressor in the heat chamber. To determine their suitability, the components are tested under real installation conditions and high ambient temperatures. Cold-start applications and the legal requirements for the thawing times of the AdBlue system were tested and optimized on two tractors at extremely cold temperatures on the climatic roller test bench at the MAN plant in Munich.

Fig. 5. Testing the boundary conditions on the viscous damper, cold start, belt drive, etc.

The exhaust gas aftertreatment was subjected to targeted low-load testing at the engine plant in Nuremberg. For this purpose, a tractor was operated by the internal Logistics department and equipped with a data logger. At regular intervals, the depositing tendency of the AdBlue® mixer was tested using an endoscope (see Fig. 6) and the load on the DPF was determined by weighing.

Fig. 6. Optimizing the deposits in the AdBlue preparation (mixer)

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Fig. 7 illustrates the weak utilization of commuter traffic. The application was adapted according to the findings. The fact that the test vehicle is located close to the Development department means that it is much easier to implement the knowledge gained. 250 0.0

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4.2

Endurance testing

Endurance tests can also be used – and were used – to test the engine in trucks and coaches. In total, the new D15 engine ran for over 200,000 hours on a wide range of test benches. About 10% of this time was spent on off-road applications (see Fig. 8).

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Fig. 8. Test bench – scope of testing

Another important validation cornerstone is achieved by testing the engine in the vehicle. This means that the engine is running for a comparable number of test hours in the tractor. During the development phase, more than 20 tractors were validated under the toughest conditions (see Fig. 9).

Fig. 9. Vehicle – scope of testing

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In order to select the correct endurance program, the actual measured field cycles were compared in advance with the internal endurance cycles. While creating the test program, real application cases were specifically examined and evaluated with regard to different criteria. The following criteria, among others, were used for evaluating the design: ─ ─ ─ ─

Work over the runtime (average engine load) Average speed (number of load cycles) Thermo-mechanical damage using exhaust manifold as an example Switching cycles of mechatronic components

Fig. 10 compares the measured field cycles with the test cycles. Taking into the account the number of load cycles and the engine load over the runtime (e.g., 8,000 hours), it becomes clear that it is almost impossible to test these factors during an endurance run. In comparison, it is possible to adequately map the thermo-mechanical damage by using generic cycles. The endurance cycles, however, were not optimized with regard to the dynamic load of the actuators, as these were tested during the component tests.

Fig. 10. Comparison of real and test cycles

5

In-service monitoring (ISM)

5.1

Legal provisions in the EU

For the first time, emissions from internal combustion engines of mobile machines and equipment must be measured in real operating conditions for EU Emissions Stage V within the legal framework of the European Union. Engines in performance categories

11

NRE-v-5 and NRE-v-6 must be measured with regard to their exhaust tailpipe emissions in real operating conditions. With an output of more than 300 kW, the MAN D1556LE engine falls into category NRE-v-6. 5.2

Attaching the measuring equipment to the exhaust gas aftertreatment

The MAN D1556LE engine has a modular exhaust gas aftertreatment consisting of DOC/DPF and SCR catalytic converter. Once the exhaust tailpipe has been dismantled, the measuring tube of the PEMS measuring equipment (see Fig. 11) is mounted onto the outlet of the SCR catalytic converter. The exhaust gas is then extracted via a heated line to measure the exhaust gas volume flow. In addition to the gas emissions, data from the engine control unit is also recorded in accordance with the legal framework and output via the measuring equipment.

Fig. 11. PEMS measuring equipment

5.3

Challenges and difficulties in off-road use

In contrast to measurements from passenger cars or trucks, the off-road sector presents extensive challenges and difficulties for which the mobile exhaust-gas measuring equipment had to be adapted. For example, the PEMS equipment had to be designed as a modular system that can be mounted onto every possible machine in the off-road sector. While the design was relatively simple for tractors, additional installation space – within the legal maximum dimensions of a vehicle – had to be created for measuring vehicles such as harvesters to be able to install the mobile equipment. The power supply was deliberately not supplied by the vehicle or a power unit. This prevented coming into contact with the legal 1% rule as well as cross-influences from the unit’s exhaust gases affecting the emission measurement. The power was supplied

12

by a battery pack tailored to the needs of measuring equipment. This pack ensures that the measurements are carried out during at least one day and can be recharged overnight. Further difficulties included enclosing the measuring equipment in order to completely rule out influences from dust (e.g., harvest dust), humidity, and splashing water. Moreover, a measuring instrument that measures the recirculated air independently via synthetic air was used for analyzing the exhaust-gas emissions. The measuring equipment also had to be cooled as the maximum measuring temperature of 40 °C is quickly exceeded, especially during summer harvests. Fig. 12 shows how the PEMS measuring equipment is attached onto the front hydraulics of a tractor. All of the aforementioned issues are taken into account to ensure that the exhaust tailpipe emissions are measured correctly.

Fig. 12. Attaching the PEMS measuring equipment onto the front hydraulics of a tractor

In addition, the PEMS measuring tube is insulated to prevent excessive surface temperatures and thermal incidents during harvesting. One of the biggest differences to on-road applications is the dependency on weather conditions and harvest seasons. In Europe, for example, tractors can only be used for agricultural tasks under certain conditions and harvesters can only be operated during harvest seasons. As a result, there are only short time windows in which PEMS measurements can be carried out.

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5.4

Measurements and results

The length of the measurement varies greatly to achieve the legally required length of 5–7 times the cycle work of a Nonroad Transient Cycle (NRTC). For example, the measurement can take only approx. 1 hour for high-load tasks and up to 3.5 hours for low-load applications. Since the cold start is not yet a fixed component of the ISM measurements, several measurements can be carried out during the day. ISM measurements that have already been carried out on various vehicles in a wide range of applications have shown that the results of the determined conformity factors vary greatly, depending on the load and what the vehicle is used for. However, it can been proven that results from the engine test bench are reproducible under real operating conditions and that conformity factors smaller than 1 are also possible in highly utilized applications such as ploughing with a tractor.

6

Summary

The new 9-liter D15 engine with its compact dimensions and high torque perfectly complements the MAN engine portfolio. Thanks to the EU Stage V, EPA Tier 4 Final, and EU Stage IIIa certificates, it serves the global market. The SCR-only concept paired with state-of-the-art common rail injection technology and efficiency-optimized VTG turbocharging provides a solid basis for future developments. During testing, the results of real driving cycles were compared with results from the endurance program, thus optimizing the scope of validation and improving the understanding of component stress. Statutory ISM measurements must be carried out for the first time for emission stage EU V. The results vary greatly depending on the load and type of vehicle use. The results from the engine test bench are reproducible under real operating conditions.

CatVap® – a new heating measure for exhaust aftertreatment system Robert Szolak1 and Bernd Danckert2 Dr. Alexander Susdorf1, Paul Beutel1, Katharina Pautsch1, Christian Ewert1, Florian Rümmele1, Anand Kakadiya1, Dr. Achim Schaadt1 1

Fraunhofer ISE, Heidenhofstraße 2, 79110 Freiburg, Germany 2 ICCL – Integrated Consulting Company Ltd., Cyprus

Abstract. Today's challenge is to almost comply with the emission limits 100% in Real Drive Emission (RDE) behavior, especially in the cold season, the socalled "cold city rides", which are characterized by the fact that during the stop and go rides in the city at no time a sufficient temperature for an efficient exhaust aftertreatment is reached. The future Euro6+ or certainly the expected new Euro7 limit values will therefore only be achieved with suitable effective thermal heating, heat retention as well as regeneration measures for the exhaust aftertreatment (EAT) system. CatVap® is an easily scalable system for both diesel and gasoline engines that provides highly efficient thermal energy, simply from fuel (diesel, gasoline, synthetic fuels), coupled into the exhaust system and thus provides the thermal energy required for exhaust gas conversion. It is fast, highly dynamically controllable, can be used as a catalytic burner for rapid heating during a cold start of the engine system. Furthermore, CatVap® can then be used as a light-off accelerator (by fuel tailoring to synthesis gas (syngas) and shorter alkenes) for the Diesel Oxidation Catalyst (DOC) and the Selective Catalytic Reduction (SCR) catalyst. CatVap can also be used for the diesel particulate filter, for its effective active regeneration or temperature control for continuous regeneration. The CatVap system is an important key enabler for near zero-emission internal combustion (IC) engines that can be demonstrated in the near future. Keywords: CatVap, Fast Heating and Warming Measure, Exhaust Aftertreatment System, Real Driving Emissions, Enabler Near Zero-Emission IC Engines

1

Background

1.1

Upcoming emission legislation changes and motivation

At low engine loads, i.e. during inner-city driving and especially at low ambient temperatures below 0°C, the exhaust gas temperature is to low and not sufficient for efficient chemical catalytic reactions. For example, a reliable selective catalytic reaction for an adequate NOx-conversion requires an exhaust gas temperature of at least 220°C. In the situations described above, © Springer Fachmedien Wiesbaden GmbH, ein Teil von Springer Nature 2020 W. Siebenpfeiffer (Hrsg.), Heavy-Duty-, On- und Off-Highway-Motoren 2019, Proceedings, https://doi.org/10.1007/978-3-658-31371-5_4

2

however, this condition is only reached after a very long driving time or in extreme cases, e.g. with inner-city driving cycles in the cold seasons, possibly never at all. If the engine is running at low load, the exhaust gas temperature, which is used for efficient exhaust aftertreatment, is already too low to comply with the current limit values in this driving condition, as described above. Therefore, so-called internal and external engine thermal management measures must be applied, such as engine intake and exhaust throttling, late combustion phasing, cylinder deactivation, variable valve control (e.g. Miller-Cycle – internal EGR), late injection, as well as low and high pressure exhaust gas recirculation (with high recirculation rates), HC dosers, electrical exhaust heaters and various burner systems, etc. Furthermore, the increasing improvements in engine efficiency and more and more upcoming hybridization of the powertrain will lead to even lower exhaust gas temperatures in the future. The latest emission legislation might requires that the emission limit values down to -7°C for roller bench tests are permanently adhered to, and compliant with the limit values in real driving emission (RDE) operation. In addition, ambient temperatures of 0 degrees Celsius are under discussion. It can also be assumed that low-load operation will also be included in the driving cycle of type tests. For example this has already been done for SCR retrofitting of public transport buses. However, these classical heat management measures have a limited potential or are even unsuitable for the future requirements of maintaining emissions at low loads and low outside temperatures within the shortest possible time and also permanently. The exhaust gas recirculation technology in particular will reach its absolute limits here, strong soot formation and deposits will be the result, which will make the systems susceptible to failure. In addition, high EGR rates worsen the burn-through behavior of the combustion, unsteady running, especially in idling and low load operations, of the engine (most frequent urban operation) and soot, partial and unburned fuel components are transported with the excessively cold exhaust gas through the various exhaust aftertreatment system components. This causes deposits on exhaust gas flaps and valves as well as e.g. exhausts gas recirculation coolers and also causes condensate to form in the intake manifold. At the low emission certification temperatures required in the future, e.g. EGR in city operation and at low outside temperatures will only be applicable to a very limited extent or will hardly represent a measure for the reduction of NOx emissions. Shifting the focus of combustion to a late phase worsens thermal efficiency (leading to an increase in CO2) and increases the risk of incomplete combustion (and unburned hydrocarbons increase the risk of thermal overload of the DPF during regeneration) as well as increased particle loadings of the DPF (which leads in turn to increased necessary DPF regeneration cycles). The following picture shows the thermal losses of the exhaust gas downstream of the exhaust system with exhaust aftertreatment when extra fuel is introduced into the combustion chamber to heat up the exhaust gas. It becomes clear how inefficient the process and how low the thermal efficiency of the exhaust system component to be heated are and underlines the need for a better way to heat the exhaust system quickly and efficiently and to keep it in the right temperature range.

3

Fig. 1. Schematic representation of an engine with exhaust aftertreatment and Sankey diagram of thermal losses1

Furthermore, the increase in the exhaust gas temperature at idling speed and in the lowest load range (stop and go) caused by cylinder deactivation and optimization of the valve lift curve (internal EGR, e.g. by Miller cycles) is very low and the effort required for this must be regarded as too high. Due to these challenges, new requirements arise for the developments of more effective EAT heating measure: A sufficiently high light-off temperature for the exhaust aftertreatment components (e.g. for DOC and SCR components and catalysts) has to be achieved very quickly and immediately after engine start and maintained sufficiently high for a good emission reduction performance under all operating conditions (especially in the low load operating conditions described above). Since the exhaust gas temperature cannot meet these requirements due to the operating condition, an additional energy source must be used to increase the temperature. The CatVap® system, whose function is described below, fulfils these requirements in a simple and effective way.

4

2

New Heating Measure CatVap®

The fuel processing technology CatVap® is a new heating measure for exhaust aftertreatment systems. The heat will be provided thermally by adding extra fuel with a regulated dosing pump depending on the actual power requirement into the CatVap® system. The CatVap® system has a very compact design (see Fig. 2 compared to a 0.33 l can) and a high power density1 (45 kW/l) as well as a wide modulation range from at least 1/18 (power range from 1 kWth. for heat retention with very low load to 18 kWth. for DPF regeneration and fast EAT heating).

Fig. 2. Functional prototype of the fuel processing technology CatVap®

The CatVap® system is very easily scalable due to its stack construction. Two combined CatVap® systems as shown in Fig. 2 are already sufficient for medium and heavy duty applications (see Fig. 4, CatVap® system integrated in the exhaust pipe for medium and heavy duty applications). The CatVap® system contains an electric heating catalyst (with common industrial coating) in which the chemical reactions take place. The diesel fuel reacts with the oxygen contained in the exhaust gas and generates heat. The new heating technology CatVap® can be easily highly dynamically controlled in performance and can be operated in two different operating modes, which are briefly described below (and which will be described in more detail later): 1. CatVap® burner operation mode for cold start and fast DOC heating until light-off temperature (lean operation; CatVap ≥ 1) 2. CatVap® DOC operation mode under rich conditions (CatVap < 1) for tailoring of diesel fuel to syngas (CO and H2) and short chain alkenes which will be oxidized

1

Power density CatVap®: kW/l thermal power per liters CatVap® volume, thermal power is based on the lower heating value of diesel

5

in the DOC. By supplying syngas to the DOC the light-off temperature is lowered significantly. The possibility of fuel tailoring shows a significant reduction in the DOC light-off temperature with one of the important advantages of the CatVap® DOC mode being that high performance can be achieved and the heat is generated precisely where it’s needed in the DOC catalyst. The proof of concept was carried out on a catalyst screening test bench, hot gas test bench and engine test bench on a passenger car diesel as well as on an industrial heavy-duty diesel engine (results will be shown in the course of this article). The operational strategy of CatVap® burner mode and DOC mode allows a rapid heating of the EAT. The DOC catalyst is approx. 1/5 of the total EAT catalyst mass. Due to the reduction of the light-off temperature and the relatively low catalyst mass, the DOC light-off temperature can be achieved very quickly. After switching into DOC mode the whole EAT will be heated up with high power which is resulting in a rapid attainment of the SCR light-off (see simulation results chapter 3). 2.1

Integration new heating measure CatVap® in a diesel series engine

Fig. 3 shows the integration of the CatVap® system in a series diesel engine. The required exhaust gas for the CatVap® system is generated in the exhaust gas bypass flow. The bypass flow can be generated via a turbocharger bypass (wastegate, see position 1) or downstream of the turbocharger in the exhaust system, e.g. via an exhaust throttle valve, as most commercial vehicles and industrial engines for thermal management or engine braking purposes already have today (see position 2).

6

Fig. 3. Integration of CatVap® in a series engine (schematic). The CatVap® system is in bypass of the exhaust gas (figure above). Position 1: Integration CatVap® in the wastegate line (left). Position 2: Integration after turbocharger, use of exhaust flap (right)

Following Fig. 4 shows the new generic CatVap® design for medium- and heavy-duty applications integrated downstream of the turbocharger with exhaust pipe, exhaust flap and one possible mixing zone (see position 2 at Fig. 3). The CatVap® system is easily scalable by stack constructions and enables the customization to different applications, from passenger cars to heavy-duty on- / off-highway applications. The assembly as shown in Fig. 4 allows a very compact design (total volume of approx. 0.8 liters, diameter approx. 120 mm). The new stack design can be run with at least 9 kWth. in the “burner mode” and with 36 kWth. power in the “DOC-mode” (see chapter 2).

7

Fig. 4. Integrated CatVap® system for medium- and heavy-duty applications with exhaust flap and mixing zone.

The CatVap® system is designed for low pressure drop. The pressure drop at CatVap® burner mode is approx. 38 mbar and can be reduced at DOC mode to approx. 6 mbar. Further reduction of the pressure drop in the burner mode is possible by optimization of the flow conditions in the CatVap® system. The CatVap® operating modes, the burner and DOC mode and their effects are described in more detail below. 2.2

CatVap® Burner Mode:

The CatVap® system can be operated as a catalytic burner with exhaust gas coming from the diesel engine. The burner mode is very robust. Several tests carried out at Fraunhofer ISE under transient conditions with strongly fluctuating oxygen concentration and mass flow variations have shown a very stable operating behavior. The CatVap® system requires no extra air (no air compressor). Due to the catalytic combustion no additional NOx will be produced during the cold start phase.

8

The CatVap® burner mode can be initialized immediately after starting the engine. The hot exhaust gas from the CatVap® system then heats up the DOC very quickly to light-off temperature. The thermal power of the CatVap® system depends on the exhaust gas bypass mass flow. The experiments have shown a high thermal performance of up to 12 kWth. Even higher power at burner mode seems to be possible. The CatVap® outlet temperature is regulated to maximum temperature of 900 °C to avoid catalyst deterioration. 2.3

CatVap® DOC Mode:

After having reached the DOC light-off temperature, the CatVap® system switches in the DOC mode. During DOC mode, a significant increase in thermal power output is possible. The CatVap system then runs under so-called rich conditions. Under the influence of heat, the incoming fuel reacts to synthesis gas (hydrogen and carbon monoxide) and shorter alkenes, such as propene and ethene (see Fig. 5). This product gas composition leads to a significant reduction in the DOC light-off temperature. The proof of concept was demonstrated on a catalyst characterization test rig, a hot gas and an engine test bench for a passenger car engine and an industrial heavy-duty diesel engine. Fig. 5 shows the composition of the educts (negative y-axis; diesel fuel and simulated exhaust gas with 14 vol.% oxygen) and the measured product gas composition (positive y-axis). The diesel fuel used in the experiments is conventional fuel from the petrol station. It consists of a mixture of alkanes, alkenes, aromatics and oxygenated hydrocarbons (biodiesel). At standard temperature and pressure, the alkanes, as shown in the following diagram are composed of gaseous (C1-C5-species) and liquid alkanes (> C6-species). The diesel fuel that is fed into CatVap® consists only of liquid, longchain alkanes. With the exhaust gas flowing into the CatVap®, the diesel fuel reacts with the residual oxygen contained in the exhaust gas and generates heat and reaction products such as hydrogen and carbon monoxide (syngas). Furthermore, the diesel fuel will be cracked to mostly short-chain alkenes such as propene and ethene. The product gas contains almost no condensable (at room temperature) liquid alkanes resulting in a significant reduction in the DOC light-off temperature, as shown in Fig. 6.

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Fig. 5. Feed streams into the CatVap® system (educts, negative y-axis) and the composition of the product gas at CatVap outlet (positive y-axis)

Several investigations were carried out to measure the light-off temperature behavior on a DOC catalyst with the CatVap® product gas. Fig. 6 shows the light-off temperature behavior of the CatVap® product gas tested at an OEM on a passenger car diesel engine. The exhaust gas flow was set to 50 kg/hr and maintained at 140 °C. It displays the DOC and DPF temperatures as a function of time. The CatVap® product gas was introduced after approx. 40 seconds. After introducing the CatVap® product gas, a fast temperature rise at DOC was achieved. The thermal power, introduced by CatVap was 6 kW.

Fig. 6. Stationary diesel engine test and DOC light-off behavior as well as DPF temperature.

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3

Simulation CatVap® under real WHTC conditions for heavyduty on-road applications

The simulated power demand related to a WHTC cycle2 for a heavy-duty on-road application is shown in Fig. 7. The simulations have been carried out in order to optimize the cold start behavior and to achieve the SCR light-off as soon as possible related to the CatVap® power range up to 36 kWth. and an EAT catalyst mass of approx. 15 kg. Fig. 7 shows additionally the measured SCR temperature without any thermal management measures compared to the simulated SCR temperature with the theoretical power supply of 4 kWel. respectively 8 kWel. No heat losses are taken into account. Furthermore, the simulations do not take into account exhaust pipe masses as well as the connectors as these depend on the way of implementation at a given EAT. However, by comparing different heat measurements under the same boundary conditions, the trend should be the same in percentage terms. Without any thermal management measures, the SCR light-off temperature was achieved hardly after approx. 1,600 seconds. Additional heating measure is required for approx. 70% of the WHTC-cycle.

Fig. 7. Total exhaust mass flow (1st diagram), the simulated power input time-optimized regarding fast start up related to the power range of the CatVap® system and compared to electrical heating 4 kW and 8 kW (2nd diagram) and the resulting SCR temperature with CatVap®, 4 kWel., and 8 kWel.heating (3rd diagram red line) and the measured outlet temperature without thermal management measures (3rd diagram blue line).

With CatVap®, the SCR light-off temperature (in our case 250°C) is already reached after 80 seconds! With the same EAT configuration and the electric heater installed in 2

World-wide Harmonized Transient Cycle

11

front of the DOC, the SCR light-off temperature is reached with 8 kWel. after 175 seconds or 415 seconds with 4kWel.. With a high thermal power input at the beginning of the cycle, the SCR light-off temperature can be reached very quickly, which is one of the most important prerequisites for the emissions (related to NOx) to become almost zero.

Fig. 8. Section of the first 480 seconds of the WHTC cycle (cold) and comparison of different heating measure (CatVap® and electrical heater) and the resulting SCR outlet temperature.

The following Fig. 9 shows the integral of the heating power over the entire WHTC cycle. The simulation results consider the initial electrical start of the CatVap®-system which is needed for a very short time. The intention of the operation strategy of CatVap® is to achieve the SCR light-off as quickly as possible. The default settings of DOC mode with CatVap® in burner mode (9 kWth.) until reaching DOC light-off (approx. 160 °C) and DOC mode with maximum CatVap® power (36 kWth.) until reaching the SCR light-off were specified in the simulations. Because of the high power input with CatVap the averaged heating power is compared to the electrical heating measure higher (see Fig. 9). Comparing the extra fuel which will be required for the EATheating based on the average power input, the CatVap® system is much more efficient than the electrical heating measure because of the relatively low engine and alternator efficiency. The additional required fuel consumption for EAT-heating with CatVap® could be reduced by 67 % (compared to 8 kWel. heater) and over 50 % (compared to 4 kWel.heater) despite of a higher power input. Another advantage of EAT heating with CatVap is that it can be used as a heating measure for DPF regeneration at all engine loads.

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Fig. 9. Comparison of the average power input at WHTC cycle and comparison of the extra fuel introduction for EAT-Heating.

4

Experimental investigation of CatVap® under WHTC conditions for heavy-duty on-road applications

To investigate the CatVap® system under nearly realistic transient conditions, a dynamic test rig was built at Fraunhofer ISE (Fig. 10). The experiments are carried out under transient conditions based on real engine data (e.g. WHTC). The CatVap® system is tested with highly precise and fast reacting mass flow controllers with fluctuating O2 concentrations and with constant or fluctuating bypass mass flows, whereby the exhaust gas composition is represented by corresponding gas mixtures (see Fig. 11). Additionally highly dynamic fuel pumps from automotive applications are used for the experiments.

Fig. 10. Dynamic test rig at Fraunhofer ISE.

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Fig. 11 shows the total exhaust mass flow at WHTC conditions for a medium duty application during the first 300 seconds after engine start. Furthermore, the exhaust bypass flow to the CatVap® system is shown, considering two control options. The dashed red line shows the CatVap® bypass flow with an adjustable exhaust flap, keeping the bypass flow constant. The solid red line shows the CatVap® bypass flow with a fixed exhaust flap position and highly fluctuating CatVap® exhaust mass flow. The highly fluctuating oxygen concentration can set in both cases as described before.

Fig. 11. Total exhaust mass flow (1st diagram), exhaust bypass flow through the CatVap® system related to a fixed exhaust flap position (fluctuating mass flow) and a regulating exhaust flap (constant mass flow) and oxygen concentration exhaust gas.

Fig. 12 shows experimental results from the dynamic characterization test rig with CatVap® first run in burner mode with 12 kWth. (bypass flow 40 kg/h) and then switched in the DOC mode with 36 kWth (bypass flow 14 kg/h). The oxygen concentration varies depending on the real WHTC data. The SCR temperature was calculated based on the experimental CatVap® results. Heat losses of the CatVap® system, the CatVap® reactor mass and evaporation enthalpy of diesel were taken into account. The total EAT catalyst mass in the calculations was 15 kg. The DOC light-off temperature was reached based on the experimental results after approx. 80 seconds. The CatVap® system then switched to DOC mode by increasing the diesel flow and reducing the bypass flow. The SCR light-off temperature was reached after approx. 110 seconds. The SCR and DOC temperatures were calculated without taking any heat losses into account.

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Fig. 12. Experimental results with CatVap® at burner mode with 12 kWth. power (bypass flow 40 kg/hr) and DOC mode with 36 kWth. (bypass flow 14 kg/h).

The CatVap® product gas temperature at the outlet of the CatVap® system reacts very quickly to mass flow and oxygen changes as shown in Fig. 13. The CatVap® thermal output was approximately 16 kWth. Fig. 13 also shows the CatVap® cold start behavior after the initial electrical pre-heating.

Fig. 13. Experimental CatVap® results with fluctuating bypass flow and oxygen concentration.

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The experiments have shown that the system is very robust against mass flow and oxygen changes. The CatVap® bypass flow changed between 30 kg/h and 90 kg/h and the oxygen concentration between 7 and 20.5 % by volume. The experiments on the dynamic test bench under almost realistic transient conditions, based on engine data (WHTC), have shown that CatVap® can very quickly switch between burner and DOC mode with strongly fluctuating oxygen concentration and/or fluctuating bypass mass flow and oxygen concentration. In addition, high synthesis gas concentrations (hydrogen and carbon monoxide) were achieved in DOC mode. Due to the very compact design and the robustness against exhaust gas bypass mass-flow and oxygen fluctuations, the CatVap® system is very well suited for efficient EAT heating. The high power range (2 kWth. and 36 kWth.) and the two possible operating modes (burner mode and DOC mode) enable a very fast SCR light-off, which is, together with synthetic fuels (PtL3 and BtL4), important for near zero emission internal combustion engines.

5

Conclusion and Outlook

The article describes the function of an innovative, compact and highly dynamically controllable exhaust aftertreatment heating, tempering and regeneration system called CatVap®. The new, stricter exhaust gas legislation, which will soon come into force for a wide range of applications in passenger cars and commercial vehicles, requires rapid heating and thus rapid activity of exhaust aftertreatment systems after a cold start, especially in RDE city operation, and this will be necessary in significantly colder ambient conditions. The article impressively describes the highly dynamic function of the CatVap® system. The system can be used as an integrational unit for future new OEM Euro7 or as an upgrade or supplement to existing Euro6 exhaust aftertreatment systems. CatVap® also has the potential to reduce the complexity of existing, very complex exhaust aftertreatment systems of combustion engine or hybrid drives, i.e. to design them less complex. In concrete terms, this can lead to the saving of existing, less effective exhaust aftertreatment system components, e.g. to the partial or complete replacement of existing exhaust gas recirculation technologies. Compared to competitor systems, the system has significantly lower additional fuel consumption and thus a considerable CO2 advantage. The CatVap® system is suitable for all combustion processes of all different combustion engines (diesel, petrol and gas engines), provides high thermal performance in a small space, is highly dynamically controlled and easily scalable (from passenger cars to larger heavy-duty applications) and can be operated with almost all known fuels (especially with new, future fuels such as OME (oxymethylene dimethyl ether), FischerTropsch diesel, biofuels like octanol and blends). The effect of these synthetic fuels on engines and their exhaust aftertreatment system is currently investigated in the BMWi 3 4

Power to Liquid Biomass to Liquid

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funded project C3-Mobility with many partners including Fraunhofer ISE. Furthermore, this technology is particularly suitable for hybrid drives, as the combustion engines are regularly switched off and thus cool down quickly. CatVap® also makes a considerable contribution to making combustion engines, especially diesel engines, more economical and, above all, virtually emission-free. Especially because CatVap® has been designed for the operation of various fuels, this technology represents more than just a bridge technology for future emission-free mobility, especially when using new CO2 neutral fuels. Fraunhofer ISE is also very successful in the research of “alternative and synthetic fuels” and “Power to X technologies” (see also Key Note Speak “The Role of Synthetic Fuels in an Integrated Energy System”) and offers promising synthesis technologies for these new fuels which are assessed in terms of technology, economy and ecology (Life Cycle Assessment, LCA). In the interaction of different industrial sectors, future potential for combustion engines or combined hybrid drives can be demonstrated effectively by appropriate sector coupling, here in particular with solar technologies to a further economic and ecological sustainable continued utilization of new combustion engine concepts, even in the distant future. In cooperation with the OEMs and their power train units, the Fraunhofer ISE Institute in Freiburg is currently conducting a series of further experiments with the CatVap® system. The CatVap® system, which is currently in the functional prototype stage, will be brought into the next phase of industrialization in cooperation with wellknown and established partners from the 1st and 2nd tier supplier industries. Further promising results will soon be presented to the market in spring 2020.

References 1. Gao et al., An analysis of energy flow in a turbocharged diesel engine of a heavy truck and potentials of improving fuel economy and reducing exhaust emissions, Energy Conversion and Management, Jan. 2019 2. Szolak, R. et al, “Active filter regeneration with fuel vapor and syngas”, Heavy-Duty, Onand Off-Highway Engines 2012, 7th International MTZ Conference 6. – 07.11.2012

A novel low-cost aftertreatment solution for lean‑burn gas engines Matthew Keenan1, Jacques Nicole2 and Ben Rogers1 1

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Ricardo UK, BN43 5FG, UK Ricardo Inc., Santa Clara, CA 95054, USA [email protected]

Abstract. Lean burn methane fueled combustion engines give low tailpipe CO2 emissions compared to diesel and stoichiometric natural gas applications. Optimised engine efficiency by operating lean combined with a low carbon containing fuel lead to the lowest possible tailpipe CO2 for a non-hybridized system. However, due to the lean combustion mode, both NOx and CH4 become a challenge for exhaust emissions control. Traditionally, urea based SCR is used to control NOx and a highly loaded precious metal based methane oxidation catalyst (MOC) is used to attempt to achieve low tailpipe CH4 emissions. Therefore, due to the number of catalyst and the associated control requirements, the cost of the combined aftertreatment solution becomes a significant proportion of the total engine cost. In addition to the high cost of the methane control catalyst, the MOC only becomes highly efficiency in the region of 500 oC, which for lean operating engines is rarely reached under normal engine operation. Due to the high temperature stability of methane, a novel approach has been taken to develop a catalyst system which is able to oxidise methane at low temperatures, via the use of alternative oxidising agents. Dioxygen (O 2) is relatively stable, whereas ozone (O3) is highly reactive and is a significantly stronger oxidising agent compared to O2. Synthetic gas reactor experiments were performed using O3 as the oxidising agent and methane as the hydrocarbon feed. A current production iron based SCR catalyst was used and was found to oxidise methane at 220oC with an efficiency of >60%. Experiments are continuing with a multi staged fixed bed reactor with the aim of demonstrating >95% conversion, leading to the potential to eliminate the expensive PGM based MOC. These results will be presented at the conference. There is a cost associated with the onboard O3 generator and the associated power consumption needs to be included in the engine efficiency. However, the solution is anticipated to deliver a significant GHG (CO2 + CH4 equivalent) reduction compared with diesel and stoichiometric NG solutions but at an aftertreatment cost that is lower than diesel and equivalent to stoichiometric NG. This paper will discuss the novel approach to emissions control utilising an alternative oxidising agent for low temperature and low cost emissions control, including the advantages and disadvantages of the system for different applications. Keywords: Methane, Ozone, Low Temperature.

© Springer Fachmedien Wiesbaden GmbH, ein Teil von Springer Nature 2020 W. Siebenpfeiffer (Hrsg.), Heavy-Duty-, On- und Off-Highway-Motoren 2019, Proceedings, https://doi.org/10.1007/978-3-658-31371-5_5

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1

Introduction

1.1

Natural Gas Challenge

Reducing the tailpipe CO2 from industrial engines is a significant challenge and the choice of fuel plays an important role. The use of natural gas (NG) can yield lower tailpipe CO2 as it contains 75% carbon by mass compared to 86% carbon by mass for Diesel. Therefore, combusting 1kg of NG yields 2.75kg of CO2 whereas, combusting 1kg Diesel yields 3.14kg of CO2. Considering heat of combustion (LHV in the case of reciprocating engines) then NG yields ~58gCO2/MJ relative to Diesel ~75gCO2/MJ. In other words, 28% more CO2 is emitted with Diesel compared to NG on a fuel heat of combustion basis. Hence, operating an industrial engine on natural gas (NG) has it merits. To further reduce the CO2 impact, the engine can be operated lean which leads to the lowest possible tailpipe CO2 for industrial engines without the use of hybridisation. However, the challenge with all NG engines is controlling the emissions of methane. Methane has a significant greenhouse gas impact (GHG) of 28 times that of CO2 for a 100 year greenhouse warming potential (GWP). Methane is the most stable hydrocarbon and does not catalytically oxidise until 500oC in the exhaust. This is due to the strong C-H bond in methane, which requires 416kJ/mol of dissociation energy to break the bond compared to 346 kJ/mol for a C-C bond for longer chain hydrocarbons. Hence high temperatures are required for methane combustion compared to higher hydrocarbons. During stoichiometric combustion exhaust temperatures can be significantly greater than 500°C but under lean combustion the exhaust temperatures can be significantly below 500°C. Hence for precious metalbased exhaust catalytic methane control there is a significant challenge in controlling emissions in a lean operating engine. Traditional precious metal based three-way catalysts (TWC) for stoichiometric operation and methane oxidation catalysts (MOC) for lean operation, are highly loaded with 150 –200 g/ft3 PGM, leading to high costs for the aftertreatment system [1]. In additional to the stability of methane, oxygen is not a very strong oxidising agent, which compounds the challenge. Hence the focus of this research was to investigate stronger oxidising agents which can react with hydrocarbons at low temperatures, with the focus being ozone (O3). Ricardo calls this low temperature emissions control approach using strong oxidising agents KINETA. 1.2

Industrial Applications – Methane Emissions Legislation

For European, On Highway Heavy Duty Engines, the Euro VI legislation has a specific limit of 500mg/kWh CH4 over the WHTC only for positive ignition engines. For European Stage V, Off Highway Machinery operating on fully and partially gaseous fueled engines, the HC limit is replaced by a value calculated as follows and includes CH4 HC = 0.19 + (1.5 x A x GER) where A is the A Factor = 1.10 and where GER is the average gas energy ratio over the appropriate cycle. Where both steady-state and transient test cycles apply, the GER is determined from the hot-start transient test cycle. Gas Energy Ratio (GER) is the value

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of the energy content of the gaseous fuel divided by the energy content of both fuels. In the case of single-fuel engines, GER is either 1 or 0 depending upon the type of fuel. Where 100% NG is 1 and 100% diesel is 0. If the calculated limit for HC exceeds the value of 0.19 + A the limit for HC is set at 0.19 + A. Hence for mono-fuelled natural gas Off Highway application, the HC limit (including CH4) = 1.29g/kWh. The methane emission limit is significantly higher for Off Highway compared to On Highway for the most current legislation [2]. For the example of a lean burn application having an engine out methane emission of 3g/kWh, this would lead to methane efficiency requirements of 83% for On Highway and 57% for Off Highway industrial applications. Fig. 1 shows engine out data for a range of speed and load operating points (Key Points KP), for a Ricardo developed lean burn dedicated NG operating engine where the WHTC CH4 was 3g/kWh. The methane range in terms of ppm was 700 – 1700. Hence this sets targets for the methane control requirements for the research and development project of >57% efficiency over a range of emissions from 700 – 1700ppm, in realistic exhaust conditions containing high concentration of H2O.

Fig. 1. Engine out methane emissions for a Lean Burn Industrial engine (KP = Key Points)

The Large Engine (marine and powergen) sectors do not currently have methane emission legislation in place although there are indicators that this legislation will come in future. An example of the potential for methane emission legislation comes from the European Commission’s Large Combustion Plant Directive BREF, where “in order to reduce non-methane volatile organic compounds (NMVOC) and methane (CH4) emissions to air from the combustion of natural gas in spark ignited lean-burn gas engines Best Available Technology, BAT is to ensure optimised combustion and/or to use oxidation catalysts” [3]. Regardless of direct methane emission legislation, there

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are already in place GHG reduction targets which indirectly drive for methane emission reduction. An example of a significant indirect, GHG based driver is the IMO (International Maritime Organisation) resolution to “reduce the total annual GHG emissions by at least 50% by 2050 compared to 2008, while, at the same time, pursuing efforts towards phasing them out entirely” [4]. Taking the GHG example, methane emissions have a 28-36 GWP (Global Warming Potential factor) in a 100 year timeframe [5], therefore even relatively small quantities of methane emission from a NG engine can cancel out the fuel’s inherent CO2 advantage and even lead to a net GHG increase, relative to a Diesel fueled engine. Hence there is a driver in the Large Engine market for methane aftertreatment to address both the direct and indirect impacts of methane emissions. 1.3

Properties of Ozone

Table 1 gives a summary of the properties of ozone and oxygen. The standard reduction potential is a measure of how strong an oxidising agent the compound is, with the larger number being a stronger oxidising agent. Hence ozone is a significantly stronger oxidising agent compared to oxygen. However, energy is required to form ozone at 143kJ/mol, which requires energy from the system. Due to different generator designs and efficiency, ozone production ranges from 7-15kW per kg/h O3 [6, 7]. It is difficult to make ozone and then store as it is generally unstable hence the approach for this research is to produce ozone via a corona discharge then use the formed ozone directly in the simulated exhaust gas. Ozone is thermally unstable at elevated temperatures, by 350°C ozone has completely dissociated and reverted to oxygen. Hence, due to its lack of stability, ozone must be ustilised under relatively low temperature conditions which coincide with the exhaust temperatures found in a lean operating exhaust. Table 1. Properties of Ozone and Oxygen.

Formula Molar Mass Solubility (H2O 25°C mg/l) Boiling Point (°C) Enthalpy of formation (kJ/mol) Standard Reduction Potential (E° acid – V) Complete Dissociation Temperature (°C)

Ozone O3 48 109 -112 143 +2.08 350

Oxygen O2 32 8 -183 0 +1.23 n/a

Equation (1) shows with the requirement of an energy input, the formation of free oxygen (O*), which further reacts with O2 in reaction (2) to form O3. O2 → O* + O*

(1)

O* + O2 → O3

(2)

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Equations (3) and (4) show the dissociation of O3 back to O2, which can occur thermally of catalytically.

1.4

O3 → O2 + O*

(3)

O* + O3 → 2O2

(4)

Ozone Reaction with Methane over Fe-BEA catalyst

Previously, it has been shown that methane can be fully and partially oxidised at temperatures in the region of 200°C when using ozone as the oxidising agent over a current production Fe-BEA catalyst [8, 9]. Fig. 2 shows the conversion efficiency as a function of temperature for reaction of methane with ozone over and Fe-BEA catalyst and the reaction of methane with oxygen over a traditional precious metal-based catalyst. With O2 as the oxidising agent, methane conversion efficiency starts at approximately 350°C with high efficiency (>95%) occurring at approximately 500°C in a fresh catalyst and >550°C for an aged catalyst. With ozone as the oxidising agent, a peak efficiency of 50% at 220°C was achieved in the presence of water. The use of ozone shows a step change in the temperature at which methane can be control over a simple catalyst which is currently used in Heavy Duty applications as an SCR catalyst. The products of the reaction are both CO and CO2 depending on the reaction conditions. Postulated reactions include partial oxidation and complete oxidation. CH4 + O3 → CO + 2H2 + O2

(5)

CH4 + 4/3O3 → CO2 + 2H2O

(6)

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Fig. 2. Methane conversion efficiency over Fe-BEA with O3 (100ppm CH4 and 2.5% H2O for experiment with water) and over a PGM catalyst with O2

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Experimental Approach

All experiments were conducted in a bench scale continuous flow reactor, comprising of an in-line heater, gas injection ports, water injector and vaporizer, in-line mixer, sample ports, sample holder, and exhaust. The catalyst samples were all cordierite cores with 300 cells per square inch cell density. The synthetic gas reactor was modified to house multiple catalyst sections each with their own ozone feed. Up to 3 catalysts and ozone feeds were used in the series of experiments. The gases were mixed by a helical stainless steel in-line static mixer directly upstream of the catalyst sample. The ozone was generated by air-cooled, corona discharge ozone generators. The ozone generators were fed with dry air or oxygen. The 1 to 2 SLPM ozone containing outlet stream of the ozone generator was fed to the reactor directly upstream of the individual catalysts. Outlet gas samples were analysed by an FTIR gas analyser. Up to 3 ozone injectors were used in the suite of experiments. The ozone generators used were only 6.5% efficient and had a power consumption of 12.7kWkg-1h-1 O3. Fig. 3 shows the schematic of the multi-staged reactor used in the low temperature methane control experiments with individual catalyst sections each with its own ozone feed.

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Fig. 3. Schematic of the multistage reactor

3

Results

The experimental set up was designed to determine the impact of multi-staging the ozone feed to the catalysts. Fig. 4 shows the impact of increasing the number of ozone feeds to the reactor on the methane conversion efficiency at a fixed temperature of 230°C and a methane feed of 100 ppm, without water. 3:0:0 is where 100% of the ozone feed was added at the first stage and led to an efficiency of 71%. 1:0:0 is where one third of the ozone feed was added at the first stage and gave 33% efficiency. 1:1:0 is where one third of the ozone was added at the first stage and another one third of the ozone feed was added at the second stage, which led to an efficiency of 51%. 1:1:1 shows 91% efficiency where equal thirds of the ozone were added over three stages. There was a 20% advantage in methane conversion efficiency when supplying ozone in three stages compared to adding the ozone at a single injection point. This shows the advantage of multi-staging the ozone feed to increase the conversion efficiency of the reaction.

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Fig. 4. Impact of multiple ozone feeds on methane conversion efficiency

All future experimental data shown are with all 3 ozone feeds active. Fig. 5 shows the methane conversion efficiency as a function of temperature in the absence of water with a methane feed of 1250ppm. At 100°C > 50% methane conversion efficiency was achieved. However, peak efficiency of 51% was achieved at 220°C with total consumption of ozone occurring at 140°C with 3 feeds of ozone. Hence showing that the reaction is limited by ozone but leads to the capability of using ozone as an oxidant but not emitting tailpipe ozone. To control any ozone slip below 140°C the ozone feed would be reduced or a low loaded precious metal catalyst would decompose any slipped ozone.

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Fig. 5. Methane conversion efficiency (1250ppm CH4 no H2O, 3 ozone injection points)

Fig. 6 shows the impact of adding 10% water to the feed-gas. The addition of water increased the conversion efficiency to a peak of 63% at 220°C, compared to 51% in the absence of water. 10% addition of water is consistent with the water content expected in the exhaust gas of hydrocarbon-based fuel and hence is considered a realistic quantity for the experimental set up.

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Fig. 6. Methane conversion efficiency (1250ppm CH4, 10% H2O and 0% H2O, 3 ozone injection points)

Fig. 7, Fig. 8 and Fig. 9 show results of experiments targeting high flows of methane and determining the impact on the catalyst system in terms of methane conversion. The feed-gas conditions used were 300 – 2500 ppm of methane, with the water content ranging from 9 to 12% and a fixed ozone feed of approximately 1800 ppm split equally over 3 injection points. The experiments were performed at a steady state temperature of approximately 220°C. Fig. 7 shows the removal of methane as a function of methane feed. Increasing the methane feed increased the amount of methane removed for all water concentrations. High concentrations of water do not impact the reactions taking place showing the approach is robust to realistic exhaust conditions.

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Fig. 7. Increasing methane removed with increasing methane feed at 220°C

Fig. 8 shows the ozone utilization which can be defined as moles methane converted divided by the moles of ozone consumed. Increasing the methane feed increased the ozone utilization by greater than a factor of 5 from 300 to 2500 ppm methane feed. The greater utilization was seen for the higher water content in the feed-gas.

Fig. 8. Increasing ozone utilization with increasing methane feed

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Fig. 9 shows the methane conversion efficiency for the different water concentrations. Under low methane feed conditions and all water concentrations, >80% methane conversion was achieved. Under high methane feed conditions and all water concentrations feed and >40% methane feed conversion was achieved. The peak efficiency of 93% at 220°C, was achieved at a methane feed of 600ppm and a water content of 9%.

Fig. 9. Methane conversion efficiency with and without water in the feed-gas

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Discussion and Conclusions

A step change in the temperature at which methane can be controlled in realistic exhaust conditions has been shown with the use of ozone as the oxidant and using a low cost current production PGM free catalyst. The active temperature range ~180-350°C of the KINETA technology is very compatible with existing exhaust aftertreatment installation locations and an advantage relative to the circa 500°C requirement for precious metal based catalysts. The fact that water does not impact the reaction for methane control is an interesting and promising result for this technology. Many catalytic processes are impacted by water and hence lead to a reduced efficiency. One hypothesis is that the water stabilises the ozone in the cage structure of the zeolite and hence allows the ozone to survive long enough to react with the methane rather than thermally or catalytically decompose to oxygen and an oxygen radical. It has also been shown that the ozone utilization increases with methane concentration. This can be explained by an increased collision frequency between

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ozone and methane and hence leading to an increased reaction potential. Further reactor design would target increasing the ozone utilization to limit the production demands of ozone. The efficiency of the ozone generators would need to be improved to reduce the power demand of the generator. The ozone generators used were only 6.5% efficient with a power requirement of 12.7kWkg-1h-1 ozone. However, at a 6.5% ozone generator efficiency, the cost per mole of ozone in terms of Diesel consumption would be $0.112. Whereas for comparison, the cost of ammonia per mole for NOx control via the SCR reaction is $0.127. Increasing the ozone generator efficiency to 10%, which is commonly quoted for many generators, would lead to a per mole ozone cost of $0.073. The calculation assumes a Diesel cost of $2.57/gallon and a DEF cost of $5.60/gallon. However, the technology does allow methane control for lean operating natural gas applications which historically has been a significant challenge due to the low exhaust temperatures and the high methane light off temperatures. Those applications are discussed below.

Fig. 10. Methane Conversion Efficiency Requirements for Different Industrial Applications

Fig. 10 shows the regions where methane can be controlled and for the range of exhaust conditions for an Off Highway industrial engine. Initial investigations have shown that >57% efficiency could be achieved to meet the Off Highway efficiency targets. Further development would be required to meet the 83% efficiency requirement for On Highway applications.

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Fig. 11. Large Engine GHG Benefits (applies a 28 times GWP factor)

Fig. 11 shows the GHG benefit of KINETA for Large Engine applications is also significant, bringing SI Gas, lean burn engines up to levels similar to HPDI, without the need for high pressure gas systems. Notably, the net GHG benefit of the SI Gas, lean burn engine with KINETA would be better than a Lambda1/TWC engine whilst retaining the lean burn CO2 and cost of power advantages, Dual Fuel engines, preferred in the marine market for safety reasons, tend to have relatively high methane emissions with a net GHG negative. The potential benefit of KINETA could bring these engines to a net positive. Going further, Fig. 12 shows the potential for Dual Fuel engines to achieve a 50% reduction in GHG using a combination of KINETA and a combustion system upgrade, with just 15% required from other measures, such as blending in biogas.

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Fig. 12. Dual Fuel engine potential walk to 50% GHG reduction

References 1. 2. 3. 4. 5. 6. 7. 8. 9.

Raj, A., Johnson Matthey Technol. Rev., (2016) 60(4):228–235 www.emleg.com https://eippcb.jrc.ec.europa.eu/reference/BREF/LCP/JRC107769_LCP_bref2017.pdf http://www.imo.org/en/MediaCentre/PressBriefings/Pages/06GHGinitialstrategy.aspx (2018) https://www.epa.gov/ghgemissions/understanding-global-warming-potentials www.lenntech.com/otozone.htm Magara, Y., Itoh, M. & Morioka, T., Progress in Nuclear Energy, Volume 29, Supplement, 1995, pp175-182. https://doi.org/10.1016/0149-1970(95)00041-H Keenan, M., Nicole, J. & Poojary, SAE 2018-01-1702 Keenan, M., Nicole, J. & Poojary, D. Top Catal (2019) 62: 351. https://doi.org/10.1007/s11244-018-1098-8

CFD simulation of particle deposition in exhaust gas treatment systems Dorian Holtz1 and Conrad Gierow1 and Robert Bank1 and Dirk Kadau2 and Flavio Soppelsa2 2

1 FVTR GmbH, Joachim-Jungius-Str. 9, 18059 Rostock, Germany Winterthur Gas & Diesel Ltd., Schützenstr. 1-3, 8401 Winterthur, Switzerland [email protected]

Abstract. Selective catalytic reduction (SCR) is an established technology for treatment of exhaust gases from combustion engines. However, a major drawback of typically used monolithic catalytic converters is their susceptibility regarding high particle loads. Clogging is a critical issue limiting space efficient design and operation time, particularly under high-dust operation from HFO fired large bore Diesel engines. In order to evaluate the clogging risk of exhaust gas treatment systems (EGTS) during the design process, a novel simulation approach for deposit formation has been developed and is presented in this paper. An Eulerian-Lagrangian treatment has been utilised for simulation of the particleladen exhaust gas flow, where drag, gravitation, buoyancy, turbulent dispersion, Saffman’s lift force as well as Brownian motion are considered as governing particle transport mechanisms. For the modelling of the gas flow (CFD), the monolithic catalyst is represented by a porous media formulation. The grid used for the CFD is extended by a collision subgrid providing more detailed geometry data for particle-wall contact consideration. The sticking probability of an impacting particle is calculated based on an elastic-adhesive contact approach. The simulation procedure has been developed using ANSYS Fluent. The underlying subgrid as wells as the particle-wall contact model have been implemented to the flow solver via multiple User-Defined Functions (UDF). The model development has been accompanied by extensive experimental investigations on a test rig with a 4-stroke single cylinder research engine of type 1VDS18/15-CR. The simulation procedure achieved good agreement with respect to experimental data of particle capture efficiency as well as reduction of open frontal area (OFA). Keywords: Exhaust Gas Treatment Systems, Particle Deposition, CFD, Discrete Particle Method, User-Defined Functions (UDF).

1

Introduction

Since 2016 IMO Tier III limits are valid for emission control areas (ECA). To be able to comply with these stricter limits, typically internal or external NOx reducing technologies are used. In order to develop an efficient Exhaust gas treatment system (EGTS) with a maximum performance index and a minimum of required space or pressure drop © Springer Fachmedien Wiesbaden GmbH, ein Teil von Springer Nature 2020 W. Siebenpfeiffer (Hrsg.), Heavy-Duty-, On- und Off-Highway-Motoren 2019, Proceedings, https://doi.org/10.1007/978-3-658-31371-5_6

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an optimal design has to be found. Long service intervals and ability to operate on residual fuel oils are challenging factors. The EGTS has to remain in full operation and sufficient efficiency even with changing fuel qualities. To guarantee operational capability a sophisticated knowledge about the effects of the exhaust gas and its components on the EGTS is necessary. Amongst others, this includes the knowledge of particle deposition in the particular EGT components to prevent system failures. Particularly under high-dust operation, clogging of SCR systems caused by adhering particles is a critical issue regarding system optimization in terms of volume and compactness. Only very few studies exist that are dealing with simulation of deposit formation and clogging in monolithic SCR catalysts in order to gain better understanding of the underlying processes. A comprehensive CFD model for deposit formation in EGTS has been published by [1]. The concept is based on the discrete particle method involving particle transport due to drag, turbulent dispersion, gravitation, buoyancy, Saffman’s lift force, Brownian motion and electrostatic forces. For the description of the deposit build-up process, a quasi-stationary treatment of the particle-wall contact based on a critical velocity formulation [2] was used. The authors concluded that inertial impaction, turbulent dispersion and gravitational settling are the dominating transport phenomena with respect to deposit formation. Moreover, they observed a trend to higher deposit formation rates at the inlet section of the catalyst. A similar procedure has been proposed in [3], where particle transport by thermophoresis is also considered. The simulation results indicate an increased particle capture efficiency for turbulent flow compared to laminar flow conditions. Another main finding was, that particles with larger Stokes number, and thus stronger dominated by inertial forces, tend to be deposited at the inlet section of the catalyst. The abovementioned studies [1,3] aimed for detailed insight and understanding of the deposit formation within individual channels of the brick. For this reason, these processes have been resolved to greater extend. Thus, in [1] and [3] small segments of a catalyst brick are considered only. However, this is not sufficient for the case of a complex EGTS structure containing multiple bricks or complicated flow phenomena brick upstream. Indeed, such conditions are typical for exhaust gas ducts of large marine Diesel engines. For such systems, an appropriate resolution of individual channels and the flow field upstream would lead to very high grid densities and a computation time which is not affordable. Thus, for engineering application a simplified approach is necessary. This study aims for the introduction of new simulation concept that has been designed for such purposes. The model development has been supported and accompanied by experimental investigations on a single cylinder test bed. Prior to discussion of the modelling approach, the experimental system and conditions considered shall be explained.

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Experimental Investigation

Experimental investigation of deposit formation in a catalytic converter has been conducted on a test rig with a 4-stroke single cylinder engine of type 1VDS18/15-CR. The engine is equipped with a common rail system allowing injection pressures up to

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1,600 bar. For the analysis of the particle number, particle size distribution and particle mass an AVL Particle Counter APC Advanced 489 and a TSI Engine Exhaust Particle Sizer Spectrometer TSI EEPS 3090 are used. The Particle Counter is already equipped with a two-stage dilution for the exhaust gas to allow measurements in exhaust gases with high particle loads. The APC detects the particle number concentration of nonvolatile particles (#/cm³) in the diluted sample gas. The EEPS allows detection of particle number as a function of particle size. This size distribution covers the range of 5.6 to 560 nm of particle diameter and has a time resolution of 10 Hz. The brick was installed in a housing, positioned in the high-pressure part of the exhaust gas system. The housing is made of stainless steel and equipped with an inflow and an outflow funnel. Upstream of the brick a small trap is installed inside the housing for deposit capturing and subsequent analysis. Upstream the housing the exhaust gas path is equipped with a bypass system allowing for reduction of the brick inflow velocities. The experimental setup described is depicted in Fig. 1.

Fig. 1. Schematic of the experimental setup and photography of the catalyst housing as well as bypass system

The engine was operated in low load using HFO complying with IFO 380 and with sulphur content of 2.27%. In order to achieve particle loads as for 2-stroke marine SCR applications, 2-stroke lube oil was added to the fuel. An overview with respect to the test conditions is given in Table 1. Table 1. Overview of the test conditions

Parameter Engine load Temperature upstream SCR Pressure upstream SCR Exhaust gas mass flow (SCR) Avg. inflow velocity SCR Operating time

Unit % °C bar kg/h m/s h

Value 37.5 358 1.94 100 1.14 5

We varied operating conditions and cell densities of the SCR bricks in order to obtain sufficient particle deposition. For very severe test conditions and high cell densities,

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after a test run of 5 hours operation a significant deposit build-up was observed particularly at the frontal faces of the brick and the channel inflow areas. A qualitative comparison between the clean monolithic catalyst prior to experiment and the fouled brick after experiment is depicted in the following Fig. 2.

Fig. 2. Clean monolithic catalyst prior to experiment (a), close-up of the frontal area after experiment (b) and deposits downstream the trap after experiment (c)

In addition to a distinct blackening that is obvious from Fig. 2b, a representative number of channels was partially or almost completely clogged at the frontal area or the inflow section, respectively. This is also evident from Fig. 2b. An optical post processing of the deposit structure by means of colour thresholding indicated an average loss in open frontal area (OFA) of approx. 48%. The minimum OFA reduction of an individual channel was about 36% and the maximum was found to be 89%. The massive deposit formation was also reflected by the measurement data of the TSI particle sizer illustrated in Fig. 3, showing a clear reduction in particle number over a broad range of size classes when comparing data upstream and downstream the SCR brick.

Fig. 3. Exemplary results of the TSI Particle Sizer upstream (us) and downstream (ds) the catalyst for test run no. 4

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Apparent from Fig. 3 the deposit build-up cannot be attributed to particles of a specific size class. However, the comparison of the measurement data upstream and downstream the brick indicates a small shift from smaller to larger particles on the pathway through the catalyst. According to the AVL Particle Counter the overall reduction of particle number amounted to approx. 32%. Experimental results of OFA reduction and particle number reduction are used for subsequent model validation.

3

Simulation Methodology

3.1

Basic Model Concept

The particle-laden exhaust gas flow through the monolithic catalytic converter is described by an Eulerian-Lagrangian treatment. Thus, the fluid phase is solved by a set of continuous balance equations for mass, momentum and energy. The SST-k-ω model is used for turbulence modelling. The particle freight is considered as discrete phase, represented by a defined number of particle parcels tracked through the simulation domain. Urea dosing is not considered in the current work, and only particles emitted by the engine are taken into account. These are assumed to be rigid, spherical and inert particles, following the density distribution of [4]. The particle injection is based on a random spatial distribution of the parcels at the domain inlet. As the volume fraction of the particle load is comparable small and only low thermal gradients exist, the influence of the particulate material to the continuous fluid phase can be neglected. For limitation of the computation effort, the particle tracking and subsequent deposit modelling are hence performed as a post processing step. All simulations have been performed using ANSYS Fluent. 3.2

Particle Model

As mentioned previously, a fixed number of particle parcels is tracked through the computation domain. Each parcel summarizes a defined number of particles which have the same properties in terms of size, mass and composition. The particle size distribution follows the measured data brick upstream in Fig. 3. Movement of each parcel is determined by a momentum balance on the particle scale which reads (1) Here, 𝑚𝑝 and 𝑣 𝑝 are the particle mass and velocity. The equation of motion contains a combined force of gravitation and buoyancy (𝐹⃗𝐺 ), drag (𝐹⃗𝐷 ) as well as Saffman’s lift force (𝐹⃗𝐿 ) and Brownian motion (𝐹⃗𝐵 ). Due to low thermal gradients in the systems considered, thermophoresis is neglected. The combined force of gravitation and buoyance is given as (2)

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Here, 𝜌𝑝 and 𝜌𝑓 are the densities of the particulate material and the fluid. Drag force is determined by (3) where 𝜇𝑓 denotes the dynamic fluid viscosity, 𝑅𝑒𝑝 is the particle Reynolds number and 𝑑𝑝 is the particle diameter. Besides, 𝐶𝐷 stands for the drag coefficient, calculated by the Stokes-Cunningham drag law for sub-micron particles. It must be noted, that the drag force in equation (3) also introduces the effect of turbulent dispersion to particle motion. This is achieved by usage of the instantaneous fluid velocity 𝑢 ⃗⃗𝑓 , rather than the mean fluid velocity obtained by Reynolds averaging. The instantaneous fluid velocity encompasses the mean value and a fluctuating component. The latter is derived by turbulence quantities using the Discrete Random Walk (DRW) model. Considering turbulent flows, particle motion is mostly governed by drag, turbulent dispersion and gravitation, though a certain effect can be attributed to other movement mechanisms in areas of low turbulence or laminar flow. This is especially valid for submicron particles which may be sensitive to weaker transport phenomena. In order to achieve a broad applicability of the simulation concept, shear-induced lift (Saffman’s lift force) as well as Brownian motion are taken into account. The former is given as

(4) Here, K represents a modelling constant, 𝜈𝑓 is the kinematic fluid velocity and 𝑑𝑖𝑗 is the deformation tensor. For more information about this formulation of Saffman’s lift force please refer to [5]. The Brownian motion is modelled based on a Gaussian white noise process with the spectral intensity (5) In equation (5) 𝛿𝑖𝑗 is the Kronecker delta function and 𝑆0 is defined as

(6) where 𝑘𝐵 denotes the Boltzmann constant and 𝑇𝑓 is the absolute fluid temperature. Furthermore, 𝐶𝑐 is the Cunningham correction, calculated by the free mean path λ and the particle diameter as follows: (7) Based on equation (5) the components of the Brownian force are determined by

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(8) Here, 𝜉𝑖 is a zero-mean, unit-variance-independent Gaussian random number and ∆𝑡 is the time step size. For further information on particle movement theory and its treatment in ANSYS Fluent please refer to [6]. 3.3

Deposit Model

A particle transported to a system wall may either adhere or reflect depending on the conditions of the particle-wall contact. In the used model, the sticking probability of a single particle 𝑝𝑝 is determined by a comparison of its kinetic energy prior to wall contact 𝐸𝑘𝑖𝑛 and the dissipation of energy 𝐸𝑐 through adhesive interaction during the contact event. If the kinetic energy is high enough to overcome the contact energy, the particle is rejected. Otherwise, it adheres. This is expressed by a particle sticking probability as follows: (9) The equation for calculation of the contact energy reads [7, 8]

(10) 𝑊𝑎 denotes the adhesive work, 𝑅 is the contact radius and 𝐸𝑒𝑓𝑓 is the effective Young’s modulus. The work of adhesion is determined by (11) where 𝛾𝑝 and 𝛾𝑤 are the surface free energies of the particle freight and the wall. The effective Young’s modulus is obtained by Young’s modulus and Poisson’s ratio of particles and boundary surface. (12) The used contact parameters are related to the composition of the particle freight observed during experiment. 3.4

Brick Model

In order to avoid extensive computation time for the fluid modelling because of high grid densities, the brick is treated as an anisotropic porous medium, as mentioned previously. ANSYS Fluent’s porous media model (PMM) appears as an additional source

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term within the fluid momentum balances. Considering the i th momentum balance, the source term reads (13) The permeability 𝛼𝑝𝑚 and the inertial loss coefficient 𝐶𝑝𝑚 are fitted to meet the pressure drop of the test bed application. It is assumed that the solid material of the brick is in thermal equilibrium with the fluid phase. In order to account for the higher gas velocities resulting from the reduction of the effective flow cross section, the physical velocity formulation of the PMM is used. Although the PMM allows a description of the global fluid flow through the catalytic converter, a more detailed description of the brick geometry is required for deposit modelling. Hence, a collision subgrid is superimposed to the CFD grid in order to provide necessary geometry data for particle-wall contact treatment. The schematic in the following Fig. 4 illustrates the basic idea of the subgrid approach.

Fig. 4. Explanation of the subgrid approach used for modelling or deposit formation inside the brick

The geometry of the brick is approximated by two grid functions, one aligned with the vertical axes of the cross sections and another aligned with the horizontal direction. However, the concept depicted in Fig. 4 is valid for both directions. The distinction between the free stream of a channel and the channel wall is achieved by the grid function 𝑝(𝑥). It determines the probability of a particle-wall contact in dependency of the current particle position and the particle size. Only in case of a contact event and prevailing particle velocities too low for overcoming the adhesion effect, deposit build-up is probable. This can be expressed by a global sticking propensity.

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(14) Both, the grid function for contact detection as well as the deposit model previously described in section 3.3 are introduced to ANSYS Fluent using multiple compiled UserDefined Functions (UDF). Additionally, an export routine has been implemented to ANSYS Fluent via UDF to allow for subsequent channel accurate analysis of the formed deposit structure as a part of the post processing.

4

Simulation Results and Discussion

The simulation domain encompasses a part of the high-pressure part of the exhaust gas duct including the bypass valve, the housing and the brick. The CAD model is pictured in Fig. 5.

Fig. 5. Simulation domain

Fig. 6 illustrates the simulated streamlines in the computation domain under conditions summarized in Table 1. Since particle tracking is performed as a post processing step without recoupling to the fluid flow, the streamlines are representative for the particle tracks. It is conspicuous, that high turbulence arises from the bypass throttle. The confuse flow structure is further supported by the deposit trap brick upstream on the lower right corner of the housing. This behaviour leads to irregular distribution of gas flow and particle load at the brick inlet. Furthermore, the confuse flow field brick upstream leads to numerous particle-wall contacts causing deposit formation especially at the channel inflow areas. The deposit distribution at the frontal area of the brick is also depicted in Fig. 6 in terms of the deposited mass. Following the inhomogeneous flow field and particle-load distribution brick upstream there is an uneven deposit build-up at the brick too. On the right-hand side of the catalyst, a high peak of deposit formation occurs due to a flow redirection resulting from the trap. Contrary, simulation indicates no deposit formation right behind the trap. Although a heterogeneous deposit structure

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was also apparent from experiment, this result is slightly deviating from test bed investigation where no deposit-free area was observed. It is believed that these uncertainties between simulation and test bed investigation arise due to transient flow phenomena that are not resolved by the steady state approach.

Fig. 6. Streamlines and deposited mass at the brick frontal area after 5 hours operation

In general, simulation achieved good agreement with experiments in relation to the particle capture efficiency. According to measurement data the reduction of particle number amounts to approx. 32%, while simulation delivers a value of 28%. Moreover, optical analysis of a brick after experiment reveals an average OFA reduction of 48%, as already explained under chapter 2. Simulation indicates a value of 55% and is thus also in good accordance to experiment. The simulation result is obtained by channel accurate analysis of the deposit structure based on the subgrid approach, see Fig. 7.

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Fig. 7. Channel accurate deposit structure at a representative section of the catalyst inflow area obtained by post processing of the CFD result

Each particle deposited or passing the brick during the CFD simulation is assigned to a specific channel of the monolithic catalyst according to the subgrid. This procedure is used for a more sophisticated analysis of deposit formation within the catalytic converter and the underlying influential parameters. Fig. 8 illustrates the integral spatial distribution of deposits along the brick axis. Here, the entire brick is decomposed into discrete zones of 2 mm length. The share of channel wall surface at the frontal brick area is not considered, as it is markedly higher than the observed maximum in Fig. 8.

Fig. 8. Integral particle deposit distribution along the brick axis

From Fig. 8 can be concluded that deposit formation within a channel mainly appears at the inflow section, what is in accordance to literature [1]. This behaviour is caused by high turbulence brick upstream and the complex inflow structure leading to a high number of particle-wall contacts by inertial impaction and turbulent dispersion. With decreasing turbulence along the path through the brick, the effect of turbulent dispersion decreases. Furthermore, the smoothed flow field inhibits a high amount of contacts by inertial impaction. Hence, the highest risk of clogging results at the frontal faces and the channel inflow area. Moreover, it can be concluded that critical deposit formation

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is mainly driven by inertial impaction and turbulent dispersion. A certain effect might also be attributed to gravitation. Weaker transport phenomena, namely shear-induced lift as well as Brownian motion show a minor influence under conditions considered. Subsequently, the influence of increased fluid velocities has been investigated. The related results are illustrated in Fig. 9.

Fig. 9. Relative number of channels with a specific cross section reduction per hour in a high cell density brick under varying gas velocities (superficial)

As Fig. 9 reveals, the fluid velocity shows a minor influence with respect to deposit formation rate. There is only a slight shift to higher OFA reduction with increasing velocities. This effect can be reasoned by a higher degree of turbulence leading to more particle-wall contacts. The velocity range considered, represents typical operating conditions of the real application. In this conjunction, the gas and respective particle velocities are not high enough to help more particles overcoming the contact event. However, under progressed fouling close to channel clogging and a very inhomogeneous flow field that behaviour might tip over at certain areas, as high gas velocities resulting from increased OFA reduction will help the particle freight to overcome the sticking limit. Rebounding will then most probable start with larger particles, since these provide a higher kinetic energy to get over the adhesive bonding. Resolution of such processes requires a recoupling between particle deposition, the PMM and the respective flow calculation. It can thus not be performed as a pure post processing step. Regardless, the basis of such an approach is already implemented and the complete integration is central part of future model improvements. However, this procedure is beyond the scope of the current paper. Finally, deposit formation in a catalyst with lower cell density was investigated and compared to results of the high cell density brick. The results are illustrated in Fig. 10. It shows that using a monolithic catalytic converter with much higher cell density under the same conditions will tend to produce critical deposits faster. Consequently, a more progressed OFA reduction appears after the same period of time or deposit induced failure occurs after shortened operation period. Vice versa, for the low cell density brick the deposit model predicts the same OFA reduction as observed in the experiments after

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approx. 13 hours, rather than five hours operation in case of the high cell density system. The major reason for increased deposit formation at higher cell densities can be seen in the increase of potential surface for deposit build-up and the decreased mean free path between a particle and the next channel boundary. For the design of SCR systems installed in high-dust exhaust gases from HFO fired marine diesel engines, knowledge of such circumstances is valuable in order to prevent fast clogging or at least extend the operation time to service intervals.

Fig. 10. Relative number of channels with a specific cross section reduction per hour in a brick with low and with high cell density

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Summary and Outlook

This study presented a new and efficient methodology for deposit modelling in CFD simulation of a monolithic catalytic converter. The procedure is based on an EulerianLagrangian approach and a porous media formulation. Simulation bas been conducted using ANSYS Fluent. Multiple UDFs have been integrated to the flow solver for the resolution of the deposit formation process inside the brick. As opposed to the porous media approach for the gas flow, the UDFs allow for channel accurate resolution of deposit build-up. For that, a collision subgrid has been superimposed to the CFD grid. Interactions between the particle freight and the brick walls are considered as elasticadhesive contacts. This procedure provides the basis for calculation of the particle sticking propensity. With the concept described, OFA reduction and clogging of individual channels can be estimated. The method is in overall good accordance to experimental investigation on a test bed. This is valid for both, the observed integral OFA reduction after certain operation time as well as particle capturing efficiency. Subsequently, the validated concept has been used for variation calculations with respect to inflow gas velocity and catalyst cell density. Future model improvements may comprise a more detailed consideration of the particle freight in relation to flight properties and sticking behaviour. On the other hand, a

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recoupling between the deposit build-up and the porous media as well as the flow structure could be addressed. Furthermore, cleaning effects. i.e. particle removal by incoming particles or flow peaks could be considered.

References 1. Heiredal, M. L., Jensen, A. D., Thøgersen, J. R., Frandsen, F. J., Friemann, J.-U.:PilotScale Investigation and CFD Modeling of Particle Deposition in Low-Dust Monolithic SCR DeNOX Catalysis. American Institute or Chemical Engineers Journal 59 (6), 19191933 (2013). 2. Heinl, E., Bohnet, M.: Calculation of particle-wall adhesion in horizontal gas-solid flow using CFD. Powder Technology 159, 95-104 (2005). 3. Feng, H., Wang, C., Huang, Y.: Particle deposition behaviors of monolithic De-NOx catalysts for selective catalytic reduction (SCR). Korean Journal of Chemical Engineering 34 (11), 2832-2839 (2017). 4. Maricq, M., Xu, N.: The effective density and fractal dimension of soot particles from premixed flames and motor vehicle exhaust. Journal of Aerosol Science 35, 1251-1274 (2004). 5. Li, A., Ahmadi, G.: Dispersion and Deposition of Spherical Particles from Point Sources in a Turbulent Channel Flow. Aerosol Science and Technology 16, 209-226 (1992). 6. Ansys, Inc.:ANSYS® CFD Premium Help System, Release 15.0. 7. Johnson, K. L., Pollok, H. M.: The role of adhesion in the impact of elastic spheres, Journal of Adhesion Science and Technology 8 (11), 1223-1332 (1994). 8. Reza Razmavar, A., Reza Malayeri, M.: A simplified model for deposition and removal of soot particles in an exhaust gas recirculation cooler. Journal of Enginnering for Gas Turbines and Power 138 (2016)

Variably honed cylinder liners, iron-based cast pistons and variably coated piston rings as PCU system for friction loss and TCO reduction Dr.-Ing. Daniel Hrdina (Speaker)1 Dipl.-Ing. Marco Maurizi2, Dipl.-Ing. Bartek Lemm3, Dipl.-Ing. Hakan Kahraman 4, Dipl.-Ing. Guilherme Soares de Faria5 (Co-Authors) 1

Head of PCU Technology, MAHLE GmbH, Pragstr. 26- 46, 70376 Stuttgart (Phone: 0711 501 – 12370, Mail: [email protected]) 2 Head of PCU Technology Liner, Pin, Systems 3 Global Product Expert Liner and Pin 4 Global Product Expert MHD Piston 5 Product Expert Piston Rings

The recently achieved agreement within the European Union introduces very challenging limits for CO2 emissions for heavy duty vehicles. Until 2025 a reduction of 15% in average and by 2030 of 30% has to be achieved. To ensure a low-CO2 and lowemission, up to locally emission-free, mobility, the drive concepts of the future will necessarily increase the variety with the introduction of electric drives and fuel cells but also keeping a good share of internal combustion engines (ICE) being powered by Diesel or even Hydrogen. Thus it is extremely important and necessary to continue to work on improvements in terms of friction reduction especially for ICE. Significant improvements in efficiency can be achieved through direct and indirect tribological system optimizations. Specifically, the functional friction reduced PCU in combination with low HTHS oils are considered in this work. Significant friction reductions (up to 5%) can be achieved by reducing the losses connected with the viscosity of the oils. Most of the OEMs already moved to low viscosity oils (HTHS ≤ 2.9 cP) and further reductions are planned. To enable this transition an upgrade of the single components in terms of wear and scuffing resistance is often necessary. Intelligent modifications of the components could also result in an additional contribution to the friction losses reduction. Due to the high price sensitivity and the demand for lower total cost of ownership, cost-reduced technology routes focus on providing added value where it is specifically needed in the respective applications. Liner: The new liner concept “STRATUM honing” developed at MAHLE considers all these boundary conditions.On the one hand, targeting the increase of robustness when using low viscosity oils (HTHS ≤ 2.9) and on the other hand, contributing directly to a further measurable friction losses reduction. The optimization and differentiation of the honing parameters along the piston stroke respects the different piston speed conditions, promoting hydrodynamic lubrication in the middle stroke by smoothing the running surface. Additionally, it targets the reduction of polished areas, typically caused © Springer Fachmedien Wiesbaden GmbH, ein Teil von Springer Nature 2020 W. Siebenpfeiffer (Hrsg.), Heavy-Duty-, On- und Off-Highway-Motoren 2019, Proceedings, https://doi.org/10.1007/978-3-658-31371-5_7

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by increased boundary friction close to the reversal points, by maintaining sufficient oil retaining properties in these areas. To validate results out of this development different tools were considered: Preliminary FEA was able to define potential friction benefits of ~15% of the liner’s friction portion; tribological rig tests for scuffing confirmed the validity of the honing parameter variation concept and different engine tests validate the concept and the calculated friction reduction, which corresponds to a reduced fuel consumption of approximately 0,2-0,3%. Piston: Depending on the application and its boundary conditions, aluminium pistons may reach their technological limits whereas steel pistons are well overengineered. Therefore, an additional material class comes into consideration. Nodular cast iron pistons offer advantages compared to steel pistons made of forging blanks. The material strength of a nodular cast iron piston covers the transition area between steel and aluminium pistons and thus avoids the aforementioned potential overengineering by using a forged steel piston. Additionally, a nodular cast iron piston enables a high degree of freedom in terms of the design features, where for example, friction reduction by low compression height can be achieved as known from steel pistons such as MAHLE’s Monotherm®. Furthermore, the number and differentiation of variants within an engine platform can be reduced by implementing a nodular cast iron piston for all power ratings. This reduces the number of different components in case of having aluminium and steel pistons alongside in the engine family. On the one hand, nodular cast iron pistons create an interesting alternative, which transfers performance advantages of the iron based material technology while enabling cost reduced synergies in engine design and on the other hand, implement the advantages of short compression heights and related friction and packaging benefits. Rings: The MAHLE DLC thick-at-tip (T@T) Top Ring – Technology is developed to increase robustness of the current DLC monolayer concept, providing increased durability exactly in the most critical area of the Top Ring, which is in the ring tips area. Such a solution avoids unnecessary coating application and thus cost by increasing the thickness only where it is really needed around the ring circumference. Besides the technical benefits this brings also some cost benefit in comparison to a thicker coating all over the contact face. New Generation V-Shaped DLC – This new oil ring is developed in order to cope with the challenging lower friction requirements and CO2 emissions, by improving land width's precision and circularity, allowing higher surface pressure even with very low tangential loads, translating into excellent lube oil consumption (LOC) performance right from the beginning by an improved running-in effect. In this context, DLC coating will increase even more the robustness over lifetime of this high surface pressure by keeping the land width lower for an extended period due to its superior wear resistance. In terms of cost, all of these performance benefits will be brought up with an almost cost neutral impact.

3

1

Liner STRATUM honing concept and roughness analysis

As friction reduction combined with long time performance in HD combustion engines is getting more important, the product aim was the development of a robust and friction reduced HD liner honing as an enabler for the use of low viscosity oils with HTHS ≤ 2.6cP. Thus, compared to today’s EUVI series honing specifications the cylinder liner has an optimized, differentiated surface roughness along the piston stroke. For reduced ring and liner wear, polished areas and scuffing tendency close to TDC / BDC a large Rvk as oil reservoir with better lubrication is used in zone 1. The reduced friction on middle stroke (zone 2) on the other hand is gained by a smooth honing in this area, which reduces the oil film thickness and improves the hydrodynamic lubrication of the fast running piston / ring package. By using low viscosity oils, STRATUM honing allows an increased usability of present ring surface technologies and can be applied on top-, middle- and low stop liners. Fig. 1 shows a section with corresponding tactile roughness measurements and WLI measurements of a STRATUM honed liner.

Fig. 1. STRATUM Honing sectional model with a) corresponding tactile roughness measurements and b) WLI measurements

White light interferometer (WLI) measurements with short coherence as light source can give additional two-dimensional information about the honed cross hatch structure. With this contact-free, nondestructive, optical measurement with µm resolution, further friction and wear influencing surface parameters such as pores, folded metal or groove distribution can be recorded and controlled.

2

FE Simulation

An HD simulation of Ø131mm series steel pistons with a corresponding ring package shows ~ 12% reduction in friction losses (FMEP) of the liner’s friction portion by using a smooth honing compared to conventional EUVI plateau honing. This corresponds to

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a fuel saving of approx. 0.25%. Fig. 2 shows the results of a calculation taking into account full load conditions at 1900 rpm. A reduction of Rk from 0.8 μm to 0.5 μm mainly reduces the friction of the oil control ring due to the reduced oil film thickness. Since only the middle stroke range has hydrodynamic lubrication, STRATUM honing takes into account only smooth honing in zone 2, in which there is no risk of seizure due to boundary lubrication.

Fig. 2. STRATUM Honing: FMEP calculation results for Ø131mm series PCU with steel piston

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Scuffing SRV rig tests and new E4025 material for increased wear resistance

SRV tests for detailed analysis of scuffing resistance were performed using a matrix of: ─ 2 Honing variants (Z1 "rough" and Z2 "smooth") ─ 2 oil viscosities (HTHS 3.5cP and 2.6cP) ─ 3 ring surface technologies (Cr-Ceramic, PVD, DLC) For a test duration of 240 minutes, a ring segment is oscillating under defined load, temperature and oil feed on a liner specimen. If the coefficient of friction exceeds a preassigned value (> 0.3), the test stops and its duration is counted as the time until scuffing, otherwise the test is completed positively. The main findings of the SRV tests are that Cr-Ceramic and PVD rings (as expected) scuff earlier with smooth honing and have higher wear than with rough honing. DLC coated rings showed no seizure at all, depending on the surface roughness they could work better with smooth honing. The test time to seizure shows that the influence of the oil viscosity appears to be greater than of the ring surface technology. Increased scuffing resistance is required primarily in the reversal areas and was confirmed for rougher honing of Z1 on the SRV test bench, as shown in Fig. 3.

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Fig. 3. a) Test rig SRV®5 from Optimol with main test parameters; b) results of scuffing sensitivity test for three different ring surface technologies, two liner surface roughness and 2 oils specifications

Additionally to the honing pattern and ring coating, the liner material also has a strong influence on wear resistance. Grey cast iron (GCI) having a bainitic matrix is known as market leading in regards of strength and hardness, but is also the most expensive material compared to pearlitic GCI. MAHLE therefore developed a similar pearlitic material (called E4025) with comparable strength and hardness of bainitic materials (UTS ≥ 400MPa, 250 – 320 HBW 30). The new material offers significant cost advantages by keeping similar mechanical properties and wear resistance in combination with different ring coatings as shown in Fig. 4.

Fig. 4. a) Wear test results of 3 different materials (1 bainitic and 2 pearlitic) with PVD ring technology; b) wear test results of 3 different materials (1 bainitic and 2 pearlitic) with DLC ring technology

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Floating liner results STRATUM vs EU VI honing

In order to quantify friction benefits, a floating liner engine (FLE) was used for measurements. Baseline was a standard EUVI honed cylinder liner, which was compared to a liner with STRATUM honing. Fig. 5 shows technical data of the single

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cylinder engine, which uses an aluminum piston with Ø=130mm, a stroke of 150mm and HTHS 2.6cP oil.

Fig. 5. a) FLE schematic concept; b) surface roughness measurements of the investigated liner

In operation, the engine map of the STRATUM honed liner shows a FMEP reduction of ~ 13% at high loads compared to EUVI honing (see Fig. 6). With increasing speed, the FMEP advantage also decreases slightly. This friction reduction corresponds very well to an SFC reduction of ~ 0.25%, as already calculated in the FEA.

Fig. 6. FMEP engine map of the STRATUM honed liner

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Engine results

Different engine runs with HTHS 2.9cP and 2.6cP oils with steel pistons and different ring coatings as well as running times up to 2000 h showed a completely visible honing pattern in all zones. Wear, blow-by and LOC were at a similar level to the baseline EUVI honing. Fig. 7 shows a representative example of a 1000-hour thermal shock run with DLC rings and HTHS 2.9cP oil.

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Fig. 7. Example of 2 liner scans after a 1000-hour thermal shock run with DLC rings and HTHS 2.9 oil.

The further development of the STRATUM honing concept with the aim of coping with a 3D shape in addition to the differentiated roughness along the stroke honing has already been initiated.

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MonoCast® piston concept

With the so called MonoCast® piston concept, MAHLE presents a piston with the known advantages of friction reduction due to the low compression height and a thermal expansion coefficient similar to the grey cast iron block with additional benefits of cost reduction. Nodular cast iron MonoCast® piston complements the piston portfolio of existing and well known aluminium and steel technologies to offer the best cost/benefit ratio for all application purposes. Generally, the MonoCast® piston is a one piece concept with an as cast closed cooling gallery, as shown in Fig. 8. A major benefit is the high degree of freedom in terms of combustion bowl and cooling gallery shape due to the flexibility of the core design that allows also undercuts or geometries, which are not feasible by machining processes.

Fig. 8. MonoCast® nodular cast iron piston design

Compared to existing steel piston concepts, the closest similarity can be found to the laser welded short compression height MonoLite® piston (Fig. 9). MonoLite® also

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offers a high degree of freedom in terms of the cooling gallery design due to the bowl rim section being welded into the lower part of the piston. However, such closed gallery piston concepts made of steel are usually based on a two piece raw part design. Therefore, an additional joining process is needed in the manufacturing route, adding process steps and cost. Furthermore, with machined cooling galleries, the design limits are given by the reachability with the tooling. This is also valid for the one piece Monotherm® piston design, as can be easily seen from Fig. 9 as well.

Fig. 9. Steel piston concepts; a) One piece Monotherm®, b) Two piece MonoLite® (laser welded)

Internal tests have also shown, that a closed cooling gallery has general benefits and can be optimized further, e.g. by specific flow guiding shapes. These measures improve the performance in terms of the cooling performance significantly and enable further potential for CO2 reduction. Generally, aluminium pistons still represent the majority of applications in the medium duty range. However, the development for new emission legislations or completely new platforms is usually also combined with increased demands in power output or higher peak cylinder pressures for better economy. Taking a view on current series engines and engine programs for medium duty in development, first steel piston applications are visible (Fig. 10).

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Fig. 10. Overview of engines and piston types for series and development

From Fig. 10, it is also obvious, that the range between 19-21MPa is showing a clear threshold, above which only steel solutions are applied. For aluminium, especially beyond 20MPa, it is important to pay a high attention to the overall combustion development and boundary conditions, as thermomechanical effects are causing a high load for the bowl rim and thus only small differences can lead to critical failures. So far, only few applications at 21MPa with an aluminium piston are in series and their combustion system is operating specifically at rather lower temperatures. Of course, reinforcement measures such as remelting can increase the thermomechanical loadability of aluminium pistons to a certain extent. However, comparing a high end aluminium witha forged steel piston shows a clear benefit in robustness for the iron based material and this is reflected in a strong trend towards steel also in the medium duty segment, following hereby heavy duty. Certainly, also the specific power plays a significant role. Most of the applications mentioned are operating at max. 35-40kW/L, which is also comparable to heavy duty engines. Unfortunately, the positive aspects of steel are also related to cost and therefore, the steel piston offering still good performance even at high firing pressures beyond 23MPa might appear over engineered for lower output applications. Exactly this area is targeted with the nodular cast iron piston, combining aspects of both worlds. Process wise, the manufacturing technology is more similar to aluminium pistons. By that, nodular cast iron combines benefits such as compression height reduction and therefore engine friction reduction up to 0,5%, and the design of the concept allows to achieve an almost identical weight like steel pistons as shown in Fig. 11.

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Fig. 11. Weight comparison of piston assembly. From left to right: TopCast Aluminium, Monotherm® steel piston and MonoCast® nodular cast iron

For higher specific targets, the existing steel solutions with the Monotherm® or even MonoLite® piston design cover the engine platform applications completely (see Fig. 12).

Fig. 12. From left to right: TopCast Aluminium, MonoCast ® nodular cast iron, Monotherm® and MonoLite® steel piston concepts

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7

Introduction DLC T@T coating for HDD Upper Compression Rings

The DLC coating has been used to improve the tribological conditions between the upper compression ring (UCR), low viscosity oils and the cylinder liner of modern Heavy Duty Diesel engines. Under harsh lubrication conditions the DLC coating can provide lower friction, higher wear and scuffing resistance, especially in combination with oils with lower viscosity (HTHS < 2.6cP) and engines with increased peak combustion pressure (PCP > 250bar). Usually for HDD applications 20µm minimum coating thickness is enough to withstand the whole engine life without the risk of coating wear-out, especially at the ring tips where the wear level is higher than along the ring periphery. In global platforms, different fuel quality, high level of EGR, excessive soot and abrasive particles can accelerate wear, resulting in DLC run out on the UCR tips region (see Fig. 13).

Fig. 13. DLC coated ring with base material exposure at the ring’s tip 0 and 360° position of contact face, after 1000h durability test

The exposure of base material by wear-out of the DLC coating does not necessarily generate functional problems. Nevertheless it should be avoided in order to prevent any possible scuffing occurrence, especially in combination with harsher environments with poor lubrication conditions, higher EGR rates and soot content, higher temperatures and peak cylinder pressure. The overall increase of the coating thickness would represent additional costs because of the higher deposition cycle time, needed to increase the thickness and further unnecessary challenges for the manufacturing process. MAHLE has developed the DLC Thick at the Tips coating (T@T), where the DLC thickness is increased only at the ring tips area up to 30µm minimum (see Fig. 14), avoiding reinforcements along the full top ring circumference where it is not needed, with a positive contribution to the costs.

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Fig. 14. DLC T@T Configuration

Concerning performance over lifetime, already performed engine durability tests with the same process technology for PVD coatings have demonstrated that the Thick at Tips approach achieves similar results in terms of lube oil consumption and Blow-by as a uniform coating thickness (Fig. 15). Similar tests in combination with the DLC coating are scheduled and expected to confirm these results.

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Fig. 15. Durability Test 1000h with Thick at Tips concept in a 13L Application

8

New Generation V-Shaped DLC Introduction

MAHLE V-Shaped oil control rings are widely used in HDD applications, coated with PVD or DLC. Even uncoated variants are in mass production where boundary conditions allow to utilize the wear properties of the nitrided stainless steel only. This technology is successfully tested up to almost 2 million kilometers in the field with very good results for oil consumption. For the new generation of HDD engines, the future CO2 requirements will be very challenging and in order to fulfill the targets, extreme low total ring pack tension is being proposed in combination with low friction coatings, smooth honing and low viscosity oils (Fig. 16).

Fig. 16. Reduction of Total Tangential load in HDD

14

To confirm the benefits in terms of friction, fuel savings and thus CO2by reducing the OCR tangential load by 50%, a variant was tested in a high precision friction test cell. The results show advantages (Fig. 17) in the friction mean effective pressure (FMEP) for the whole engine map, which translates into a reduction of 3,2 g/km of CO2 representing 0,37% in fuel savings. The calculation is based on an average of nine HDD Long Haul cycles including VECTO, Heavy Heavy-Duty-Diesel Truck (HHDDT) Cruise and Transient and at different trailer load scenarios, respectively.

Fig. 17. Potential CO2 saving by reducing OCR FT demonstrated at Friction Test Cell

Especially for the oil control ring, which plays the most important role to keep the lubricant oil consumption (LOC) in target, a very low tangential load means that the contact pressure must be kept high, around 2.0MPa, in order to maintain a healthy system concerning LOC. The new designed V-Shape oil control ring was especially developed for the best interaction between ring and cylinder liner. It maximizes contact pressure by maintaining very precise and low land width values of ~ 0,10mm. The improved roundness also ensures that the ring has precise contact in both lands around the circumference. Engine tests have demonstrated that the running in, especially for the coated OCRs is significantly improved as stable values of LOC are already reached during the very first hours of test (Fig. 18).

15

Fig. 18. New Gen. V-shape performance in Run-in and Durability Cycle with 13L Engine

Regarding cost, the new V-Shape oil control ring design has been developed including an optimized process route. This feature allows tolerance improvements without cost increase, contributing to better performance and maintenance of the total cost of ownership.

9

Conclusion

In order to meet the future limits for CO2 emissions, MAHLE has developed robust PCU solutions that offer lowest friction properties and enable the use of low viscosity oils without detrimental performance in terms of wear or durability. Additionally, the system approach allows to add value in technology where needed and offer cost improved solutions where possible. STRATUM liners with differentiated honing patterns show friction advantages of 12-15% in FE simulations and motor-validated friction measurements compared to today's EUVI honings with almost neutral cost profile. This translates into approx. 0,25% fuel consumption improvement. In order to achieve the best TCO performance, the honing is a crucial part of the overall tribological system in the engine offering even further potential for friction reduction. New developments focus on the right honing parameters and liner material in conjunction with the ring technology, oil viscosity and piston installation clearance. On piston side, MonoCast® pistons offer cost reduction potential compared to various steel pistons such as MonoTherm® or even MonoLite®, still contributing up to 0,5% fuel consumption reduction and engine performance for certain load levels. The nodular cast iron technology specifically fills the technical gap between aluminium and steel pistons by combining the key aspects of each technology.

16

DLC T@T upper compression rings provide the right technological feature to locate the needed robustness in the area of highest loads and wear requirements. Especially at unfavorable conditions in terms of lubrication or peak cylinder pressures, while improving wear and scuffing resistance. For the oil control ring, the new generation of the V-shape design enables the reduction of the tangential load by simultaneously keeping low oil consumption and low emissions. The key factor is an improved manufacturing process technology to achieve the right performance even in a green engine or after significantly reduced run-in time. This enables a reduction of tangential load contributing with up to 0,37% fuel consumption reduction as confirmed by bench test on a representative 13L HD engine.

Parameter study of the appearance and allocation of small oil aerosol particles at the piston, piston ring and cylinder liner surfaces in the engine blow-by and the evaluation of countermeasures Magnus Lukas Lorenz Daimler AG

Prof. Dr. sc. techn. Thomas Koch Institute for Internal Combustion Engines (IFKM), KIT Karlsruhe Institute of Technology Springer Heidelberg, Tiergartenstr. 17, 69121 Heidelberg, Germany [email protected]

Abstract. In addition to particle emissions from the aftertreatment box, a further known source is crankcase ventilation. Using a modern separation system, the particle mass limit is adhered to, thus enabling open crankcase ventilation into the atmosphere in heavy duty combustion engines. Since closed crankcase ventilation influences the powertrain negatively due to compressor coking, contamination and high ash emission, the open crankcase configuration is very common. The Euro VI legislation introduced a total particle number limit. This brought the focus onto smaller particles stemming from the separation system, in addition to the bigger particles that are of concern when meeting the particle mass limit. Efforts to reduce the engine particle raw emissions prior to the separation system have been successful, as has the improvement of the separation system itself. This paper focuses on the drivetrain as the suspected main origin of micron and submicron oil aerosol particles. A single cylinder engine with real-time piston temperature measurement equipment enables further insights into parameters for particle formation on the one hand, and the analysis of different countermeasures on the other. This article points out the engine oil temperature, the engine load and the piston oil nozzle flow rate as main formation parameters. Additionally, the connection between specific oil aerosol diameter modes to the combustion process and the piston surface temperatures is shown. Drivetrain tribology and heat management are highlighted as major countermeasure categories, and several ideas for reducing emissions are discussed. Finally, the influence on emission reduction of a piston filled with a heat transfer medium is demonstrated. Keywords: oil aerosol; particle emission; crankcase ventilation

© Springer Fachmedien Wiesbaden GmbH, ein Teil von Springer Nature 2020 W. Siebenpfeiffer (Hrsg.), Heavy-Duty-, On- und Off-Highway-Motoren 2019, Proceedings, https://doi.org/10.1007/978-3-658-31371-5_8

2

1

Introduction

Reducing the emissions of commercial vehicle diesel engines is one of the most important and most challenging tasks for original engine manufacturers. As a result of combustion, blow-by passes the piston rings in every internal combustion engine. To avoid increasing crankcase pressure, the system needs to be vented. There are two basic options: open crankcase ventilation and closed crankcase ventilation. Since the blowby is a mixture of air, soot, metal abrasion, oil, oil aerosol and water, it needs to be cleaned before it can be vented it to the environment or into the intake air path. For separation in internal combustion engines, two solutions are most common and exhibit a proven 99.9% efficiency down to a specific particle diameter. Both utilize the inertia of bigger particles to separate them from the gas flow. These systems can be divided into actively driven systems and passive systems. The passive systems use a specific configuration of baffle plates and are in general able to separate particles bigger than 3 µm in diameter. The active systems use the engine oil flow to power a centrifuge, for example, and are in general able to separate particles bigger than 1 µm in diameter. With these systems, the particle mass (PM [mg/kWh]) limit of different exhaust emission legislations is safely met. Since closed crankcase ventilation leads to negative consequences for the powertrain, such as compressor coking, contamination and higher ash emission, open crankcase configurations are very common [1]. The implementation of the total particle number limit of 8𝑥1011 𝑝𝑎𝑟𝑡𝑖𝑐𝑙𝑒𝑠/𝑘𝑊ℎ in the WHSC test with EURO VI legislation measured down to 0.021 µm in diameter added a new dimension to this [2]. Starting with PEMS stage e in 2021, this limit will also apply to real driving emissions (RDE) measurements. Besides particle emissions from aftertreatment boxes, another significant source is open crankcase ventilation. Therefore, the effort to reduce the particle mass using separation systems must be expanded to reduce the particle number as well. In additional to the efforts to achieve more efficient separators, this paper will focus on the origin of small oil aerosols (see Fig. 1)

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Fig. 1. Schematic spectrum of oil aerosols for internal combustion engines and definition of the particle size in focus

Particles smaller than 3 µm but bigger than 0.3 µm are very interesting because they are still relevant concerning particle mass while also being relevant to the particle number. Investigations indicated that the main source of particles in the size range of concern is the drive train, and oil evaporation is the main formation mechanism [3,4,5]. Therefore, the next step is to set up a single cylinder engine with real-time piston temperature measurement equipment on a test bench. The single cylinder excludes the influence of variations between the cylinders and of the turbocharger, compared to a full engine. This assembly is used to measure the influence of different operation parameters. It also enables a variety of hardware tests. The idea is to gain insight into particle formation due to oil evaporation.

2

Measurement setup

2.1

Particle counter

The particles in the size range in question with diameters ranging from 0.3 and 3 µm are counted using an optical particle counter. The selected counter is Promo 2000HP by PALAS, which features a particle diameter measurement range of 0.2 to 10 µm, a measurement frequency of 1 Hz and a maximum concentration of 106 𝑝𝑎𝑟𝑡𝑖𝑐𝑙𝑒𝑠/𝑐𝑚3. A dilution system is necessary because the single cylinder engine has no separation system, on account of which the expected particle concentration is higher than the maximum concentration of the measuring unit. Therefore, a two-stage dilution process invented and verified by the KIT is used. The conversion between particle number and particle mass concentration is calculated using the engine oil density and a specific form factor for the particle shape.

4

2.2

Investigation object

Table 1 shows the key figures of the highly compressed single cylinder engine. Table 1. Key figures of highly compressed single cylinder engine

Displacement Stroke Bore Peak cylinder pressure Piston material

2.13 liters 156 mm 132 mm 230 bar and more Steel

The single cylinder is set up with a high pressure indication and real-time piston temperature measurement equipment (see Fig. 2).

Fig. 2. Piston real-time temperature measurement setup and measuring position overview

On the basis of different temperature measurement positions at 1-mm depth on oilexposed piston surfaces, the areas where the engine oil most likely evaporates are investigated. These areas are inside the piston cooling channel and below the piston bowl. All maximum temperature values are transmitted from the chip to the antenna when the piston is at bottom dead center. In this setup, the oil reservoir is next to the engine to avoid splashing and to enhance oil temperature and pressure variation. Furthermore, the oil supply of the piston oil nozzle is separated from the oil circuit to manipulate the flow rate.

5

The extraction point for the dilution system is in the middle of the blow-by flow at the top of the crankcase, next to the cylinder head. The crankcase pressure pulsation at one cycle in this engine is up to 100 mbar, caused by the piston movement. However, the optical particle count is very sensitive to pulsation, especially if the flow direction turns. Therefore, a resonator was developed using the Helmholtz formula (1) [3]. 𝑓=

𝑐

𝜋∗𝑟²

2∗ 𝜋 √𝑉∗(𝑙+𝜋∗ ) 𝑟 2

(1)

By applying this theory, several configurations of tubes and barrels were designed and tested. The best solution is able to reduce the pulsation by more than 95% (see Fig. 3).

Fig. 3. Reduced crankcase pressure pulsation using a specially designed Helmholtz resonator at 900 rpm and 200 Nm output torque

Furthermore, the dilution system also reduces the flow pulsation in the measuring volume. Using the resonator and the dilution system, reproducible measurements with a coincidence of less than 5% are possible. A thermodynamic reference test shows no significant change in the engine performance values caused by the resonator.

6

3

Parameters for the formation of small-sized oil aerosol particles

3.1

Operation conditions

A complete engine acceptance and performance test with defined boundary conditions is conducted. To gain insight into the effect of different parameters, the screen design of experience (DoE) method is applied (see Fig. 4). In this figure, all values are plotted relative to the total particle emission at 1100 rpm and an output torque of 200 Nm. The relative total particle number concentration is defined in formula (2). max=3 𝜇𝑚

𝑅𝑒𝑙𝑎𝑡𝑖𝑣𝑒 𝑡𝑜𝑡𝑎𝑙 𝑝𝑎𝑟𝑡𝑖𝑐𝑙𝑒 𝑛𝑢𝑚𝑏𝑒𝑟 𝑐𝑜𝑛𝑐𝑒𝑛𝑡𝑟𝑎𝑡𝑖𝑜𝑛 =

∑min=0.3 𝜇𝑚 𝑝𝑎𝑟𝑡𝑖𝑐𝑙𝑒 𝑛𝑢𝑚𝑏𝑒𝑟 𝑐𝑜𝑛𝑐𝑒𝑛𝑡𝑟𝑎𝑡𝑖𝑜𝑛 𝑟𝑒𝑓𝑒𝑟𝑒𝑛𝑐𝑒 𝑡𝑜𝑡𝑎𝑙 𝑝𝑎𝑟𝑡𝑖𝑐𝑙𝑒 𝑛𝑢𝑚𝑏𝑒𝑟 𝑐𝑜𝑛𝑐𝑒𝑛𝑡𝑟𝑎𝑡𝑖𝑜𝑛

(2)

The relative total particle mass concentration is defined in formula (3). max=3 𝜇𝑚

𝑅𝑒𝑙𝑎𝑡𝑖𝑣𝑒 𝑡𝑜𝑡𝑎𝑙 𝑝𝑎𝑟𝑡𝑖𝑐𝑙𝑒 𝑚𝑎𝑠𝑠 𝑐𝑜𝑛𝑐𝑒𝑛𝑡𝑟𝑎𝑡𝑖𝑜𝑛 =

∑min=0.3 𝜇𝑚 𝑝𝑎𝑟𝑡𝑖𝑐𝑙𝑒 𝑚𝑎𝑠𝑠 𝑐𝑜𝑛𝑐𝑒𝑛𝑡𝑟𝑎𝑡𝑖𝑜𝑛 𝑟𝑒𝑓𝑒𝑟𝑒𝑛𝑐𝑒 𝑡𝑜𝑡𝑎𝑙 𝑝𝑎𝑟𝑡𝑖𝑐𝑙𝑒 𝑚𝑎𝑠𝑠 𝑐𝑜𝑛𝑐𝑒𝑛𝑡𝑟𝑎𝑡𝑖𝑜𝑛

(3)

The hottest point inside the piston cooling channel is selected as the piston temperature.

Fig. 4. Parameters for the formation of small-sized oil aerosol particles in a screening DoE at 1100 rpm and 200 Nm for constant blow-by flow measurements

The above figure can be divided into two sections. The first two parameters change the blow-by flow as they vary. The remaining parameters were investigated with a constant blow-by flow. When the engine load is changed from half to full load, the cylinder peak pressure, the blow-by and the component temperatures increase. In the cooling channel, the temperature is increased by 27%. As a result, the total particle number concentration is

7

more than doubled and the total particle mass concentration is increased by 47%. More engine oil is evaporated because the oil nozzle flow rate only depends on the engine speed in a real engine. In addition, the higher blow-by flow carries away more particles. Therefore, the engine load is a major influencing factor on the particle emission. The influence of the engine speed at a constant output torque is minor compared to the engine load. There is no change in the piston temperature and only minor changes in the total particle number and mass concentration. The root cause is a balance between higher peak pressures at higher engine speeds for the same output torque and an increased blow-by flow. At the single-cylinder test bench, the coolant temperature is independent of the engine oil temperature. By increasing the coolant temperature, the total particle number and mass concentration is also increased. This is a result of the higher component surface temperatures with decreasing heat transfer to the coolant. The slight change in the piston temperature of 3% shows no influence on the engine blow-by. The engine oil temperature is one of the major influencing parameters. Increasing it from 70 °C to 115 °C results in the total particle number concentration rising by 77% and the total particle mass concentration more than doubling. The particle formation drivers are the 35 K hotter piston cooling surface temperature and the smaller gap between the engine oil temperature and its evaporation temperature. At the single-cylinder test bench, the oil pressure is independent of the oil nozzle flow. Increasing the oil pressure under constant boundary conditions lowers the component surface temperatures and therefore the total particle emission. However, the cooling effect is very small within the applied pressure range. The oil nozzle flow is the third major influencing parameter on the total particle emission. Decreasing the oil nozzle flow reduces the cooling oil amount on the surfaces. Consequently, the surface temperatures increase and more oil evaporates. Halving the oil nozzle flow results in a 62% higher total particle number concentration and 66% higher total particle mass concentration. Increasing the intake air temperature in the cylinder results in higher particle emissions due to higher component surface temperatures. However, the influence is very small compared to the main drivers. In the direct comparison, the engine load is the main driver of the total particle number concentration. The engine oil temperature is the major influencing parameter on the total engine mass concentration. The oil nozzle flow is an important parameter as well. 3.2

Influence theory

The next step is a deeper analysis of the different levers affecting the particle number and mass concentration. For purposes of comparison, the oil aerosol diameter spectrum under consideration is investigated (see Fig. 5).

8

Fig. 5. Levers acting on particle number concentration formation in different oil aerosol modes at 1100 rpm

When using the two modes defined in Fig. 5, a significant difference in the two levers is outlined. The number of oil aerosol particles between 0.2 and 0.5 µm in diameter increases 6-fold when varying from half to full load. The variation of the engine oil temperature has only a minor influence on this. In contrast, the oil aerosol particle concentration of particles 0.5 and 0.8 µm in diameter increase significantly as the engine oil temperature varies. By plotting the total particle number concentration of each mode for the two engine loads and the six different engine oil temperatures, an increasing factor can be defined for easier comparability. The increasing factor for mode 1 particles is nearly three times higher than the factor for mode 2 particles, between half and full load. This is a strong indicator that mode 1 particles are connected to the combustion process and the resulting blow-by flow. Using the same factor logic for the engine oil temperature variation, mode 2 particles change 50% more than mode 1 particles. Therefore, it is hypothesized that mode 2 particles are more strongly associated with piston cooling. The correlation between diameter and volume is: 𝑉 ~ 𝑑³

(4)

The engine oil temperature is the bigger lever acting on the total particle mass concentration because of its higher influence on bigger particles (compare Fig. 4).

4

Countermeasures

4.1

Engine raw emission reduction

In addition to the development of more efficient separation systems, this paper focuses on reducing the engine raw emission. As a result of the parameter study in chapter 3, the levers are divided into measures influencing the tribology system and measures dealing with heat management at the piston (see Fig. 6).

9

Fig. 6. Clustered overview of engine particle raw emission reduction measures at the drivetrain components

With regard to heat management, the distribution of heat flux is a lever to influence the formation of small-sized oil aerosol particles. For example, a separate oil temperature level for the piston oil nozzles or a different heat transfer rate through the cylinder liner could be designed. Another idea would be to use insulating material to maintain the combustion temperature in the exhaust gas and to thereby lower the piston temperature on the oil side. Furthermore, the temperature distribution uniformity on the piston oil side could be improved by using a heat transfer medium in a closed piston cooling channel. Another lever is the tribology system consisting of the piston, piston rings and cylinder liner. Every measure to reduce the blow-by flow needs to be effective. At the piston, a smaller installation clearance is one possible measure. Moreover, increasing the tension of the oil control ring or adding another piston ring would lower the blowby flow. A different cylinder liner deformation behavior could also influence the blowby. All measures reducing the oil layer thickness on the cylinder liner are interesting as well. This could be achieved by changing the cylinder liner honing, the tension of the piston rings or the piston ring design.

10

This is only a small selection of the possible options to influence the two levers. However, the whole system needs to be taken into account because individual measures also influence other performance parameters, such as friction power. 4.2

Results of piston with heat transfer liquid

One countermeasure is the idea to increase temperature uniformity on the piston oil side using a heat transfer medium inside a closed piston cooling channel. This setup is realized using a specific oil inside of the piston (see Fig. 7).

Fig. 7. Influence of increased piston temperature uniformity on the oil side, realized with a heat transfer medium at 1100 rpm and an output torque of 200 Nm

Using the heat transfer medium, the piston oil side temperature is increased by 17 K and the blow-by is reduced by 39%. This results in 21% less total particle concentration and 24% less total particle mass concentration. The reduced blow-by flow is the main driver of the improvement. Especially significant is the decrease in the number of

11

mode 1 particles (compare Fig. 5). Comparative improvements could also be found for other speed and load combinations. Therefore, this is an effective measure to reduce the engine raw particle emission.

5

Summary and outlook

In this paper, different formation parameters for micron and submicron oil aerosol particles were investigated. Using a single cylinder engine with independent oil and coolant circuits in combination with real-time piston temperature measurement equipment, several boundary conditions could be varied. As described, the challenge for the optical particle counter was crankcase pressure pulsation caused by the piston movement. A special resonator using the Helmholtz formula was designed and tested and the pulsation was nearly erased. As shown in this paper, the following conclusions can be drawn: The main drivers of the particle formation process are the engine load, the engine oil temperature and the piston oil nozzle flow rate. By dividing the particle diameters into two modes, a strong link between very small particles and the combustion process was outlined. The second, more mass-relevant mode could be linked to piston oil surface temperature management. Based on these drivers, an overview of the countermeasures concentrating on the tribology system and heat management was presented. Various measures related to the piston, piston rings and cylinder liner were discussed. Different reduction strategies that rely on heat distribution, insulation and local temperatures were shown. Finally, a piston filled with a heat transfer medium in the closed cooling channel reduced the mass emission up to 24% due to decreased blow-by flow. Further investigations will focus on the continuing reduction of engine raw emission. Ultimately, all improvement potentials will need to be confirmed for an engine under full load to ensure functional and reliability validation.

References 1. E. Greuter, S. Zima, W. Hoffmann, “Motorschäden: Schäden an Verbrennungsmotoren und deren Ursachen” Vogel Communications Group GmbH & Co. KG, ISBN 9783834331939, 2011. 2. Economic Commission for Europe of the United Nations (UN/ECE), “Regulation No 49 – Uniform provisions concerning the measures to be taken against the emission of gaseous and particulate pollutants from compression-ignition engines and positive ignition engines for use in vehicles,” Brussels, 2013. 3. D. Ueberschaer, D. Schulmeyer, J.S. Parr, et al., “Messung des Ölverbrauchs von Abgasturbolader und Luftpresser,” ATZ Extra Nr. 19: 2014 doi: 10.1365/s35778-0141244-2 4. H. Bockhorn, “Soot Formation in Combustion,” Springer Verlag, Berlin, Heidelberg, New York 1994.

12 5. S. Wurster, J. Meyer, G. Kasper, “On the relationship of drop entrainment with bubble formation rates in oil mist filters,” Separation and Purification Technology 179 (2017) 542–549.

Application of virtual sensors for stress-related design and operation-specific lifetime prognosis Martin Diesch, M.Sc.1 Dr.-Ing., Thomas Bubolz2, Dr.-Ing., Martin Dazer 1 Dipl.-Ing. Kevin Lucan1, Prof. Dr.-Ing. Bernd Bertsche1 1

University of Stuttgart – Institute of Machine Components, Stuttgart, Germany [email protected] 2 MTU Friedrichshafen GmbH, Maybachpl. 1, 88045 Friedrichshafen, Germany [email protected]

Abstract. In this article, a procedure for the application of virtual sensors for load-related sizing and lifetime prognosis of components of large diesel engines is presented. Due to high costs for testing to assure the lifetime of a large diesel engine, the knowledge about the load is of utmost importance. The approach consists of several steps beginning with the identification of the relevant loads on specific components and ends with either a prognosis of the remaining useful life (RUL) or the use of load spectrums for the development. Virtual sensors are models that use given input variables to estimate an undetected output variable of the system. Thus the load of the system during the installation is determined. The load-time signal can be used for both a stress appropriate design as well as a prediction of the remaining useful life. The necessary steps for deriving requirements for the development process as well as the lifetime prognosis are described. Keywords: virtual sensors, reliability, development methods, PHM

1

Introduction

The development of a reliable product is one of the main goals in mechanical engineering. The reliability is defined as “the probability that a product does not fail under given functional und environmental conditions during a defined period of time” (Bertsche). Due to increasing cost pressure and lightweight design of components and products, safety margins are decreasing. In order to avoid unforeseen failures the knowledge of the relationship between the strength and the stress or load on the component, which is commonly known as stress-strength interference (SSI), is of utmost importance. The probability of failure is determined by linking the stress and strength of the component. Geometry, the manufacturing process and the material of the part influence the strength, whereas the load is determined by environmental conditions, the user and interactions within the system. Due to the variation of the individual parameters, the exemplary distribution of load and strength can be seen in Fig. 1. Depending on the knowledge of

© Springer Fachmedien Wiesbaden GmbH, ein Teil von Springer Nature 2020 W. Siebenpfeiffer (Hrsg.), Heavy-Duty-, On- und Off-Highway-Motoren 2019, Proceedings, https://doi.org/10.1007/978-3-658-31371-5_9

2

the operating conditions the actual and assumed load distribution may differ considerably from each other. This leads either to oversized components or to increasing failed parts. [1]

Fig. 1. Strength-stress-interference (SSI) [1]

Knowing the actual load of a component is not only important for the development of new products, but can also be used for a Prognostics and Health Management (PHM) approach. PHM describes a field of research, which focuses on estimating the health of a system or part to predict the health state and thus leading to an estimation of the remaining useful life (RUL). The information about the RUL can either be used for maintenance or an active change of the operational characteristics of the system and thus resulting in an increasing RUL. [10] It is clear that knowing the load is a key-factor for development and RUL estimation of a product. This is why this paper focuses on an approach to determine relevant loads using virtual sensors. The procedure includes the selection of failure-relevant loads, their modelling by means of virtual sensors and the corresponding further processing of the data for development and RUL estimation.

2

Approach

The approach consists of 11 or 12 steps, depending on the result. At the beginning the considered system has to be analysed. In the first step a boundary of the system has to be defined, over which mass, energy and information flow are transmitted. This way all input and output parameters are described, which characterize the function of the system. With the physical values of these parameters the state as well as changes to the system can be described. Depending on the complexity of the system it might be necessary to further divide the main function into sub-functions.

3

Fig. 2. Approach for stress appropriate construction and operation-specific lifetime prognosis

In the next step a Failure Mode and Effect Analysis (FMEA) is used to determine possible failure modes, the cause of the corresponding failures and its impact on the system. In general, a FMEA is divided into the steps system analysis, functional analysis, failure analysis, risk assessment and optimization. Thereby the focus is on the risk assessment and the failure analysis, as each parameter from step one can be assigned to a cause of failure. Thus each component of the system is characterized by the cause of failure, the damaged location and a parameter, causing the failure. This results in a large number of failures, which can be described using the influence parameter. These parameters can be used in a later stage in both the development of new components and the prediction of the RUL. Since not every parameter and the associated failure are relevant for later use, they are prioritized on the basis of the risk assessment in the FMEA. The prioritisation marks the end of the weak-point analysis. []

4

The next step is to check to which extent the selected parameter is or can be recorded. Is the parameter non measured, a virtual sensor approach can be used to model the parameter. Thereby virtual sensors describe a model, which is implemented in a software structure and uses available input variables to determine a non-measured output variable. In the case of this approach the output variable is referred as the load on the specific part. This way a load-time-signal during the operation of the system can be obtained. In the step “field data analysis” the load-time-signal is further processed, as usually the signal cannot be saved on an engine control unit (ECU). Thus to reduce the amount of data, the signal can be discretised. Depending on the failure mechanism a counting method has to be selected, as these methods greatly reduce the information of the time signal and only by prior selection is it possible to draw conclusions about the failure mechanism. If a suitable counting method was selected, a load spectrum can now be determined. [4,5] After the load spectrum for the selected parameter is determined, two paths of actions are possible. The first path describes the evaluation of a health state of the component or system and prognosis is made according to future loads. First the load spectra of the specific customer is used and the stress on the system is calculated. Then the stress is used for the estimation of the health state, whereby a lifetime model is required. The prediction of the future load is applied for a prognosis according to the present health state. Either the load spectra are collected in a data base and are further evaluated, if there is a certain amount available. The necessary amount depends on the total number of systems in operation. Also it has to be analysed whether the system is used for different applications. As not every load spectra of each costumer is taken into account for the development of the component a representative load spectrum has to be determined using the data. This one load spectrum can then be used for definition of requirements and testing in the development phase. Therefore certain quantiles of the load spectra are determined, which describe a certain percentage of customers. The representative load spectrum is then selected from these quantiles and used in the development. In the following chapters The approach is now described using an exhaust turbocharger of a large diesel engine.

3

Weak-point analysis

To determine the input and output parameters of an exhaust turbocharger a system boundary and all energy, mass and information flows are defined according to Pahl et. al 2007. The turbocharger thereby consist of the three subgroups turbine, bearing, compressor, see Fig. 3. The different flow types between the subgroups are also characterized, as there are a lot interactions between the parts, which can affect the failure mechanisms. The subgroups are characterized by the function of the specific group. So exhaust is provided to the turbine, which extracts energy from the exhaust and provides it to the compressor over the shaft.

5

Fig. 3. Input and output variables and interactions between the subsystems of a turbocharger

In the next step a FMEA is performed, which describes all possible failure mechanisms of the system. Information can be added to the FMEA as for example the location of the failure. Is the influencing parameter and the failure mechanism equal to each other the same load spectra can be used. As for the turbocharger relevant failure mechanisms are thermo-mechanical fatigue (TMF), structural-mechanical fatigue (LCF) and wear. TMF is the cause of failure for the turbine housing whereas LCF is relevant for the compressor wheel. Thereby the varying temperature of the turbine housing results in restriction of the expansion, thus damaging the component. 3.1

Virtual sensors

In general virtual sensors are models, which predict a none-measured variable with available variables. The application of these models has the following reasons or advantages:      

High costs of hardware sensors Degradation or damage of hardware sensors during operation No time delay of measurement (e.g. exhaust emission measurement) Additional information and thus better monitoring Fault detection Damage of a component due to the attachment of the hardware sensor

The modelling generally can be divided into the five steps of data selection, data filtering, model structure, model estimation and model validation. First it is necessary to have experimental data, which contains the output variable. Before a model is selected outliers and other implausible measurements have to be removed from the data set. A Principal Component Analysis (PCA) with Jolliffe parameters or a Hampel identifier

6

are commonly used methods to get rid of outliers. The sampling rates of the input and output variables of the virtual sensors have also taken into account. In order to use the output variables in further steps of the approach, the sampling frequency has to be in a range through which relevant effects can be observed. [2,3] The third step of the modelling is to select a model structure. The model structure can be categorized as black-box, grey-box and white-box-models, Fig. 4. Black-Box models use only experimental data to map the input variables of the model to the output variable. Thereby no knowledge about the behaviour of the system is necessary. Neuronal Networks are one example of black-box models. Whereas a modelling of physical behaviour of the system is called white-box-modelling. For these kind of models a deep understanding of the system is necessary. Due to the high complexity of many technical systems, modelling is very time-consuming.

Fig. 4. Black-, white-, grey-box models [2]

The connection of a white-box and black-box model is described as grey-box model. These models tend to combine the advantages of both approaches. As black-box models can only learn the system behaviour due to the experimental data, the generalization is often a problem as well as overfitting. The model complexity of white-box models however increases very fast. Therefore, there is usually not enough computing power available to apply such models to control units. The selection of the input variables of the model is a next important step, since these must contain a high amount of information regarding the output variable. Once a model has been selected and adapted to the specific problem, the validation is considered the final step of the virtual sensor modelling. The temperature of the turbine housing is to be estimated with a virtual sensor. As testing costs for large diesel engines are high, only a small amount of experimental data with a measurement of the component temperature is available. According to the approach the collected data is filtered. Due to the small amount of data as well as insufficient representation of possible operating ranges a black-box-model wasn’t applied on the data set. Therefore a physics-based approach is chosen to determine the temperature of the component. To model the component temperature the heat transfer in the turbine

7

housing has to be described using the physical laws. Fig. 5 shows the heat fluxes in the turbine housing. To calculate the temperature, the heat fluxes must be balanced. However, simplifications and assumptions are necessary to use this physical model as a virtual sensor during system operation. [6,7,8]

Fig. 5. Heat transfer in the turbocharger [8]

Not every input parameter of the physical model is known during operation. For instance to estimate the mass flow a data driven approach was chosen. The other parameters are either material properties or measures. Since all required inputs variables are determined, the model can be used to estimate the turbine housing temperature. In Fig. 6 the virtual sensor was applied to measurement data of a specific engine. In the upper half the signal of the component temperature is displayed along with the exhaust temperature. In the lower half the virtual sensor model and the engine power are presented. The engine power shows a dynamic behaviour, which nevertheless leads to high component temperatures. Also a discontinuity (green box) can be detected in the signal, which arises from a stop of the engine. The stop must be detected, otherwise this will lead to an error in the interpretation of the measured values. Thus a virtual sensor provides significantly more information about the operating behaviour and the load of individual components. This is possible without measuring the relevant parameter directly thus avoiding the need for additional sensors.

8

Fig. 6. – Application of the virtual sensor on measurement data

The displayed load-time signals can be used to generate a load spectrum. Depending on the failure mechanism an appropriate counting method has to be selected, as the memory of an engine control unit is limited. Even if a data logger is available the signals might be stored on a server, but several thousand hours of operating time are not easy to analyse. So in the first step the data has to be filtered and discretized. This is intended to reduce noise for example, which would otherwise be included in the counting. For fatigue analysis the most common method is rainflow-counting. This method detects closed hysteresis in the load signal, which damage the component. Thus a rainflowcounting was used to create the load spectra. In Fig. 7 a load with 12 different levels of the component temperature and the number of hysteresis is displayed. In contrast to the SSI a load spectrum describes multiple load levels. After this step, the procedure is divided into the two paths RUL and design.

Fig. 7. Load spectrum of the virtual sensor

9

3.2

Design

One individual load spectrum is limitedly suited for the development and production in series, as only one customer is taken into account while designing a new product. This issue is also described in Fig. 1, as this way specific customer cannot be related to the entire population. Therefore it is not obvious whether the customer uses the product above average or below average. So the load spectra should be saved in a data base and then be analysed. This way quantiles can be calculated, which describe the operating behaviour of a certain percentage of customers. In Fig. 8 two quantiles of the load spectra of the virtual sensor are displayed. When analysing such quantiles of a specific system the design, for which engines it is used as well as the type of application has to be considered.

Fig. 8. –quantiles load spectra of the virtual sensor

Thereby a representative load spectra of the system can be determined and used for definition of requirements and testing in the development phase. Thus a component or system can be designed that meets the requirements of a certain percentage of customers. This leads to a reduction of oversizing and thus to a reduction of material costs. Furthermore, the test program can be adapted by knowing the load in the field. 3.3

Remaining Useful Life

In the second path a customer-specific load spectrum is used to determine the remaining useful life of the component. Therefore the load has to be converted into stress on the component. Afterwards a lifetime model can be used to determine the current health of the component. In this case the stress is calculated using the Finite Element Method and an S-N curve is used to determine the current level of damage of the component. If the stress is then compared to the S-N-curve a number of load cycles until failure with a survival probability of 98% can be calculated. In order to predict the RUL the past load has to be considered, as only a number of cycles and the current health of the component simply isn’t enough for a useful prediction. 3.4

Discussion

The possibilities of virtual sensors as well as the application for a turbine housing are described. However due to the physic based modelling of the turbine housing, several assumptions are necessary to enable the application of the model. In order to increase

10

the accuracy of the virtual sensor the current model could be followed by a blackbox model. Even different black-box models are possible, which allow the estimation of the conditions at different locations. On the other hand a physics based approach can easily adapted for different use cases. In order to calculate the RUL of the component a FEM simulation is required. Due to the complex geometry and conditions in the turbocharger a very time consuming simulation is necessary to calculate the stress of the component. With this approach the load in the field can be analysed and thus allows the further use for the development of new product generations. This leads to the reduction of over sizing and thus to reduced component costs. Furthermore, the load can be used with a corresponding lifetime model for a prognosis of the remaining useful life. With the prognosis maintenance intervals can be scheduled better or an increased lifetime can be achieved by intervening in operation of the system.

4

Summary and Outlook

In this paper, a virtual sensor was applied to estimate the time load on a component. Thereby the system of the exhaust turbocharger was analysed in more detail and the turbine housing was selected. The turbine housing mainly fails due to TMF. This is caused by the changing temperature of the component. A physical-based approach was chosen to model the load on the component, whereby several assumptions were necessary for the calculation. In the next step calculated load was used to generate a load spectrum. These were collected in a database and quantiles were determined. Thereby the quantiles can be used in the development of new products. In future work the load will be converted into the stress of the component and then be compared with an S-N-curve for lifetime modelling. The stress is calculated using FEM-simulation and as lifetime model an S-N-curve is used, which was determined at the corresponding component temperature. For the verification of the RUL prognosis failures during field operation will be used.

References 1. B. Bertsche: “Reliability in Automotive and Mechanical Engineering.” Berlin, Springer, 2008. 2. L. Fortuna: “Soft Sensors for Monitoring and Control of Industrial Processes.” Springer London 2007. 3. F. Souza, et. al.: “Review of soft sensor methods for regression applications. “ Chemometrics and Intelligent Laboratory Systems 152, pp. 69-79, 2016. 4. M. Köhler: “Load Assumption for Fatigue Design of Structures and Components”, Berlin, Springer, 2012. 5. P. Johannesson: “Guide to load analysis for durability in vehicle engineering” John Wiley & Sons, Ltd; 2014. 6. M. Nakhjiri: “Physikalische Modelle und Skalierungsmethoden zur effizienten Applikation von Turboladern”, Darmstadt, 2014.

11 7. J.R. Serrano: “A model of turbocharger radial turbine appropriate to be used in zero- and one-dimensional gas dynamic codes for internal combustion engines modelling.” Energy Conversion and Management 49, 2008. 8. A. Romagnoli, et. al: “A review of heat transfer in turbochargers.”, Renewable and Sustainable Energy Reviews 79, pp. 1442-1460; 2017 9. G. Pahl,: “Engineering Design: A Systematic Approach”, Berlin, Springer, 2007. 10. K, Goebel: “Prognostics” CreateSpace Independent Publishing Platform, 2017

Model-based injector deposit detection Michael Hinrichs1, Prof. Dr. Dr. h.c. Rolf Isermann2 and Prof. Dr. Peter Pickel1 1

European Technology Innovation Center, John Deere, 67657 Kaiserslautern, Germany 2 Institute of Automatic Control, TU Darmstadt, 64283 Darmstadt, Germany Abstract. Solenoid coil injectors are widely spread in the agricultural and heavyduty vehicle sector. These types are installed due to their higher development status, longer service life and robustness. Nevertheless, injector problems are one of the more frequent causes of malfunctions and failures in diesel combustion engines. Due to ever smaller diameters and more precise coordination of the individual components, injector deposits can have a major influence on the function of injectors. Deposits change the injected fuel mass, the emissions and the reliability of the engine. Reliability is the by far the most important customer requirement of commercial machinery. Especially internal deposits can lead to a complete failure of the engine. For example, the injector can stick in closed or even in open position and may cause engine damage. This paper presents a method for detecting internal and external deposits in the injector. The basis for the detection method is a high sampling rate of the rail pressure sensor. The results allow a qualitative statement about the injector condition. Internal and external deposits can be distinguished. If the deposits are too big, a service recommendation for injector replacement is issued. Keywords: Predictive maintenance, injector deposits, biogenic fuels, engine diagnosis fuel system

List of abbreviations ECU IDID EDID IID SIQ MIQ SIP EIP SOPD EOPD INO INC SOD

Engine control unit Internal diesel injector deposits External diesel injector deposits Injector impulse duration Setpoint injection quantity Modeled injection quantity Start of injection pulse End of injection pulse Start of pressure drop End of pressure drop Injector needle open Injector needle closed Setpoint opening delay

SCD SSIP SEIP SNO RP MSD MOD MCD Δ𝑂𝐷 ΔC𝐷 ΔNO ΔA Δm

Setpoint closing delay Setpoint start of injection pulse Setpoint end of injection pulse Setpoint needle open time Rail pressure Modeled sensor delay Modeled opening delay Modeled closing delay Difference opening delay Difference closing delay Difference needle open time Difference cross-section nozzle Difference infection quantity

© Springer Fachmedien Wiesbaden GmbH, ein Teil von Springer Nature 2020 W. Siebenpfeiffer (Hrsg.), Heavy-Duty-, On- und Off-Highway-Motoren 2019, Proceedings, https://doi.org/10.1007/978-3-658-31371-5_10

2

1

Introduction

The reliability of a machine is, beside its economy, a very important feature in the commercial vehicle sector. This applies in particular for agricultural machinery. Vehicle diagnosis and predictive maintanance try to reduce the down time and have been growing for many years [2]. The ultimate aim is “No failure during operation”. Many predictive models have been created for this purpose in order to detect and replace worn components early [4][5]. Another reason for predictive maintance is that machines are serviced periodically, although a needs-based maintenance is better from an economic and ecological perspective [1] When a machine breaks, detailled diagnosis helps technicians to repair the machine quicker without having them search for a long time for the causes of a certain fault. This is an important criterion for the customer, especially in seasonal work peaks such as harvesting. Electrical problems, defective turbochargers or faulty exhaust gas recirculation valves often cause engine malfunctions, but defective injectors also account for a large proportion of faults. Troubleshooting, repair time and down-time of the machine is more expensive as additional injectors. Often all injectors are replaced at the same time without beeing necessary. Hence, an predictive injector-specific diagnosis is very desireable. The reasons for injector damage are often internal diesel injector deposits (IDID) or external diesel injector deposits (EDID).

Fig. 1. Overview of Denso G4S Injector [3]

3

External deposits are described as deposits which are visible from the outside of the injector which may reduce the cross-section of the nozzle holes. This leads to a reduced injection mass and thus to a reduced peak performance of the engine. Furthermore, external deposits often lead to a deterioration of the fuel spray and consequently to higher emissions. [3] Deposits inside the injector are called internal deposits. The most sensitive components for the function of the injector are the inlet orifice, outlet orifice, the control plate and the control valve (Fig. 1). Even small deposits on these components can have a large effect on injector function such as a deviating opening and closing delay of the injector. The opening delay is defined as the time from triggering the injector to the opening of the injector needle. The closing delay is the time from stopping the injector pulse to a closed injector needle. Delays are also affected if there are deposits in the area of the nozzle needle guidance. These deposits increase the actuation forces for opening and closing the nozzle needle. Internal deposits can also indirectly influence the injected fuel mass and thus the performance of the engine due to changed needle opening times. At low rail pressures increased actuation forces are particularly present with IDID since the hydraulic forces are not high enough to move the needle instantly [3]. In the worst case the injector remains in a closed or open position. If the needle remains open, catastrophic engine damage can be the result, as uncontrolled large quantities of fuel are injected into the combustion chamber. However, a not opening injector leads to engine malfunction too, especially to non-smooth operation. In this paper all deposits that change the timing of the nozzle needle are defined as internal deposits. All deposits that change the flow through the injector (with the same needle opening time) are defined as external deposits. In order to ensure the reliability of the machine the presented model classifies the injectors of the engine and indicates at an early stage that injectors should be replaced. Investigations were carried out on a John Deere 4045 4.5 L diesel engine of the EU exhaust stage 5. The fuel system is a Denso high-pressure system (HP3 high-pressure pump and G4S solenoid injector).

2

Model Data Acquisition

The maximum possible sampling rate of the used control unit is 2 kHz (T0 = 500 µs). Since many processes in the injector have high dynamics, a higher sampling rate is needed to receive information about the injector behavior. Pre investigations show that the rail pressure signal should be sampled with at least 100 kHz. Certain phenomena, like differences in the opening delay between new and injectors with deposits, require even higher sampling rates. However, a sampling rate of 100 kHz is enough to reliably classify injectors in the vehicle with series sensors like the rail pressure sensor. To reach that sampling rate with the used control unit, the signal will be sampled repeatedly.

4

Railpressure

ECU schedules injection (pulse in e.g. 3513 μs) 1. run 2. run 3. run 50. run

I

Injectionpulse e.g. 1200 μs

10 μs

13 μs 3500 μs

Time Fig. 2. Repeated sampling principle

Only one injector is examined at the same time. For this injector, multiple injections are deactivated for the duration of the deposit detection. To get to the required resolution of 100 kHz, 50 injections are recorded in a row. In every run the signal is sampled at 2 kHz but shifted compared to the injection pulse by 10 µs. A run is defined as the time from scheduling the injection up to the next pumping process of the highpressure pump (Fig. 2 and Fig. 3). After the ECU has calculated the start time for the injection pulse (in the shown example the start of the next injection pulse will be in 3513 µs), the start time for recording the signal is set. In the example shown in Fig. 2, the recording will start exactly 13 µs after the injection scheduling (a sampling rate of 2 kHz means one sampling point every 500 µs, so the 7th sampling point of the first run will be 3513 µs after the scheduling, where the injection pulse is starting). In the 2nd run (next injection of examined injector), the start of the sampling is shifted by 10 µs relative to the start of injection. This is carried out until the 50th run, so that the injection signal of the examined injector is sampled with an even distance of 10 µs. The data of the respective run is recorded until the next pumping process of the high-pressure pump takes place (Fig. 3). The requirement for repeated scanning is an identical signal. Since the injection process in common rail systems is independent of the crank angle, an identical injection process can be influenced by the ECU. This can be done by a constant rail pressure setpoint and an identical injector impulse duration (IID). However, for an identical IID and thus an identical injection mass (pulse at same rail pressure), delivered engine torque must be (almost) stationary. Since only one injector at a time is examined for deposits, slight torque fluctuations can be compensated by lowering or increasing the IID of the remaining injectors. In this case a difference of injection quantity between injectors of ±10% is allowed. In an agricultural machine, this leads to barely noticeable additional vibrations. In addition, on machines with an infinitely variable transmission, torque delivered by the engine can be actively influenced by changing the transmission ratio. IID is set very precisely controlled by the ECU. The rail pressure, on the other hand, is subject to slight fluctuations.

5

Railpressure

𝑝𝐻 2 𝑝𝐻 𝑝𝐻 1

Inj. 1

Inj. 4

Inj. 3

Inj. 1 Inj. 2

Data 2. run

Data 1. run Time Fig. 3. Rail pressure fluctuations

These fluctuations can result from an inaccurate metering unit of the high-pressure pump, speed fluctuations or unequal injections of the remaining injectors. This effect is shown schematically in Fig. 3. In areas with an increasing rail pressure, pumping processes of the high-pressure pump take place. In areas with a decreasing rail pressure, injections of the individual injectors take place. The pressure 𝑝𝐻1 in the 1st run of injector 1 is much lower than the 𝑝𝐻2 of the same injector in the 2nd run. Since the signal should be identical (same starting pressure at every injection), all recorded data of the 50 runs are linear shifted to the same starting pressure ̅𝑝̅̅𝐻̅. For this purpose, all 𝑝𝐻𝑖 of the 50 runs are averaged: ̅𝑝̅̅𝐻̅ =

∑50 𝑖=1 𝑝𝐻𝑖 50

(1)

Subsequently, all recorded data is shifted so that each recorded run has the rail pressure ̅𝑝̅̅𝐻̅ before an injection (Fig. 3). Although all real injections do not have the identical rail pressure, the measurements can be combined. This can be done because there is only a small difference between real and shifted rail pressure. The differences in the injector behavior (e.g. higher injection quantity, because of higher rail pressure) are so small that they can be neglected. However, 𝑝𝐻𝑖 is not constant too. In order to compensate rail oscillations or measuring noise, the mean value from all data points before an injection is calculated for one run. In the shown example that would be 6 data points for the first run (7th data point is at start of injection). In practice operation approximately constant load points are very often present. This means the detection program can be carried out during normal engine operation without affecting the work of the machine. However, fixed rail pressure and IID of one injector, and the deactivation of multiple injections can worsen exhaust emissions. This should be considered in series production to stay inside the emission regulations. Since the effects of injector deposits are particularly well detectable with large injection quantities, the detection mode is only started with an injection quantity of at least 50 mg/str.

6

3

Deposit Detection Model

The ECU is connected to a speed sensor so that the current crankshaft position can be calculated at any time. If a certain crankshaft-based injection angle is specified, it will be converted into a time-based injection in the ECU. The ECU contains tables for the opening and closing delay of the injector, based on rail pressure and IID. For a desired start of injection angle, the ECU calculates the setpoint start of injection pulse (SSIP). Furthermore, the duration of the pulse is determined based on the desired injection quantity and the current rail pressure. Since each injector is slightly different, correction factors are stored in the ECU. These are provided by the manufacturer for each injector. This compensates differences in the opening delay, closing delay and the injector flow rate, as long as the injectors are new. With the stored data the setpoint end of injection pulse (SEIP) can be calculated. The influence of the fuel temperature on the density and injection quantity is not considered in series production ECU operation. However, since the fuel temperature has an influence on the injector behavior, the presented tests are carried out at a conditioned low-pressure fuel temperature of 30°C. 105

Railpressure (MPa)

Needle Closed

Needle Open

Needle Closed

SIP EIP

100

Railpressure Injectorpulse Calculated by Model

IID.=1200 us Fueltemp.=30 °C mfuel=51.6 mg

95 0

INO

1000

SOPD

INC

3000 2000 Time (µsec)

EOPD

4000

5000

Fig. 4. High-resolution rail pressure signal with new injector

The high-resolution rail pressure signal generates a lot of information on the injector behavior which is not recorded in series operation. It is noticeable that the rail pressure rises before it drops during the injection. The reason for this is the high current of the injector actuation. High currents have a large disturbance influence on the rail pressure signal. This interfering signal cannot be prevented without changing the standard wiring harness and the ECU. The typical injector current curve of a solenoid coil injector can be seen (Fig. 4 red). In a fault-free condition SIP=SSIP and EIP=SEIP applies. With setpoint opening delay (SOD) stored in the ECU, the time when the injector needle should be open (INO) can be calculated as

7

𝐼𝑁𝑂 = SSIP + 𝑆𝑂𝐷.

(2)

The setpoint closing delay (SCD), the time at which the needle is closed (INC), can be calculated analogously from: 𝐼𝑁𝐶 = SEIP + 𝑆𝐶𝐷

(3)

It is obvious that the time from the end of the pressure drop (EOPD) does not match INC (Fig. 4). The reason for this is a delay between the rail pressure sensor and the real pressure drop at the injector. Inertia and position of the sensor, and the speed of sound of the fuel, introduce a delay into the measurement. To classify the injector later the opening delay, closing delay, total needle opening time, and the injected fuel mass must be determined. For this, SOPD and EOPD and the sensor delay must be calculated first. The individual calculation steps are described below: Step 1: The start (SIP) and end (EIP) of the injector pulse can be determined by the maximum and minimum slope of the rail pressure signal in the range of an injection. Step 2: By comparing the setpoint times (SSIP, SEIP) of the injector control with SIP and EIP, an incorrect control of the injector can be detected. A tolerance of ±20 µs is allowed for the start and end points. If the deviation is greater, an error is detected. This error is stored separately and has no influence on the deposit detection, since SIP and EIP are used. A reason for a large deviation between SSIP/SEIP and SIP/EIP could be e.g. engine tuning with enlarged pulse durations, or other electrical problems. Step 3: To be able to determine the SOPD despite the injector interference signal, an filtered line is created in the undisturbed injection area after the injection pulse (Fig. 4 green). To calculate the gradient during an injection, the measured values 30 µs after EIP up to INC are used. Step 4: SOPD can be determined by the time at which the equalizing line reaches the value ̅𝑝̅̅𝐻̅. Step 5: The average pressure after injection (𝑝 ̅̅̅) 𝐿 is calculated. To calculate this, all measured values between 1500 µs after INC up to the end of data (high-pressure pump delivering fuel) are averaged. This ensures that the mean value is calculated after EOPD and that any rail vibrations that occur do not falsify the value. Step 6: EOPD is calculated by the time at which the equalizing line reaches the value 𝑝𝐿 ̅̅̅. Step 7: The total needle opening time can be calculated with:

8

𝑡𝑁𝑒𝑒𝑑𝑙𝑒𝑂𝑝𝑒𝑛 = EOPD − SOPD

(4)

Relative calculations in this context are not entirely accurate since different rail pressures at SOPD and EOPD have an influence on the speed of sound of the fuel and thus on the sensor delay. However, the difference can be neglected in comparison to the sampling rate of 10 µs. The calculated needle opening time (1920 µs) for the new injector in Fig. 4 corresponds well with the setpoint needle opening time (1908 µs) from the ECU tables. Step 8: With the help of the fuel temperature in the rail 𝑇𝑅𝑎𝑖𝑙 and the pressures ̅𝑝̅̅𝐻̅ and 𝑝𝐿 the density of the fuel can be calculated before ρbefore inj and after an injection ̅̅̅, ρafter inj . Since the high-pressure volume VHPS is approximately constant and known, the fuel mass that has left the rail can be calculated. If the leakage mass mleakage i of the injector (Fig. 1 bottom left) is subtracted from this mass, the modeled injection quantity (MIQ) can be calculated: 𝑀𝐼𝑄 = (ρbefore inj − ρafter inj ) ⋅ VHPS − mleakage i

(5)

Densities and mleakage i are calculated from LookUp tables. The calculated mass for the new injector was 52.0 mg. This is at the same level as the measured actual injection mass of 51.6 mg. With 51.4 mg, the injection quantity setpoint was slightly below the measured value. The slight deviation between the setpoint and the measured fuel consumption could be due to the lack of temperature adjustment in the ECU, or measurement noise. Step 9: Using the sensor response time 𝑡Resp , the sound velocity 𝑣Sound of the fuel (depending on pressure and temperature), as well as the distance of the investigated injector to the pressure sensor (𝑠𝑖 ), modeled sensor delay (MSD) can be calculated approximately: 𝑀𝑆𝐷 = 𝑡Resp +

si vSound

(6)

Step 10: Modeled opening delay (MOD) of the injector can be calculated: 𝑀𝑂𝐷 = SOPD − MSD − SIP

(7)

For the new injector in the shown example an opening delay of MOD=300 µs is calculated. This is very accurate compared to the theoretical opening delay of 302 µs. The closing delay is calculated analog by: 𝑀𝐶𝐷 = EOPD − SD − 𝐸IP

(8)

The model calculates a closing delay 𝑡𝐶𝐷 = 1000 𝑢𝑠𝑒𝑐. The theoretical closing delay is 1010 µs, so the model is one sampling point off.

9

Step 11: In the last step, the differences between computed parameters and setpoints are calculated. The setpoint opening delay (SOD) is subtracted from the modeled opening delay: Δ𝑂𝐷 = MOD − SOD

(9)

The difference in closing delay is Δ𝐶𝐷 = MCD − SCD ,

(10)

wherein the setpoint closing delay (SCD) is subtracted from the modeled closing delay. The difference in the needle open time can be calculated analogously by subtracting the needle open time setpoint (SNO): ΔNO = MNO − SNO

(11)

The change of the injection quantity is set as a percentage of the setpoint injection quantity (SIQ): Δm =

MIQ−SIQ SIQ

(12)

EDID are not detected based on Δm. There are theoretically injector states where the effect of EDID is compensated by IDID (EDID reduces injected fuel and longer total needle opening times increases injected fuel, so that the overall injected mass is the same). ΔA as the change in the cross-section of the nozzle holes of the injector, ΔNO

ΔA = Δm − SNO

(13)

is introduced. It is assumed that the injector opening time is proportional to the injection mass and that ΔA has a linear effect on the injection mass too. This assumption can be made if the nozzle holes are considered simplified as an orifice. For an orifice (according to Bernoulli and the law of conservation of mass) applies 2

𝑄 = α ⋅ A ⋅ √(𝜌 ⋅ (𝑝1 − 𝑝2 )),

(14)

where Q is the volume flow, α is a flow coefficient, A is the cross-section, ρ is the density of the fluid, 𝑝1 is the pressure before and 𝑝2 is the pressure after the orifice. ΔA is considered as the reduction of A, so ΔA has a linear influence on the volume flow. Assumed that the density does not change during injection, an extended needle opening time can be considered just like ΔA as a linear effect on mass flow. After calculating ΔOD, ΔCD, ΔNO and ΔA, they are stored on the control unit. If 10 different stationary load points have been saved, the oldest one will be overwritten. To classify the injector, mean values over 10 evaluated load points of ΔOD, ΔCD, ΔNO und ΔA are calculated. This compensates the influence of relatively inaccurate recording resolution. Especially at the opening delay, a recording value shifted by one measuring point can result in a relatively large deviation.

10

4 classes are distinguished. If all mean values lie within a tolerable range, the injector is classified as “Injector new”. As soon as individual variables deviate, a distinction is made between "Injector still good", "Change / clean injector at next service" and "Replace injector as quickly as possible" depending on the deviation. Limit values for the respective averaged parameter are shown below: Table 1. Injector Classification features

Class

Injector Status

Difference opening delay Δ𝑂𝐷 (µs)

Difference closing delay ΔC𝐷 (µs)

Difference needle open time ΔNO (µs)

A B C

Injector new Injector still good Change/clean injector at next service Replace injector as quickly as possible