3D Nano-Thermistors for Thermal Probing | ACS Applied Materials

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Functional Nanostructured Materials (including low-D carbon)

3D Nano-Thermistors for Thermal Probing Jürgen Sattelkow, Johannes E. Fröch, Robert Winkler, Stefan Hummel, Christian H. Schwalb, and Harald Plank ACS Appl. Mater. Interfaces, Just Accepted Manuscript • DOI: 10.1021/acsami.9b04497 • Publication Date (Web): 03 Jun 2019 Downloaded from http://pubs.acs.org on June 8, 2019

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ACS Applied Materials & Interfaces

3D Nano-Thermistors for Thermal Probing

Jürgen Sattelkow1,†, Johannes E. Fröch2,3†, Robert Winkler1, Stefan Hummel4, Christian Schwalb5, Harald Plank 1,2,6,*

1

Christian Doppler Laboratory DEFINE, Graz University of Technology, 8010 Graz, Austria

2

Graz Centre for Electron Microscopy, 8010 Graz, Austria

3

Institute of Biomedical Materials and Devices, University of Technology Sydney, Ultimo, New South

Wales 2007, Australia 4

Physics of Nanostructured Materials, University of Vienna, 1090 Vienna, Austria

5

GETec Microscopy Inc., 1120 Wien, Austria

6

Institute of Electron Microscopy and Nanoanalysis, Graz University of Technology, 8010 Graz, Austria

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ACS Applied Materials & Interfaces 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

† both

authors have equally contributed to this work

* Corresponding author e-mail: [email protected]


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1 Abstract Accessing thermal properties of materials or even full devices is a highly relevant topic in research and development. Along with the ongoing trend towards smaller feature sizes, the demands on appropriate instrumentation to access surface temperatures with high thermal and lateral resolution also increase. Scanning thermal microscopy is one of the most powerful technologies to fulfil this task down to the sub-100 nm regime, which, however, strongly depends on the nano-probe design. In this study, we introduce a 3D nano-probe concept, which acts as a nano-thermistor to access surface temperatures. Fabrication of nano-bridges is done via 3D nano-printing using focused electron beams, which allows direct-write fabrication on pre-structured, self-sensing cantilever. As individual branch dimensions are in the sub-100 nm regime, mechanical stability is first studied by a combined approach of finite element simulation and scanning electron microscopy assisted in situ atomic force microscopy (AFM) measurements. After deriving the design rules for mechanically stable 3D nano-bridges with vertical stiffness up to 50 N.m-1, a material tuning approach is introduced to increase mechanical wear resistance at the tip apex for high-quality AFM imaging at fast scan speeds. Finally, we demonstrate the electrical response in dependence of temperature and find a negative temperature coefficient of –(0.75 ± 0.2) 10-3 K-1 and sensing rates of (30 ± 1) ms.K-1 at noise levels of ± 0.5 K, which underlines the potential of our concept.

Keywords: 3D nanoprinting, additive direct-write manufacturing, focused electron beam induced deposition, scanning thermal microscopy, nanomechanics, nanoelectrics, nanothermics



2 Introduction In research and development, thermal influences play a central role as physical, chemical and functional properties often rely on local temperatures, its spatial distribution and the temporal evolution in 3D space 1,2 . Accessing these details quantitatively is comparably easy at macroscopic dimensions, gets more complicated at the micro-range and becomes very challenging when entering the nanoscale. The latter, however, becomes more and more relevant due to the still ongoing trend to smaller structures 3 . Therefore, to specifically improve material design, the accessibility of thermal properties on the lowest nanoscale is not only important but often decisive for further optimization of thermoelectric nano-materials 4 . Among the comparably small pool of techniques, which allow thermal probing in the sub-100 nm range, scanning thermal microscopy (SThM) is the most powerful technology to access laterally resolved temperatures and further correlate the findings with surface morphology

1,5–7

. Impressive examples have been demonstrated, such as Joule self-heating of

graphene 8 , Seebeck coefficient determination 9 , and heating of plasmonic structures 10 just to name a few. The success of such studies, however, is strongly related to suitable nano-probes with respect to their dimensions, sensitivities and stabilities 3,7 . Commercial SThM tips can basically be divided into three different types. Wollaston probes use bridged metal wires (often Pt based) attached to a prestructured cantilever basis 1,6,9 . Although tip radii of down to 20 nm 7,11,12 have been reported using very special treatments, typical apex radii are in the range of some hundreds of nanometers to a few micrometers 3,13,14 , which are strongly limiting achievable resolution and positioning accuracy. The second approach uses U-shaped cantilevers, consisting of highly doped Si side wall elements and low doped tip areas in between 1,15 . Although tip radii below 100 nm can be achieved 16,17 , the “active areas” are fully connected with the entire cantilever, representing a huge heat sink. By that, small temperatures are more complicated to access when operated in SThM mode 1,5,18 . A combination of both approaches led to the introduction of surface modified cantilevers: lithography based methods are used to selectively fabricate metal structures across the tip, which can be understood as structured, conductive atomic force microscopy (C-AFM) tips. This approach enables small “active areas” down to


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20 nm gap widths, however, with end radii in the range of about 50 nm

1,5,19–21

. More advanced

approaches use carbon nanotubes 11 , single dot assisted regions 22 or direct current imaging through the samples using the contact potential as “sensing element” (restricted to conductive samples) 23 . In this study, we introduce a 3D nano-probe (3DNP) concept, which combines several aforementioned advantages in a single probe. In brief, a freestanding, multi-legged, 3D bridge architecture is used, which is electrically operated to access temperature dependent resistance changes (thermistor) 24 once in contact with the sample surface. Fabrication is done via focused electron beam induced deposition (FEBID), which has matured in recent years 25 . Aside of strongly improved predictability / reliability 26,27 , FEBID was also used for 3D nano-applications including plasmonics 28 , magnetics 29,30 , and sensors 31 as it enables direct-write fabrication of complex, freestanding 3D nanostructures 27,32 with branch diameters below 100 nm and tip end radii in the range of 10 nm without any further processing. By that, such freestanding 3DNPs provide very small active volumes in contrast to the large heat sinks for planar thermal nano-probes, which is a decisive element for fast and sensitive temperature sensing. On the other hand, the sharp apex is essential when aiming for highest possible lateral resolution, which exploits the full potential when applied in vacuum conditions due to the absence of any resolution diminishing, convection based heat transfer 3 . Aside from the 3D capability, FEBID allows direct-write fabrication on practically any given material and surface morphology without specific pre- or post-treatment steps at the region of interest. Hence, this nano-printing technique is ideal for fabrication and / or modification of already finished devices such as AFM self-sensing cantilevers, as relevant in this study, which eliminates more complicated fabrication procedures during for functional probe formation. While the long-term goal of our activities is the application of FEBID based 3DNPs for thermal highresolution mapping via SThM in vacuum conditions, this first study exclusively focuses on the detailed characterization of the actual 3D nano-thermistor elements. This starts with the 3D design, nanofabrication and optimization, goes over property tuning for stable AFM operation and ends with the temperature dependent, electric response in static and dynamic conditions as originally intended.



By that, this study delivers the proof-of-principle for the application of FEBID based 3D nano-structures as thermal nano-thermistors for local temperature probing and lies the foundation for the aforementioned SThM related follow-up activities.

3 Results and Discussion For the proof-of-principle concerning the application of freestanding 3D multi-pods as AFM compatible, thermal nano-probes, the study is split into three parts, as schematically summarized in Figure 1. First, finite element simulations (FES) are used for upfront confinement of 3DNP architectures and dimensions, complemented by scanning electron microscopy (SEM) assisted in situ AFM experiments for quantitative stiffness validation. In a second step, we focus on imaging quality of such 3DNPs during AFM operation and introduce a material tuning approach to increases the mechanical wear resistance. In the final step, we study the temperature dependent electric response in static and dynamic conditions to deliver the intended proof-of-principle of our 3DNP concept for thermal nanoprobing.

3.1 Simulation Based Design Selection The first focus lies on mechanical properties of freestanding 3DNPs in dependence on overall architectures and related dimensions. This is of essential relevance due to the unavoidable vertical and lateral force load during AFM operation. Therefore, both, vertical and lateral stiffness (𝑘𝑉 and 𝑘𝐿, respectively) have to be maximized by proper design, which is studied in this section. Instead of an exhaustive trial-and-error approach between 3DNP fabrication via FEBID and mechanical characterization using AFM, FES were used to derive design-to-stiffness relationships for the upfront confinement of most promising architectures and related dimensions. The therefor required Young


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modulus has recently been extracted by Arnold et al. 31,33 , who found a value of (14 ± 2) GPa for freestanding FEBID nano-pillars. The application of this value as FES input parameter is justified as 3D structures in the mentioned as well as in the present study were fabricated in the same dual beam microscope using the same FEBID parameter (30 keV, 21 pA), precursor (MeCpPt(IV)Me3) and technical setups. In this first step, stationary modelling was performed in COMSOLs Structural Mechanics Module. As no information on the Poisson ratio 𝜈 for nano-crystalline Pt in a disordered hydrocarbon matrix are available in literature, we conducted an upfront study to evaluate its influence on final stiffness. For that, we compared simulated stiffness for a 5 µm high and 50 nm wide nano-pillar with analytical solutions, both for increasing Poisson ratios from 0 to 0.5. For 𝜈 close to 0.5 the biggest deviations were 0.1 % and 1 % in vertical and lateral direction, respectively (see Supporting Information 1). This overall negligible influence of 𝜈 can be renationalized by the high aspect ratio of the geometry, which implies that an axial volume change will result in an almost negligible lateral volume variation. For further simulations, 𝜈 was therefore fixed to 0.2, in agreement with literature values for carbon-based materials 34,35 . Next, 𝑘𝑉 and 𝑘𝐿 were calculated by 𝑘 =

𝐹

∆𝑙 with the applied force 𝐹 and the obtained

compression ∆𝑙 for different pillar lengths (1 – 5 µm) and diameters (20 – 100 nm). The results are shown by 3D plots in Figure 2a (𝑘𝑉) and 2b (𝑘𝐿), revealing stiffer behavior for larger diameters (linear scaling with cross-sectional areas) and shorter overall heights (linear scaling with reciprocal height; both effects are shown in Supporting Information 1). More importantly, however, is the finding that kV exceeds kL by typically 4 orders of magnitude (see stiffness scales in Figure 2). For example, FES predict vertical and radial stiffness of 13 N.m-1 and 10-3 N.m-1, respectively, for a 3 µm high and 60 nm wide nano-pillar. To evaluate FES reliability, we compared the results with the analytical solution 36 using:

∆𝑙𝑉 =

∆𝑙𝐿 =

4 𝐹𝑉ℎ 𝜋𝑑2𝐸

64 𝐹𝐿ℎ³ 3𝜋𝑑4𝐸


equation 1

equation 2


with h as pillar height, d as diameter and E as Young’s modulus. The results revealed a deviation of less than 1 % over the studied range (see Supporting Information 1), confirming the reliability of our FES approach. Figure 2c shows a 60 nm wide and 3 µm high Pt-C single-pillar under axial (top) and radial (bottom) force of 10 nN together with the spatial stress distribution (see qualitative color bar). While axial compression leads to spatially homogeneous stress, radial force loads generate highest stress at the fixation point at the bottom, which becomes relevant later. In the second step, we changed the design to bi-, tri- and tetra-pods with a constant pillar diameter of 60 nm and simulated 𝑘𝑉 and 𝑘𝐿 in dependency on total heights h and inclination angles  (measured against the substrate surface). Figure 2 shows the stress distribution in bi- (b) and tetra-pods (c) under vertical (upper) and lateral (lower) force loads (tri-pods are shown in Supporting Information 1). In contrast to single-pillars (c), multi-pods reveal highest stress at the topmost merging area under vertical force load (compare upper models in c-e), which indicates that the tip apex will experience highest stress during vertical force load in AFM. Beside these qualitative findings, Figure 3a presents quantitative 𝑘𝑉 scaling plots for all architectures (abscissa) and selected variations in h,  and d. As reference, multi-pod heights of 1 µm, inclination angles of 60° and branch diameters of 40 nm are shown by black squares. Red circles reveal implications of structure heights (1 µm  5 µm), blue triangles indicate steeper angles (60°  80°), and purple inverted triangles show the effect of broader branches (40 nm  80 nm). The first detail is the linear stiffness increase with the number of branches, except for single-pillars, which are always vertical ( = 90°). In agreement with single pillar results (Figure 2a), taller structures lead to decreased stiffness (red arrow), while thicker branches increase the values for 𝑘𝑉 (purple arrow). The stiffer character for larger inclination angles is also in agreement with single pillar findings as increasing  (towards 90°) can be understood as convergence to vertical pillars with highest possible stiffness. The radial stiffness behavior, shown in Figure 3b, is more complex due to the dependency on force load directions relative to multi-pod orientation. This becomes most pronounced for bi-pods, when ACS Paragon Plus Environment

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comparing force loads perpendicular (𝑘𝐿 ⊥ ) and parallel (𝑘𝐿 ∥ ) to the bi-pod plane (see indications in the lower part of Figure 2b), resulting in a strong stiffness splitting (duplets in Figure 3b, connected by solid / dashed lines). While the former is identical to lateral flexing of single-pillars (see bottom models in Figure 2c and d), the latter resembles an on-axis compression. This leads to 𝑘𝐿 ⊥ and 𝑘𝐿 ∥ values, which differ by more than 3 orders of magnitude, which immediately points out the need for a radially resolved 𝑘𝐿 study. The red curve in Figure 3c shows the result for a bi-pod structure, which actually is a very small ellipse. This finding immediately excludes bi-pods as potential AFM probes, as stable XY scanning becomes practically impossible. While the radial 𝑘𝐿 homogeneity strongly increases for tripods (blue curve in Figure 3c), the situation becomes very homogenous with a 4-legged tetra-pod (TP) architecture, shown by the almost circular green curve in Figure 3c. Radial stress distributes homogeneously across TPs, as shown in the lower part of Figure 2e, thus makes it the most favorable architecture. Analogous to 𝑘𝑉, Figure 3b shows quantitative 𝑘𝐿 scaling plots for the same geometrical variations. As mentioned before, the bi-pod splitting stems from the 𝑘𝐿 ⊥ / 𝑘𝐿 ∥ anisotropy, where higher values are always connected to force loads parallel to the bi-pod plane. While increasing heights and diameters lead to the same qualitative shifts for 𝑘𝐿 as for 𝑘𝑉, the dependency on inclination angles  is inverted: high angles are beneficial for vertical stiffness, while dramatically reducing its lateral counterpart. This implies that a compromise has to be found for the inclination angles to provide highest 𝑘𝑉 with acceptable 𝑘𝐿 values. The green diamonds in Figure 3a and 3b show such a compromise taking realistic pillar diameters, fabrication capabilities (topmost merging volume) and geometrical boundary conditions on self-sensing cantilever (minimum footprint) into account as well. In detail, TP geometries with 60 nm branch diameters, 2 µm overall heights and 70° inclination angles were chosen for further experiments due to radially homogenous behavior and comparably high 𝑘𝑉 and 𝑘𝐿 values of 60 N.m-1 and 4 N.m-1, respectively. The former value exceeds the target stiffness of the cantilevers by 1 – 2 orders of magnitude (nominally 4 N.m-1 and 8 N.m-1 for two different cantilever types). In the next step, real experiments are presented for quantitative validation of our FES studies.



3.2 Experimental In Situ Characterization To evaluate FES results, real compression experiments were performed with a SEM integrated in situ AFM 37 instrument (AFSEM® by GETec Microscopy Inc.) to study the morphological behavior during force load. For that, self-sensing cantilever were prepared via focused ion beam (FIB) processing by removing the front part for real time SEM inspection (see Supporting Information 2). The apex of the original tip has been removed with a pre-tilt to produce a parallel contact area between AFM probe and sample surfaces as shown in Figure 4a by a SEM side view image. Next, TP arrays were fabricated on Si-SiO2 substrates, using the simulated design to evaluate the scaling behavior for comparison to FES. First compression tests, however, revealed an unexpected peculiarity by means of a twisting behavior. A particularly clear example is shown in Figure 4, starting with the first AFM approach (b) followed by a vertical compression of 350 nm (c) and after force release (d) (see also Supporting Movie 1). Although small in this example, the TP suffered an irreversible damage as indicated by the red arrows in (d) (compare to (a)). During a large series of compression experiments, many different deformation types during and after compression were found (examples can be found in Supporting Information 2). As such twisting behavior was not predicted by FES, we subjected the TPs to a closer SEM inspection, which revealed spatial mismatches in the topmost merging zones. Figure 5 shows an example by means of a vertical displacement (a) and an even stronger lateral mismatch in the central merging zone (b). When including real mismatch values derived from SEM in the FES model (inset in c), FES was able to mimic the deformation mode as found by experiments as shown in Figure 5d. Please note, the latter SEM image was taken after compression and actually shows an irreversible deformation, which is identical with the bending direction during dynamic compression observed by live SEM imaging. These findings clearly point out that highest spatial precision is indispensably required to avoid uncontrolled deformation. The solution for this problem is a slightly adapted design in the merging region, as described later in more detail as it goes along with another morphological adaption. For completeness, we want to mention that a non-vertical compression (lateral flexing) due


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to experimental inaccuracies was considered via FES as well, which, however, could not reproduce the observed twisting effects (see Supporting Information 2). During the same experimental series, non-twisting TPs were furthermore subjected to AFM based quantitative mechanical characterization. For this, the pre-processed cantilever were first calibrated against Si-SiO2 surfaces to recalculate the cantilever stiffness 𝑘𝐶𝐿. Next, ramp cycles with limited Z movement were performed on tetra-pods, while monitoring the exerted force. Figure 6a shows a set of response curves using a cantilever with a stiffness of (7.2 ± 0.2) N.m-1. The steep, black dotted line gives the reference response on Si-SiO2. The colored solid lines give the raw data of a multi-cycle response for the same TP (see legend), which immediately reveals the astonishing result of a saturation behavior. As evident, a widely linear range is observed for small displacements (shaded green), followed by a weak kink (blue arrow) and another widely linear regime (shaded blue), which eventually changes into a strong saturation behavior (shaded red). This means, that at a certain point (blue  red) the required force to further lower the cantilever strongly decreases and saturates in almost zero efforts for further compression (red region). This remarkable result stems from the cycles actually lying on top of each other, clearly indicating a reversible compression even after 10 cycles. While the example in Figure 6a reveals a practically identical onset, offset behavior was found as well for some tetra-pods, which indicates irreversible, plastic deformation after compression cycles, excluded from these analyses. When considering the TP with a collective single spring cTP, equivalent to kV, one can use the formalism for serial springs to recalculate this value from the first linear region (shaded green), suggesting a cTP of (18 ± 2) N.m-1 in this case. While FES do not predict the saturating behavior, a closer look reveals another morphological deviation by means of non-straight side branches of the TP, as shown by a tilted SEM image in Figure 6b. The deviation from straight substrate-tip geometries (dashed yellow lines) leads to a special form of Euler buckling with both ends fixed 38 although a certain degree of free translation parallel to the substrate can occur. To simulate such a behavior, we first introduce the take-off angle TO together with the intended angle eff connecting start- and end-point of the side branches, both shown in Figure 6b. Next, we modelled the branch bending by a 2nd order polynomial,



which fits well real circumstances (see Supporting Information 2). The results of the adapted FES are summarized in Figure 6c, showing the influence for increasing  (defined by TO - eff) on vertical stiffness (ordinate). Starting with a cV of 20 N.m-1 for perfectly straight side branches (green indications) without any spatial mismatch at the apex, even a small deviation of only 5° leads to a stiffness drop by a factor around 4 (indicated blue). This fast decay can be understood as a transition from axial compression along the branches to lateral flexing for increasing bending angles, which clearly points out the high relevance for straight side branches. To evaluate the validity of the FES results, TP arrays with varying heights and target angles were fabricated, initially characterized via SEM to access real heights, eff and TO values. The same TPs were then subjected to quantitative AFM compression tests, where only non-twisting TPs were used for analysis. The results can be seen in Figure 7 by a direct quantitative comparison between experiments (a) and simulations (b). As evident, a very good agreement with relative deviations of less than 10 % were found, which clearly reveals the non-linear force-to-compression behavior as a consequence of non-straight branches. Finally, we simulated the dynamic response during force load by including the cantilever in the simulation as well, which showed good agreement with real experiments as well (see Supporting Information 2). This validates our FES approach, including the initially used value of Young’s modulus, to predict spatial stiffness for FEBID based multi-pod architectures and explains undesirable artifacts such as radial twisting and non-linear behavior that occur during AFM compression experiments.

3.3 Design Optimization While ideal architectures (4-legged TPs), including relevant overall heights and inclination angles, can be derived from FES, further design optimization is needed to minimize twisting and non-linearity effects. As described before, the former occurs due to lacking precision in the merging apex region, which typically stems from mechanical stage drift and non-ideal boundary conditions during FEBID fabrication, both discussed by Winkler et al. 25 . Although these effects can technically be minimized, it


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is difficult to achieve reliable accuracy in the lowest nanometer range in 3D space. To overcome this challenge, we have increased the volume in the merging region, making the apex more rigid, thus less sensitive to very small deviations and less prone to twisting effects. Figure 7 shows SEM top view (c) and tilted images (d), where the slightly increased merging region is evident, realized by additional patterning points. The second unwanted effect, namely the non-linear response under force load, is a consequence of non-straight branches. As recently revealed, this artifact has its origin mainly in local heating during FEBID 39 . As branch growth proceeds, the temperature at the beam impact region increases, resulting in a reduction of precursor coverage and therefore reduced growth rates. While optimization towards minimized bending was done by the introduction of additional refresh times in this study, we are currently working on the integration of a temperature compensation module in the CAD software 3BID 26 . Figure 7d shows the results of this adapted process strategy, leading to straight single branches. When compressing such idealized geometries, both, twisting and irreversible damage are eliminated and minimized, respectively (see Supporting Movie 2) Figure 7d shows a direct comparison of a TP before (left) and after multi-cycle compression (right). Prior to a final design selection, the boundary conditions for TP footprints had to be determined. This follows the electrode layout on our self-sensing cantilever, which currently requires a leg-to-leg distances of at least 500 nm. From the stiffness plots in Figure 7b, it becomes evident that shortest TPs with steepest sidewalls are most promising. Now choosing a realistic side-wall angle around 70° together with the aforementioned demand on the base width, TP heights between 700 nm and 1100 nm should lead to vertical and radial stiffness of higher than 50 N.m-1 and 5 N.m-1, respectively. Those values are one to two orders of magnitude higher than the planned cantilevers spring constants and therefore should allow stable AFM operation as discussed in the following section. In short summary, a 4-legged tetra-pod geometry is chosen as it provides radially homogeneous stiffness, while the opening angle is dictated by the targeted axial stiffness and the applied footprint. Although higher leg numbers would further increase mechanical stiffnesses, this approach was not followed as it would also increase the active volume, which can impact both, response times and temperature sensitivity.



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3.4 Property Tuning As base area for the optimized TPs, the original tip of a self-sensing cantilever was FIB processed, resulting in a flat top plateau with 800 nm side length. Next, FEBID was used for TP fabrication as shown in Figure 8a by a tilted SEM image. Such 3D nano-probes (3DNP) were then used for AFM measurements on a FIB processed test layout. Figure 8b gives a representative 3D height image, obtained in contact mode with a constant force load of 250 nN and a scan speed of 20 µm.s-1. Dedicated scan-speed tests revealed stable operation up to 200 µm.s-1 (see also Supporting Movie 3 and Supporting Information 3). While these experiments confirm the general applicability of 3DNPs as AFM tips, image quality is evidently poor by means of high noise, line jumps and blurred edge features (Figure 8b), which continuously worsened during scanning for more than one hour. Figure 8 shows tilted SEM images before (c) and after AFM operation (d) of another 3DNP, which reveals the decaying image quality as mechanical wear effect at the tip apex. This finding is in agreement with the FES results of highest local stress at the tip apex for multi-leg architectures (see Figure 2e). The origin of this low wear resistance can be attributed to the inner structure of FEBID based materials, which consist of nano-sized platinum grains (typical 2 to 4 nm in diameter) embedded in a carbon matrix with typical C-contents around 85 at.%

40,41

. The latter mainly stems from incompletely

dissociated precursor molecules and non-volatile fragments during FEBID and is unavoidable for this precursor material 40,42 . In a recent study, Arnold et al. 31 focused on the mechanical properties of PtC nano-pillars, demonstrated their tunability and gave an explanation for its underlying reasons. In brief, Pt-C based FEBID materials were subjected to post-growth electron-beam curing (EBC) 43 , which entails two effects. First, incompletely dissociated precursor molecules are further fragmented, leading to slight particle growth with strong implications on electric properties as initially demonstrated by the work group around Michael Huth in several studies

29,30,44–46

. Complementary, Arnold et al.

demonstrated a strong increase in the Young modulus during EBC by a factor of 5. The origin of this


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behavior was suggested to rely on the electron-beam induced, chemical modification of the carbon matrix from loosely bound sp2 networks over glassy carbon towards amorphous glass 47 . Based on these findings, we applied the same procedure to our TPs with a typical curing dose of 150 nC.µm-2 performed at 30 keV and 44 pA to minimize the risk of co-deposit by transmitting low energy electrons and high beam currents 27 . Figure 9a shows a series of SEM images of a 3DNP right after fabrication (left), after EBC with a dose of 150 nC.µm-2 (center) and after long time AFM scanning (right) as discussed in the following. As can be seen, the TP slightly shrank after EBC (~ 30 nm) in agreement with literature

48

and revealed a small increase of the apex radius from ~ 6 nm to ~ 9 nm. Prior to AFM

measurements, force-ramp experiments, as presented in Figure 6, were repeated, which indicated a linear behavior up to 250 nN with spring constants higher than 40 N.m-1 (see Supporting Information 3). Figure 9b shows a 3D AFM height image acquired with the 3DNP shown in the central SEM image of (a) at identical AFM operating conditions used for the image in Figure 8b for as deposited 3DNPs. As evident, the image quality has significantly improved and provides lateral resolution below 10 nm. Next, the scan speed was incrementally increased to 160 µm.s-1 with 20-minute scan periods each, summarized by Figure 9c. As can be seen, the image quality slightly deteriorates at scan speeds higher than 80 µm.s-1, but can be restored when going back to 20 µm.s-1 (bottom right). Using the former scan speed with a resolution of 512 x 512 pixels, a 4 x 4 µm² scan takes only 50 s. This is the same time scale as the acquisition of high-resolution SEM images, hence, demonstrating the high-speed capability of our 3D nano-probes. In the following, we subjected such EBC treated 3DNPs to long time measurements by scanning a total distance of about 55 cm in less than 4 hours at a constant force load of 250 nN and a scan speed of 40 µm.s-1. Although the image quality was still good (see Supporting Information 3) we could recognize slight feature broadening. This tip quality loss is not unusual for such operating conditions and total scan lengths. The right image in Figure 9a shows the same 3DNP after both, scan-speed variation (Figure 9c) and long-time scanning (Supporting Figure S 20), revealing a small increase of the tip apex radius (~ 11 nm) in agreement with the aforementioned feature broadening. The main finding, however, is the absence of strong wear effects for EBC treated 3DNPs even after long time scanning (right image in Figure 9a) compared as-deposited 3DNP after just one ACS Paragon Plus Environment


hour (Figure 8d). Hence, electron-beam curing has been proven as successful material improvement approach to increase mechanical wear resistance for stable AFM operation as originally intended.

3.5 Thermal Response The final goal of our approach is the fabrication of 3D nano-probes on self-sensing cantilever (SS-CL) for application with a novel in situ AFM platform, called AFSEM™ (GETec Microscopy Inc., Vienna, Austria). A central element of this system are piezoelectric self-sensing elements, integrated close to the cantilever-chip contact region as shown in Figure 10a by a tilted SEM image. This approach enables electric read-out of cantilever deflection and twisting, thus replacing space consuming optical laserdetector systems 44 . By that, the entire system is highly compact for seamless integration in space confined SEM, FIB or dual beam microscopes (DBM) to enable complementary in situ AFM characterization. The target concept for thermal probing using 3D nano-probes is based on electric thermistors, which effectively change electric resistivity in response to their temperature, broadly used as temperature sensor concept for macroscopic applications (e.g. Pt100 element) but also for SThM probes 3,49,50 . Here, the boundary conditions are two electrodes, which are bridged by our 3DNP with 2 legs, each. Applying a constant current (explained in detail later) while measuring the voltage drop across the bridge gives information about resistivity changes and by that indirectly about the temperature of the 3DNP. As indicated in Figure 10a, the original tip is covered by a 100 nm thick, prestructured Au electrode. To modify the Au covered tip for thermal nano-probe fabrication, a FIB based two-step procedure is applied, followed by the FEBID fabrication of 3DNPs. First, the tip region is removed to form a 700 nm wide plateau with a pre-tilt of 11° to establish parallel orientation to the substrate during AFM operation. Next, the Cr/Au electrode is split across the entire tip region by the formation of a 100 nm wide trench, as can be seen in Figure 10b-d by a tilted and re-colored SEM image. Electrical in situ measurements were used to accurately stop FIB milling once the path is opened, indicated by a sudden resistance increase in the instrument compliance (R > 800 M. Finally, 3DNPs were fabricated with 600 nm wide, squared footprints and inclination angles around 70°, resulting in heights around 1200 nm, with branch diameters around 85 nm. The 3DNPs were placed in ACS Paragon Plus Environment

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such a way that two legs were connected at each electrode as mentioned above and shown in Figure 10 by SEM side (c) and top view (d) images. After fabrication, 3DNPs are subjected to post-growth EBC to improve mechanical wear resistance as discussed in the previous section. As mentioned before, electron-beam curing also affects electrical properties as the slight grain growth reduces grain-to-grain distances for the here applied doses, which increases the tunnel probability and by that the electric conductivity. As this process eventually stops, electric in situ measurements can be used to control the curing process for defined and reproducible electric properties of 3DNPs. For our architectures, we typically observed first saturation tendencies around 10 nC.µm-2 followed by the second linear regime around 70 nC.µm-2, indicating the emergence of the graphitization dominated regime (see Supporting Information 4)

31

. EBC was always stopped after a total dose of 150 nC.µm-2, resulting in bridge

resistances between 10 k and 20 k , depending on initial 3DNP dimensions and now denoted as 3D thermal nano-probes (3DTNP). Figure 10e gives an I/U curve of the shown 3DTNP with a linear behavior even for small voltages as evident by the double logarithmic inset. Taking into account typical bridge resistances around 10 M before EBC, this means an improvement by almost 3 orders of magnitude in well agreement with literature 40,45,48 . For completeness, we want to mention that the shown curve are uncorrected raw data with a serial pre-resistor (RPRE) of 1 M. Next, the electric resistance response of 3DNPs is studied in dependency on their temperature. To provide a reliable experimental setup, calibrated MEMS heater chips (Wildfire S3 ships, DENS Solutions, Delft, The Netherlands) were used as substrate with an in-house built DBM chip holder, which provides full compatibility with our in situ AFM equipment (see Supporting Information 4). Experiments were conducted on silicon-nitride coated chips to prevent electric contact to the underlying Pt heating elements. For heat response measurements, 3DNP modified SS-CL with spring constants of 4 – 8 N.m-1 were used, electrically operated by a constant current source at 10 nA while the voltage drop was constantly measured. Please note, current sweeps up to ± 1 µA did not reveal deviations from a linear I/U behavior, which would indicate Joule self-heating. All experiments were performed in quasi-static conditions, realized by a scan range of 5 x 5 nm, with a force load of 50 nN.



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Figure 11a shows the relative resistance variation R/R0 (left ordinate) during a slow temperature profile up to 80°C and back to room temperature (solid blue line, right ordinate). As reference, unprocessed Au covered SS-CL were used, which revealed no resistance change as evident by the red line and circles. In contrast, 3DTNPs showed a clear and reversible response to temperature variation as illustrated by the green curve and triangles in Figure 11a. Please note, the shown data are unprocessed, spline-less raw data with a symbol distance of 30 values. Apart from this general proofof-principle, the decreasing resistance requires a closer look (see inverted left ordinate). In nanogranular FEBID materials electric transport takes place in the correlated variable-range-hopping (cVRH) regime and do not behave like a classical metal. As shown by Huth and coworkers, electric conductivity increases with temperature as inter-granular (co-)tunneling is thermally assisted 45 . This leads to a negative temperature coefficient (NTC) compared to positive temperature coefficients (PTC) for most metals. By that, the here found decreasing resistance with increasing temperatures is consistent with theory. This behavior also explains why a constant current approach was chosen: the dissipated power by the measurement circuit is given by = 𝑈.𝐼 . Substituting U by 𝐼.𝑅 gives 𝑃 = 𝐼2.𝑅, which makes clear that decreasing resistances for higher temperatures lead to lower dissipation and by that prevents selfheating. In contrast, constant voltage measurements would lead to 𝑃 =

𝐼2

𝑅 with the consequence

that decaying resistances increase power consumption, which can lead to unwanted self-heating. NTC values for the here studied 3DTNPs were found to -(7.5 ± 0.2) × 10-3 K-1, which ultimately depends on volume and shape of the topmost merging region. A closer look on the noise level in the top plateau at 80°C is given by the inset in Figure 11a by unprocessed raw data, revealing a recalculated temperature noise of less than ± 0.5°C derived from the maximum variation. The here extracted, absolute NTC value of ranges in the same order of magnitude as for most metals (3.6 × 10-3 K1

to 6.5 × 10-3 K-1) and / or alloys (0.2 × 10-3 K-1 to 6.3 × 10-3 K-1) 51 . Compared to literature reported,

resistive SThM probes, the absolute NTC values of our 3DTNPs are also in the same order of magnitude 52

; e.g. Wollaston bridges (1.6 × 10-3 K-1)

53

or Pd coated small scale thermistors (3.6 × 10-3 K-1)

54

.

Further comparison to other concepts is partly complicated as many approaches use e.g. Wheatstone


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/ Thomson bridges and / or AC readout approaches such as 3

3,50,53,55,56 ,

which amplify the sensitivity

in comparison to our approach. Wielgoszewski et al. demonstrated the smart adaption of read-out electronics and achieved a sensitivity increase from 0.91 µV.K-1 (Thomson) over 5.43 µV.K-1 (Wheatstone) up to 7.05 µV.K-1 (modified Wheatstone) 55 . Recalculating the values for our 3DTNPs gives intrinsic sensitivity values of 0.8 – 1.5 µV.K-1 (depending on the final geometry) without any bridge electronics, which is very promising once configured in advanced read-out electronics as planned for the future. As nano-granular FEBID materials are known to change their electric properties by compressive and tensile deformation 44,48 , mechanical compression influences were tested as well. Gradual force increase up to 75 nN at different temperatures revealed voltage offsets of less than 1 rel. %, while NTC values at constant pressures remained the same with the error margin of 0.2 × 103

K-1 (see Supporting Information 4). In the final step, we maximized the MEMS heating ramp to

evaluate the achievable sensing rate for 3DTNPs. This is of particular relevance, as one of the main arguments for our 3D concept are the small active sensing volumes, which are a basic prerequisite for fast thermal response during both, heating up and heat dissipation 3 . Figure 11b shows the time resolved response for 25°C – 30°C and 20°C – 50°C ramping profiles, again by unprocessed raw data. While the dotted blue line shows the temperature feedback from the MEMS heater (right ordinate), the green curve depicts again the relative resistance variation (left ordinate) acquired in 66 ms steps. As evident, the slopes of both signals practically lie over each other, which reflects an immediate electric response during local temperature changes. Based on the slopes, a sensing rate of (30 ± 1) ms.K-1 is found, while the temperature noise of ± 0.5°C in the plateau (see inset) is similar to the slow ramping experiments (compare to inset in Figure 11a). Please note, that even faster response rates for the 3DTNPs are very likely as both curves are very similar and the MEMS heater eventually limited the temperature ramping speed. Although faster MEMS heaters have to be applied in future to determine the achievable sensing rates, the presented data clearly confirms that our 3DTNP concept allows for fast, quantitative and reversible temperature sensing, even without bridge electronics or more sophisticated AC read-out approaches.



4 Conclusion In this study, we have delivered the proof-of-principle for the application of FEBID based 3D nanobridges as thermal nano-probes with quantitative, reversible and fast sensing properties, suitable for stable AFM operation. First, a FES model was developed and validated by experimental correlation, which identified the implications of lacking spatial precision by means of twisting, bending and nonlinearity during force load. Optimization of both, architecture and geometric dimensions, led to 3D nano-probe design with 4 legs (tetra-pod) revealing vertical and radial stiffness higher than 50 N.m-1 and 5 N.m-1, respectively. Furthermore, highest local stress in 3DNPs during force load was predicted in the tip apex region, which was verified by AFM experiments leading to strong wear effects after scanning in contact mode. Post growth electron-beam curing with a dose around 150 nC.µm-2 was introduced as counterstrategy, which significantly increased the wear resistance of the apex region. Fully optimized 3DNPs revealed high AFM image qualities up 80 µm.s-1 scan speeds and minor wear effects after continuous scanning for several hours (~50 cm total scan length). In the last step, 3DNPs were fabricated on self-sensing cantilevers, electrically operated in constant current mode and tested on calibrated heating chips. The results reveal a negative temperature coefficient of resistance of – (7.5 ± 0.2) × 10-3 K-1, which compares very well to alternative resistive probe concepts. This also holds for the noise level and dynamic sensing rates of ±0.5 K and 30 ms.K-1, respectively. By that, the concept of freestanding, FEBID based 3D nano-thermistors for temperature sensing, now denoted as 3D thermal nano-probes, is fully validated. Finally, two more advantageous details should be highlighted in comparison to alternative resistive probe concepts: 1) the achievable apex radii are in the range of 10 nm without any further process steps, which is an essential prerequisite towards high-resolution imaging; and 2) effective process times can be as low as 10 minutes for both, initial direct-write 3D fabrication and electron-beam curing, which eliminates more complicated multi-step procedures during cantilever / nano-probe manufacturing. As mentioned in the introduction, current activities focus on the integration of our 3DNPs in SThM applications, including advanced electronic read-outs as well as morphological and thermal resolution tests to explore the full potential of our nano-probe


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approach in combination with GETec’s novel in situ AFM equipment for seamless integration in SEM, FIB and DBM systems.

5 Methods 5.1 Finite Element Simulations Simulations were conducted using the stationary studies option of the COMSOL Structural Mechanics Module. All models were generated within the COMSOL environment, assuming homogeneous material parameters. Density and Young’s Modulus as input parameter were determined in a previous study to be 4 g.cm-3 and 14 GPa 31,33 . After a brief investigation on the influence of the Poisson ratio, this parameter was set to 0.2, as it would not change the final stiffness by more than 1 % assuming a value lower than 0.5 (see Supporting Information 1). The mesh settings were chosen to “finer”, which indicated a change of less than 1 % from “extra fine”, as determined by initial simulations on a pillar test geometry. Structures were deflected in certain directions using an applied force of 10 nN. For better visibility, the deflection of structures was increased in the settings to emphasize and understand the qualitative bending behavior. Stiffness values from simulations were extracted as the ratio of simulated deflection over applied force.

5.2 Nanofabrication Focused ion beam (FIB) processing and 3D-nanoprinting via Focused Electron Beam Induced Deposition (FEBID) were performed in a NOVA 200 Dual Beam Microscope (Thermo Fisher Scientific). All deposits used trimethyl(methylcyclopentadienyl)platinum(IV) (MeCpPt(IV)Me3 ; CAS: 94442−22−5) as precursor, using a standard gas injection system (GIS) mounted in a high angle port at 38°, arranged 100 µm above the sample surface and 340 µm radial distance to the beam center. The precursor was always heated to 45°C for at least 3 hours prior to first experiments. Prior to any deposition, the beam



was corrected for stigmatism and focused by initially deposited single dots. Before 3D fabrication was initiated, the GIS valve was opened for at least 3 minutes to establish a thermodynamic equilibrium. The base pressure of the microscope before deposition was (5 – 7) 10-6 mbar, which increased to (0.9 – 1.3) 10-5 mbar after opening the GIS valve. 3D FEBID was conducted using exposure files created via 3BID 26 software for a 16-bit patterning generator simply by building a computer aided design within the software. While 30 keV and 21 pA were used for fabrication, 30 keV and 44 pA were used for postgrowth electron-beam curing in top view arrangement. For compression experiments, 1 x 1 cm² silicon substrate with 3 nm oxide layer were used 57 . AFM imaging tests used the same substrates, which were sputter coated with 100 nm Au, further FIB processed at 30 kV and 10 pA using bitmap files. AFM tips were truncated via FIB, also using 30 kV and 10 pA. Before any SEM inspection, a gas clearance time of at least 1 hour was introduced. Imaging parameters were 5 keV and 98 pA with lowest possible dwell times to minimize unwanted effects.

5.3 AFM Characterization In-situ AFM: Compression and thermal studies used the AFSEM™ platform by GETec Microscopy Inc. (Vienna, Austria), which is designed for seamlessly integration in SEMs, FIBs and dual beam microscopes without limiting the main functionalities. A key feature are piezo-electric self-sensing cantilevers, which eliminate space consuming, optical detection systems. All experiments in this study were performed with cantilevers showing stiffness of 6 – 8 N.m-1. Truncating of the original tip was performed via FIB as described in the previous chapter. The AFM is operated by a specially modified ANFATEC controller (Oelsnitz, Germany), which allows manipulation and extraction of electric signals towards the tip electrodes, as relevant for thermal measurements. Compression experiments used the classical ramp mode with controlled Z movement depending on the studied 3D multi-pods. Prior to any ramping experiment, the cantilever were calibrated against the Si-SiO2 substrates to enable quantitative force measurements. Thermal response experiments were carried out in quasi-static scan situations using a 5 x 5 nm scan range and a force load of 50 nN. ACS Paragon Plus Environment

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Ex-situ AFM: Imaging experiments of 3D nano-probes, including wear studies, were performed with a FastScan Bio system, operated by a Nanoscope V controller (Bruker Nano, Santa Barbara, USA) using Tap150A cantilever (Bruker AFM Probes, Camarillo, USA). The cantilever had a nominal spring constant of 5 N.m-1 and were initially truncated via FIB for subsequent 3D nano-probe fabrication as described in the previous chapter. Prior to each experiment, the cantilever were calibrated against Si-SiO2 substrates to allow for quantitative force measurements. All imaging experiments in this study were carried out in contact mode with a force load of 250 nN, while gain parameters were optimized for highest quality possible. Analyses were done by Nanoscope Analysis 1.5 software (Bruker Nano Surface, Santa Barbara, USA) or by Gwyddion v.2.51 58 . If not stated otherwise, images were only subjected to a 1st order plane tilt.

5.4 MEMS Heating Thermal response measurements for 3D nano-probes were performed on calibrated MEMS heater chips (Wildfire S3 Chips, DENS Solutions, Delft, The Netherlands). To operate the heater chips in our NOVA 200 dual beam microscope, a sample holder was designed, fabricated and tested in house. Heating and temperature sensing was performed with the original equipment by DENS Solution using the software Digiheater (V.3.2, DENS Solutions, Delft, The Netherlands), which allows controlled ramping, holding and on-demand temperature profiles, while reading out actual temperatures. As mentioned before, 3D nano-probes, fabricated on self-sensing cantilever, were used in quasi-static conditions (5 x 5 nm scan range) with a constant force load of 50 nN.

6 Acknowledgements HP, JS, JF, SH and RW thank Prof. Ferdinand Hofer for scientific discussions and for financial support concerning instrumentation. The same authors thank DI Franz Hofbauer and DI Anna Weitzer for



support with AFM electronics and manuscript preparation, respectively. Particular gratitude goes to Dr. Ernest Fantner for the long lasting collaboration and for bringing alive the “Christian Doppler Laboratory for Direct-Write Fabrication of 3D Nano-Probes” in joint efforts. The financial support by the Austrian Federal Ministry of Science, Research and Economy and the National Foundation for Research, Technology and Development is gratefully acknowledged. Financial support was also received from the FFG Austria in the frame of the “Beyond Europe” initiative (Project AIM, Nr. 11056459).

7 Supporting Information The following aspects are covered by the supporting information: 1) finite element simulation related aspects concerning i) Poisson’s ratio, ii) comparison to analytical solutions, iii) stiffness scaling behavior in dependency on heights and cross-sectional areas, and iv) the full set of stiffness simulation for different architectures, heights and branch diameters. 2) compression related data concerning i) AFM tip preparation for FEBID modification, ii) morphological twisting modes via experiments and simulations and iii) non-linear force distance curves via experiments and simulations. 3) AFM imaging detail concerning i) high-speed scanning with as-deposited tetra-pods, ii) stiffness and linearity for fully optimized tetra-pods and iii) long time / distance scanning with ideally fabricated and post-treated tetra-pods. 4) aspect during thermal response studies concerning i) resistance changes during postgrowth electron-beam curing, ii) MEMS heater setup and iii) cross-influences of force load on electric resistances. In addition, we provide 3 supporting movies, which show: 1) twisting effects during in situ compression experiments, 2) compression behavior of optimized tetra-pods and 3) real time SEM imaging during AFM scanning in vacuum using tetra-pod modified AFM tips.


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Figure 1

Figure 1: The conceptual route towards 3D thermal nano-probes. The basis are self-sensing cantilevers (grey), which are pre-structured with Au electrodes (yellow). The actual 3D nano-probe (3DNP, bright blue) is placed on top of a truncated tip region and electrically bridges two electrodes. Thermal probing is based on temperature dependent resistance changes through the 3DNP (thermistor), once in contact with the surface of interest. The modular study starts with the overall 3DNP design (Step 1) and then focuses on material modification to enable stable AFM operation (Step 2). Finally, the electric response of the 3DNP thermistor is studied in thermally static and dynamic conditions (Step 3) using a variable heating stage (red dashed circle).



Figure 2

Figure 2: Mechanical stiffness scaling and local stress distribution. FES based scaling of mechanical stiffness during vertical (a) and lateral (b) force loads for a single pillar in dependency on its diameter and height (note the very different stiffness scales). (c) – (e) gives the qualitative, local stress distribution under vertical (top row) and lateral (bottom row) force load for a single pillar (c), a bi-pod (d) and a tetra-pod (e). The inclination angle  and the different force directions with respect to the bi-pod plane (𝒌𝑳 ⊥ and 𝒌𝑳 ∥ ) are indicated in the top and bottom parts of (d), respectively.


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Figure 3

Figure 3: Architecture dependent stiffness scaling and radial homogeneity. (a) gives the vertical stiffness (𝒌𝑽) vs. the architecture type for selected heights, diameters and inclination angles (indicated in Figure 2d) to see individual influences. Particular relevance is the final design shown by green diamonds as described in the main text. (b) gives the same plot as shown in (a), however for radial stiffness (𝒌𝑳). For bi-pods, there are always two points, which represent force loads perpendicular (𝒌𝑳 ⊥ ) and parallel (𝒌𝑳 ∥ ) to the bipod plane (see Figure 2d), corresponding to the lower and higher values respectively. (c) shows the normalized, lateral stiffness in dependency on the force load direction. For the bi-pod (red ellipse), the force directions are indicated in agreement with indications in Figure 2d. While the tetra-pod (green) is circular, the tri-pod shows slight deviations of less than 4 rel.% according to its three-fold symmetry (dotted blue).



Figure 4

Figure 4: Tetra-pod twisting during force load. SEM assisted AFM compression tests used FIB pre-processed cantilever to generate a parallel plane with respect to the substrate (a). (b) – (d) show a compression series of a tetra-pod starting with the approach (b) over the compression (c) and further release (d) of the force load (see also Supporting Movie 1). In this case, both the strong twisting during compression and slight irreversible deformation after compression are observed (see red arrows and also Figure 5).


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Figure 5

Figure 5: Identification of twisting due to spatial fabrication mismatch. SEM inspection of tetra-pods in a tilted (a) and top view arrangement (b), which both reveal spatial inaccuracies in the central merging region. To study the implications of this mismatch, FES were conducted, which use different types of misalignments (see inset in (c)). This approach could well mimic the observed twisting effects during real experiments (d). Please note, that (d) shows an irreversibly deformed tetrapod from top after multi-cycle compression, which is qualitatively identical to dynamic deformation during force load.



Figure 6

Figure 6: Identification of non-linear compression effects due to non-straight branches. (a) shows AFM based force vs. displacement curves with the reference curve of the cantilever against the substrate by the dotted, black line. All curves are unprocessed raw data. The solid curves illustrate compression behavior on tetra-pods revealing two linear (green and blue) regimes followed by a strong saturation behavior (red). Recalculation of the vertical tetra-pod stiffness suggests (18±2) N.m-1 following the indicated equation. (b) shows a tilted SEM image of a tetra-pod, which reveals the non-straight character of the side branches (see eff). (c) The red dotted line indicates the take-off angle TO, which was used for calculating the angle deviation  (see definition on top). These data were then used for FES accessing the vertical stiffness in dependency on , revealing a fast drop by a factor of 4 for  ~ 5°. The top insets show the according models, which reflect the non-straight side branches.


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Figure 7

Figure 7: Design adaption and FES validation. (a) shows a vertical stiffness simulation using SEM based values for inclination angles, heights and deviations from straight branches whenever found. (b) shows the same plot but with real experiments using the same tetra-pods, which reveals good qualitative and quantitative agreement with vertical stiffness values up to 70 N.m-1. Experimental errors are ± 2 N.m-1. (c) shows a SEM top view image of an optimized tetra-pod, where the slightly increased merging region by the introduction of additional pattern points can be seen (blue arrows). This becomes also evident in a SEM side view in the left image of (d), which shows a tetra-pod before AFM compression. Furthermore, the process adaption by additional refresh times during 3D growth leads to very straight side branches, also evident in (d). The right part of (d) shows the same tetra-pod after multi-compression cycles, which leads to slight sideward bending, while the overall geometry is still maintained (see also Supporting Movie 2).



Figure 8

Figure 8: AFM operation and wear effects. (a) shows a tilted SEM image of a FIB truncated AFM tip (yellow arrow) for further fabrication of a 3D tetra-pod as 3D nano-probe (blue arrow). Such tips were then used in AFM contact mode on a FIB processed test structure. (b) shows a representative image, which was taken with 250 nN force load and a scan speed of 20 µm.s-1 (see also Supporting Movie 3). This image shows raw data, except for a 1st order plane tilt. The poor image quality could be traced back to a wear effect of the tip apex region, shown by a direct comparison before and after AFM operation in (c) and (d), respectively.


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Figure 9

Figure 9: Material tuning and AFM performance. (a) shows tilted SEM images of 3D nano-probes after fabrication (left), after electron-beam curing (center) and after AFM operation (right). As evident, electron-beam curing (EBC) leads to small, vertical shrinkage and slight increase of the tip apex radius. After contact mode AFM for more than 4 hours at a scan speed of 40 µm.s-1 and a constant force load of 250 nN, the 3D nano-probe maintains its overall shape and reveals only minor increase of the apex radius. (b) shows an AFM height image, taken with the probes shown in (a) using 20 µm.s-1 and a force load of 250 nN. The comparison to Figure 8b clearly reveals the impact of EBC by the strongly improved image quality. (c) shows a speed test series from 20 µm.s-1 up to 160 µm.s-1 and back to 20 µm.s-1 (bottom right). As evident, the image quality starts to decrease at 80 µm.s-1 but can be restored to the original quality when going back to 20 µm.s-1, which indicates high wear resistance in agreement with the right SEM image in (a). AFM image processing was limited to a 1st order plane tilt.




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Figure 10

Figure 10: Thermal nano-probe fabrication and basic characterization. (a) shows the self-sensing cantilever including the two self-sensing elements and the pre-structured Au electrodes. (b) – (d) are SEM close ups of a 3D thermal nano-probe, showing the FIB based fabrication of a flat plateau and the electrode splitting together with the 3D tetrapod, which electrically bridges the electrodes. (c) and (d) are partly colored to indicate the different layers, while still providing the original SEM image at the left. (e) shows I vs. U curves after EBC with a dose of 150 nC.cm-2, revealing a linear behavior. The small inset is a double logarithmic plot to demonstrate the linearity for small voltages as well. Recalculation of the resistances gives ~ 8.6 M before and ~ 10.8 k bridge resistance, meaning a resistivity reduction by a factor of ~ 800 after EBC. Please note, the presented curves are raw data using a 1 M serial pre-resistor.



Figure 11

Figure 11: Thermal response in static and dynamic conditions. (a) shows the electric response of 3D thermal nano-probes (3DTNP) for increasing temperatures performed on a heating MEMS chip. The red curve shows the response of an untreated, Au covered self-sensing cantilever as reference together with the MEMS temperature read-out given by the solid blue line (right ordinate). While Au is not changing the resistance (R/R0, left ordinate), 3DTNPs reveal reduced resistances for higher temperatures due to thermally assisted tunnelling transport in the nano-granular material. Please note, the green curve are raw data with a symbol spacing of 30 points. The negative temperature coefficient of resistance is found to –(7.5 ± 0.2) 10-3 K-1 while the noise level corresponds to ± 0.5°C (see raw data in the inset). (b) is a similar plot as (a) but performed with highest possible heating rates to evaluate 3DTNP sensing rates. Comparing the temperature read-out of the MEMS (dotted blue, right ordinate) with the electric response signal for 3DTNPs (solid green raw data, left ordinate), the fast response character becomes evident. The response rate is estimated with (30 ± 0.2) ms.K-1, while the temperature noise level is found to be similar to (a) as shown by the inset (T ± 0.5 °C). All measurements were performed in quasi-static scanning conditions using a scan range of 5 x 5 nm with a constant force load of 50 nN.


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8 References [1]

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